Light Metals 2011

Light Metals 2011

Check out these new proceeding volumes from the TMS 2011 Annual Meeting, available from publisher John Wiley & Sons: 2nd International Symposium on High-Temperature Metallurgical Processing Energy Technology 2011 : Carbon Dioxide and Other Greenhouse Gas Reduction Metallurgy and Waste Heat Recovery EPD Congress 2011 Friction Stir Welding and Processing VI Light Metals 2011 Magnesium Technology 2011 Recycling of Electronic Waste II, Proceedings of the Second Symposium Sensors, Sampling and Simulation for Process Control Shape Casting: Fourth International Symposium 2011 Supplemental Proceedings: Volume 1 : Materials Processing and Energy Materials Supplemental Proceedings: Volume 2: Materials Fabrication, Properties, Characterization, and Modeling Supplemental Proceedings: Volume 3: General Paper Selections To purchase any of these books, please visit www.wiley.com. TMS members should visit www.tms.org to learn how to get discounts on these or other books through Wiley.

Light Metals 2011 Proceedings of the technical sessions presented by the TMS Aluminum Committee at the TMS 2011 Annual Meeting & Exhibition, San Diego, California, USA February 27-March 3, 2011

Edited by Stephen J. Lindsay

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Copyright © 2011 by The Minerals, Metals, & Materials Society. All rights reserved. Published by John Wiley & Sons, Inc., Hoboken, New Jersey. Published simultaneously in Canada. No part of this publication may be reproduced, stored in a retrieval system, or transmitted in any form or by any means, electronic, mechanical, photocopying, recording, scanning, or otherwise, except as permitted under Section 107 or 108 of the 1976 United States Copyright Act, without either the prior written permission of The Minerals, Metals, & Materials Society, or authorization through payment of the appropriate per-copy fee to the Copyright Clearance Center, Inc., 222 Rosewood Drive, Danvers, MA 01923, (978) 750-8400, fax (978) 750-4470, or on the web at www.copyright.com. Requests to the Publisher for permission should be addressed to the Permissions Department, John Wiley & Sons, Inc., I l l River Street, Hoboken, NJ 07030, (201) 748-6011, fax (201) 748-6008, or online at http:// www.wiley.com/go/permission. Limit of Liability/Disclaimer of Warranty: While the publisher and author have used their best efforts in preparing this book, they make no representations or warranties with respect to the accuracy or completeness of the contents of this book and specifically disclaim any implied warranties of merchantability or fitness for a particular purpose. No warranty may be created or extended by sales representatives or written sales materials. The advice and strategies contained herein may not be suitable for your situation. You should consult with a professional where appropriate. Neither the publisher nor author shall be liable for any loss of profit or any other commercial damages, including but not limited to special, incidental, consequential, or other damages. Wiley also publishes books in a variety of electronic formats. Some content that appears in print may not be available in electronic formats. For more information about Wiley products, visit the web site at www.wiley.com. For general information on other Wiley products and services or for technical support, please contact the Wiley Customer Care Department within the United States at (800) 762-2974, outside the United States at (317) 572-3993 or fax (317) 572-4002. Library of Congress Cataloging-in-Publication Data is available.

ISBN 978-1-11802-935-0 Printed in the United States of America. 10 9 8 7 6 5 4 3 2 1

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TABLE OF CONTENTS Light Metals 2011 Preface About the Editor Program Organizers Aluminum Committee

xix xxi xxiii xxix

Alumina and Bauxite Bauxite Resources and Utilisation New Development Model for Bauxite Deposits P. ter Weer

5

Study on the Characterization of Marginal Bauxite from Parâ/Brazil F. Silva, J. Sampaio, M. Medeiros, andF. G arrido

13

Resource Utilization of High-sulfur Bauxite of Low-median Grade in Chongqing China J. Yin, W. Xia, andM. Han

19

Development of Bauxite and Alumina Resources in the Kingdom of Saudi Arabia A. Al-Dubaisi

23

Digestion Studies on Central Indian Bauxite P. Raghavan, N. Kshatriya, and Dasgupta

29

Effects of Roasting Pretreatment in Intense Magnetic Field on Digestion Performance of Diasporic Bauxite Z Ting-an, D. Zhihe, L Guozhi, L. Y an, D. Juan, W. Xiaoxiao, andL. Y an

33

Bayer Process I Application of Operation Integrity Management in the Alumina Industry C Suarez, D. Welshons, J. McNerney, andJ. Webb Influence of Solid Concentration, Particle Size Distribution, Ph And Temperature on Yield Stress of Bauxite Pulp C. Barbato, M. Nele, and S. Franca

41

47

A New Method for Removal Organics in the Bayer Process B. Yingwen, L. Jungi, S. Mingliang, andZ. Fei

51

Alunorte Expansion 3 - The New Lines Added to Reach 6.3 Million Tons per Year D. Khoshneviss, L. Correa, J. Ribeiro Alves Filho, H. Berntsen, andR. Carvalho

57

One Green Field Megaton Grade Large Alumina Refinery with Successful Engineering & Operation Experience...63 L. Xianqing, and Y. Xiaoping Advanced Process Control in the Evaporation Unit C. Kumar, U. Giri, R. Pradhan, T. Banerjee, R. Saha, and P. Pattnaik

v

69

Improvements in Smelter Grade Alumina Quality at Clarendon Alumina Works R. Shaw, A. Duncan, andM. Crosciale

75

Red Mud Application of Nanofiltration Technology to Improve Sea Water Neutralization of Bayer Process Residue K. Taylor, M. Mullett, L. Fergusson, H. Adams on, andJ. Wehrli

81

Caustic and Alumina Recovery from Bayer Residue S Gu

89

Investigation on Alumina Discharge into the Red Mud Pond at Nalco's Alumina Refinery, Damanjodi, Orissa, India B. Mohapatra, B. Mishra, and G Mishra Production of Ordinary Portland Cement( OPC) from NALCO Red Mud G Mishra, D. Yadav, M. Alii, and P. Sharma

93 97

Recovery of Metal Values from Red Mud P. Raghavan, N. Kshatriya, andK. Wawrynink

103

Red Mud Flocculants used in the Bayer Process S. Moffatt, F. Ballentine, andM. Lewellyn

107

Reductive Smelting of Greek Bauxite Residues for Iron Production A. Xenidis, G Zografidis, I. Kotsis, andD. Boufounos

113

Precipitation, Calcination and Properties Effect of Technological Parameters on PSD of Aluminum Tri-Hydroxide from Seed Precipitation in Seeded Sodium Aluminate Solutions 121 Y. Wu, L. Mingchun, and Q. Yanping Methods to Reduce Operating Costs in Circulating Fluidized Bed Calcination G Klett, M. Miss alla, B. Reeb, andH. Schmidt

125

Pressure Calcination Revisited F. Williams, and G Misra

131

Dynamic Simulation of Gas Suspension Calciner (GSC) for Alumina B. Raahauge, S. Wind, M. Wu, and T. Jensen

137

Physical Simulation and Numerical Simulation of Mixing Performance in the Seed Precipitation Tank with a Improved Intermig Impeller Z Ting-an, L. Y an, W. Shuchan, Z Hongliang, Z. Chao, Z Qiuyue, D. Zhihe, andL. Guozhi

145

Two Perspectives on the Evolution and Future of Alumina L. Perander, J. Metson, and G Klett

151

Significant Improvement of Energy Efficiency at Alunorte's Calcination Facility M Missalla, H. Schmidt, J. Ribeiro, andR. Wischnewski

157

Attrition of Alumina in Smelter Handling and Scrubbing Systems S. Lindsay

163

vi

Energy and Environment Perspective on Bayer Process Energy D. Donaldson

171

Optimization of Heat Recovery from the Precipitation Circuit R. Singh, S. Hial, andM. Simpson

175

Alunorte Global Energy Efficiency A. Monteiro, R. Wischnewski, C. Azevedo, and E. Moraes

179

Opportunities for Improved Environmental Control in the Alumina Industry R. Mimna, J. Kildea, E. Phillips, W, Carlson, B. Reiser, andJ. Meier

185

Alumina Refinery Wastewater Management: When Zero Discharge Just Isn't Feasible L. Martin, andS. Howard

191

High Purity Alumina Powders Extracted from Aluminum Dross by the Calcining—Leaching Process L. Qingsheng, Z. Chunming, F. Hui, andX. Jilai

197

Effect of Calcium/Aluminium Ratio on MgO Containing Calcium Aluminate Slags W. Bo, S Hui-Lan, G Dong, andB. Shi-Wen

201

Study on Extracting Aluminum Hydroxide from Reduction Slag of Magnesium Smelting by Vacuum Aluminothermic Reduction W. Yaowu, F. Naixiang, Y. Jing, H. Wenxin, P. Jianping, D. Yuezhong, and W. Zhihui

205

Application of Thermo-gravimetric Analysis for Estimation of Tri-hydrate Alumina in Central Indian Bauxites—An Alternative for Classical Techniques 211 Y. Ramana, andR. Patnaik Determination of Oxalate Ion in Bayer Liquor Using Electrochemical Method S. Turhan, B. Usta, Y. Sahin, andO. Uysal

215

Alternative Alumina Sources - Poster Session The Effect of Ultrasonic Treatment on Alumina Leaching from Calcium Aluminate Slag S. Hui-lan, W. Bo, G Dong, Z. Xue-zheng, andB. Shi-wen

221

Theory and Experiment on Cooling Strategy during Seeded Precipitation Z Liu, W. Chen, and W. Li

227

Extraction of Alumina from Red Mud by Divalent Alkaline Earth Metal Soda Ash Sinter Process S. Meher, A. Rout, andB. Padhi

231

Dissolution Kinetics of Silicon from Sintering Red Mud in Pure Water X. Li, K. Huang, and H Zhu

237

The Effect of Cooling Rate on the Leachability of Calcium Aluminate Slags W. Bo, S. Hui-lan, Z. Xue-zheng, and B. Shi-wen

241

Preparing Polymerized Aluminum-ferrum Chloride with Red Mud L. Guilin, Y. Haiyan, andB. Shiwen

245

Adsorption of Polyethylene Glycol at the Interface of Dicalcium Silicate - Sodium Aluminate Solution Y. Haiyan, X Pan, Z Lu, and T. Ding

251

vu

Production of Hematite Ore from Red Mud P. Raghavan, N. Kshatriya, and K Wawrynink

255

Aluminum Reduction Technology Enviroment- Emissions/ Anode Effect I HF Measurements Inside an Aluminium Electrolysis Cell K. Osen, T Aarhaug, A. Solheim, E. Skybakmoen, andC. Sommerseth

263

LasIRTM-R - The New Generation RoHS-Compliant Gas Analyzers Based on Tunable Diode Lasers J. Gagne, J. Pisano, A. Chanda, G. Mackay, K. Mackay, and P. Bouchard

269

Use of Spent Potlining (SPL) in Ferro Silico Manganese Smelting P. von Krüger

275

Reduction of PFC Emissions at Pot Line 70 kA of Companhia Brasileira de Aluminio H. Santos, D. Melo, J. Calixto, J. Santos, andJ. Miranda

281

Towards Redefining the Alumina Specifications Sheet - The Case of HF Emissions L Perander, M. Stam, M. Hyland, andJ. Metson

285

Design of Experiment to Minimize Fluoride and Particulate Emissions at Alumar E. Batista, P. Miotto, E. Montoro, andL. Souza

291

Innovative Distributed Multi-Pollutant Pot Gas Treatment System G. Wedde, O. Bjarno, and A. Sorhuus

295

Fluoride Emissions Management Guide (FEMG) for Aluminium Smelters N. Tjahyono, Y. Gao, D. Wong, W. Zhang andM. Taylor

301

Enviroment- Emissions/ Anode Effect II On Continuous PFC Emission Unrelated to Anode Effects X. Chen, W. Li, J. Marks, Q. Zhao, J. Yang, S. Qiu, and C. Bayliss Monitoring Air Fluoride Concentration around ALUAR Smelter in Puerto Madryn (Chubut Province, Argentina) J. Zavatti, C. Moreno, J. Lifschitz, and G. Quiroga

309

315

Reduction of Anode Effect Duration in 400kA Prebake Cells W. Zhang, D. Wong, M. Gilbert, Y. Gao, M. Dorreen, M. Taylor, A. Tabereaux, M. Soffer, X. Sun, C. Hu, X Liang, H. Qin, J. Mao, andX Lin

319

Sustainable Anode Effect Based Perfluorocarbon Emission Reduction N. Dando, L. Sylvain, J. Fleckenstein, C. Kato, V. Van Son, andL. Coleman

325

The Initiation, Propagation and Termination of Anode Effects in Hall-Héroult Cells G. Tarcy, and A. Tabereaux

329

Towards Eliminating Anode Effects A. Al Zarouni, B. Welch, M. Mohamed Al-Jallaf, and A. Kumar

333

Correlation between Moisture and HF Formation in the Aluminium Process C. Sommerseth, K Osen, T Aarhaug, E. Skybakmoen, A. Solheim, C. Rosenkilde, and A. Ratvik

339

Vili

Particulate Emissions from Electrolysis Cells H. Gaertner, A. Ratvik, and T. Aarhaug

345

Investigation of Solutions to Reduce Fluoride Emissions from Anode Butts and Crust Cover Material G. Girault, M. Faure, J. Bertolo, S. Massambi, and G. Bertran

351

PFC Survey in Some Smelters of China W. Li, X Chen, Q. Zhao, S. Qiu, andS. Zhang

357

Considerations Regarding High Draft Ventilation as an Air Emission Reduction Tool S. Broek, N. Dando, S. Lindsay, and A. Moras

361

Cells Thermal Balance Increasing the Power Modulation Window of Aluminium Smelter Pots with Shell Heat Exchanger Technology ...369 P. Lavoie, S. Namboothiri, M. Dorreen, J. Chen, D. Zeigler, andM. Taylor New Approaches to Power Modulation at TRIMET Hamburg T Reek

375

Some Aspects of Heat Transfer Between Bath and Sideledge in Aluminium Reduction Cells A. Solheim

381

Towards a Design Tool for Self-heated Cells Producing Liquid Metal by Electrolysis S. Poizeau, andD. Sadoway

387

Heat Recovery from Aluminium Reduction Cells Y. Ladam, A. Solheim, M. Segatz, and O. Lorentsen

393

Effects of Composition and Granulometry on Thermal Conductivity of Anode Cover Materials H. Wijayaratne, M. Hyland, M. Taylor, A. Grama, and T Groutso

399

Restart of 300kA Potlines after 5 Hours Power Failure X. Zhao, B. Gao, H. Han, J. Liu, J. Xiao, J. Qian, J. Yan, andD. Wang

405

Multiblock Monitoring of Aluminum Reduction Cells Performance J. Tessier, C Duchesne, andG. Tarcy

407

Cells Technology, Development and Sustainability High Amperage Operation of AP18 pots at Karmoy M Bugge, H. Haakonsen, O. Kobbeltvedt, andK. Paulsen

415

Aluminium Smelter Manufacturing Simulation - Can These Bring Real Cost Savings? M Meijer

421

Simultaneous Preheating and Fast Restart of 50 Aluminium Reduction Cells in an Idled Potline - A New Soft Restart Technique for a Pot Line 425 A. Mulder, A. Folkers, M. Stam, andM. Taylor SWOT Perspectives ofMidagePrebaked Aluminium Smelter P. Choudhury, and A. Sharma

431

Integrated Approach for Safe and Efficient Plant Layout Development R. Pires, R. Baxter, L. Tikasz, andR. McCulloch

437

IX

New Progress on Application of NEUI400kA Family High Energy Efficiency Aluminum Reduction Pot ("HEEP") Technology 443 D. Lu, J. Qin, Z. Ai, and Y. Ban Improving Current Efficiency of Aged Reduction Lines at Aluminium Bahrain (Alba) A. Ahmed, K. Raghavendra, H Hassan, andK. Ghuloom

449

Development of NEUI500kA Family High Energy Efficiency Aluminum Reduction Pot ("HEEP") Technology ..455 D. Lu, Y. Ban, X Qi, J. Mao, Q. Yang, andK Dong

Cells Process Control Current Efficiency for Aluminium Deposition from Molten Cryolite-alumina Electrolytes in a Laboratory Cell....461 G. Haarberg, J. Armoo, H Gudbrandsen, E. Skybakmoen, A. Solheim, and T Jentoftsen

Improvement in Cell Equipment and Design Retrofit of a Combined Breaker Feeder with a Chisel Bath Contact Detection System to Reduce Anode Effect Frequency in a Potroom J. Verreault, R. Gariépy, B. Desgroseilliers, C. Simard, X. Delcorde, C. Turpain, S. Simard, andS. Déry

467

Anode Dusting from a Potroom Perspective at Nordural and Correlation with Anode Properties H. Gudmundsson

471

The Application of Continuous Improvement to Aluminium Potline Design and Equipment W. Paul

477

Alcoa STARprobe™ X. Wang, B. Hosier, and G. Tarcy

483

Active Pot Control using Alcoa STARprobe™ X. Wang, G. Tarcy, E. Batista, and G. Wood

491

Technology & Equipment for Starting Up & Shutting Down Aluminium Pots under Full Amperage Y. Tao, L. Meng, C. Bin, and Y. Xiaobing

497

Study on Solution of A1203 in Low Temperature Aluminum Electrolyte H Kan, N. Zhang, andX Wang

503

Applications of New Structure Reduction Cell Technology in Chalco's Smelters F. Liu, S. Gu, J. Wang, andK Yang

509

Transport Numbers in the Molten System NaF-KF-AlF3-Al203 P. Fellner, J. Hives, andJ. Thonstad

513

Cells Process Modeling Development and Application of an ANSYS Based Thermo-electro-mechanical Collector Bar Slot Design Tool M Dupuis Impact of Amperage Creep on Potroom Busbars and Electrical Insulation: Thermal-Electrical Aspects A. Schneider, D. Richard, and O. Charette

519 525

Modern Design of Potroom Ventilation A. Vershenya, U. Shah, S. Broek, T. Plikas, J. Woloshyn, andA. Schneider

531

A Preliminary Finite Element Electrochemical Model for Modelling Ionic Species Transport in the Cathode Block ofa Hall-Héroult Cell 537 F. Gagnon, D. Ziegler, andM. Fafard CFD Modelling of Alumina Mixing in Aluminium Reduction Cells Y. Feng, M. Cooksey, and P. Schwarz

543

Bubble Transport by Electro-Magnetophoretic Foces at Anode Botttom of Aluminium Cells V. Bojarevics, andA. Roy

549

Anodic Voltage Oscillations in Hall-Héroult Cells K. Einarsrud, andE. Sandnes

555

Energy Savings by Cell Design Improvements Electrical Conductivity of the KF-NaF- A1F3 Molten System at Low Cryolite Ratio with CaF2 Additions A. Redkin, A. Dedyukhin, A. Apisarov, P. Tin'ghaev, and Y. Zaikov

563

Study of ACD Model and Energy Consumption in Aluminum Reduction Cells T. Yingfu, and W. Hang

567

Modeling of Energy Savings by Using Cathode Design and Inserts R. von Kaenel, andJ. Antille

569

Experimental Investigation of Single Bubble Characteristics in a Cold Model of a Hall-Héroult Electrolytic Cell S. Das, Y. Morsi, G Brooks, W. Yang, andJ. Chen

575

Large Gas Bubbles under the Anodes of Aluminum Electrolysis Cells A. Caboussat, L. Kiss, J. Rappaz, K. Vékony, A. Perron, S. Renaudier, and O. Martin

581

Initiatives to Reduction of Aluminum Potline Energy Consumption Alcoa Poços de Caldas/Brazil A. Abreu, M. Salles, and C. Kato

587

Overview of High-Efficiency Energy Saving for Aluminium Reduction Cell X. Canming, and Y. Xiaobing

591

Cell Voltage Noise Reduction Based on Wavelet in Aluminum Reduction Cell B. Li, J. Chen, X Zhai, S. Sun, andG. Tu

599

Poster Session Human Factors in Operational and Control Decision Making in Aluminium Smelters Y. Gao, M. Taylor, J. Chen, andM. Hautus

xi

605

Aluminum Rolling Session I An Investigation of Deformation Behavior of Bimetal Clad Sheets by Asymmetrical Rolling at Room Temperature L. Xiaobing, Z Guoyin, and D. Qiang

615

Coil Build Up Compensation during Cold Rolling to Improve Off-line Flatness L. Almeida Neto, and T. Ayhan

621

Through Process Effects on Final Al-sheet Flatness S. Neumann, andK. Karhausen

625

Cast Shop for Aluminum Production Casthouse Productivity and Safety New Casthouse Smelter Layout for the Production of Small Non-Alloyed Ingots: Three Furnaces/Two Lines J. Berlioux, A. Bourgier, andJ. Baudrenghien

635

Use of Process Simulation to Design a Billet Casthouse G. Jaouen

641

Optimizing Scrap Reuse as a Key Element in Efficient Aluminium Cast Houses T. Schmidt, J. Migchielsen, D. Ing, andK Grab

647

Implementation of an Effective Energy Management Program Supported by a Case Study R. Courchée

653

Molten Metal Safety Approach through a Network C. Pluchon, B. Hannart, L Jouet-F* astre, J. Mathieu, R. Wood, J. Riquet, F. Fehrenbach, G. Ranaud, M. Bertherat, andJ. Hennings

657

Improved Monolithic Materials for Lining Aluminum Holding & Melting Furnaces A. Wynn, J. Coppack, T. Steele, andK. Moody

663

Direct Chill Casting Cold Cracking during Direct-chill Casting D. Eskin, M. Lalpoor, andL. Katgerman

669

Surface Defects Structures on Direct Chill Cast 6xxx Aluminium Billets M. Erdegren, and T. Carlberg

675

Effect of Cooling Water Quality on Dendrite Arm Spacing of DC Cast Billets S. Mohapatra, S. Nanda, and A. Palchowdhury

681

Mould Wall Heat Flow Mechanism in a DC Casting Mould A. Prasad, andL Bainbridge

687

Productivity Improvements at Direct Chill Casting Unit in Aluminium Bahrain (ALBA) A. Noor, S. Chateeriji, and A. Ahmed

693

Xll

The Coupling of Macrosegregation with Grain Nucleation, Growth and Motion in DC Cast Aluminum Alloy Ingots M. Zaloznik, A. Kumar, H. Combeau, M. Bedel, P. Jarry, andE. Waz

699

Investment Casting of Surfaces with Microholes and Their Possible Applications T. Ivanov, A. Buehrig-Polaczek, U. Vroomen, C. Hartmann, A. Gillner, K Bobzin, J. Holtkamp, N. Bagcivan, andS. Theiss

705

Using SEM and EDX for a Simple Differentiation of?- and ?-AlFeSi-Phases in Wrought Aluminum Billets M Rosefort, C. Matthies, H. Buck, and H. Koch

711

Dross Formation, Control and Handling Oxidation of AlMg in Dry and Humid Atmospheres A. Kvithyld, D. Stevens, S. Wilson, and T. Engh

719

Study of Early Stage Interaction of Oxygen with Al; Methods, Challenges and Difficulties B. Fatela, G BrooL·, M. Rhamdhani, J. Taylor, J. Davis, andM. Lowe

725

Quality Assessment of Recycled Aluminium D. Dispinar, A. Kvithyld, and A. Nordmark

731

Melt Quality Control In-Line Salt-ACD™: A Chlorine-Free Technology for Metal Treatment P. Robichaud, C. Dupuis, A. Mathis, P. Coté, andB. Maltais

739

The Effectof TiB2 Granules on Metal Quality M. MohamedAl-Jallaf, M. Hyland, B. Welch, A. Al Zarouni, andF. Abdullah

745

Thermodynamic Analysis of Ti, Zr, V and Cr Impurities in Aluminium Melt A. Khaliq, M. Rhamdhani, G. BrooL·, andJ. Grandfield

751

Current Technologies for the Removal of Iron from Aluminum Alloys L. Zhang, J. Gao, andL. Damdah

757

Electromagnetically Enhanced Filtration of Aluminum Melts M. Kennedy, S. Akhtar, R. Aune, andJ. Bakken

763

A Review of the Development of New Filter Technologies Based on the Principle of Multi Stage Filtration With Grain Refiner Added in the Intermediate Stage 769 J. Courtenay, S. Instone, andF. Reusch Wettability of Aluminium with SiC and Graphite in Aluminium Filtration S. Bao, A. Kvithyld, T. Engh, andM. Tangstad

775

Study of Microporosity Formation under Different Pouring Conditions in A356 Aluminum Alloy Castings L. Yao, S. Cocker oft, D. Maijer, J. Zhu, and C Reilly

783

Grain Refinement Alloying, Solidification and Casting Hycast Gas Cushion (GC) Billet Casting System 7. Steen, and A. Hakonsen

793

Xlll

Studies of Fluid Flow and Meniscus Behavior during Horizontal Single Belt Casting (HSBC) of Thin Metallic Strips D. Li, J. Gill, M. Isac, andR. Guthrie

797

Development of Alba High Speed Alloy A. Ahmed, J. Hassan, G. Martin, andK Ghosh

803

Dissolution Studies of Si Metal in Liquid Al under Different Forced Convection Conditions M. Seyed Ahmadi, S. Argyropoulos, M. Bussmann, andD. Doutre

809

Modification and Grain Refinement of Eutectics to Improve Performance of Al-Si Castings M. Felberbaum, and A. Dahle

815

Production of Al-Ti-C Grain Refiners with the Addition of Elemental Carbon and K2TiF6 F. Toptan, I. Kerti, S. Daglilar, A. Sagin, O. Karadeniz, and A. Ambarkutuk

821

Effect of Mechanical Vibrations on Microstructure Refinement of Al-7mass% Si Alloys T. Tamura, T. Matsuki, andK. Miwa

827

Predicting the Response of Aluminum Casting Alloys to Heat Treatment C Wu, andM. Makhlouf

831

Electrode Technology for Aluminium Production Anode Baking Determination of Coke Calcination Level and Anode Baking Level - Application and Reproducibility of L-sub-c Based Methods 841 S. Rorvik, L. Lossius, and A. Ratvik Operation of an Open Type Anode Baking Furnace with a Temporary Crossover E. Cobo, L. Beltramino, J. Artola, J. Rey Boero, P. Roy, andJ. Bigot

847

Recent Developments in Anode Baking Furnace Design D. Severo, V. Gusberti, P. Sulger, F. Keller, andM. Meier

853

Sohar Aluminium's Anode Baking Furnace Operation S. AlHosni, J. Chandler, O. Forato, F. Morales, C. Jonville, andJ. Bigot

859

Meeting the Challenge of Increasing Anode Baking Furnace Productivity F. Ordronneau, M. Gendre, L. Pomerleau, N. Backhouse, A. Berkovich, andX. Huang

865

Wireless Communication for Secured Firing and Control Systems in Anode Baking Furnaces N. Fiot, and C. Coulaud

871

Full Control of Pitch Burn during Baking: It's Impact on Anode Quality, Operational Safety, Maintenance and Operational Costs D. Maiwald, D. Di Lisa, and P. Mnikoleiski High Performance Sealing for Anode Baking Furnaces P. Mahieu, S. Neple, N. Fiot, I. Ofico, andM. Eufrasio

xiv

875 881

Anode Raw Materials and Green Carbon Property Profile of Lab-scale Anodes Produced with 180°C Mettler Coal Tar Pitch W. Boenigk, C Boltersdorf, F. Lindner, andJ. Stiegert

889

Quality and Process Performance of Rotary Kilns and Shaft Calciners L. Edwards

895

Sub-surface Carbon Dioxide Reaction in Anodes D. Ziegler

901

Paste Quality Improvements at Alcoa Poços de Caldas Plant B. Vry, C Kato, J. Araujo, F. Ribeiro, and A. Abreu

907

Prebaked Anode from Coal Extract (2) - Effects of the Properties of Hypercoal-coke on the Preformance of Prebaked Anodes M. Hamaguchi, N. Okuyama, N. Komatsu, J. Koide, K. Kano, T. Shishido, K. Sakai, and T. Inoue The New Generation of Vertical Shaft Calciner Technology J. Zhao, Q. Zhao, and Q. Zhao

913 917

Petroleum Coke VBD Historical and Future Challenges with the Vibrated Bulk Density Test Methods for Determining Porosity of Calcined Petroleum Coke J. Panchal, M. Wyborney, andJ. Rolle

925

Prediction of Calcined Coke Bulk Density M. Dion, H. Darmstadt, N. Backhouse, F. Cannava, and M. Canada

931

Calcined Coke Particle Size and Crushing Steps Affect Its VBD Result F. Cannava, M. Canada, andB. Vitchus

937

Bulk Density - Overview of ASTM and ISO Methods with Examples of Between Laboratory Comparisons L. Lossius, B. Spencer, andH. 0ye

941

Improving the Repeatability of Coke Bulk Density Testing L. Edwards, M. Lubin, andJ. Marino

947

ASTM D7454 Vibrated Bulk Density Method - Principles and Limitations F. Laplante, andL. Duchesneau

953

Vibrated Bulk Density (VBD) of Calcined Petroleum Coke and Implications of Changes in the ASTM Method D4292 B. Spencer, L. Johnsen, D. Kirkpatrick, D. Clark, andM. Baudino

959

Anode Quality and Rodding Processes Multivariate Monitoring of the Prebaked Anode Manufacturing Process and Anode Quality J. Lauzon-Gauthier, C Duchesne, J. Tessier, K. Cantin, and I. Petit

967

Characterization of a Full Scale Prebaked Carbon Anode using X-Ray Computerized Tomography D. Picard, H. Alamdari, D. Ziegler, P. St-Arnaud, andM. Fafar d

973

xv

FEM Analysis of the Anode Connection in Aluminium Reduction Cells S. Beier, J. Chen, M. Fafard, andH. Fortin Development of Industrial Benchmark Finite Element Analysis Model to Study Energy Efficient Electrical Connections for Primary Aluminium Smelters D. Molenaar, K. Ding, and A. Kapoor

979

985

Real Time Temperature Distribution During Sealing Process and Room Temperature Air Gap Measurements of a Hall-Héroult Cell Anode 991 O. Trempe, D. Larouche, D. Ziegler, M. Guillot, andM. Fafard Effects of High Temperatures and Pressures on Cathode and Anode Interfaces in a Hall-Heroult Electrolytic Cell L. St-Georges, L. Kiss, J. Bouchard, M. Rouleau, andD. Marceau

997

New Apparatus for Characterizing Electrical Contact Resistance and Thermal Contact Conductance N. Kandev, H. Fortin, S. Chénard, G. Gauvin, M. Martin, andM. Fafard

1003

Carbon Anode Modeling for Electric Energy Savings in the Aluminium Reduction Cell D. Andersen, and Z Zhang

1009

Cathode Design and Operation Preheating Collector Bars and Cathode Blocks Prior to Rodding with Cast Iron by Passing an AC Current Through the Collector Bars 1017 E. Jensen, H Bjornstad, andJ. Hansen Development and Application of an Energy Saving Technology for Aluminum Reduction Cells P. Jianping, F. Naixiang, F. Shaofeng, L. Jun, and Q. Xiquan

1023

Study of Electromagnetic Field in 300kA Aluminium Reduction Cells with Innovation Cathode Structure B. Li, X. Zhang, S. Zhang, F. Wang, andN. Feng

1029

Evaluation of the Thermophysical Properties of Silicon Carbide, Graphitic and Graphitized Carbon Sidewall Lining Materials Used in Aluminium Reduction Cell in Function of Temperature 1035 A. Khatun, andM. Desilets Advanced Numerical Simulation of the Thermo-Electro-Mechanical Behaviour of Hall-Héroult Cells under Electrical Preheating D. Marceau, S. Pilote, M. Désilets, L. Hacini, J. Bilodeau, and Y. Caratini

1041

Influence of Technological and Constructive Parameters on the Integrity of the Bottom of Aluminum Reduction Cells during Flame Preheating 1047 A. Arkhipov, G. Arkhipov, and V. Pingin Creep Behaviors of Industrial Graphitic and Graphitized Cathodes during Modified Rapoport Tests W. Wang, J. Xue, J. Feng, Q. Liu, L. Zhan, H. He, andJ. Zhu

1053

Cathode Materials and Wear Measurement of Cathode Surface Wear Profiles by Laser Scanning E. Skybakmoen, S. Itervik, A. Solheim, K. Holm, P. Tiefenbach, and O. Ostrem

1061

Coke Selection Criteria for Abrasion Resistant Graphitized Cathodes R. Perruchoud, W. Fischer, M. Meier, and U. Mannweiler

1067

xvi

Determination of the Effect of Pitch-Impregnation on Cathode Erosion Rate P. Patel, Y. Sato, and P. Lavoie

1073

Simplifying Protection System to Prolong Cell Life M. Mohamed Al-Jallaf, M. Hyland, B. Welch, and A. AlZarouni

1079

Aluminate Spinels as Sidewall Linings for Aluminum Smelters X. Yan, R. Mukhlis, M. Rhamdhani, and G. Brooks

1085

A New Ramming Paste with Improved Potlining Working Conditions B. Aliard, R. Paulus, and G. Billat

1091

Towards a Better Understanding of Carburation Phenomenon M. Lebeuf, M Coulombe, B. Allard, and G. Soucy

1097

Characterization of Sodium and Fluorides Penetration into Carbon Cathodes by Image Analysis and SEM-EDS Techniques 1103 Y. Gao, J. Xue, J. Zhu, K. Jiao, and G. Jiang

Inert Anodes and Wettable Cathodes Pressureless Sintering of TiB2-based Composites using Ti and Fe Additives for Development of Wettable Cathodes H. Heidari, H. Alamdari, D. Dubé, andR. Schulz Furan Resin and Pitch Blends as Binders for TiB2-C Cathodes H. Zhang, J. Hou, X. Lü, Y. Lai, andJ. Li Influence of Cobalt Additions on Electrochemical Behaviour of Ni-Fe-Based Anodes for Aluminium Electrowinning V. Singleton, B. Welch, andM. Skyllas-Kazacos Effects of the Additive Zr0 2 on Properties of Nickel Ferrite Cermet Inert Anode X. Zhang, G. Yao, Y. Liu, J. Ma, andZ. Zhang

1111 1117

1123 1129

Effect of Sintering Atmosphere on Phase Composition and Mechanical Property of 5Cu/(10NiO-NiFe2O4) Cermet Anodes for Aluminum Electrolysis 1135 Z Zou, C. Wei, Z Tian, K. Liu, H. Zhang, Y. Lai, andJ. Li

Poster Session - Electrode Influence of Ultrafine Powder on the Properties of Carbon Anode Used in Aluminum Electrolysis X. Jin, D. Songyun, L. Jie, L. Yanqing, andL. Yexiang

1143

Preparation NiFe 2 0 4 Matrix Inert Anode Used in Aluminum Electrolysis by Adding Nanopowder Z Zhang, G. Yao, Y. Liu, andX Zhang

1149

Cold Water Model Simulation of Aluminum Liquid Fluctuations Induced by Anodic Gas in New Tape of Cathode Structure Aluminum Electrolytic Cell 1155 Y. Liu, T. Zhang, Z Dou, H. Wang, G. Lv, Q. Zhao, N. Feng, andJ. He Effects of Physical Properties of Anode Raw Materials on the Paste Compaction Behavior K. Azari, H. Ammar, H. Alamdari, D. Picard, M. Fafard, andD. Ziegler

xvii

1161

Furnace Efficiency - Energy and Throughput Session I Furnaces Designed for Fuel Efficiency D. White

1169

Latest Trends in Post Consumer and Light Gauge Scrap Processing to include Problematic Materials such as UBC, Edge Trimming and Loose Swarf 1173 F. Niedermair, and G. Wimroither Investigation of Heat Transfer Conditions in a Reverberatory Melting Furnace by Numerical Modeling A. Buchholz, andJ. Rodseth

1179

Oxyfuel Optimization using CFD Modeling T. Niehoff, and S. Viyyuri

1185

Operational Efficiency Improvements Resulting from Monitoring and Trim of Industrial Combustion Systems J. Oakes, andD. Bratcher

1189

New Technology for Electromagnetic Stirring of Aluminum Reverberatory Furnaces J. Herbert, and A. Peel

1193

Evaluation of Effects of Stirring in a Melting Furnace for Aluminum K. Matsuzaki, T. Shimizu, Y. Murakoshi, andK. Takahashi

1199

Business Analysis of Total Refractory Costs C. Belt

1205

Improved Furnace Efficiency through the Use of Refractory Materials J. Hemrick, A. Rodrigues-Schroer, D. Colavito, andJ. Smith

1211

Study on the Energy-saving Technology of Chinese Shaft Calciners G. Lang, C. Bao, S. Gao, R. Logan, Y. Li, andJ. Wu

1217

Author Index

1221

Subject Index

1227

xvin

PREFACE As editor it is my pleasure to present to you these contributed proceedings of TMS's 140th Annual Meeting and Exposition in San Diego, California. The volumes included here represent a large collective undertaking. All of it has been volunteered by the hundreds of authors, dozens of Session Chairpersons, and more than a dozen Symposium Chairpersons and Vice-Chairpersons that have created Light Metals 2011. As members we owe them all a debt of gratitude for the time and effort that they have donated. Although our industry has faced trying times in recent years contributions to this volume from industry and academia have been generous. The authors of these technical papers represent approximately fifty universities. There are almost twice that number of contributions from private industry, individual contributors, and research institutes combined. Many of these individuals or organizations have prepared technical papers on varied topics and across multiple symposia. Such support from our membership is what makes our annual meetings both productive and successful. This represents the best of TMS, a professional, diverse, and growing organization that embraces and promotes both pure and applied sciences. The volumes of Light Metals represent a large fraction of the accumulated knowledge of our industry that is in the public domain. It is often used as a primary source of reference information in the preparation of new contributions to the technical literature. Each year the wealth of information in the accumulated volumes of Light Metals grows and 2011 is no exception. Yet, it is not enough to rest on these laurels. Our future is being shaped now by forces that our industry could not have anticipated even a decade ago. We look to grow, to include academic and industrial papers from countries, universities, and enterprises that have yet to be represented in Light Metals along with those from more established contributors. We hope to allow future authors to see further, if not from standing upon the shoulders of the giants that have preceded them in our industry. I encourage our members to not only participate in annual meetings, but also to get actively involved. TMS committees are all composed of volunteers. Authors that have contributed in the past are most likely to contribute again. However, they, as I, would like to hear from members who may have been tempted to write a technical paper but never have done so. The strength of our organization is built upon new ideas and insights that come from all quarters of academia, research groups, and industry. New authors are always welcomed. On behalf of the organizers for Light Metals 2011 allow me to thank the TMS staff including Maria Boots, Chris Wood, and Christina Raabe Eck and the TMS Light Metals Aluminum Committee for their support. I would also like to recognize the contributions of John A. Johnson for his guidance and for his organization of the Plenary Session celebrating 125 Years of the Hall-Héroult Aluminum Reduction Process. I especially would like to recognize the 2011 Subject Chairpersons: Mohammed Mahmood, Abdullah Habib Ahmed Ali, Dr. Alan Tomsett, Dr. James Metson, Dr. Geoffrey Brooks, Kai Karhausen and Thomas Nieoff for their dedication and leadership in preparation for our 2011 Annual Meeting. Stephen J. Lindsay

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EDITOR'S BIOGRAPHY

STEPHEN J. LINDSAY LIGHT METALS 2011 EDITOR Stephen Joseph Lindsay holds a B.S. in Chemical Engineering from Clarkson College of Technology and an M.A. in Applied Behavioral Science from Bastyr University's Leadership Institute of Seattle program. During his time with Alcoa he has held numerous positions with responsibilities in anode, cathode, pollution control systems, and reduction technology. He has specialized in areas including emissions control, metal purity, alumina and electrolytes. In these areas he supports Alcoa's Primary Products division worldwide. His wife, Dr. Margarita Merino de Lindsay, an author, poet, and artist in her own right is his muse. She has encouraged Steve to contribute regularly to technical literature and education, plus control of pollution. It is for her sake that the colors of Light Metals 2011 are those of Spain, her home country. A member of TMS since 1985, Steve has regularly authored or co-authored in Light Metals. He has received the Best Paper Award in Reduction Technology in 2006 and again in 2009. He served as the Subject Chair for Reduction Technology in 2006, has instructed in various short courses, and has served under the direction of Dr. Halvor Kvande in the TMS Industrial Electrolysis Courses held since 2005. He has also authored or co-authored technical papers appear in the proceedings of the 8th and 9th Australasian Smelting Technology Conferences, the 7th and 8th International Alumina Quality Workshops, the International Committee for Study of Bauxite, Alumina, & Aluminium 2010, and the International Beryllium Research Conference 2007. Steve served on the Aluminum Association's Industrial Hygiene sub-committee for beryllium contributing to the understanding its mass balance in aluminum smelters. He has participated as an instructor on a regular basis in courses organized by the University of Auckland's Light Metals Research Centre, the University of New South Wales, and Alcoa's own Process Engineering Training Program. Steve is based at Alcoa's Tennessee Operations near Knoxville, Tennessee and works with Alcoa's Technology, Innovation and Center of Excellence group. His is a manager in Primary Metal's Best Practices group.

xxi

PROGRAM ORGANIZERS ALUMINA and BAUXITE Jim Metson graduated with PhD in Chemistry from Victoria University of Wellington, New Zealand, before taking up a position at Surface Science Western, University of Western Ontario Canada. He then moved to the University of Auckland, New Zealand, where he is a Professor, the Associate Director of the Light Metals Research Centre and Head of the Department of Chemistry. He is a Director of the New Synchrotron Group Ltd, a councillor of the Australian Institute of Nuclear Science and Engineering and chairs the Research Infrastructure Advisory Group (RIAG) for the New Zealand Government. His research interests are in materials and particularly surface science, with an emphasis on applications in the aluminium industry including alumina calcination and evolution of microstructure, smelting technology and in particular the impacts of alumina properties, and the surface science of aluminium metal. He has had more than 20 years of engagement with the aluminium industry and has been a regular participant at the Annual TMS meeting. He is a past Light Metals Award winner and has co-ordinated a course "Alumina from a Smelter Perspective" held as part of the 2004 TMS meeting and was a presenter in the 2009 course "Alumina Refinery Fundamentals and Practice". Carlos Suarez has been associated with the alumina and bauxite industry for 30 years and has been a member of TMS since 1984. Carlos attended the University of Oklahoma where he obtained a degree of Science in Chemical Engineering. He also attended the University of Phoenix where he obtained a Master in Business Administration. Carlos has been involved in all aspects of alumina refining for employers such as Bauxilum, Nabalco, Vialco and Gramercy Alumina in the areas of Process Safety, Quality, Training and Development, Technical Sales, Plant Operations Research and Development, Commissioning and Start-Ups, Knowledge Management, Organizational Development, Technology Transfer and Business Development. He has been a Process Consultant for Hatch since 2004 where he has served as process lead and project manager for different alumina plant projects around the world. Carlos has been an active member of TMS. He has contributed with several technical papers and was one of the instructors for the first Alumina Refinery Fundamentals and Practice course sponsored by TMS in 2008.

xxiii

ALUMINUM REDUCTION TECHNOLOGY Mohammed Mahmood holds Master degree in Process Engineering from Strathclyde University in Scotland in 1989. He began his career with Aluminium Bahrain (ALBA) more than thirty years ago, rose through the ranks to various managerial positions, from Manager of Potlines, Manager Process & Quality Control to Manager Human Resources & Development and then to General Manager Metal Production from 2004 - 2009 and finally in 2009 to his present position as Chief Operating Officer. Among the major milestone in his career has been the retrofitting of pot lines 1 -3 that increased the production by 21 %, lead the team to further improvement and achieve 2.7% higher productivity and improve pot operation age by 16%. Being a prominent figure in Bahrain, Mohamed is very often invited to speak at International Conferences both Technical and People Development related. He is the head of the Alba Community Service Committee where his role encouraged the spirit of philanthropy amongst Alba employees and enhanced kingdom wide appreciation of Alba's corporate social responsibility initiatives. His main passion is the development of youth to become future leaders. Abdulla Habib Ahmed, Manager Research & Development in Aluminum Bahrain (Alba), joined Alba in March of 1995 after completing his degree in Chemical Engineering with first honor class as Process Engineer. Abdulla was involved in many projects and studies to maximize Aluminum Production in Alba. He gradually climbed the success ladder of Alba hierarchy to become in charge of Metal Production as Reduction Line Superintended in year 2000. On November 2004, Abdulla completed his Master degree in University of New South Wales in Australia with first honor class. In July, 2007, Abdulla embarked on doing his Ph.D. in the same University to be the first Bahraini doing the Ph.D. in Aluminium Smelting technology and among few people in Middle East specialized on the Aluminum field. In September, 2009 he has been promoted to become first R&D Manager in Alba. Abdulla is looking after the innovations; process improvements in Reduction, Carbon and Casthouse in Aluminum Bahrain (Alba). Charles "Mark" Read is Bechtel Senior Specialist - Primary Aluminium Processes. Mark is currently Bechtel's Area Manager, Reduction for the Ma'aden "Ras Az Zawr Aluminium Smelter Project, Kingdom of Saudi Arabia. Previous Bechtel roles included Engineering Manager for green-field and brown-field aluminium smelter projects, and technology and engineering oversight of studies for major Middle Eastern, North American and Russian aluminium Smelters. Mark has 33 years experience in business and technology management in the Metals Industry, over 25 years of which were in the aluminium industry including in-depth technical experience of Hall-Héroult cell design and operation, pre-baked carbon products processing and performance, and aluminium casting operations. Mark joined Bechtel's Montreal-based "Aluminium Centre of Excellence" in late 2003. Prior to joining Bechtel, Mark held various technology management positions with Elkem Metals, Kaiser Aluminium & Chemical Corporation and Alcan Inc. Mark is a graduate of Sheffield Hallam University, England. He holds a B.Sc. degree in Metallurgical Engineering and M.Sc. in Industrial Metallurgy.

xxiv

CAST SHOP for ALUMINUM PRODUCTION Geoffrey Brooks, B.Eng. (RMIT), B.A. (SUT), PhD (Melb.) F. LEng. Aust, has been a Professor in the Faculty of Engineering and Industrial Sciences at Swinburne University of Technology since 2006, where he leads the High Temperature Processing research group. He also the leader of a cluster of researchers from Australian and New Zealand Universities focussed on improving Aluminium smelting. Previously, he was a Senior Principal Research Scientist at CSIRO (2004-2006), an Associate Professor in Materials Science and Engineering at McMaster University (2000-2004) and a Senior Lecturer at the University of Wollongong (1993-2000). In the 17 years since completing his PhD at University of Melbourne, he have published over 100 papers and run many large research projects with funding from many major companies and government agencies. He is currently active in work on dross formation in aluminium processing, controlling minor elements in the casthouse, sidewall materials in aluminium cells, development of sensors for bubbling in high temperature operations, modelling of injection processes and distribution of elements in magnesium production. He has been a key reader for Metallurgical and Materials Transactions since 1998 and is a Fellow of the Institute of Engineers (Australia). Geoff has been a member of the TMS since 1990. Dr. John Grandfield is director of Grandfield Technology Pty Ltd, a consulting and technology firm. John has a Bachelor of Applied Science in Metallurgy (RMIT), a MSC in Mathematical Modelling (Monash University) and a PhD in Materials Science (University of Queensland). John has 25 years experience in light metals cast house research in industry and government laboratories (Rio Tinto Alcan, CASTcrc and CSIRO). He has developed new technology for aluminium and magnesium DC casting, and open mould conveyor ingot casting. He conducts problem solving and research projects, presents cast house technology training courses around the world, participates in in-house innovation workshops and conducts R&D program reviews. John has four patents and has published more than 50 conference and journal papers. He is chair of the Australasian Aluminium Casthouse Technology conference.

xxv

ELECTRODE TECHNOLOGY for ALUMINUM PRODUCTION Alan Tomsett has over twenty years experience in carbon anode and cathode technology. He received his BSc and PhD in Chemical Engineering from the University of New South Wales in Sydney, Australia. He joined Rio Tinto Alcan at the R&D centre in Melbourne in 1987. His activities with R&D Group have included leadership of the global and regional Carbon R&D program, provision of technical support for the RTA Australasian smelters, carbon raw material evaluation and carbon plant technology selection for brownfield and greenfield expansions. Since 2008, Alan has been the Technical Manager - Carbon for Rio Tinto Alcan Primary Metal Pacific. Alan has been a member of TMS since 1996. He is the coauthor of several TMS papers and is a previous TMS session chair. He is also a regular contributor to the Australasian Smelting Conference. Barry Sadler has been involved in the Aluminium Industry for more than 25 years in a range of positions but always focusing on anode carbon technology. His career started in 1982 at the Comalco (Now Rio Tinto Alcan) Research Centre in Melbourne, Australia. In 1989 he moved to Comalco's New Zealand Aluminium Smelter as Carbon Plant Manager. After a stint as General Manager Organisational Effectiveness for Hamersley Iron, in 1989 Barry took up the position of Technical General Manager at Comalco Aluminium's corporate headquarters in Brisbane, Australia. Leaving Rio Tinto/Comalco in 2002 to establish Net Carbon Consulting Pty Ltd, Barry now provides consulting advice, training, and support to clients on improving plant performance, with emphasis on the practical application of statistical thinking to process management. Barry has been a regular contributor at TMS meetings for over 20 years as an author, session chairperson, and Electrodes subject organiser.

xxvi

ALUMINIUM ROLLING Kai Friedrich Karhausen is department manager for process technology at the central Rolled Products R&D of Hydro Aluminium in Bonn, Germany. Dr. Karhausen earned his doctorate at the RWTH Aachen and worked in the industrial aluminum research for 15 years both in Norway and Germany. His principal work is focused on the modeling and optimization of materials behavior in industrial production processes. Dr. Karhausen has issued 75 scientific presentations and publications. In 2003 he was awarded the Georg-Sachs-Preis of the German Materials Society (DGM) for important achievements in the field of integrated modeling of metal forming and materials behavior.

xxvn

FURNACE EFFICIENCY ENERGY and THROUGHPUT Thomas Niehoff currently Head of Non Ferrous and Mining at The Linde Group, Div. Linde Gas is based in Munich, Germany. Graduated from RWTH Aachen in Germany in mechanical engineering in 1992. Thomas has 18 years experience in combustion and metallurgical applications related to industrial gases. In his global role Thomas now overlooks the R&D activities for Linde Gas. He has in depth experience in metallurgy of aluminum, iron and steel - combustion processes and emissions from combustion. He did his PhD at RWTH Aachen on coke fired cupola process optimization with oxyfuel.

xxvni

ALUMINUM COMMITTEE 2011-2012 Chairperson John A. Johnson Johnson's Consulting Group Krasnoyarsk, Russia

Light Metals Division Chairperson John N. Hryn Argonne National Laboratory Illinois, USA

Vice Chairperson Stephen J. Lindsay Alcoa Inc. Tennessee, USA

JOM Advisor Pierre P. Homsi Rio Tinto Alcan

Past Chairperson Geoffrey Paul Bearne Rio Tinto Alcan Victoria, Australia

Secretary Charles Mark Read Bechtel Corp. Quebec, Canada

MEMBERS THROUGH 2012 Hussain H. Alali Retired, Aluminum Bahrain Manama, Bahrain

Stephen J. Lindsay Alcoa Inc. Alcoa, Tennessee, USA

Martin Iffert Trimet Aluminum AG Essen, Germany

MEMBERS THROUGH 2013 Gilles Dufour Alcoa Canada Quebec, Canada

Everett Phillips Nalco Company Illinois, USA

Pierre Le Brun Alcan Voreppe Research Center Voreppe Cedex, France

Barry Sadler Net Carbon Consulting Pty. Ltd. Kangaroo Ground, Australia

MEMBERS THROUGH 2014 John G rand field Granfield Technology Pty. Ltd. Victoria, Australia

Ketil A. Rye Alcoa Mosjoen Mosjoen, Norway

Charles Mark Read Bechtel Corp. Quebec, Canada

Carlos Suarez Hatch Associates Inc. Pennsylvania, USA

xxix

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINA and BAUXITE

ORGANIZERS

James Metson University of Auckland Auckland, New Zealand Carlos Suarez Hatch Associates Inc. Pittsburgh, Pennsylvania, USA

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINA and BAUXITE

Bauxite Resources and Utilisation SESSION CHAIR

Shawn Kostelak Gramercy Alumina Louisiana, USA

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

NEW DEVELOPMENT MODEL FOR BAUXITE DEPOSITS Peter-Hans ter Weer1 'TWS Services and Advice, Imkerweg 5, 1272 EB Huizen, The Netherlands; [email protected] Keywords: Bauxite, Alumina, Project Development, Alumina Technology, Economics

1.

comprises a string of equipment which together performs the desired process step, e.g. digestion with feed tank, heat exchangers, pumps, digester vessel(s), flash vessels, etc. Such a string of equipment is often referred to as a "train", "unit" or "circuit" (e.g. digestion unit, precipitation train, mill circuit). Alumina refinery design generally takes the digestion area as plant bottleneck due to its high unit capital cost and its requirement for constant flow for optimum performance. The design / initial refinery production capacity of greenfield projects has evolved over time from about 0.5-1.0 Mt/y alumina 25-30 years ago (e.g. Worsley, Alumar, Aughinish) to 1.4-3.3 Mt/y alumina for more recently constructed and future planned projects (e.g. Lanjigarh, Yarwun, Utkal, GAC). Figure 1 illustrates this trend.

Abstract

Developing a greenfield bauxite deposit nowadays generally includes constructing an alumina refinery. Economics have resulted in ever-increasing production capacities for recently-built and future planned greenfield refineries. Rationale: economy of scale. As a result the complexity of a greenfield project has significantly increased and its capital cost has grown to several billion USD. Important consequences: • Project owners aim at risk reduction through project financing and formation of joint ventures, further complicating project implementation. • Globally only a limited number of (large) companies have the human andfinancialresources to develop greenfield bauxite & alumina projects. • Only a limited number of engineeringfirmshave the required skills and experience to successfully implement these mega projects. • Only large bauxite deposits get developed.

Greenfield Refinery Design Capacity as function of Start-up Year j

;

♦

This paper proposes an alternative development model for bauxite deposits resulting in a more efficient use of resources and a lower threshold to develop bauxite & alumina projects.

2.

Actual

♦■

;^ !

o

Bauxite Deposit Development

♦

♦ T

Planned

♦ ♦

♦

♦

*

«

i

The development of bauxite deposits is sometimes limited to the mining of bauxite for export purposes, which may or may not include drying the bauxite to a certain moisture percentage. Examples are the Boke and Kindia mines (both in Guinea), and the Bintan mine in Indonesia (now closed). In other cases the mine supplies both a local / in-country refinery, as well as exporting bauxite, e.g. the Trombetas mine (Brazil), and the Gove and Weipa mines (both in Australia). In most recent cases the projected greenfield development of a bauxite deposit includes directly or indirectly the construction of a captive alumina refinery. Examples: Utkal (India), GAC (Guinea), Aurukun (Australia), CAP (Brasil), Ma'aden (Saudi Arabia). In some cases the project may be executed in two stages: a first stage of establishing the bauxite mine with (temporary) export of bauxite, and a second stage including the construction of an alumina refinery. A recent example is the Darling Range project of Bauxite Resources Ltd in Australia as stated in press releases. How have greenfield production capacities and more specifically greenfield alumina refinery design capacities developed over time, and did this have a bearing on project implementation?

3.

♦

' 1990

' 2000

i

1

2010

Start-up Year

Figure 1 - Refinery Design Capacity vs Start-up Year Note that actual refinery production capacities increase over time as a result of de-bottlenecking, improved process efficiencies and operations performance, etc. In a paper presented at the ICSOB A 2008 conference [1] R. den Hond even suggests a doubling of design capacity by exploiting overdesign and post start-up installation of novel technology. What has been the rationale for this trend of ever-increasing design production capacities for recently built and future planned greenfield refineries and what are its consequences? 3.2 Economy of Scale The rationale offered for this trend is the economy of scale: an increased alumina production capacity improves the economics (NPV, IRR, VIR1) of a greenfield bauxite and alumina project2. In the context of alumina refinery projects, economy of scale aspects may be applied to Operating Cost and Capital Cost.

Alumina Refinery Capacity Evolution

3.1 Overview An alumina refinery consists of a number of unit operations such as grinding, digestion, evaporation, etc. A unit operation generally

1 NPV=Net Present Value; IRR=lnternal Rate of Return; VIR=Value over Investment (capital efficiency) ratio. 2 Reference [2] provides an overview of bauxite & alumina project economics.

5

3.2.1 Effect on Operating Cost3 To better assess the effects of the economy of scale on Operating Cost, we should consider its major components: • Variable costs: In $/year these costs vary with plant production, at least within certain plant production rates (typically ± 10-15%), examples: bauxite, caustic soda, coal, fuel oil, lime. The overall plant on-line time of an alumina refinery with more than one train / unit / circuit, e.g. a digestion train, is higher than a plant with one train only, as a result of moreflexibilityin equipment operation and maintenance. The effect on plant on-line time is generally limited (indie. 0.2-0.5% abs), however may vary widely and in a specific case could be significant (>1% abs). As a result the plant operates with less interruptions and operating efficiencies (e.g. bauxite, caustic soda, energy consumption) improve, albeit generally to a limited extent (indie. 0.5-3%). • Fixed costs: In $/year these costs do not vary with plant production, at least within certain plant production rates (typically ± 100,000 t/yr), examples: labour, maintenance materials, administration, other fixed costs. This is the area on which the economy of scale potentially has the largest effect, i.e. a drop in cost per tonne of alumina produced, due to the "dilution" of "fixed" annual expenses by a larger production volume. This applies particularly to labour and otherfixedcosts. If the increase in production capacity includes an increase in the number of trains, this positive effect is dampened because not just the size of the equipment involved increases, but also its number. In addition, the requirements of complex and large alumina refineries may result in disproportional increases of overhead costs. The example provided in Table 1 may illustrate the above. In this example the larger refinery capacity is based on an increase in the number of operating units in several areas, resulting in a limited improvement only of thefixedcosts per tA.

1.4

3.2

Variable Costs, $/tA

85

83

Fixed Costs, $/tA

40

34

Total Operating Cost, $/tA

125

117

* Mt/y = million tonne alumina per annum

3.2.2 Effect on Capital Cost4 Economy of scale has the following main effects5 on Capital Cost: • In general larger size equipment, particularly tanks and vessels, is more cost effective per tonne alumina (tA) produced because larger tanks have a smaller surface area over volume ratio than smaller tanks, hence are cheaper in material cost per m3 stored volume. This effect is sometimes known as the "0.6 factor rule"6, and potentially represents a significant drop in capital cost per tA (note: this factor may be different for different equipment types and unit operations). Although technological improvements have resulted over time in a general increase in equipment size available for most processing equipment (vessels, tanks, pumps, mills,filters,etc), there are physical, technical and/or economic limitations to the size of all equipment. In addition, design considerations may favor in specific cases a large number of small equipment over a small number of large equipment. • Infrastructure (both shared and non-shared) costs are diluted (e.g. piperacks, water supply, power distribution), and spare equipment may be shared in case of a larger production capacity resulting in the construction of more units. Both of these result in a lower capital cost per tA produced. As an illustration: for a refinery with two digestion trains, shared facilities represent indicatively 20-25% of its capital cost (includes raw materials handling, general facilities, shared spares, etc). Here too there are limitations: both with respect to sharing of spare equipment and because capacity increases in infrastructure are required at some stage. The overall effect is a drop in capital cost per tA produced at higher production capacities. A straightforward power factor relationship between these would look like Figure 2.

Table 1 - Effect of Economy of Scale on Opex - 1 Refinery Production Capacity, Mt/y*

The conclusion from the above is that the primary effect of economy of scale on Operating Cost is on fixed costs (expressed per tA), and particularly if a capacity increase is the result of an increase in equipment size rather than equipment number.

Power factor relationship of Greenfield Refinery Capex as function of Design Capacity

Table 2 provides an example in which the capacity increase involved an increase in equipment size rather than the number of operating units, illustrating in that case a more pronounced effect onfixedcosts per tA.

Capex * (Capacity),actor

a

Table 2 - Effect of Economy of Scale on Opex - 2

Design Capacity, kt/year

Refinery Production Capacity, Mt/y*

2.8

:::::.3.3::;::

Variable Costs, $/tA

84

84

Figure 2 - Refinery Capex vs Design Capacity - Power Factor

Fixed Costs, $/tA

50

42

Total Operating Cost, $/tA

134

126

In many cases however plant (and thus project) capacity increases are a combination of increases in equipment size and in equipment numbers (e.g. as a result of an increase in operational

* Mt/y = million tonne alumina per annum

4

Reference [4] provides an overview of Capital Cost A second-order effect is an increased plant on-line time as a result of a plant consisting of more than one train resulting in a slightly lower capex per annual tA. 6 Theoretically the factor is 0.67. 5

3

Reference [3] provides an overview of Operating Cost

6

units / trains). In addition, an increased project scope also adds (at some stage disproportionally) to its complexity. As a result, actual capital cost per tA produced may deviate from a smooth curve as shown in Figure 2. In fact Canbäck and others [5] refer to Bain who found in a study of twenty industries that at the plant level, beyond a minimum optimum scale few additional economies of scale can be exploited. Available information suggests for the alumina industry that with respect to the relationship of refinery capital cost and design capacity, a differentiation can be made in two design capacity ranges as illustrated in Figure 3: • Up to about 1.5 Mt/y: a power factor of -0.7. • Above about 1.5 Mt/y: a power factor of -0.9.

Table 3 - Effect of Capacity on Overall Project Economics Refinery Capacity

Capital Cost*, M$ Mine 115 Refinery 1,635 Infrastructure 500 (railway, port, town) Total Capital Cost*, M$ 2,250 $/AnntA 1,500

Greenfield Refinery Capex as function of Design Capacity

3 Mt/y

L5Mt/y

200 3,000 680 1,293

3,880

Operating Cost, $/tA (incl. Infrastructure opex)

137

125

Sustaining Capital, $/tA

8

8

Economics* (indie.) NPV(8%), M$ IRR, % Payback period, y

-139 7 10.5

369 9 9

* Basis W Europe, Mid 2010 US$ Alumina price at 325 $/tA

#

«ex -s- Capacity""0!7

Table 3 shows that, despite the Refinery capex per annual tA for the two options following the trend illustrated in Figure 3, the overall project economics flip from a significant negative NPV (with IRR 7% and payback period 10.5 years) to a significant positive NPV (with IRR 9% and payback period 8.5 years). A major contributor is the disproportional increase in $/tA of the Infrastructure capex. To underpin that: had the delta in capital cost between the two project options expressed in $/Annual tA remained unchanged from the delta between the two refineries, the economics of the 1.5 Mt/year project (in that case at a total capex of 1,383 $/AnntA) would have looked as follows: NPV(8%) = -12 M$; IRR = 8%; Payback period = 9.5 years. On re-considering the trend shown in Figure 1, the reasoning could be turned around: a disproportionate increase in project scale is required to result in acceptable economics. In a similar context, A. Kjar in his paper presented at the TMS 2010 Annual Meeting [6] discusses in general terms the uncompetitive capital cost of recent Western-developed greenfield alumina projects as a result of (among other reasons) large project size and increased project complexity.

Capex * Capacity 0 β

appro«. 1.5 M t / y

Design Capacity, kt/year

Figure 3 - Refinery Capex vs Design Capacity From Figure 3 it would appear that although further gains in capital cost per tA are possible at design capacities above -1.5 Mt/y, these will be limited. A design capacity of about 1.5 Mt/y for an alumina refinery might perhaps be the "minimum optimum scale" referred to by Canbäck. Note that 1.5 Mt/y is meant to be indicative only. This raises the question how this result can be reconciled with the design capacity of some future planned projects which are well above 1.5 Mt/y (refer Figure 1).

3.3 Consequences The indicated increase in the design / initial capacity of greenfield (bauxite mine and) alumina refinery projects over the past decades has had the following major consequences: • The complexity of these mega projects7 has increased significantly, especially in terms of project planning and management. Significant infrastructural works are often required, involving extensive government involvement, adding to project complexity. • Project capital cost has grown to several billion USD, and project owners reduce risk through projectfinancingand the formation of multi-party joint ventures. This is perfectly reasonable, however it complicates project implementation (e.g. with respect to decision making processes). • Due to thefinancialcommitments involved, globally only a limited number of (very) large companies have the financial and human resources to develop greenfield bauxite & alumina projects.

3.2.3 Infrastructure Costs & Overall Economics The explanation for the above result is that greenfield projects have infrastructural requirements which may include access roads and bridges, a railway line, port facilities, and employee living facilities. In case of extensive infrastructural requirements, the related capital cost is significant and has a disproportional bearing on the economics of a smaller capacity greenfield project. An example may illustrate the above for two greenfield project options at the same location: option 1 at 1.5 Mt/year alumina production design capacity, and option 2 at 3 Mt/year. Assumed infrastructural requirements for this location: • 100 km railway line. • Jetty and wharf, and ship loading/unloading facilities at the alumina export port. • Employee housing and living facilities. Table 3 provides indicative numbers for capital, operating and sustaining capital costs for the two options considered in this example and their economics.

7

7

Typically projects over 1 billion US$.

•

•

Table 4 - Greenfield 1.5 Mt/y Aa Refinery Capital Cost (typ.)

For the same reasons (project scope, complexity), only a limited number of engineering firms have the required engineering, construction and project management skills and experience to successfully implement these projects. Typically a project life of 30+ years is (implicitly) applied to justify the significant investment of a greenfield bauxite & alumina project. Reason: an alumina refinery can operate effectively for decades (refer e.g. Paranam, Gove, Kwinana, QAL). For greenfield bauxite & alumina projects with a captive refinery this means that the bauxite deposit on which a project is based should be able to sustain refining operations for such a period. Therefore only (very) large bauxite deposits are developed, indicatively 200-300 Mt and more.

Cost Item Direct Costs Equipment* Commodities* Total Direct Costs, M$ Indirect Costs Freight EPCM Temp. Construction, start-up, Commissioning, etc Owner's Engineering & Other Costs Total Indirect Costs, M$ Contingency, M$

In summary, worldwide only a small number of companies develop mostly very large greenfield bauxite and alumina projects, which often take a decade and more to develop.

Total Refinery Capital Cost*, M$

231 539

770

78 256 180 190 704 161 1,635

* Incl. steam & power generation, sub stations, residue disposal, water supply, communication & info systems # Incl. concrete, steel, mechanical bulks, piping, wire and cable, etc & Basis W Europe, Mid 2010 US$

3.4 Where from here? With an objective to lower the threshold for the development of bauxite and alumina projects, the question may be asked if the underlying trend, viz. ever-increasing alumina refinery design capacities, is inevitable, or if viable alternatives exists. The basic reason for the trend being economics (refer section 3.2), the question could be reformulated as follows: is it possible to develop smaller greenfield bauxite and alumina projects at acceptable economics? A. Kjar addresses this question and some of the issues discussed above, albeitfroma different perspective, in his earlier mentioned paper. He indicates that as a means to overcome some of these issues, attempts were made by others: 1. To gain improved control over the project execution process; and 2. To increase the level of pre-assembly to reduce total costs of on-site construction labor and low productivity - refer also a paper by R. Valenti and P. Ho [7]. A. Kjar proposes the use of replication of a modern plant design, and small increments of capacity (without quantifying a capacity), in order to quickly and more cost-effectively build a large plant / project. Although A. Kjar's paper has a different angle (viz. building a large plant at lower capital cost), there are overlaps with the subject of the current paper (investigating the possibility to lower the threshold for the development of- smaller - bauxite and alumina projects). To further explore the subject, a more in-depth look at the makeup of a greenfield project's capital cost is required.

4.

1.5 Mt/y

4.1.2 Commodities and Plant Layout Aspects Table 4 illustrates that the Commodities represent a very significant element in the refinery capital cost. Commodity amounts and their related capital costs reflect plant design including plant layout. Current alumina refinery layouts are designed to accommodate additional (future) digestion units (and all of the other required process units - e.g. precipitation, evaporation). The consequence is that plant design is not optimized for its initial production capacity. Plant layout is characterized by an "open architecture", at best compromising between on the one hand the limited layout requirements for the initial / design capacity and on the other hand the more extensive requirements to accommodate future additional process units. And in the worst case consisting of a layout of a large-capacity plant of which part is built, resulting in an inefficient plant layout for the design / initial capacity. In addition, in some cases plant design includes equipment which at some future stage might be used to its full capacity, but operates (well) below design for a considerable part of its lifetime. 4.1.3 Alternative Approach - Dedicated Plant Capacity A. Kjar's proposal to use replication means that a design is developed for a dedicated production capacity. Or putting it differently, this alternative design approach aims at designing an alumina refinery for a dedicated production capacity, i.e. without provisions for future expansions. This approach enables optimizing plant layout for the targeted production capacity, e.g. with respect to positioning similar equipment close to each other, use of common spares, etc. This more "closed" layout architecture results in a more efficient plant layout, reflected for example in the design of main plant piperacks. This is illustrated in Figure 4 which shows the main piperack layout for a typical (current design) 1.5 Mt/year capacity refinery ( i.e. in the expectation that additional production lines in the various areas will be added in the future), and the layout for a dedicated 1.5 Mt/y capacity alumina refinery (same scale). The alternative approach with its more closed layout design impacts positively on commodity volumes: for the same production capacity, commodity volumes for a greenfield plant designed along this alternative approach are similar to that of a brownfield expansion of an existing refinery. This is illustrated in Figure 5 which shows the total length of piping of greenfield and

Capital Cost Make-up

4.1 Refinery Capital Cost 4.1.1 Overview The capital cost of a greenfield alumina refinery may be split up as shown in Table 4. In this table typical numbers are shown for a low-temperature digestion alumina refinery with a 1.5 Mt/y production capacity. Note that actual numbers may deviate significantly as a result of bauxite quality, technology choices, plant location, etc.

8

reflected in lower Commodities costs, resulting in lower Direct Capital Costs, in turn lowering Indirect Capital Costs. The overall effect on the capital cost of a greenfield dedicated lowtemperature digestion alumina refinery of 1.5 Mt/y is illustrated in Table 5 (indicative numbers). As can be seen in this table, the alternative approach improves the total refinery capital cost indicatively by over 10%. In fact the capital cost expressed per annual tonne of alumina capacity is lower than that of the current-design refinery at 3 Mt/y capacity (976 vs 1,000 $/Ann tA - refer Table 3).

brownfield projects as function of plant production capacity, and the requirement of a dedicated plant of 1.5 Mt/y capacity. This approach also stimulates focusing on a "lean" design and exploit any potential overdesign right from start-up (refer the comment made in section 3.1). Layout Main Piperacks of Dedicated 1.5 Mt/y Alumina Refinery

Table 5 - Comparison of Refinery Capital Costs (indie.) 1.5 Mt/y Refinery Capacity

Cost Item

Current-design

Typical Layout Main Piperacks of 1.5 Mt/y Alumina Refinery

Direct Costs Equipment* Commodities Total Direct Costs, M$

231 539

Indirect Costs Freight EPCM Temp. Constr., start-up, Comm. Owner's Eng. & Other Costs Total Indirect Costs, M$

78 256 180 190

Contingency, M$

1

Total Capital Cost*, M$

$/AnntA 1,090

Dedicated

770

224* 459 682

704

69 227 175 168 640

161

142

1,635

976

1,464

* The more efficient plant layout enables slightly lower equipment cost as a result of a more efficient use of common spare equipment Basis W Europe, Mid 2010 US$

#

4.1.5 Compact Refinery - Simple & Limited Scope Along the lines of A. Kjar's paper (although he does not quantify "small increments of capacity"), applying the proposed dedicatedcapacity approach to a compact alumina refinery capacity of 0.4 Mt/y results in a project with a simple and much more limited scope. Available data suggest that as a result some Indirect capital cost items decrease more than proportionately, particularly costs related to temporary construction and start-up support, camp and other construction related items, and owner's costs. Table 6 illustrates the capital cost for a 0.4 Mt/y alumina refinery based on a dedicated design (indicative numbers). The table shows that the capital cost per annual tonne alumina (1,295 $/AnntA) is higher than that of the much larger 1.5 Mt/y dedicated plant (976 $/AnntA - refer Table 5), however is at a level which could result in a project with acceptable economics, provided Infrastructure capital cost is limited (compare with the 1,293 $/AnntA for the overall project capital cost of a 3 Mt/y refinery - see Table 3). Table 6 also shows that the total capital cost is at a level which would enable many more (relatively small) companies to develop such a project without necessarily requiring the formation of multi-party joint ventures, simplifying overall project management and thus enabling to lower costs (effect not included in Table 6). Note that the 0.4 Mt/y refinery production capacity used here is notfixedbut is meant to typify a capacity range of ~ 0.3-0.6 Mt/y. The higher end of this range is limited by the objective to end up with a total project capital cost well below 1 billion US$, the lower end is determined by logistical limitations (e.g. with respect to caustic soda and fuel oil shipments) and may vary for different locations.

h Figure 4 - Main Piperack Layout Comparison 400,000

Total Piping as function of Plant Capacity

Greenfield

Brownfield Plant with dedicated layout design

Plant Capacity, t/y

Figure 5 - Total Piping as function of Plant Capacity 4.1.4 Effect Alternative Approach on Commodities Cost A dedicated greenfield plant design results in lower amounts (in some cases significantly lower amounts) per annual tA produced of commodities such as steel, concrete and piping. This is

9

project with a simple and limited scope, further improving capital cost per tA produced. To ensure acceptable economics, Infrastructure capital cost should be limited. At the same time such a project has few infrastructural requirements , especially if located close to an existing port.

Table 6 - 0.4 Mt/y Refinery Capital Cost (indie.) Cost Item

Direct Costs Equipment Commodities Total Direct Costs, M$ Indirect Costs Freight EPCM Temp. Constr., start-up, Comm. Owner's Eng. & Other Costs Total Indirect Costs, M$

0.4 Mt/y Refinery Capacity (dedicated design) 95 177

28 91 37 33

Contingency, M$ Total Capital Cost*, M$ $/AnntA

* Basis W Europe, Mid 2010 US$

5.2 Main Advantages The main advantages of the new development model are: • Due to the significantly smaller project capital expenditure involved (lower risk), this approach enables the development of bauxite & alumina projects by smaller companies without a need to form multi-party joint ventures, i.e. it increases the number of companies potentially interested in developing bauxite deposits. In other words competition increases, which should result in more efficient use of resources, both in terms of capital resources and in terms of global bauxite deposits. • Due to the decreased complexity of compact alumina refining projects, the number of engineering companies potentially able to develop these projects increases, again resulting in more competition and the potential for a more efficient use of resources. • Small and simple projects carry less risks and require less time to develop, implement and start-up, all of which has a positive impact on economics. • A long term alumina refining project based on the new model requires only a relatively small bauxite deposit (a deposit of ~40 Mt could support a 0.4 Mt/y project for 30 years). This means that worldwide the number of bauxite deposits that lend themselves to development increases. • The new development model may be applied also to the development of part(s) of a large deposit. • This approach may in some cases lower the threshold to increase value creation through alumina refining rather than being limited to bauxite export sales. This is attractive both to host countries and to companies developing potential bauxite & alumina projects. • In some cases, an adapted version of this new development model may enable bauxite deposit development even in locations with little existing infrastructure, albeit at a larger than compact scale (refer e.g. to Table 5 for a dedicated 1.5 Mt/year capacity project).

272

189

!

57 1,295

518

4.2 Infrastructure Capital Cost As mentioned above, in order to realise acceptable economics for a project based on a compact dedicated production capacity, Infrastructure capital cost should be limited. Conversely a project based on a compact plant capacity has very limited infrastructural requirements and has several advantages over a large plant, particularly if the project is located close to an existing port, e.g. it may be allowed closer to residential areas (i.e. is closer to existing infrastructure); the existing infrastructure may be sufficient for a small plant, but not for a big plant; a suitable location for a small residue disposal area is easier to find than for a large one, etc. As outlined in section 5.3 several such locations exist worldwide. 4.3 Refinery Technologies Note that the alternative approach proposed above is independent of the selected refinery technologies, while at the same time stimulating to focus on improvements, e.g. positioning similar equipment close to each other, the use of common spares, etc. 4.4 Replication and Indirect Costs A. Kjar indicates in his paper that the use of replication of a modern design at small capacity increments has as one of its main advantages far lower indirect capital costs, comprising Project management; Procurement; and Technology & EPCM fees. Although no direct quantification is mentioned in the paper, this appears consistent with the results discussed above for a dedicated plant design at a compact production capacity. Some of the replication-related cost savings mentioned by A. Kjar may come on top of the cost improvements indicated in this paper.

5.

5.3 Possible Locations Following are some examples of bauxite deposits that may lend themselves to development via the proposed alternative approach (between brackets the potential alumina export port): • Haden, Queensland, Australia (Brisbane). • Bindoon, Western Australia (Fremantle). • El Palmar, Venezuela (Ciudad Guayana). • Trelawny, Jamaica (Discovery Bay). • Kibi, Ghana (Tema). The above list is not exhaustive and meant to be illustrative only. In addition some bauxite deposits which in view of their size could support the current development approach with largecapacity alumina refining projects, may also lend themselves to stage-wise development through the proposed alternative

New Bauxite Deposit Development Model

5.1 New Development Model The bauxite deposit development model proposed in this paper as detailed above is based on the development of a dedicated compact alumina refinery in the range ~ 0.3-0.6 Mt/year. The dedicated refinery design has no provisions for future expansions, enabling optimizing plant layout and resulting in lower capital cost per tonne of alumina (tA) produced compared with current plant design. The compact capacity results in a

10

approach. In this case these deposits would be able to support several (smaller) greenfield bauxite and alumina projects as outlined in the last bullet point of section 5.2 above. Example: some of the Eastern Ghats deposits in Orissa and Andhra Pradesh, India, e.g. the Kutrumali deposit (with Visakhapatnam as potential alumina export port).

6.

References

1. R. den Hond, "Technology Choices for Greenfield Alumina Plants" (paper presented at ICSOB A 2008, Bhubaneswar), pp 267-270. 2. P.J.C, ter Weer, "Greenfield Dilemma - Innovation Challenges" (paper presented at Light Metals 2005, San Francisco, California), pp 17-22. 3. P.J.C, ter Weer, "Operating Cost - Issues and Opportunities" (paper presented at Light Metals 2006, San Antonio, Texas), pp 109-114. 4. P.J.C. ter Weer, "Capital Cost: To Be or Not To Be" (paper presented at Light Metals 2007, Orlando, Florida), pp 43-48. 5. S. Canbäck, P. Samouel, and D. Price, "Do Diseconomies of Scale Impact Firm Size and Performance - A Theoretical and Empirical Overview", Journal of Managerial Economics, 2006, Vol. 4, No. l,pp 27-70. 6. Anthony Kjar, "A Case for Replication of Alumina Plants" (paper presented at Light Metals 2010, Seattle, Washington), pp 183-190. 7. R. Valenti and P. Ho, "Rio Tinto Alcan Gove G3 Experience on Pre-Assembled Modules" (paper presented at the Alumina Quality Workshop 2008, Darwin), pp 1-5. For further information, please contact P.J.C, ter Weer at [email protected] or visit www.twsservices.eu.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

STUDY ON THE CHARACTERIZATION OF MARGINAL BAUXITE FROM PARA/BRAZIL Fernanda A.N.G. Silva1'2, Joäo A. Sampaio2, Francisco M. S. Garrido1, Marta. E. Medeiros1 universidade Federal do Rio de Janeiro, Instituto de Quimica, Avenida Athos da Silveira Ramos, 149, Cidade Universitaria; Rio de Janeiro, RJ, 21941-909, Brasil. 2 Centro de Tecnologia Mineral / CETEM-MCT, Avenida Pedro Calmon, 900, Cidade Universitaria; Rio de Janeiro, RJ, 21941-908, Brasil. Keywords: Marginal Bauxite, Mineralogical and Chemical Characterization In this context, the aim of this work was to: ore dress, provide the chemical, structural and mineralogical characterization of crystallized amorphous bauxite (CAB) and compare its behavior and characteristic with the bauxite processed by the Bayer process, such as crystallized bauxite (CB). The techniques applied in this study were: chemical analysis by potentiometric titration and flame atomic absorption spectroscopy, X-ray fluorescence, Xray diffraction, infrared spectroscopy, scanning electron microscopy and thermal analysis.

Abstract Bauxite from Para is divided into five different layers. However, only one is processed. The crystallized-amorphous (CAB) phase is considered a marginal bauxite because it presents a high quantity of Si02reactive and its use depends on special technologies. CAB was characterized and the results were compared with the bauxite used nowadays in the alumina plant. Characterization was performed by XRD, IR, XRF, chemical analysis, TGA and SEM. XRD determined the mineral content: such bauxite is gibbsitic and has been associated with kaolinite and hematite. IR data supported the XRD results. XRF was used to determined the sample's chemical composition. The chemical content of Al203avaiiabie and Si02reactive w a s determined by potentiometric titration and FAAS. The results found for the Bayer process sample were 41.7% and 7.1%, respectively. TGA observed the bauxite decomposition and SEM supplied chemical and thermal analysis. Thus, based on stoichiometric relations of the bauxite components decomposition, it was possible to confirm the presence of the following phases: gibbsite and kaolinite.

I

I

j

j Clay layer (Laterite Soil)

Nodular bauxite

Introduction Crystallized Nodular Bauxite

Bauxites are usually considered to be of two major types: (1) lateritic (sometimes called equatorial) and (2) karst, both being weathered products from the underlying parent rock: lateritic bauxites. [1] Lateritic bauxites, formed in equatorial climates, comprise 90% of the world's exploitable bauxite reserves [1]. The weathering process has resulted in a typical profile in which the valuable aluminous material lies atop of an aluminosilicate base (often clay) and has formed from it through the leaching of silica. The main silicate mineral is kaolinite which is often associated with goethite as the iron mineral. Aluminous minerals are predominately gibbsite and to a lesser extent boehmite [1].

I

Crystallized Bauxite

I

Crystallized Amorphous Bauxite

Amorphous Bauxite

Bauxite from NE Para is divided into five layers: (1) nodular (NB), (2) crystallized nodular (CNB), (3) crystallized (CB), (4) crystallized amorphous (CAB) and (5) amorphous (AB) (Figure 1) [3]. The NB, CNB, CAB and AB layers present a high quantity of iron minerals, reactive Si0 2 and others impurities. [2-3]. These bauxites are considered marginal and their ore dressing requires special technologies arisingfromtheir content of impurities.

Figure 1. Bauxite geological profile from Northern Parâ/Brazil. Materials and Methods

Bauxite mining methods vary according to the nature of the mineralized field bodies, but in most of the cases a strip or block of bauxite is exposed and surface-mined [3]. Although the mining process is selective and only the layer of crystallized bauxite is removed, for a bauxite to be considered economically useful for the Bayer process, the available A1203 content should be between 45-55% whilst the reactive Si0 2 content should be between 4-6%. [4-6].

1 - Sample Preparation The bauxite ore was crushed in a jaw crusher and the product was classified with the use of a sieve (1.65 mm). The coarse fraction was crushed (crushing rolls) in a closed circuit. The -1.65 mm fraction was classified to remove the -37 μιη particles (sludge). The +37 urn fraction was homogenized with the product of the crushing rolls and two samples of 20 kg and 5 kg, respectively, were separated for wet granulometrie analysis. For grinding,

13

samples of 20 kg were homogenized and separated into piles of lkg. Figure 2 shows the block diagram used in the bauxite beneficiation [7].

2 - Chemical and Mineralogical Characterization The bauxite ore was submitted to chemical and mineralogical analysis with the use of: X-ray diffraction, infrared spectroscopy and X-ray fluorescence.

Samples (1 kg) of the final product of the preparation stages were wet ground in a stainless steel mill bar with 10 stainless steel bars of 20 mm diameter. The slurry in water was prepared using a bauxite solid concentration of 1000 g L"1. Grinding time varied from 0 to 30 minutes. A wet granulometrie analysis was carried out after each grinding to adjust the sample to the necessary conditions for the Bayer process [7].

During the ore dressing tests, four different samples were obtained: a work sample (WS) obtained after crushing, two work sample fractions (833 and 208 μπι) and the sludge (-37 μιη fraction obtained after desliming of the -1.65 mm fraction). These samples were submitted to the same characterization techniques of the crude sample.

The granulometrie analysis was carried out with samples of lkg, according to the damp method [8]. A vibratory sifter (684.5 rpm), equipped with a group of sieves with openings from 3.350 mm to 37 μπι, was used according to the Tyler series. All the granulometrie analysis fractions obtained in these tests were dried (100°C) and weighed.

The BP sample (the sample ground to the Bayer process requirement) was characterized with the same techniques used in the crude sample followed by potentiometric titration, flame atomic absorption spectrometry, thermal analysis and scanning electron microscopy. 2.1 - X-Rav Diffraction (XRD)

Feed Aliquot 80 kg (Archive)

«-

1 Homognization

1

Pile

1

Samples were examined by XRD in a Bruker-AXS D5005 diffractometer, with Co Ka (35 kV/40 mA) radiation, 0.02° goniometer velocity and 2Θ by path with 1 s by path counting time and data collected from 5 to 80° 2Θ.

1

2.2- Infrared Spectroscopy (IR)

Crushing (Jaw Crusher)

1

I

Fraction +1.65 mm |

Fraction -1.65 mm

2.3 - X-Rav Fluorescence (XRF)

1

*

Granulometrie -37 μπι «Analysis (Sludge) (3.32 mm to 37 μπι)

Crushing (Rolls Crusher)

1

I

Fraction f -1.65 mm |

Fraction +37 μπι

1

Infrared spectra (FTIR) was performed in a Nicolet Magna 750 Fourier transform spectrometer, from 4000 to 400 cm"1, with resolution of 4 cm'1, with sample mounting using KBr discs.

1 Homogenization

The samples were melted with lithium tetraborate at 1100°C in the proportion of 1:6 sample/fluxing agent. The melted bead was analyzed in an energy dispersive X-ray fluorescence spectrometer (Bruker-AXS model S4-Explorer), equipped with Rh tube. To obtain the semiquantitative chemical analysis, the sample spectrum was evaluated by a Spectra plus v. 1.6 software, in the standardless method mode, without a specific calibration curve.

+ r~ ON 3 |3

1

2.4 - Thermal Analysis (DTA/TGA) Thermal analysis was carried out in a Shimadzu TA-50WSI equipment (thermogravimetric analysis), DTA-50 (differential thermal analysis) in a heating grade of 10°C/min, from room temperature to 1200°C under a flow of air.

Pile

1

1 ws

1

1 (Work Sample) |

1

2.5 - Determination of Available Alumina and Reactive Silica

▼

Granulometrie Analysis (3.32 mm to 37 μπι)

[►

WS833μm WS208μm

The method to determine the amount of available alumina (the amount that will be refined to obtain A1203 in the Bayer process) and the amount of reactive silica (kaolinite), consisted of bauxite digestion in alkaline medium (NaOH) under controlled pressure and temperature, simulating the Bayer process. For the determination of available alumina, a sodium gluconate solution was added to the supernatant to form an aluminum hydroxide gluconate complex. The excess of NaOH, used in the digestion step, was neutralized with the addition of an HC1 solution. Then, a KF solution was added and back titration was carried out. Afterwards, an excess of HC1 in the standardized solution was

1

Grinding - 30 min Bayer Pulp (PB)

Figure 2. Block diagram of the stages used in the beneficiation of bauxite samples.

14

vibration [9, 12,13] and the bands near 3695 are attributed to OH stretching of kaolinite [14].

titrated with a NaOH standardized solution. The solid phase, resulting from ore digestion stage, was dissolved in an HNO3 solution [5]. The concentration of reactive silica was determined by flame atomic absorption spectrometry (FAAS).

In order to compare the mineralogical phases between CAB and CB, it is possible to observe that in CB there are more phases related to iron minerals, such as aluminum-goethite (oc(Fe,Al)OOH), goethite (FeOOH) and hematite (aFe203) [7] than in CAB. These results are associated with the distribution of the layers in the bauxite's profile since the CB layer is near the NB layer, which is composed essentially of iron minerals. For the dissolution of gibbsitic bauxite, the temperature in the Bayer process must be between 140-150°C. At this temperature the iron minerals are inert and do not react during the process.

2.6- Flame Atomic Absorption Spectrometry (FAAS). Chemical analysis was performed by flame atomic absorption spectrometry in an AA6 Varian equipment with 248.3 nm wave number, 0.5 nm slit and with air/acetylene. 2.7 - Scanning Electron Microscopy (SEM) Bauxite morphology was determined by scanning electron microscopy in high vacuum SEM (Leica/F440). The sample was embedded in resin epoxy and polished. The resin was covered again by vaporized carbon to be used as a conductor. Results and Discussion

3526

In this work, the fraction of crystallized amorphous bauxite (CAB) was characterized and the results were compared with crystallized bauxite (CB). The mineralogical phases that compose the crystallized bauxite (CB) were determined by X-ray diffraction (XRD), Figure 3. Therefore, this bauxite is essentially gibbsitic and is associated with kaolinite (Al4(SÌ4Oi0)(OH)8 and hematite (aFe203).

c

o CO

-Ω

<

9P 4000

3500

3000

— 1 — 2500

2000

1500

1000

500

Wavenumber (cm" ) Figure 4. Infrared spectra of CAB bauxite In order to determine the sample's chemical composition and how its content varied after ore dressing, XRF analysis was carried out, Table I. Table I. Chemical composition of the bauxite ore

^JuJ - 1

-

10

-r— 20

Samples 30

—r—

—1—

Bauxite Ore WS WS 833 μπι WS 208 μπι Sludge

50

40

2Θ

OGibbsite

■ Kaolinite

· Hematite

Figure 3. XRD of crystallized amorphous bauxite (CAB)

Samples

Infrared spectroscopy was used as a support to the XRD results. The infrared spectra analysis of crystallized amorphous bauxite (CAB), Figure 4, shows the bands related to the stretching vibration of Al-O-H groups observed at around 3620 cm"1. This band is related to gibbsite and kaolinite minerals [9]. The bands around 3526, 3460 and 3390 cm'1 are assigned to O-H stretching of gibbsite. The band at 914 cm'1 is due to Al-O-H group deformation vibration [10], the bands observed around 1.1001.030 cm'1 are related to the deformation vibrations of OH and are characteristic a gibbsite bauxite ore [11]. Bands appearing at around 450, 540, 660, 1100 cm'1 are assigned to the Si-O

Bauxite Ore WS WS 833 μπι WS 208 μπι Sludge

A1 2 0 3 46.06 53.10 56.54 50.70 40.94 Ti0 2 1.84 1.79 1.37 3.92 2.03

Com ponents (% in mass) MnO Si0 2 Fe 2 0 3 14.64 0.03 9.11 0.04 9.08 8.77 5.86 2.25 10.77 0.15 11.70 35.70 12.95 Com ponents (% in mass) PF Zr0 2 WO3 0.22 28.10 0.25 26.97 0.36 0.36 27.26 0.77 21.99 0.28 7.97 -

| SO3

-

|

0.13

| Total 100.00 100.00 100.00 100.00 100.00

WS: sample obtained after crushing (final product of ore dressing); Sludge: -37 μηι fraction obtained from the removal of the -1.65 mm fraction

The analysis of the results in Table 1, indicated a higher content of Si0 2 in the finer fractions than in the coarse fractions, whereas

15

the opposite behavior was observed in the content of AI2O3. The same behavior could be observed in the CB bauxite. Nevertheless, the values of the chemical content are different because CAB has a higher amount of kaolinite than CB, according to its geological profile [7]. However, in order to evaluate such behavior in relation to the contents of available alumina and reactive silica responsible for the consumption of NaOH and the formation of dessilication product (DSP) by the Bayer process, a chemical analysis by potentiometric titration and flame atomic absorption spectroscopy was carried out. These analyses were done after adjusting the CAB sample to the Bayer process.

i. The first event was related to Al(OH)3 dehydroxylation, with the formation of a AIO(OH) and χ-Α1203 mixture [7].

The grinding study showed that after a period of 20 min., the bauxite particle size distribution was similar to the optimum particle size for the Bayer process. The grinding curve, Figure 5, showed that around 90% of the particles had grain size lower than 0.208 mm and 50% under 0.043 mm.

Table II contains the data related to the samples' mass loss obtained to the Bayer process sample (BP) to both bauxites. Results analysis showed a bigger mass loss related to the first and third events. These results can be explained based on explanations i and iii [7].

ii. The second one corresponded to aFe203.nH20 dehydroxylation, since among the mineralogical phases identified through XRD, this was the only phase where it could be decomposed at such temperature [7]. iii. The third event was related to a AIO(OH) dehydroxylation combination formed in the first event and (Al4(SÌ4O10)(OH)8) dehydroxylation [7].

In order to evaluate Table II, it can be observed that the CAB bauxite had a larger mass loss related to the third event (kaolinite) when compared with the CB bauxite. An opposite behavior was observed for the mass loss related to alumina (first and third events), in which CB presented a larger mass loss than CAB. These results were in agreement with those obtained through XRD, IR and XRF, which showed that the DTA/TGA analysis could be used to observe the relationship between gibbsite and kaolinite in bauxites. 295 SC

15 20 grinding time (min)

Figure 5. Grinding curve of CAB bauxite. The time used to mill CAB was shorter than the time used to mill CB, at 20 and 30 min., respectively [5]. This result was in agreement with the results obtained from XRF, since this analysis showed that in CAB there was a larger amount of finer particles when compared with the CB. Considering that the finer fractions were composed essentially of kaolinite, according to the XRD and IR results, the time of mill was short because this mineral presented particle size smaller than 43 μπι.

365 QC —1— 200

508 QC

— i

■

400

1

600

«—

Temperature (°C)

After adjusting the CAB sample to the Bayer process requirement (20 minutes of grinding), this bauxite was characterized by thermal and chemical analysis. The chemical analysis used the following techniques: potentiometric titration and flame atomic absorption spectroscopy.

- 1 — 800

1000

1200

Figure 6. DTG/TGA of Bauxite Ore. Table II. Mass loss related to thermal events of CAB and CB samples.

Thermal analysis of the bauxite varied according to the sample's origin, impurities and crystalinity. Thus, thermal analysis (DTA/TGA) was carried out in this work as a complementary technique to XRD, IR, XRF and chemical analysis, in order to obtain a complete characterization as well as a better knowledge of the rock that composes the geological profile of NE Para [4, 5, 7].

Samples CABBP CBBP

1st

2nd

Al(OH)3 18.4 21.0

Fe 2 0 3 *

2.0 1.2

BP: sample ground to Bayer process. *To CABA aFe 2 0 3 .nH 2 0 *To CB (a(Fe,Al)OOH)

According to the results obtained through DTA/TGA for the CAB sample to the Bayer process (Figure 6) three events were observed:

16

Mass Loss (%) Events 3 rd | (AlOOH) Al4[Si4O10](OH)8 1

3.7 4.2

2.1 1.8

I

Thermal analysis results were in accordance with those obtained through the chemical analysis for available alumina and reactive silica, 41.7% and 7.1%, respectively (Table III). The ratio between A1 2 0 3 available and Si0 2 reactivem the CB bauxite was 8:1. When this ratio is below the conditions recommended by the Bayer process, 10:1, it causes a reagents and products loss during the process of bauxite dissolution. Therefore, CAB cannot be used in the Bayer process because its ratio is 6:1 and it provides a high loss to the process.

AI

7*

Table III. Chemical contents of A1 2 0 3 available and Si0 2 reactive of CAB and CB samples by potenciometric titration and atomic absorption.

Samples CABBP CBBP

A1 2 0 3

Al· -

Chemical Analysis (% in mass) available

(s<0.3%) 41.7 47.5

BP: sample ground to Bayer process.

(A) 1 1

I/7

^ 1 U 2 reactive

(s<0.1%)

7.1 5.9

>mJl·—«m

i è'

i ft a , ; ;

>!>>>··:

r—

s

■ T-

1

(B)

Scanning electron microscopy (SEM) was used to support the aforementioned analyses. Figure 7 illustrates the SEM of the Bayer process sample. This sample was divided into two different particles: pure gibbsite with smooth surface (A) and gibbsite associated with kaolinite particles with rugous surface (B).

Figure 8. Energy Dispersive Spectrum to the CABA BP sample: (A) pure gibbsite and (B) gibbsite associated with kaolinite. After the ore dressing stages, it was possible to observe that the prepared sample for the Bayer process had kaolinite particles associated with gibbsite through a physico-chemical interaction as well as in the finest fractions that comprised the sample. Conclusion The ore dressing stages aimed at the adaptation and characterization of the fractions that compose the bauxite sample of NE Para that is not currently used in the Bayer process. According to the results, the sample was adjusted to the Bayer process (BP) requirements after crushing stages, screening, removal of fine fractions and 20 minutes of grinding. XRD indicates the mineralogical composition of this bauxite. Therefore, the studied bauxite is essentially gibbsitic and is associated with the minerals kaolinite and hematite. XRF provided information related to the chemical composition of the sample, but did not inform the chemical content of A1 2 0 3 available and Si0 2 reactive making it necessary a chemical analysis by potentiometric titration and flame atomic absorption spectroscopy.

Figure 7. Scanning electron microscopy imaging of the CAB BP sample: (A) pure gibbsite and (B) gibbsite associated with kaolinite.

Thermal analysis helped us confirm the events related to dehydroxylation of the components presented in the sample: gibbsite, hematite and kaolinite.

With the use of EDS (Energy Dispersive Spectrometer), Figure 8A and 8B to gibbsite particle and gibbsite with kaolinite, respectively, it was possible to observe the semiquantitative composition of those minerals.

The contents of A1 2 0 3 available and Si0 2 reactive m the sample were 41.7 and 7.1%, respectively. The ratio between A1 2 0 3 available and Si0 2 reactive was below the quantity recommended for the Bayer process (10:1). Thus, the 6:1 rate caused a lot of loss in the process. This work has allowed us to understand why CAB bauxite is not used in the Bayer process. The reason is related to the high amount of reactive silica. This problem is likely to be associated with the presence of gibbsite particles associated with kaolinite

17

12. E. Mendelovici.2001. "Selective mechanochemical reactions on dry grinding structurally different silicates". Journal of Materials Science Letters, 20, 81-83.

particles, possibly, by a chemical surface interaction, even after the ore dressing steps. Acknowledgments

13. C.A. Azevedo, F.M.S. Garrido, M.E. Medeiros, 2006. "The effect of mechanochemical activation on the reactivity in the MgO-Al203-Si02 system", Journal of Thermal Analysis and Calorimetry, 83, 649-655.

We would like to thank CAPES and CNPq for their financial support. References

14. F.A.N.G. Silva "Estudo de Caracterizaçao e Beneficiamento do Caulim da Regiäo Borborema-Seridó (RN) (M.Sc. dissertation, PEMM/COPPE Universidade Federal do Rio de Janeiro, 2007), 51-52.

1. Smith P., 2009. The processing of high silica bauxites — Review of existing and potential processes. Hydrometallurgy, 98, 162-176. 2. B. Kotschoubey et al., "Caracterizaçao e Gênese dos Depósitos de bauxita da Provincia Bauxitifera de Paragominas, Noroeste da bacia do Grajau, Nordeste do Parâ/Oeste do Maranhäo." Caracterizaçao de Depósitos Minerais em Distritos Mineiros da Amazônia, eds. O.J. Marini, E.T. Queiroz, B.W. Ramos, (Brasilia, DF, Brasil: VGArte, 2005), Cap. 11, 691-782. 3. C. N. Barbato, S. C. A Franca, M. N Souza, "Study on the Rheological Behavior of Crystallized and CrystallizedAmorphous Bauxites", Light Metals, (2010), 87-91. 4. F.A.N.G. Silva, "Relatório de Estâgio realizado na Alunorte no periodo de 2/8/2007 A 21/9/2007" (Relatório RE2007-004-00, Centro de Tecnologia Mineral, Rio de Janeiro, RJ, 2007). 5. F.A.N.G. Silva, M. E. Medeiros, J. A. Sampaio, R. D. Santos, M. C. Carneiro, L. S. Costa, F. M. S. Garrido, "Technological Characterization of Bauxite from Parâ-Brazil", Light Metals, (2009), 139-144. 6. R. Gandi, M. Weston, M. Talavera, G.P. Brittes, E. Barbosa, "Design and Operation of the World's First Long Distance Bauxite Slurry Pipeline" Light Metals, (2008), 95-100. 7. F.A.N.G. Silva, R. D. Santos, J. A. Sampaio, F. M. S. Garrido, M. E. Medeiros, "Study on Ore Dressing and Characterization of Different Granulometrie Fractions that Compound Bauxite From Parâ/Brazil", Light Metals, (2010), 69-74. 8. J.A, Sampaio and F.A.N.G Silva, "Anâlise Granulométrica Por Peneiramento", Tratamento De Minérios Praticas Laboratoriais. ed J.A. Sampaio, S.C. Franca, and P.F.A. Braga, (Rio de Janeiro, RJ, Brasil: Centro de Tecnologia Minerai, 2007), Cap3, 55-74. 9. H. W. Marel, H. Beutelspacher, Atlas of Infrared Spectroscopy of Clay Minerals and their Admixtures, 1 ed., Amsterdam, Elsevier Scientific Publishing Company, 1976. 10. E. Mendelovici, "Selective mechanochemical reactions on dry grinding structurally different silicates", Journal of Materials Science Letters,(2001), 20, 81-83. 11. L. G. Shumskaya, "Directional Changes in the Properties of Aluminum Hydroxide-Oxides for Increase in Bauxite Reactivity in Hidrometallurgical Processing". Journal of Mining Science,(2002), 38, n° 03, 299-304.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

RESOURCE UTILIZATION OF HIGH-SULFUR BAUXITE OF LOW-MEDIAN GRADE IN CHONGQING CHINA Jianguo Yin, Wentang Xia, Mingrong Han Chongqing University of Science & Technology, Chongqing 401331, PR China Keywords: High-sulfur bauxite, Resource utilization, Low-median grade, Desulfurization Abstract

It is generally thought that bauxite ores containing sulfur over 0.7% cannot be used directly to produce alumina. So high-sulfur bauxite ore has been basically abandoned, or has had a little utilization mixed into other resources until now. If a process using high-sulfur bauxite to produce a suitable quality alumina can be successfully developed, the resource life of Chinese bauxite ores will be prolonged by about 10 years. So for resource utilization of high-sulfur bauxite, it is very important not only for the development of the regional economy in Chongqing but also for the China alumina industry.

Resource utilization of high-sulfur bauxite is one of the technical problems for alumina refineries in China. There are rich highsulfur bauxite ores of low-median grade in Chongqing China, which haven't been utilized until now. Sulfur in bauxite will cause many negative effects on alumina production. There are still some disadvantages for current desulfurization technologies and it is difficult to make a breakthrough progress only from the view of application of these desulfurization technologies. So it is necessary to undertake fundamental theoretical studies on the occurrence of sulfur in high-sulfur bauxites, the reaction behavior and the impact of sulfur in the production process. Combining theoretical research with experiments, it is possible to develop a feasible desulfurization process. This will provide technical support and theoretical guidance for the utilization of high-sulfur bauxites in Chongqing, and is also advantageous for prolonging bauxite resource security for the Chinese alumina industry.

Negative effect of sulfur on the alumina production Sulfur in the bauxite ores will cause a lot negative effects on the alumina production. It will increase the soda consumption and decrease the digestion ratio of alumina. Sulfur in the form of the anions S2" and S2O32" in the liquors will erode the steel and iron of the pipes and equipment, and increase the iron content in the liquors and the products, which will cause negative effects on the quality of alumina product and in the subsequent aluminum electrolysis. When sulfur accumulates to a certain concentration, it will be precipitated in certain forms and cause scaling of the spent liquor evaporator and slurry digester, which will decrease the heat transfer coefficient of these devices or even make them not work (3). So it is necessary to desulfurize in the process of alumina production from such ores.

Introduction The alumina industry in China has developed quickly in recent years. Output of alumina in the year of 2009 was up to 24 million tons, and production capacity reached 33 million tons. It may be above 40 million tons in the following 2 or 3 years. This brings serious supply problems for bauxite resources as the production capacity of alumina expands so rapidly in China. From the view of the world, the service life of bauxite ore reserves is more than 100 years based on an annual output of 80 million tons of alumina and annual consumption of 2.5 billion tons bauxite. But from the view of China, the service life is no more than 15 years, based on an annual output of 20 million tons of alumina, even if the prospective reserves are considered (1). For Chinese bauxite ores, quality will decrease rapidly or even to exhaustion in the traditional mining areas. So it is necessary for the Chinese alumina industry to cope with this serious resource problem in order to ensure the alumina industry develops continuously, stably and healthily.

Previous desulfurization technologies Desulfurization technologies for high-sulfur bauxite ores have been studied a lot at home and abroad, and can be divided into flotation desulfurization, roasting desulfurization and wet desulfurization. Classification of desulfurization technologies is shown is Fig. 1. Desulfurization Technologies Flotation

There are some bauxite resources which are difficult to be utilized until now in the main mining areas, which include high-sulfur bauxite ores. There are 11% or so of these high-sulfur bauxite ores and about 5.6 billion tons reserves in China. With the further survey of bauxite ore under the coal seams, the portion of the ore may rise. There are rich high-sulfur reserves in Nanchuan Chongqing China, the proven reserves are more than one billion tons and the prospective reserves are about 3 billion tons. Bauxite ores of Chongqing have their own characteristics, which belong to high-sulfur diaspore bauxite of low-median grade and very different from those of Guizhou. The average quality index of the bauxite ores are as follows: A1 2 0 3 44.26%-71.39%, Si0 2 9.54%19.16%, Fe 2 0 3 0.80%-20.27%, S 0.16%-2.19%, A / S (the mass ratio of alumina to silica) 3.68-7.54 (2).

Roasting

Wet II

Direct Flotation Reverse Flotation

Muffle Furnace I

Precipitation

Rotary iTube Furnacel

Lime

Fluidized Bed

Barium Salts

Zinc Species Fig.l Classification of desulfurization technologies

19

1

Wet Oxidation

hydrophobic hematite, which improves the settling performance. But if the roasted ore is too loose, it will cause red mud fines and seriously decrease its settling performance. Compared to that roasted in a muffle furnace, the ore from the rotary tube furnace has better settling performance (18). High-sulfur bauxite ore in Henan province was investigated by Xiaolian Hu et al. with lime roasting in a muffle furnace. It shows that sulfur in the form of sulfide decreases after roasting. Lime roasting has better result for sulfur-fixation, and it decreases the content of sulfur dioxide emitted into the air. In the roasting process, pyrite is converted to hematite. After treatment, the bauxite ore decreases the sulfide content of the digestion liquor and improves its digestion performance. The optimum roasting temperature is 600. When adding 1% lime and roasting, the comparative digestion ratio is 95.35 % and sulfide in the liquor is 0.16 g / L (19). In the process of lime sintering, sulfur is removed by adding anthracite. But the desulfurizing efficiency is low (20). Improving the adding way of coal can increase desulfurizing efficiency (21).

Flotation desulfurization Through differences in the physicochemical properties of mineral surfaces, the minerals can be beneficiated, and the differences can be enlarged by adding flotation reagents and manually modified, which makes the flotation method very adaptable. Moreover the flotation method has high separating efficiency, so it is one of the most extensively adopted beneficiation methods. Sulfur in the bauxite ore mainly exist in the form of pyrite. Pyrite can be floated easily by xanthate collectors, but ores containing aluminum are hydrophilic and not easy to collect. So xanthate collectors can realize the separation pyrite from the ores easily and then decreases the sulfur content of the ores. Flotation desulfurization was firstly used by the Ural Engineering Institute of the former Soviet Union. Technologies of one-roughing, twocleaning and two-scavenging stages were adopted to treat bauxite ore containing 2% sulfur and the concentrate contains sulfur less than 0.41%. But this flotation desulfurization process is too complicated. Flotation desulfurization by potential control was reported, but it was too hard to realize on production scale (3).

It can be concluded that roasting desulfurization is relatively simpler, however, it increases the energy consumption and pollutes the environment.

Xiaomin Wang et al. studied desulfurization by reversed flotation using ethyl xanthogenate, and attainted optimum conditions as follows: pH value 12, flotation reagent dosage 0.4 kg /1, agitation time 15 minutes, the content of ore slurry 10% and particle size of ore less than 0.09 mm. Under these conditions, the sulfur content of bauxite ore can be reduced from 2.08% to 0.65%, and recovery ratio of alumina reaches 91.46%, which is suitable for alumina production (4). Using butyl xanthogenate as a reversed flotation regent, good desulfurization results can also be attained (5). Using soda as pH modifier, sodium hexametahposphate as an inhibitor, sodium sulfide and copper sulfate as activators, ethyl xanthogenates and butyl xanthogenate as collectors, Wenmi Chen treated high-sulfur bauxite in Guizhou China by reversed flotation. For the concentrate, the sulfur content is 0.44% and the recovery ratio of alumina is 96% (6). Adopting a one-stage flotation process to treat high-sulfur bauxite ore in Guizhou, the concentrate contains 0.15% sulfur and alumina losses are only 6.3% (7).

Wet desulfurization 1. Desulfurization by wet oxidation Desulfurization by wet oxidation involves pumping in air to oxidize sulfur into sodium sulfate which precipitates in the evaporation process of the spent liquor (22). But increasing the sodium sulfate in the spent liquors will lead to the double salt of burkeite precipitating in the evaporation process, which accelerates the evaporator scaling and effects its operation. So the formation of the thiosulfate anion needs to be prevented, lest it hastens the corrosion of the equipment. In addition, desulfurization by wet oxidation is a little dangerous. 2. Desulfurization by precipitation Lime is commonly used as a desulfurization reagent by precipitation. Jun Lan investigated desulfurization by precipitation for a high-sulfur bauxite ore of high grade in Maochang Guizhou China. The effects of lime dosage, alkali concentration, digestion temperature and time on the desulfurization efficiency were investigated. The results show that optimum conditions are as follows: lime dosage 10%, digestion temperature 245, desulfurization time 70 minutes and alkali concentration 240 g / L. Under these conditions, the digestion ratio of sulfur is only 31.3% and that of alumina is over 81% (23). Adopting the same technology, Runde He thought the optimum conditions are as follows: lime dosage 12%, digestion temperature 260, desulfurization time 50 minutes and alkali concentration 195 g / L. Under these conditions, the digestion ratio of sulfur is only 13% and that of alumina is over 88% (24, 25).

From the view of the technology itself flotation desulfurization is matured, but from the view of alumina production it brings some negative effects as the concentrate contains water and flotation reagents. At the same time, the process needs to treat large amount of bauxite ores, waste water and tails, and add a lot of reagents and fresh water (8). Roasting desulfurization A lot of research has focused on roasting desulfurization (9-13). Muffle furnaces, rotary tube furnaces and fluidized beds were used to pretreat high-sulfur bauxite ore. Optimum roasting conditions for muffle furnace and rotary tube furnace are a temperature of 750 centigrade and time of 30 minutes. Those for fluidized bed are temperature 800 and time 10 minutes. The digestion performance of the pretreated ore from the fluidized bed is the best, with the digestion ratio of alumina reaching 93.7% under the digestion conditions of temperature 220, molecular ratio of the proportioning 1.3 and caustic concentration 220 g / L (1417). Sulfur is removed in the form of gas, and the roasted bauxite ore has better digestion performance. It is found that roasting makes the red mud loose and porous, and converts goethite to

Barium salts are another desulfurization reagents utilizing precipitation. Barium oxide and barium hydroxide were chosen to purify industrial sodium aluminate liquors (26). Desulfurization efficiency of barium salts is satisfying, and can reach 99%. But when the contents of silicate and carbonate in the liquors are high, the consumptions of barium salts will increase. As these two salts are expensive, it will increase desulfurization expense. Adopting cheap bauxite ore of high grade, and barium carbonate to generate a desulfurization reagent of barium aluminate, the desulfurizing result is similar to those of the two barium salts, but the

20

desulfurization expense can be greatly reduced (27, 28). Desulfurization ratio reaches 94.5 % when barium aluminate is adopted to purify washing water of red mud (29). It is thought that washing water of red mud is suitable to be desulfurized and purified (30). Fanghai Lu thought that spent liquors and the first washing water of aluminum hydroxide are suitable to be desulfurized (31).

Conclusion There are rich high-sulfur bauxite ores of low-median grade in Chongqing China, which are characteristic of the region and have not been utilized till now. It is necessary to undertake fundamental theoretical studies on the occurrence of sulfur in high-sulfur bauxites of low-median grade, occurrence, reaction behavior and the impact of sulfur in the production process. Combining theoretical research with experiments, it is possible to develop a feasible desulfurization process. This will provide technical support and theoretical guidance for the utilization of high-sulfur bauxites in Chongqing, and is also advantageous for prolonging bauxite resource security for the Chinese alumina industry.

Zinc oxide has also been used as a desulfurization reagent, which makes sulfur precipitate in the form of zinc sulfide, and iron was removed at the same time. But zinc species are expensive, which will increase the cost of desulfurization. Adding zinc species will also influence the quality of the products. Wet desulfurization has the advantage of high desulfurization efficiency. The key is to choose a cheap desulfurization reagent in order to decrease the total cost of the process.

References 1. Qiuxia Wang, Keren Zhang, Junwei Zhao, et al., Status, problems and countermeasures of development and utilization of bauxite resources in China, Conservation and Utilization of Mineral Resources, 5(2008), 46-50. 2. Yang Sun, Fuchun Zhou. Study on processing industry of superior mineral resource in Chongqing, Mining Safety and Environment Protection, 32(2)(2005), 27-30,54 3. Wenmi Chen, Qiaoling Xie, Xiaolian Hu, et al., Research on reverse flotation for desulfurizing of high grade bauxite containing sulfur and producing qualified alumina products, Light Metals of China, 9(2008), 8-12. 4. Xiaomin Wang, Tingan Zhang, Guozhi LV, et al., Flotation process for desulfurization of high-sulfur bauxite, Chinese Journal of Rare Metal, 33(5)(2009), 728-732. 5. Xiaomin Wang, Tingan Zhang, Guozhi LV, et al., Flotation desulfurization of high-sulfur bauxite with butyl xanthate as collector, The Chinese Journal of Process Engineering, 9(3)(2009), 498-502. 6. Wenmi Chen, Qiaoling Xie, Xiaolian Hu, et al., Experimental study on reverse flotation technique for desulfurizing of highsulfur bauxite, Mining and Metallurgical Engineering, 28(3)(2008), 34-37. 7. Fangyin Niu, Qin Zhang, Jie Zhang, Orthogonal test research on flotation desulfurization of high-sulfur bauxite, Acta Mineralogica Sinica, 27(3/4)(2007), 393-395. 8. Runde He, Si chun Hu, Zhiying Li, et al., Discussion on the method of hydrometallurgical desulfurizing during producing alumina with high grade bauxite containing sulfur, Hydrometallurgy of China, 23(2)(2004), 66-68. 9. Chongyu Yang, Technology of alumina production (Beijing: Metallurgical Industry Press, 1993), 133-140. 10. Yiyong Wang, Tingan Zhang, Xia Chen, et al., Effects of microwave roasting on leaching behavior of diaspore ore, The Chinese Journal of Process Engineering, 7(2)(2007), 317-321. 11. Hengqin Zhao, Chongyu Yang, Hangbo Liu, Investigation on kinetics of bauxite's digestion improved by preroasting process, 6(3)(1998), 28-32. 12. A. R. Padill, D. Vega, M. C. Ruiz, Pressure leaching of sulfidized chalcopyrite in sulfuric acid-oxygen media, Hydrometallurgy, 6(l/2)(2008), 80-88. 13. Pi Liu, Zijian Lv, Zhongling Han, Analysis of effect of China bauxite ore preroasting on the Bayer process and its industrial application, Light Metals of China, 4(1997), 15-18. 14. Guozhi Lv, Tingan Zhang, Li Bao, et al., Desulfurizatin of high-sulphur bauxite by fluosolid roasting and its effect on the

Our suggestion From the above discussion, for previous desulfurization technologies for high-sulfur bauxite, there are some disadvantages such as the negative effects on alumina products, increasing material or energy consumption, polluting the environment, increasing the expense or hard to realize in the production environment etc. Under the current circumstance, it seems hard to make a breakthrough progress only from the view of the desulfurization technology itself. More importantly, previous studies focused on desulfurization only for high-sulfur bauxite ore of high grade, there is little research on bauxite ore of low-median grade. In order to deal with resource utilization of the high-sulfur bauxite of low-median grade in Chongqing China, it is necessary to have fundamental theoretical studies on the occurrence of sulfur in high-sulfur bauxites, its speciation, reaction behavior and the impact of sulfur in the production process. Combining modern instrumental analysis with chemical phase analysis, it is possible to fully investigate the occurrence of sulfur in the bauxite ore first, which will guide the following work. Then simulated study of the reaction behavior of different compounds containing sulfur in the alkali liquors can be undertaken, according to this information. Investigations of the sulfur chemical species, reaction behavior and its effect on the digestion ratio of alumina in the digestion process will follow along with a study of the digestion ratio of sulfur under the conditions of controlled digestion. On the basis of previous work, combining physicochemical properties with theoretical analysis results, and controlling the chemical reaction in the whole process, a suitable way of desulfurization for high-sulfur bauxite ore of low-median grade might be attained. It is founded that total sulfur of high-sulfur bauxite in Nanchuan Chongqing is about 1.4%. Most of this exists in the form of pyrite constituting about 1%. Other forms exist such as marcasite, melnikovite or gypsum, which are disseminated through the ore deposit and hard to separate only by flotation. Because of the diversity of sulfur species for high-sulfur bauxite ores in Nanchuan Chongqing, it is hard to get good desulfurization results by only applying a single desulfurization technology. So it is necessary to undertake some fundamental studies in order to utilize high-sulfur bauxite ores of low-median grade in Chongqing China to produce alumina of metallurgie grade.

21

digestion performance, Ming and Metallurgical Engineering, 28(6)(2008), 58-61. 15. Guozhi Lv, Tingan Zhang, Li Bao, et al., Roasting pretreatment of high-sulfur bauxite, The Chinese Journal of Process Engineering, 8(5)(2008), 892-896. 16. Guozhi Lv, Tingan Zhang, Li Bao, et al., Roasting pretreatment and digestion performance of high sulfur bauxite, The Chinese Journal of Process Engineering, 9(Supplement 1)(2009), 71-75. 17. Guozhi Lv, Tingan Zhang, Li Bao, Roasting pretreatment of high-sulfur bauxite and digestion performance of roasted ore, The Chinese Journal of Nonferrous Metals, 19(9)(2009), 1684-1689. 18. Guozhi Lv, Tingan Zhang, Li Bao, et al., On the settling performance of red mud from preroasted high-sulfur bauxite, Journal of Northeastern University (Natural Science), 30(2)(2009), 242-245. 19. Xiaolian Hu, Wenmi Chen, Qiaoling Xie, Study on desulfurization of high-sulfur bauxite by calcium oxide roasting, Light Metals of China, 1(2010), 9-14. 20. Shiwen Bi, Alumina production technology (Beijing: Chemistry Industry Press, 2006), 120-122. 21. Yucheng Feng, Jinglian Zhang, Strengthened reduction and sulfur-removal with high temperature coaling for alumina clinker sintering, Light Metals of China, 1(1990), 16-19. 22. C. H. Ky3Heu;oB, Desulfurization of sodium aluminate liquors with furnace ash, Techniques of Light Metals, 1(1974), 10-14. 23. Jun Lan, Xianxi Wu, Yuancheng Xie, Research of desulfurization in the course of disposing of high grade bauxite containing sulfur, Applied Chemical Industry, 37(8)(2008): 886890. 24. Ting He, Zhiying Li, Optimal condition analysis of Guizhou high-sulfur bauxite leaching process, Nonferrous Metals, 60(2)(2008), 68-70. 25. Runde He, Zhiying Li, Nianbing Zhang, et al., Research on leaching function of Guizhou high-sulfur bauxite, Journal of Guizhou University of Technology (Natural Science Edition), 34(3)(2005), 4-7. 26. Huajun Yuan, Qinqin Li, Technological choose of Barium oxide purifying industrial sodium aluminate liquors, Light Metals of China, 4(1995), 13-17. 27. Sichun Hu, Min Guo, Hengqin Zhao, et al., Study on desulfurizing from sodium aluminate Liquors with barium aluminate, Nonferrous Metals (Extractive Metallurgy), 1(2007), 11-13. 28. Runde He, Nianbing Zhang, Sichun Hu, et al., Barium aluminate's combinative study with high grade bauxite instead of aluminum hydroxide, Light Metals of China, 4(2005), 13-17. 29. Runde He, Xifa Tan, Nianbing Zhang, et al., Barium aluminate-purifing Bayer red mud lotion desulfurization experiment, Journal of Guizhou University of Technology (Natural Science Edition), 35(3)(2006), 29-31. 30. Zhiying Li, Nianbing Zhang, Runde He, et al., Research on the new method of desulphurizing in producing alumina with exploitable sulfur-containing and high grade bauxite, Journal of Guizhou University of Technology (Natural Science Edition), 36(2)(2007), 29-31. 31. Fanghai Lu, Discussion on the desulfurizing technique of high grade bauxite containing sulfur in Maochang, Light Metals of China, 9(2008), 17-20.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DEVELOPMENT OF BAUXITE & ALUMINA RESOURCES IN THE KINGDOM OF SAUDI ARABIA AbdulGhafoor Al-Dubaisi Saudi Arabian Mining Company (Ma'aden), Al Khobar, Saudi Arabia 31952 Keywords: Saudi Arabia, Az Zabirah, Ras Az Zawr, Ma'aden, Bauxite, Alumina development of the two large phosphate and bauxite deposits discovered and utilizing the country's abundant energy resources and the developed infrastructure will generate good non-oil revenues putting Ma'aden right after Sabic in terms of revenue generation as soon as the two projects start full production.

Abstract Driven by desire to diversify its economy in an oil rich country and by the need to create jobs for the increasing number of Saudi youth, the Kingdom of Saudi Arabia established the Saudi Arabian Mining Company (Ma'aden) to develop its mineral resources. Local bauxite will be developed into an integrated mine-to-metal aluminum industry. Numerous challenges have to be overcome. Major Infrastructure has to be built; construction and operating costs have to be contained to ensure the economic viability of the project. The lack of expertise in the kingdom to run such an operation created the need to involve an international joint venture partner. The human development is no less challenging and early plans have to be in place to recruit and train a large number of Saudis to be the core of the operating organization.

The demographic changes in the kingdom are alarming. According to the Demographic Survey of 2007, the total population of the kingdom is about 24 millions including over 5 millions resident expats. With over 50% of the population under the age of 25 and a yearly growth in population over 3%, generation of new jobs becomes a high priority and warnings of social havoc due to rising unemployment rates are increasing. Mining industry- primary and secondary -offers two features that make it attractive in this regard. Firstly, this industry will contribute greatly to a balanced regional development planning—another national strategic objective—as a result of investing in remote areas where the resources are located. This will slow down and possibly reverse the migration from such areas to the already congested big cities thus helping to relax the strains, shortages, and socio-economic risks associated with such migrations.

Introduction Saudi Arabia well known for its large reserves of oil and gas contains also appreciable mineral deposits such as gold, cupper, phosphate, bauxite and others. The potential remains high as the amount of exploration for minerals done so far is limited. Most of this potential is identified to be in the Arabian Shield, the geological formation on the west of the country along the Red Sea extending inland to Najd and Qassim areas. In spite of extensive surveys conducted in Saudi Arabia in recent years, the amount of investment in exploration is very modest compared to other countries. Saudi Arabia is blessed with large reserves of phosphate, bauxite, magnesite, gold, iron, copper, and other minerals that are yet to be developed and exploited. It suffices to say that the total fund spent on exploration in the Arabian Shield in the last 30 years is less than what is spent in the Canadian Shield annually. Yet, all indications attest to the diversity of the mineral resources in the Kingdom and the tremendous potential forfindingnew resources through additional exploration.

Secondly, the industry's potential for job creation is large as the mining industry is much more labor-intensive than oil, gas and petrochemical industries. As thousands of young Saudis entering the job market each year, this industry will have the capacity to absorb a large number of job aspirants through direct and indirect hire. Location and Geology The bauxite deposit is located in a remote desert area of northern Saudi Arabia, predominantly in the province of Ha'il. Elevations range from 535 to 600 meters above mean sea level. Ma'aden's exploration license is 192 kilometers long by 35 kilometers wide, approximately centered on the town of Az Zabirah (northwest of Riyadh). Ma'aden's proposed mine targets the South Zone of the deposit and Ma'aden's mine processing facilities will be located approximately 43 kilometers northwest of the town of Qibah. The town of Az Zabirah is a similar distance to the northwest of the proposed mine processing facilities. The mine is called after a small settlement between Qibah and Az Zabirah called Al Ba'itha.

Since the discovery of oil in the late thirties and especially after the oil crisis of the seventies and the increase in oil prices, the country has enjoyed large revenues that translated in the leaps of development of both infrastructure and industry. This revenue remains however subject to fluctuating oil prices and it has seen the worst downturn cycles in the mid eighties and at the turn of the century. From the enactment of the national Five Year Development Plans in 1970, diversification of Saudi economy has remained one of the top strategic objectives of the Kingdom. Great achievement has been made in the petrochemical sector generating over $20 billion revenue in 2003 and in 2007 non-oil manufacturing contributed 10% to the GDP.

The coordinates for the approximate center of the mine processing facilities are: WGS84 latitude / longitude 27°39'52.8"N/ 44°0'38.5"E UTM Zone 38 3 060 750N / 404 600E The alumina refinery will be located on a virgin peninsula on the Arabian Gulf called Ras Az Zawr about 80 KM north of the major industrial city of Al Jubail.

Following the successful story of developing the petrochemical industry led by Sabic, the government established in 1997, the Saudi Arabian Mining Company (Ma'aden) to lead a similar development of the country's mineral resources. The

23

The coordinates of the approximate center of the processing plants (Refinery, Smelter and Rolling Mill) are: WGS84 latitude / longitude 27°30'52.4"N/ 49°10'38.5"E. UTM Zone 39 3 044 753N / 319 980E The location map below shows the location of both facilities and the railway link in between. ~*-m^mmmm k

*

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($MPT« £X$MR ftdmkmi ibpmrt MOM- ϊΆ-l jM fort l )

Since 2001 Ma'aden has continued drilling in the south and central zone to improve resource and reserve estimates. The resource in the south zone of the deposit is currently estimated at about 300Mt and the reserve is estimated at about 200 Mt at 40% TAA cutoff which represent the basis of current phase of the project. The resource in the central zone is estimated by Hatch (2004) at 135 Mt with slightly higher silica and lower iron content. Ma'aden continues through yearly drilling programs to increase resource estimate and obtain better understanding of the bauxite ore in the central zone of the deposit.

Geologic evaluation of the deposit was initiated by Riofinex through drilling exploration programs performed from 1979 through 1984. The Riofinex Report identified three distinct zones (the North Zone, Central Zone and South Zone), spaced approximately 30 kilometers apart along the strike, which exhibited economic potential. Additional exploration was performed by the French BRGM from 1987 through 1993. Ma'aden's also conducted exploration drilling at the deposit in 2002 through 2003. The deposit is located discontinuously along a strike length of approximately 105 kilometers and with a width identified by drilling of at least 5 kilometers. In the Riofinex report, Riofinex interpreted that the deposit's bauxite was formed during the Early Cretaceous as a weathering profile on Late Triassic to Early Jurassic terrigenous sediments deposited in marine or littoral environments. Riofinex tentatively assigned these sediments to the Minjur Formation (North Zone), the Marrat Formation (Central Zone), and the Dhruma Formation (South Zone). Both Riofinex and BRGM referred to these collectively in their respective reports as the "Parent Rock Sequence".

The Project The Ma'aden Integrated Aluminum project, jointly developed by Ma'aden and Alcoa, includes currently the following scope:

• •

4.25 Mtpa bauxite mine in Al Ba'itha 1.8 Mtpa alumina Refinery in Ras Az Zawr 750 Ktpa Aluminum Smelter in Ras Az Zawr 460 Ktpa Rolling Mill Common Facilities and Infrastructure

The Smelter and Rolling Mill are ahead on planned completion schedule whereas Mine and Refinery are lagging with about one year behind. The first metal from smelter is scheduled for early 2012. The Smelter will start with imported alumina until Refinery is able to supply the required alumina.

The deposit's bauxite profile is several meters thick and is unconformably overlain by Late Cretaceous fluviatile and lagoonal sediments. Riofinex and BRGM refer to these sediments as the "overburden sequence (OVB)," and the Riofinex Report characterizes the sediments as part of the Wasia Formation.

On the same site and within the Ras Az Zawr Mineral Industrial City (RAZMIC), Ma'aden is developing also a large phosphate complex that produce DAP fertilizer utilizing ore from the large phosphate deposit in Al Jalamid. The phosphate complex shares with the Aluminum Project basic infrastructure such as roads, housing and port.

The bauxite outcrops in limited places on the western edge of the deposit and uniformly dips at approximately 0.5 degree to the northeast. At the eastern edge of the deposit, the BRGM Report estimated that the overburden sequence is approximately 50 meters thick.

Large Infrastructure- The first challenge Considering the remoteness of mine sites and lack of water resources in these locations, processing plants have to be located hundreds of kilometers away. Therefore, providing a means of

24

2400 MW of electrical power and just over 1 million cubic meter per day of potable water. Most of the power generated (1350 MW) will be used by the aluminum complex through a high reliability grid connection. The power and desalination plant will also provide the required water for domestic, process and cooling duties.

transportation of the ore to potential processing locations is of utmost importance. Including development of such large infrastructure in a mining project will be fatal to its viability. Early studies of the feasibility of the Az Zabirah bauxite deposit highlighted transportation of bauxite to location of process facility as the major obstacle to develop the deposit. Saudi Arabian Sabic concluded in a study conducted in 1985 that processing of Az Zabirah bauxite is not feasible due to lack of water at the mine site and the lack of means of transport. Since inception of Ma'aden in 1997, it highlighted this problem to government agencies and requested allocation of funds to building the railway to support the feasibility of the bauxite and the phosphate projects.

The mineral railway and the deepwater port are in the final construction stages and will be operational early in 2011 to support the phosphate plant which is being commissioned. The power plant was awarded for execution and is expected to be operational late 2012 in time to support the startup of the aluminum smelter.

Another major infrastructure required for the process facilities on the Gulf is a deep water port that allow import of raw materials ( caustic, liquid pitch, calcined coke ect) and export of products ( aluminum metal, alumina, DAP, etc) from and to the world markets.

The Integration Challenge The bauxite and alumina is part of a larger development for a mine-to-metal integrated project. As outlined above the level of interfaces to develop the required government-sponsored infrastructure is by itself a challenge. It required working through different governmental and semi-governmental organizations to move things and make decisions. Coupled with that is a high level of interface management required internally to get the different elements of the project aligned. These elements (Mine, Refinery, Smelter, Rolling Mill and Integrated Infrastructure) are being designed and executed by different EPCM consultants in three continents and engineering design is spread over many countries. Yet all these elements have to coordinate activities and manage interfaces related to site, systems and procedures, corporate policies, design standards and at all execution and operation readiness levels. This presents a challenge unprecedented at this scale in the aluminum industry.

Recognizing the value of these mining projects to the development of the country, the government undertook these two major infrastructure projects to build the North South Railway (NSR) which is by itself a major railway project with a track length of over 1800 Km extending from the Al Jalamid phosphate deposit near the Jordanian borders down to Az Zabiarah bauxite deposit andfinallyto the Ras Az Zawr mineral processing facility on the Arabian Gulf. Reliable and cost-effective electric power source is vital to the success of the project. The national grid is quite stretched and will not provide the reliability and the tariff that support the project. The electric power required for the Ras Az Zawr Aluminum Complex including the three operations (Refinery, Smelter and Rolling Mill) and infrastructure, is in excess of 1300 MW. Ma'aden completed many studies for the feasibility of a power plant to support the industry but this encountered many challenges: •

Ma'aden would be the first private industry to build a major power generation plant at this scale

•

The power industry regulatory framework is at infancy stage and there are no clear regulations on issues such as power generation licensing, fuel pricing, power tariffs, power wheeling, reliability of supply etc

•

Connection to the national grid is a must not only to provide backup power but to allow sales of excess power needed by the national grid.

•

The size of investment in a large power plant became huge especially with high sulfur heavy crude oil proposed.

•

The issues above required a large level of coordination among many government organizations where decision making is very slow.

Bauxite characterization and Technology Selection Extensive sampling and assaying programs were conducted for the bauxite starting with Riofenix program. Ma'aden during a prefeasibility study in 2001-2003 carried a data acquisition and assaying program. Ma'aden undertook an intensive close-spaced drilling program in a region of the South Zone forming the initial ten-year mine plan. A number of samples from that area were then composited in a devised plan to arrive at a bulk representative sample, indicative of the first ten years mining. The drill-hole samples were assayed at the Al Amri laboratory in Saudi Arabia, and the samples were further analyzed with extensive testing by Australia's Commonwealth Scientific and Industrial Research Organization (CSIRO) Minerals laboratories in Australia. These composites were characterized in terms of chemical elemental composition, quantitative mineralogy, particle size distribution, and available alumina and reactive silica under expected refinery digestion conditions (digestion temperature and liquor caustic concentration). The predesilication behavior of the composite bauxites was investigated, particularly the extent of reactive silica to desilication product (DSP) transformation under nominated conditions, and the minimum time to achieve optimum transformation. Also On behalf Ma'aden, in 2004 Hatch engaged JKTech to determine a number of comminution parameters for Az Zabirah bauxite samples, including 'average grade - AGB', 'high grade HGB' and 'low grade - LGB'. Parameters such as the ore behavior in SAG/AG mill, the primary crushing power draw, the rod milling power draw, the ball milling power draw, the abrasion of

After much iteration, and a few years of studies, consultations and interaction with government agencies, it was decided to combine Ma'aden power requirement with water needs of the country and under special arrangement Saline Water Conversion Corporation (SWCC), a government agency, will build a combined water and electricity generation facility in Ras Az Zawr with a capacity of

25

was done. The choice of technology was governed by the following underlying principles:

mill lining, the strength of the ore sample when subjected to compressive forces without lateral constraints, were investigated. Later in 2005 and 2006 additional infill drilling was carried out in the South zone and a new composite sample representing the 30 years ROM was prepared. Further chemical and physical testing programs were undertaken between 2007 and 2010 including digestion and extraction tests, bauxite rheology tests and impurity analysis (especially of iron, fluoride, chloride and calcium). The results of all these tests confirmed findings of the earlier campaigns and were incorporated in the flow sheet design. The series of tests confirmed the following basic characteristics: • • • • • •

The alumina is highly boehmitic (monohydrate) High alumina and silica content ( average 56 %, and 9% respectively) Bauxite will require a high digestion temperature (260-280C) Due to high silica and high DSP, the circuit will therefore have a high capacity to incorporate impurities Low iron content Low organic matter

Table 1. Az Zabirah Bauxite (Average Values)

1

High grade

Representative Low grade

A1203

56.2

55.56

54.14

Fe203

12.59

11.17

10.99

Si02

5.48

7.71

9.29

CaO

1.24

1.14

1.16

Ti02

3.48

3.45

3.35

P205

0.16

0.15

0.15

MgO

0.09

0.09

0.09

S03

0.65

0.65

0.68

V205

0.1

0.09

0.09

Cl

0.03

0.03

0.03

F

0.08

0.08

0.08

LOI

19.77

19.64

19.4

TOC

-0.2

-0.2

-0.2

Design has to use public-domain technology as much as possible

•

Design has to reflect the most advance developments in Bayer process, especially in environmental, health and safety.

•

The bauxite is new and has not been processed before.

•

The refinery will have a new operation team and design must be robust and forgiving

•

The design has to allow for local harsh ambient conditions

The above considerations and the bauxite characteristics formed the basis of the process design. Although conventional Bayer circuit was selected for most areas, digestion design was given special consideration. Ma'aden operation experts along with external consultants carried out a qualitative comparative risk evaluation of digestion design options that include single stream full jacketed pipe heating, hybrid indirect and direct heater design and dual stream utilizing conventional shell and tube heaters. This evaluation was undertaken as possible means to mitigate the risk of unacceptably high scaling rates in high temperature indirect slurry heaters to determine which should be taken forward into process and engineering design. Evaluation criteria included process risks, capital and operating costs, and operability and maintainability considerations. Based on the evaluation criteria, the hybrid option was recommended for heater design in digestion. This option benefits from the jacketed pipe heating design in terms of energy efficiency and ease of cleaning. However, by shifting to the direct heating beyond 200C, the design minimizes the risk of potential scaling in the tube.

Table 1 below shows the composition of Az Zabirah bauxite

Components (wt%)

•

The Ma'aden alumina refinery consists of a two-unit greenfield plant operating the Bayer process and using public domain technology to produce 1.8Mtpa of smelter grade alumina. Approximately 1.4Mtpa of the refinery product is expected to be consumed by the nearby Ma'aden aluminum smelter. The remainder will be exported. The Refinery will nominally comprise: •

Three grinding mills, at 100% digester unit capacity each

•

Two digester units

•

Two precipitator trains

•

Two fluid-bed calciners, at 75% total plant capacity each

•

Two evaporation trains.

The Figure below shows a schematic of the Ma'aden Refinery flow sheet and the basic design conditions.

The Az Zabiarh bauxite underwent extensive bench scale testing by reputable research laboratories and results were reviewed by many experts from different leading operating companies. Following 2009 Alcoa JV partnership the process design was reviewed by Alcoa process consultants. The consistency in the results of these tests and reviews gave high confidence in the selected refinery flow sheet design. Based on the bauxite chemical and physical characterization, the selection of process technology

The Maaden Bayer circuit is unique in some aspects due to the bauxite chemistry and the country ambient conditions: 1. The flow sheet does not return any supernatent liquorfromthe residue area due to high evaporation rates

26

2. 3.

•

The Bauxite has high reactive silica and therefore high incorporation capacity of impurity in the DSP. The casticity of the liquor is relatively high ( about 0.94)

Joint venture partner discussions, negotiations, and due diligence were carried out with the major players in the industry. At last and In December 2009, a joint venture partnership was announced with Alcoa. This paved the way to progress the project into the next phase of execution and provided depth of resources in both execution and operation expertise.

All of the above predicts a fairly clean liquor circuit. Mafaden Alumina Process

The large size and complexity of the project and large equity funding ( over $10 billion investment)

6?i8kPsg

Controlling Construction Costs and Operating Costs The Saudi government extended its great support for the first major mineral development in the kingdom. The government, through Public Investment Fund (PIF) and the Saudi Arabian Railway (SAR) built a $4 billion railway link between the two deposits ( Al Jalamid and Al Ba'itha) and the process facilities in Ras Az Zawr on the Gulf coast. The railway crossing the sand dunes of An Nufude Desert involved record breaking earthwork volumes in a difficult terrain. The government through the Saudi Port Authority (SEAPA) built a major deepwater port in Ras Az Zawr, a project close to $1 billion. The government of Saudi Arabia invested in a large combined water and power project in Ras Az Zawr to provide electricity to the Aluminum complex at attractive prices. Yet, the amount of remaining infrastructure to support the project is sizable. The mine site facilities in Al Ba'itha are in remote area which lacks all infrastructures and the project has to include development of housing, water supply and access roads. The process facilities in Ras Az Zawr are proposed on a virgin peninsula on the Arabian Gulf north of the major industrial city of Jubail. Major housing, access roads, security fencing, seawater supply and return systems, and earth works have to be developed by the project. The integrated project sought possible synergies between the smelter, refinery and rolling mill to share these infrastructure facilities to reduce costs. Value engineering studies were conducted at many stages of the project to challenge some of the equipment choice and or count to eliminate any excess capital. Best execution and delivery systems had to undergo several reviews by international and reputable consultants as well as Alcoa internal peer reviews. Alcoa experience in delivering such large and complex project was brought to table. Given the dispersed execution centers, the cost of execution including owner and EPCM consultants costs have to be contained. In some of the hot market environment that passed in 2007 and 2008 when construction costs sky rocketed, this threatened the feasibility of the project.

Joint Venture Partner Since Ma'aden have no operating experience in the alumina and aluminum industry, the search for a joint venture partner was inevitable. Given the technical, political, and business challenges with establishing a vertically integrated aluminum industry and given the financial turmoil that hit the markets in the last few years, this search was long and demanding. It is true that Saudi Arabia is rich in energy and this constitutes a major driver for investment in the energy demanding aluminum industry. However from Ma'aden point of view, with a mandate to develop the local mineral resources, exploiting the local bauxite resources is equally important for the good reasons discussed above and as mandated by the country's strategic development plans. The following issues stood out in the process of search and selection of JV partner: • The magnitude of investment in the infrastructure (at earlier stages of the JV surveillance many of the infrastructure elements that are now shouldered by government were part of the project). • The legal framework of the kingdom is not advanced enough and that was perceived as a risk to external investors • Technical risks associated with processing of the new Az Zabiarh bauxites

Controlling the operating costs of each ton of bauxite, alumina and aluminum is critical to the success of the project. Efforts were made to optimize the mining plan of waste and bauxite. Considering the quality of bauxite in Al Ba'itha, the different grade control parameters, and the variability in the ore body, the mine plan went through several optimization and reviews to reduce the per ton cost. The options of contracting all or parts of the mine operation were carefully examined. Energy costs for the production of Alumina were compared against other similar operating refineries. Despite the high temperature digestion conditions (273 C), opportunities to reduce steam consumption were exploited using experience from other Alcoa energy efficient operations. Working closely with government agencies to

27

gain their support in building the Ras Az Zawr power plant is part of the effort Ma'aden management took to control energy costs essential to competitive operation.

plan was developed to ensure selecting the best students from surrounding areas of the mine and refinery facilities to support the local communities and to establish for higher retention. Students will be recruited in different intakes depending on their role and to ensure they join in time to attend on-job training before the startup of the facility. Ma'aden Human Resources and Operational Readiness personnel approached technical colleges in the area and discussed curriculum and timetables. Because this is a new industry in the region, a level of customization of college training curriculum has to be done. Ma'aden and its joint venture partners facilitated a cooperation program with overseas technical colleges with experience in training industry professionals, to ensure training program outputs respond to the specific needs of the industry.

The consumption of caustic for processing the Az Zabirah bauxite is very high averaging 200 KG/ ton of alumina due to the high reactive silica content. This puts the yearly demand of caustic for this refinery at more than 350 Ktpa on dry basis. Considering the volatile and unpredictable caustic prices, the risk of escalating operating costs due to caustic is very serious. Ma'aden explored options available including teaming up with local investors to build a caustic production facility in Ras A Zawr. Discussions and negotiations were held with the local caustic producer, Sabic. These efforts succeeded in establishing a joint venture partnership with Sahara, a local petrochemical company, to construct a facility that will provide the bulk of caustic requirements ( Plant was sized to produce 250 Ktpa of membrane grade caustic and 300 Ktpa of EDC, and planned for startup one year earlier than Refinery). The remaining quantity will be sourced from Sabic or through global procurement network of the JV partner. In parallel, efforts to reduce caustic consumption in the process continued looking at options such as lime addition, improving wash efficiency and others.

Selected roles in the mine and refinery operating organization were targeted for overseas training in an operating bauxite mines or alumina refineries. These are mainly in lead and supervisory positions. Saudis nominated for these roles will be experienced engineers, and operation/maintenance staff from existing operation who lack the specific industry experience. They are intended to be the core of the operation organization. They will be seconded to overseas operation for varying periods to give them the required exposure before commissioning the operation of mine and refinery. The plan requires them to finish in time to return to country to participate in developing operation, maintenance and training materials, coaching of the new college students and later participate in the pre-commissioning and commissioning activities. Management recognizes that staffing as early as planned is costly, however the payback in terms of readiness for operation cannot be under estimated.

To control labor costs, the mix of labor for the mine and refinery organization has to be optimized. In Saudi industry, human resources are categorized into three types: Saudi Nationals, Western Expatriates, and Eastern Expatriates. Major industries in kingdom have achieved over the years a large percentage of Saudization approaching 90%. The target set by management for Ma'aden aluminum project is 50% as minimum. Considering the level of expertise potentially available in each labor category at startup, each role was studied carefully to ensure the mix of labor supports the intent to control labor costs but does not compromise safety or quality.

Deliverables required from vendors to support training were highlighted and given high priority in the procurement process. The delivery of these was timed to support the overall training plan to ensure training material is available earlier not later. Training of locals was included in all contracts signed for technology, engineering and project management.

These efforts to contain operating costs and measures taken will result in a very competitive operation. Cash operating costs in the first quartile or lower second quartile are forecasted for the refinery and smelter operation in Ras Az Zawr.

Thinking of long term retention of developed skills - a challenge ahead.

Developing the people and the skills- Planning for operational readiness

Ras Az Zawr is a relatively remote location. It is planned to house employees on single status in company provided campus on site. Families have to stay in the nearby Jubail. The majority of our staff will be married Saudis according to the cultural norms. They will have to drive about 80 Km to come to work and back at the end of the work day.

The challenge of building the Operation Organization including people, business systems and manufacturing systems is not less than the challenge of building the physical assets. Saudi Arabia has well developed industries in oil, gas, and petrochemicals. The surrounding region is home for Saudi Aramco and Sabic, world leaders in oil& gas and petrochemical respectively. Over decades of their presence, they trained and employed large number of skilled Saudi workforce. They established English as the industry language and created a strong base of industry knowledge and support services. However, all of the skilled manpower is tied to these industries and a new industry such as Ma'aden has to work on training and developing its manpower resources. Building on the strong industrial base available, Ma'aden took manpower development as a very high priority.

The operation and maintenance of alumina refinery include labor intensive tasks not typical in the surrounding hydrocarbon production and processing facilities such as descaling of piping and equipment, continuous isolation of sections of the plant for caustic washing, etc. These physically demanding tasks coupled with harsh ambient conditions make the refinery operation a difficult environment. These features of the site and the operation will make retention of the developed skill a challenge ahead. Plans that address human resources aspects, safety and health aspects and community relations aspects will be prepared to make Ras Az Zawr an attractive place to work and maximize talent retention.

Plans for recruitment of young Saudi students were developed. Following the design of the operating organization, the skills of each of these targeted positions were identified. A recruitment

28

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DIGESTION STUDIES ON CENTRAL INDIAN BAUXITE P.K.N. Raghavan, N K Kshatriya, S Dasgupta Bharat Aluminium Co. Ltd., BALCO Nagar, Korba, India Keywords: bauxite, digestion parameters, caustic loss temperature, bauxite charge, lime quantity and digestion liquor concentration is discussed in this paper.

Abstract Indian bauxite deposits are grouped into five categories namely Eastern Ghats, Central Indian, West Coast, Gujarat and Jammu & Kashmir. Each of the bauxites has its own typical digestion and settling conditions due to variations in mineralogy. Bauxite digestion depends greatly on the temperature, pressure conditions, the recycled liquor concentration, bauxite mineralogy and charge quantity. Productivity of the entire Bayer process depends to a large extent on the digestion process. A series of experiments were conducted in an attempt to optimize the digestion conditions for Central Indian bauxite in order to achieve low caustic soda loss coupled with minimum bauxite consumption. Experiments were conducted to find out the best parameters for digestion on a laboratory scale. The effect of digestion temperature, bauxite charge, lime quantity and digestion liquor concentration is discussed in this paper.

Experimental Laboratory digestion tests were carried out in programmable, electrically heated digester, with small autoclave bombs of working volume 150 ml. Bauxite sample of -100 mesh size were dispersed in plant digestion liquor. The bauxite charge to be taken for digestion in the autoclaves was calculated to meet a defined charging molar ratio, which varied in quantity depending on the chemical assay of bauxite. Digestion tests were carried at different temperatures with varying retention time applicable for boehmitic bauxite. The above conditions were selected following the parameters for alumina production at the Bharat Aluminium Company Ltd, Korba (India). The quantity of alumina extracted was estimated by analyzing the bauxite and its corresponding red mud. The major constituents of bauxite and red mud were analyzed by XRF-PW-2440 Philips, Netherlands. The mineralogical analysis were carried out using PANlytical X'Pert Cubix Pro Series diffractometer, equipped with copper target tube, X'celerator detector and operated at 40kVand 30mA. Diffraction data were analyzed using PANlytical X'Pert High Score Plus Version 2.1. The caustic and alumina content of digested liquor were analyzed by potentiometric auto titrator, Mettler Toledo, DL55, Switzerland.

Introduction Alumina extraction from bauxite by the Bayer process requires an efficient digestion. The digestion step in the Bayer process is influenced by the relative amounts of the alumina hydrates present. Atmospheric pressure digestion technology can be used for only Gibbsitic bauxite with negligible or no boehmite content. Low pressure digestion technology on the other hand is used for the Gibbsitic bauxite which contains some boehmite. High pressure digestion (with or without catalyst addition) is exclusively used for the boehmitic bauxite or diasporic bauxite. The digestion process developed by Bayer was designed to treat European bauxite having moderate boehmite and some gibbsitic but negligible diaspore. These latter bauxites need to be digested at high temperature up to 240-2500C and high pressure up to 3035 atm. Lime addition facilitates the extraction of boehmite.

Results A set of digestion experiments was carried out at 240 °C for 18 minutes with varying target Molar Ratios, the outcome of these tests is given in Table 1. Digestion experiments were carried out at different digestion temperatures with varying target MR keeping constant digestion time and the results of the tests are given in Table 2. The results of digestion tests at different caustic concentration in the plant digestion liquor and for different digestion times are shown in Table 3. Digestion experiments with variation in bauxite quality at two targets MRs 1.42 and 1.45 are presented in Tables 4 and 5 respectively. Table 6 shows the test results obtained after bauxite digestion with various quantities of lime.

Caustic soda consumption is controlled by the digestion conditions and the amount of the kaolinite and quartz. The performance of the digestion process is a function of various parameters such as the digestion liquor concentration & Molar Ratio, the digestion temperature, the bauxite charge, residence time and lime addition. In practice, because of the fluctuating operations either all or a few of the above parameters do deviate from the laid down norms causing continuous variations in the digestion efficiency. The productivity of the entire Bayer process depends to.a large extent on the efficiency of the digestion process. In this case high temperature digestion (235-245°C) is used for Central Indian bauxites because of the boehmite content.

Discussion The digestion studies mentioned above have been carried out in order to optimize the digestion conditions to achieve the highest digestion efficiency, with lowest soda loss.

Plant efficiency requires an optimization of the digestion conditions in order to achieve low caustic soda loss coupled with minimum bauxite consumption per tonne of output alumina. In an effort to achieve the above mentioned goals, a series of experiments were conducted to find out the best parameters for conducting digestion on a laboratory scale. The effect of digestion

Laboratory digestion tests were conducted in order to determine the critical Blow off Molar Ratio beyond which the efficiency drops to a very low value. The results (Fig.l & Table 1) clearly show that with reduction in target MR the digestion efficiency decreases in all cases. But at the same time we may observe that below the target MR value of 1.40 there is a significant change in

29

the slope of the curve indicating the lower limit of the liquor to bauxite ratio. It can be seen from Fig.l and Table 1, that the digestion efficiency decreases to some extent with decrease in the target MR. But at times efficiency may be sacrificed in order to increase the throughput. This is a clear case of optimization as a techno-economic consideration.

The disadvantages are anticipated to be as follows: (a) Higher steam consumption (b) Increased scaling of live steam heated digesters (c) Increased chances of impurities going into the aluminate liquor, thus leading to increased impurity in product. The digestion studies mentioned above have been carried out in order to optimize the digestion conditions to achieve the highest digestion efficiency with lowest soda loss. Rg.2 Effect of temperature cm digestion efficiency Target MR 1.42 <

^

S 84 \

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* 83-I 240

. 242

ι 244

Table 1. Digestion tests at different target MR Bauxite Quality

Plant Liquor wt %

LOI

2235

Si0 2

3.15

Fe 2 0 3

19.43

Ti0 2

9.13

lAh
Na 2 0 c A1203

Table2. Digestion test at different temperature with varying target MR

Digestion Conditions

In gpl 188

Desili cation

90.9

Digest ion Lime

Temp,

95

Time, hrs Temp,

8

°C

°C

Bauxite Quality wt %

240

18 Time, minutes 0.75% of bauxite charge

45.00

Target MR 1.28 1.30 1.32 1.34 1.36 1.38 1.40 1.42 1.45 1.50

ML 49.81 49.76 49.17 48.99 48.76 48.33 48.40 48.36 47.97 48.46

η 79.89 81.09 81.84 82.17 82.74 82.22 83.15 83.55 84.17 84.78

SL 79.27 83.45 81.98 80.47 79.61 87.11 83.61 80.80 78.76 80.03

BXT 2.78 2.74 2.71 2.69 2.70 2.67 2.66 2.64 2.62

LOI Si0 2 Fe 2 0 3

Na2Oc AIA 19.43

Ti0 2 A1203

9.13 45.00

Temp, °C

Results

Target MR

240

ML

% % Kg

242

n

SL BXT ML H

244

ML-Mud Load (%) η - Digestion Efficiency (% TA basis) SL - Soda loss (Kg NaOH/T A1203) BXT - Dry bauxite consumption (T/T A1203 )

Plant Liquor In gpi Na2Oc 185 90.9 A1203

SL BXT ML

n

SL BXT

T

% % Kg T

% % Kg T

138 48.33 82.22 87.11 2.70 48.76 82.61 82.41 2.69 48.44 83.06 83.58 2.67

Digestion Conditions Desili Temp, °C cation Time,hrs Digest Temp, UC ion Time, minutes

95 8 240 18

1 1.40 48.40 83.15 83.61 2.67 48.55 83.60 82.41 2.66 48.18 83.93 81.78 2.65

1.42 48.36 83.55 80.80 2.66 48.12 84.04 78.29 2.64 4833 84.56 78.81 2.63

1.45 47.97 84.17 78.86 2.64 48.34 84.75 7832 2.62 48.46 85.14 77.99 2.61

Soda loss shows a declining trend when digesting bauxite for 36 minutes instead of 18 minutes (Fig.3 & Table 3).

It is well known fact that digestion efficiency increases with increase in digestion temperature. The digestion efficiency increases about 1% at 244°C (Fig.2 & Table 2), at a target MR of 1.42, which is closest to the plant operating range.

Bauxite quality plays an important role in the digestion efficiency achieved in the plant, other parameters remaining constant. Of late the bauxite quality for this operation is deteriorating, mainly with respect to recoverable alumina content, whereas the silica content is at acceptable level. In general the digestion efficiency achieved in the plant level is in direct proportion to the alumina content in the bauxite.

The advantages of operating at 244°C are: (a) Higher bauxite dissolution hence lower bauxite consumption (b) Lower soda loss

30

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Fig.3. Effect of digestion time on s o d a toss Dig T e m p - 2 4 0 ° C Target MR- 1.40

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Fe 2 0 3 Ti02

22.42 3.27 17.85 9.02

A1 2 0 3

46.57

LOI SÌO 2

Digestion liquor Na2Oc,gp MR 1 170.5 175.9 179.8 186.0 190.0

2.91 2.95 2.92 2.93 3.02

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MRUS

SL 81.61 82.81 81.61 82.09 81.10

85.48 85.93 86.01 86.13 86.32

SL 77.94 79.62 79.58 79.65 80.31

Though there are some aberrations in the experimental results, in general it can be concluded that digestion efficiency is reduced drastically when the alumina content in bauxite is below 44%. The weight ratio of A1 2 0 3 to Si0 2 known as the module of the bauxite, is more pronounced with respect to bound soda loss (Fig.4, Fig.5 & Table 4, 5) as these are more related to the silica of the bauxite processed. From the experimental results it can safely be concluded that for achieving good digestion efficiency accompanied with low bound caustic losses .The bauxite quality requirement should be as follows: A1203 > 44% , Module > 15 From Fig.8 & Table 6, it can be seen that the soda loss decreases with increasing lime charge, due to the formation of Calcium Aluminium Silicate and release of sodium hydroxide (Soda in red mud).

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as those from the coastal Gujarat are hard, highly diasporic in nature and can be best suited for Refractory applications.

TafokT.Öptiinum digestion conditions for Centrât Indian Bauxite

Fip.8. Effed of ime dosage ®n soda loss

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While carrying out the digestion studies under the conditions outlined above we can expect a lower bound on caustic loss, coupled with low bauxite consumption. We can also expect significantly lower Fe203 and Si0 2 levels in the digested liquor, which gets translated in the product hydrate and hence reports in the calcined alumina. While translating these results to plant level operation, some care must be taken on the following points:

iB

16 Saisate Poésie

Fig.6. Effect of bauxite alurrina on dkgestjon Eff. Target MR-t.AS

42

43

44

45

46

47

(1) The Bauxite used in the laboratory tests was of uniform mesh size, while that in the plant operation are of varying granulometry ranging from very coarse to fine. (2) The agitation level in the laboratory is not the same as under plant conditions. 48

(3) In the plant, flashing of the digested slurry is carried out, however this is not available in laboratory scale experiments.

Alumna« Basaste %

On careful observation of the results, one can observe that the parameters given in Table 7 would be appropriate for the digestion of Central Indian Bauxite.

Fig.?. Effect of bauxite module on soda loss Target MR-1.45

Acknowledgements

» 70 J "S 60 j co

Q2

Lksedssafs

Rg.5.£ffeet o f module on soda loss Target MR-1.42

The authors acknowledge the constant encouragement from our Chief Operating Officer, Mr. Bibhu Prasad Mishra and Chief Executive Officer and Whole Time Director, Mr. Gunjan Gupta, for their constant encouragement during the progress of this work. The Authors would like to thank the Management of Vedanta resources for allowing us to publish this paper.

1 40 + ÎG

12

14

16

18

20

Basaste til ©éufe

On Careful observations, it can be concluded that the Digestion conditions given in Table - 7 will be the optimum for high temperature - High Pressure Digestion of Central Indian Bauxites, though the digestion conditions shall vary marginally for those from the central Indian plateau towards the Eastern Ghats are softer, higher Gibbsitic in nature and hence can be digested at temperatures around 230°C, where as those in the Central India and towards the western Ghats are Bohemitic in nature and need higher temperatures and pressure.

References 1. Zhang Xiaofeng, Chen Wankun, Intensive Digestion Technique for Diasporic Bauxite, Light Metals, 199l,p 33. 2. G. Wargalla, W.Brandt, Processing of Diaspore Bauxites, Light Metals, 1981, pp 83-100. 3. V.G.Hill,R.J.Roboson, The Classification of Bauxites from the Bayer Plant Stand Point, Light Metals,1981,p 15.

The Study of Bauxite reserves from Gujarat which falls under the Western Ghats, the Bauxite Quality is mixed. We get highly Gibbsitic Bauxites (95% Gibbsite) in the regions of Kutch where

32

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECTS OF ROASTING PRETREATMENT IN INTENSE MAGNETIC FIELD ON DIGESTION PERFORMANCE OF DIASPORIC BAUXITE Zhang Ting-an1, Dou Zhi-he1, Lv Guo-zhi1, Liu Yan1, Du Juan1, Wang Xiaoxiao1, Li Yan2 Shenyang, Liaoning 110004, China \ School of Materials & Metallurgy, Northeastern University;2. Logistics Division of Northeastern University); Shenyang, Liaoning 110004, China [email protected] Keywords: Diaspore; Digestion properties; Roasting pretreatment; Intense magnetic fields

magnetic fields. The coupling effect of a magnetic field and a temperature field is able to improve the morphology of bauxite ore and chemical speciation of the main phases, leading to activation of the ore and reducing the digestion conditions needed for the bauxite. In this paper, the effect of process conditions on digestion performance of bauxite were studied during a roasting pretreatment with an intense magnetic field. This research will provide a theoretical basis of improving digestion for the alumina industry[1~51.

Abstract This paper investigates the changes of phase and apparent morphology under the combined effects of an intense magnetic field and temperature field and the effect law of different roasting conditions on the digestion performance of roasted diaspore. The results indicated that roasting pretreatment under high magnetic fields can change the microstructure and improve the digestion properties of bauxite. The reasonable roasting conditions with intense magnetic field are as follows; the roasting temperature is 550°C, roasting time is 60min and the magnetic field intensity is 6T. The digestion rate of alumina of the roasted ore is 84.17% and the digesting liquid ratio is 1.39 while the digesting temperature is 190°C with the digestion time of 60mins. The digestion rate of alumina of the roasted ore increases to higher 52.85% than that of the raw ore , but the digesting liquid ratio is decreased to lower than 0.99 of the raw ore under the above roasting conditions. The digestion temperature of roasted ore decreases by 30°C compared to the raw ore.

Experimental Diasporic bauxite from Henan Province was used. The chemical and mineralogical compositions are shown in Table I and Table II, respectively. Mineralogical analysis of the initial sample (Fig.l) was performed on a PW3040/60 X-ray diffractometer with scanning angle from 5 to 90 degree, giving the result that the bauxite primarily consists of diaspore as well as other minor minerals such as hematite, kaolinite and anatase. 18000 • ■ *

16000

Introduction Chinese reserves of bauxite include vast quantities of diaspore with a high content of aluminum and silica and low content of iron, which at present are sub-economic due to the high levels of reactive silica that require high digestion temperatures and cause expensive losses of caustic soda during Bayer processing. At present there are many intensive leaching methods for diasporic bauxite, such as roasting pretreatment - low temperature leaching technology, mining added sweetening technology and soon. In order to promote the technology of producing alumina from Chinese bauxites, Northeastern University has presented an approach to roasting pretreatment of bauxite using intense

14000

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Chemical composition

A1203

Si0 2

Fe 2 0 3

CaO

Na 2 0

LOI

Content/%

67.57

7.14

5.54

0.90

0.31

13.85 1

33

|

100

Table II Mineralogical composition of bauxite

|

Mineralogical composition

Diaspore

Hematite

Goethite

Quartz

Anatase

Kaolinite

Content%

69.4

3.1

2.7

1.0

3.7

12.1

The solution used in these experiments is a sodium aluminate liquor. The concentrations of total alkali, caustic alkali and A1203 in the solution are about 239g/L, 221g/L and 125g/L respectively.

where Na2Ok / A1203 is the mass ratio of Na2Ok and A1203. The relative digestion rate of A1203 is expressed as:

UIS)o
Roasting pretreatment experiment The graphite crucible (Φ20 χ 100mm) containing aliquots of sample was placed into the intense magnetic field generating device which was raised to the desired magnetic flux density ranging from 1 to 12T. Afterwards, the sample was heated. The temperature ranged from 300 to 600 °C. Once the roasting temperature was reached, the sample was left in the device for a certain time ranging from 5 to 60 minutes. After the roasting time elapsed, the device was cooled down to 150 °C. When the magnetic flux density was low enough, the crucible was removed from the device. The intense magnetic field roasting device is shown in Figure 2.

where (A/S)0 and (A/S)r are the ratio of alumina to silica in raw ore and red mud, respectively. Results and discussion Effect of roasting temperature The effect of roasting temperature ranging from 300°C to 600°C on the digestion performance of treated ore was evaluated. It can be seen from Fig.3 that digestion rate of alumina increased and then decreased with increasing temperature. 550°C is considered as the optimum roasting temperature where the caustic ratio, the real digestion rate and the relative digestion rate were 1.39, 84.17% and 95.08%, respectively. It would require over 220°C for the initial bauxite to get the same level of digestion rate. When fixing the digestion temperature at 190°C, the digestion rate of the roasted ore increased by 52.38%, and the caustic ratio decreased by 0.99, in comparison to that of the raw ore which displayed a relative digestion rate of 31.32%. Consequently, the optimum roasting temperature in the intense magnetic field is 550°C.

170

Fig.2 Intense magnetic field equipment (12T) Digestion experiment The digestion experiment was carried out in a WHFS-1 autoclave. Digestion time was 60min, and the content of CaO addition was 5% by weight. After digestion, the digested slurry was separated into solution and red mud byfiltration.The concentrations of A1203 and Na 2 0 K (concentration of caustic soda) in solution and the content of A1203 and Si0 2 in the red mud were determined by chemical analysis. The calculated formula for the molar ratio of the digestion solution is: 1 *AC

Na

2°k

a, = 1.645 x- Al203 k

Roast i ng t enrper at ur el °C

Fig. 3 Effect of intense magnetic field roasting temperature on digestion performance of diaspore

(1)

34

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6

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Fig. 5 Effect of magnetic field intensity on digestion performance of diaspore

Fig. 4 Effect of intense magnetic field roasting time on digestion performance of diaspore

Mechanism analysis of pretreatment of intense magnetic field Diaspore begin to dehydrate at 500°C, which leads to the breakage of the crystal chain structure and the replacement of the coordinated octahedron of Al(0,OH)6" into a suspending state. Reaction activity of this structure is clearly much stronger. The Al3+ in the centre of the octahedron is exposed because of removal of OH" and O2" through roasting pretreatment. Because the crystal structure of the diaspore is significantly degraded, digestion performance of bauxite improved measurably. The crystal lattice of alumina transforms from the Al(0,OH)6" octahedron of a Pbnm orthorhombic structure, to an A106~ octahedron of R3c trigonal structure under the condition of roasting temperature of 550°C and roasting time of 60 minutes. The digestion conditions for pure corundum are much more demanding than that of diaspore, as the activity of the former is relatively low. Since a great deal of energy is needed to support the transformation of diaspore into pure corundum, the perfect crystal structure of pure corundum could not be generated before the roasting condition achieved. We infer that the alumina phase of the roasted ore is not a pure corundum but imperfect crystallized alumina in the transition state at roasting condition of 550°C, it is extremely beneficial for improving digestion performance, because a highly activité alumina of transition state exists in the roasted ore. XRD patterns of bauxite roasted at different temperatures are shown in Fig.6. Roasting pretreatment could improve the surface form of bauxite because of the appearance of porosity and crack, which enlarges the specific surface as measured by the method of nitrogen adsorption. The results of the specific surface of the roasted ore are shown in table 3, indicating that specific surface area reaches a maximum at a roasting temperature of 550°C, and above which, the value declines gradually with increasing roasting temperature. The increase of specific surface area enlarges the contact area between ore and liquid. So the reaction dynamic conditions are improved, and the leachability of the sample is enhanced[6]

Effect of roasting time How the roasting time varying from 5 to 60 minutes impacts on the digestion rate of roasted bauxite was investigated. The diaspore was roasted in the magnetic field with intensity of 6T and roasting temperature of 550°C. The digestion condition was 190°C for digestion temperature, 60min for digestion time, 3.1 for caustic ratio of mother liquor, 220 g/1 for caustic concentration and 5% for lime addition amount, respectively. Experimental results were shown in Fig.4. The digestion rate of alumina improved rapidly with temperature increase, which reached a maximum of 95.08% at a roasting time of 60min. Meanwhile, the caustic ratio of digested liquid achieved a minimum of 1.39 under this condition. Thus, 60 minutes is the optimal roasting time. Effect of magnetic field intensity The effect of magnetic field intensities, namely IT, 3T, 6T, 9T and 12T, on the microstructure characteristics and digestion performance of diaspore at roasting temperature of 550°C and roasting time of 30min, was investigated. The digestion condition was 190°C for digestion temperature, 60min for digestion time, 3.1 for the caustic ratio of mother liquor, 220 g/1 for caustic concentration and 5% for amount of lime addition, respectively. The experimental results are shown in Fig. 5. The digestion rate of alumina achieves a maximum, while the caustic ratio of the digested liquid reaches the lowest value, when the magnetic field intensity is 6T. Above this value, digestion rate decreases with magnetic field intensity. The reason might be that the inner crystal structure of bauxite completely transformed when the magnetic field intensity is up to 6T. With the intensity going up, the orientation of the inner molecular structure is stable and the degree of induction of various compositions to the magnetic field goes to the extreme.

35

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A-rawore ; B-300°C ; C-450°C ; D-500°C ; E-550°C ; F-600°C

Table III Effect of roasting temperature on surface area of diaspore (6T) Roasting temperature/°C

Raw ore

300

450

500

550

600

Surface area/m2-g_1

1.6205

8.7339

8.4351

38.7462

49.3558

51.0172

36

1(X

84.17% and 95.08%, respectively.

Because the magnetic susceptibility and the dielectric constant vary with specific minerals, the mineral phases in the ore will be influenced in different degrees by intense magnetic field during the roasting pre-treat process. On one hand, the value of Gibbs free energy of each phase will change, resulting in an impact on the stability of the mineral phase. On the other hand, an intense magnetic field also has an effect on the kinetics of phase transformation. The morphology of mineral phases with different magnetic properties would alter, which will affect the mineral structure and properties.

Acknowledgement This research was supported by the National Natural Science Foundation of China (NO. 51004033) References 1 . CHEN Wankun, PENG Guancai, Intensified Digestion Technology ofBauxite(Beijmg: China Metallurgy Industry Press, 1998), 112-116. 2 . WANG Yiyong et al, "Effects of Microwave Roasting on Leaching Behavior of Diaspore Ore", The Chinese Journal of Process Engineering, 2(2004), 317-321. 3 .R A Hind, K S Bhargava, C Stephen, "The Surface Chemistry of Bayer Process Solids: a review", Colloids and Surfaces A: Physicochemical and Engineering Aspects, 1-3(1999), 359-374. 4 . Guo Xianjian et al, "Microwave-assisted digestion of diasporic bauxites", Nonferrous Matels, 4(1995), 55-57. 5 . Zhao Qingjie, "Investigation of new digesting process of diaspore", Light Metals, (1)2000,17-21 6 . Lv Guo-zhi et al, "Roasting pretreatment of high-sulfur bauxite and digestion performance of roasted ore", The Chinese Journal ofNonferrous Metals, 9(2009), 1684-1689. 7 . Wang Xi'ning et al, "Effects of Magnetic Field on Solid Phase Transformation", Materials Review, 2(2002), 25. 8 . C C Koch, "Experimental Evidence for Magnetic or Electric field Effects on Phase Transformations", Mater Sei Eng, 2000, 213.

For the roasting of bauxite, proper temperature increases the internal energy, accelerates the molecular thermal motion and leads to lattice distortion and phase transition. So during the process, the energy of the magnetic field will have great influence on specific magnetic species and on the direction of thermal motion of molecules. It increases the internal active energy, the degree of phase instability, the lattice distortion and makes the phase transition less complete, so the leaching ability of the ore is enhanced. When the intensity of the magnetic field reaches 6T, the temperature field and magnetic field interact with each other. The digestion rate decreases with the further increase of temperature. There are several possible reasons to explain this phenomenon. When the intensity of magnetic field is too high, the atoms and molecules will be oriented in a certain range according to the direction of the magneticfield[7~8].So this will hinder the lattice distortion and the transitions of the crystal structure. This decreases the activity of roasted ore and leads to a decrease in digestion rate. So there should be an optimum magnetic intensity, rather than an infinitely high value. Conclusions (1) The digestion performance of diasporic bauxite is enhanced effectively by a roasting pre-treat in the presence of a magnetic field. After pre-treatment under appropriate conditions, the observed leaching temperature decreases by at least 30°C compared with the raw ore. (2) The mechanism of the effect of intense magnetic field in the pre-processing is that the influence of magnetic field varies with mineral phase, due to the changes of magnetic susceptibility and dielectric constant for different minerals. It will affect the value of Gibbs free energy of different phases as well as the stability of the ore. When the intensity of magnetic field is high, the atoms, and molecules will be oriented in a certain range according to the direction of the magnetic field. In higher ranges, the activation effect is weak. (3) The optimal conditions of intense magnetic field pre-treatment are determined as temperature of 550°C, roasting time of 60 minutes, magnetic intensity of 6T. When the roasted ore was digested in the sodium aluminate solution with a mole ratio of 3.1 and alkali concentration of 220g/L at 190°C for 60 minutes, the mole ratio of digestion solution decreases to 1.39 and the absolute and relative digestion leaching rates are

37

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINA and BAUXITE Bayer Process I SESSION CHAIR

Peter-Hans ter Weer TWS Services and Advice Netherlands

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

APPLICATION OF OPERATION INTEGRITY MANAGEMENT IN THE ALUMINA INDUSTRY Carlos Suarez1, Daniel Welshons1, John McNerney2, Jim Webb2 ^atch Associates Consultants Ine; 1600 West Carson Street, Pittsburgh PA 15219, USA 2 Warren - Forthought Ine, 1212 North Velasco, Suite 207 Angleton, Texas 77515, USA Keywords: Process Safety, Operations Integrity Management, Process and Operations Improvement strategy and implementation of such a system in an alumina refinery environment.

Abstract In today's economic environment the Safety of our industry assets - People, Equipment and Processes - have become even more demanding. This paper provides means to apply Operation Integrity Management System (OIMS) to the Alumina Industry. It discusses what OIM is about, its implementation parameters such as cost and time, as well as an effective way to integrate such a system into refineries day to day operations. A series of Key Process Performance Indicators (KPPIs) are presented as well as an Electronic Knowledge Support System (EKSS) Mockingbird ® as a tool to support the implementation of OIMS.

What is OEM Operational Integrity Management (OIM) consists of an integrated set of theories, practices and techniques for ensuring that an industrial facility operates "with integrity", i.e., that the facility's performance is what it should be - no more, no less. A Brief History of Integrity Management The path of integrity management systems began in the United States, where in 1992 the Occupational Safety and Health Administration (OSHA) introduced its Process Safety Management Program. The goal of its 14 strands was to prevent or minimize the consequences of catastrophic accidents caused through the release of chemicals. The program requires a holistic approach that integrates technologies, procedures and practices, creating multiple barriers of protection.

Introduction Alumina refinery operators can use Operations Integrity Management Systems (OIMS) to improve their Process Safety and Operations Performance. Throughout this paper the authors will refer to Process Safety Management (PSM) as an equivalent system adopted by the Highly Hazardous Materials chemicals and petrochemicals operations.

For the oil and gas sector, the impetus behind the development of integrity management systems came from the Piper Alpha disaster in 1988, in which 167 people lost their lives. The inquiry into the disaster produced the Cullen Report, where the primary recommendation was that operating companies should be required to implement safety management systems that ensure safe design and operation of offshore installations.

Alumina refinery safety in the USA is mainly regulated by the Mining Safety Health Administration (MSHA). This branch of the Department of Labor focuses on employee's health and safety aspects related to surface and underground mining operations. The Occupational Safety and Health Administration (OSHA); another branch of the Department of Labor, regulates chemical and petrochemicals operations. It is important to mention this difference since OSHA regulates process safety activities and systems through PSM while MSHA does not. It is up to alumina refineries management to adopt PSM or OIMS elements to cover the safety of their process, their employees and their surrounding communities.

The report specified that such system should draw on quality assurance principles similar to ISO 9000. The Cullen Report's recommendations were accepted immediately by the British government and the new regime that resulted has influenced the development of integrity management systems around the world. The following have been reported as precursors to the Piper Alpha incident:

As the different elements of OIM are presented you will realize that a lot of them already form part of your refinery ISO, Health, Safety and Environmental systems. This makes it easier for the integration, implementation, cost and time reduction to deploy such a system.

• • •

Corporate Pride and Craftsmanship Complaisency rather than Competency Change to the Rules

The initial response by the industry included the following:

The success in implementing OIMS depends among other things on how quick the refinery culture embraces and supports it.

• • •

Once the elements are in place, the deployment, auditing and further improvements can be enhanced by using an Electronic Knowledge Support System. Mockingbird ® and its suite of applications have been successfully used by the chemical and petrochemical industry over the years. This paper discusses the

•

41

Mitigation of smoke hazards Installation of sub-sea pipeline isolation systems Improvements to the "Work Permits" management system Relocation of some pipelines emergency shutdown valves

that they perform in a manner that is consistent and compatible with Company's policies and business objectives. Evaluate and train contractors before starting any work.

The Elements of OEM Below is a brief description of the OIM elements shown in figure 1. 1.

Management establishes policy, provides perspective, sets expectations and provides the resources for successful operations. Assurance of Operations Integrity requires management leadership and commitment visible to the organization, and accountability at all levels. The adoption of such system is driven by the General Manager and his/her team of Functional Managers.

2.

Comprehensive risk assessments can reduce safety, health, environmental and security risks and mitigate the consequences of incidents by providing essential information for decision-making. Mitigating risks at the early stages of design and the continuation in the operation phase is a must.

3.

Inherent safety and security can be enhanced, and risk to health and the environment minimized, by using sound standards, procedures and management systems for facility design, construction and startup activities.

4.

Accurate information on the configuration and capabilities of processes and facilities, properties of products and materials handled, potential Operations Integrity hazards, and regulatory requirements is essential to assess and manage risk. Process Safety Information such as P&IDs, Materials Safety Data Sheets should reflect current operation.

5.

6.

7.

8.

9.

10. Effective management of stakeholder relationships is important to enhance the trust and confidence of the communities where the business operate. Emergency planning and preparedness are essential to ensure that, in the event of an incident, all necessary actions are taken for the protection of the public, the environment and company personnel and assets. Make it a point to run simulations and drills periodically so that everybody involved react according to plan. 11. Assessment of the degree to which expectations are met is essential to improve Operations Integrity and maintain accountability. Involve all levels of the organization in routine audits. Make it a point to discuss results and focus on corrective actions. OIMS11 Elements Driver

Operations

Evaluation

2. Risk Assessment and Management 1. Management Leadership. Commitment and Accountability

Control of operations depends upon people. Achieving Operations Integrity requires the appropriate screening, careful selection and placement, ongoing assessment and training of employees, and the implementation of appropriate Operations Integrity programs. Apply Management of Change when dealing with new personnel assignments, particularly in the plant areas.

3. Facilities Design and Construction 4. Information / Documentation 5. Personnel and Training 6. Operations and Maintenance

11. Operations Integrity Assessment and Improvement

7. Management of Change 8. Third Party Services 9. Incident Investigation and Analysis 10. Community Awareness and Emergency Preparedness

Operation of facilities within established parameters and according to regulations is essential. Doing so requires effective procedures, structured inspection and maintenance programs, reliable Operations Integrity critical equipment, and qualified personnel who consistently execute these procedures and practices. Make sure that Standard Operating Procedures for both Process Operations and Maintenance reflect current systems. Changes in facilities, or managed to arising from level.

Effective incident investigation, reporting and followup is necessary to achieve Operations Integrity. They provide the opportunity to learn from reported incidents and to use the information to take corrective action and prevent recurrence. Monitor and request reports from any incident, no matter how small or insignificant to the perception of the affected.

Ret·, e« mUoM

Figi. OIM 11 Elements Refinery Culture and How to Promote OIM Cultural Attributes 1. Culture is a feature of the entire organization, not just of some of the individuals within that organization. Therefore, if someone — even the general manager — leaves the organization, the culture of that organization should not change significantly.

operations, procedures, site standards, organizations must be evaluated and ensure that Operations Integrity risks these changes remain at an acceptable

2. Culture is on-going — it is not a one-time event. A facility in which everyone is continuously striving to identify and correct problems and to eliminate hazardous conditions has a strong operational integrity culture, whereas a facility which makes only

Third parties carrying out work on the company's behalf impact its operations and its reputation. It is essential

42

Loss of market share is reduced — After an incident, this loss continues until the company's reputation is restored. Adverse publicity and negative public image can have insurmountable effects

spasmodic and irregular efforts to improve such conditions does not. 3. In a strong OIM culture there is minimal disconnect between words and actions. All managers and workers 'walk the talk'; their words and deeds match.

Litigation costs are reduced — These are unavoidable after an incident and can total five times the cost of the regulatory fines.

4. The creation and maintenance of an organizational culture requires leadership from the top. Allowing lower level employees to "do their own thing" does not create a culture.

Incident investigation costs are reduced — Investigating an incident and implementing corrective actions can cost millions of dollars

5. It is difficult for any organization to truly assess the quality of its own culture. It takes an outsider to truly evaluate the quality of a company's culture. Therefore, an organization with a strong OIM culture will make frequent use of outside auditors, inspectors and reviewers to identify areas of weakness and to suggest corrective actions. Moreover, the auditors' reports will go directly to the facility managers

Regulatory penalties are reduced — For many incidents, a fine after litigation can total 1 million dollars or more Regulatory attention is reduced — A major incident usually results in increased regulatory audits and inspections Key Performance Indicators

6. A strong OIM culture is one in which employees and contract workers feel free to report on difficulties and problems, even if those employees and workers are potentially opening themselves up to criticism

Examples of KPIs for Operation Integrity Management are shown infigure2 below.

7. With regard to SHE (Safety, Health and Environmental) issues, the organization places excessive emphasis on the safety term, to the detriment of the health and environmental elements.

% P&IDs Conformance to Current Process Installations As Built Drawings Available and in Conformance Facilty Design and Built per Sound Standards % of Critical Operational and Maintenance Procedures Reviews Completed to Schedule % Compliance with Critical Procedures Material Safety Data Sheets (MSDSlAvailabilityto^^ Mandatory Training Completed to Schedule Personnel Trained on New Standard OperatingProcedures

Prepare and publish a Mission Statement that spells out the organization's stated commitment to operational integrity management principles.

2.

Develop guiding tenets that show how the OIM program is to be implemented.

3.

Develop a detailed program showing how the guiding tenets are to be achieved.

S of Risk Assessment Corrective Actions Completed to Schedule % of Risk Assessments Reviewed to Schedule

man—-

8. A strong operational integrity management culture adapts to new circumstances without its basic values being affected by issues such as economic downturns or the adoption of new technologies. It is suggested here that management can go about creating a strong operational integrity culture by following the three steps shown below: 1.

Leadership Participation in Incidents Investigation Participation in OIM Program Assessment OIM Contribution as Part of the Employees Performance Assessment

EEEEBEEEEEEa All Critical Controls for Process Safety Identified \ of Controls Inspected to Schedule ft of Controls Outside Tolerance % Compliance with Critical Procedures

% of MOC Documents Compliant with Procedures % of Temporary Changes Overdue % of MOC Physically Installed but Awaiting Completion of Documentation

llll Γ

^ ^ ^ ^ ^ ^ Μ ^ — ■

Assessment of Capabilities to Performed Work Deficiencies Corrected Effective Communication

Ü

OIM provides unparalleled capacity for enhanced risk reduction. A company's risk exposure is reduced in the following areas when well-founded process safety systems are in place.

% of Overdue Incident Investigations No of Repeat Incidents Occuring % of Follow Up Corrective Actions Completed to Schedule 1 Lessons Learned from Company and Industry Incidents

a ■ BE ',- m i mm jjuLlma-

No of Emergency Exercises /Desktop Exercises Completed to Schedule Emergency Plan Reviewed to Schedule

Lives are saved and injuries are reduced — Both the personal impact of human loss and cost of deaths or injuries are painful. A solid OIM program can help prevent these costs

% of Inspections or Tests Completed to Schedule % Compliance with Standards and Procedures

Figure 2 - KPIs for OIM Elements

Property damage costs are reduced — In the U.S., major industrial incidents cost an average of $80 million each

OEM Implementation Parameters

Business interruptions are reduced — These losses can amount to four times the cost of the property damage from an incident

43

To succeed in business a company must: • • •

Protect its license to operate Meet ever more demanding regulatory requirements Manage the sustainability of your business

• •

•

Raise stakeholder and public confidence Minimize and, where practicably possible, eliminate the risk of incidents

The cost areas for years 1-5 includes (1) the remaining cost to reach 100 % compliance and (2) the ongoing cost to maintain compliance (or quality) for the remaining years if the company reaches 100 percent compliance in that period.

How to Implement an OIM System • • • • • • • •

The cost for developing the program is described below:

Assign an OIM Manager (or team) Learn from the literature (check key references) Learn by training (from process safety professionals) Learn from other companies - align; network; participate in industry alliances Note strong synergies with ISO, TQM, RMP, Responsible Care Set some clear OIM goals (one tofiveyears) Track performance versus the goals on a regular basis Reassess OIM/plan & modify (every 3-5 yrs)

Developing an OEM Program. The cost, primarily in equivalent labor costs, to bring the OIM program (and individual element programs) from the concept stage through the final design (such as developing an MOC or MI written program that the facility personnel are confident will work). This category also includes the cost of training personnel to be proficient in various OIM activities, such as leading PHAs, leading incident investigations, leading compliance audits, writing procedures, and leading employee training.

The best OIM companies show the following attributes: •

OIM Champions who affiliate themselves with multiple disciplines (e.g., EHS, Engineering, Operations, Insurance) to work collaboratively

•

Functionally having two platforms to identify, analyze, select, implement, control & monitor process; i.e., worst case (top down) and more frequent events (bottom up)

•

Regularly conduct reviews of the OIM program against the defined elements, and

•

Implementing an OEM Program. The cost (again primarily in equivalent labor costs) to do implementation tasks, such as writing operating procedures, updating PSI, doing initial training of operators and maintenance personnel, and performing/documenting Process Hazard Analysis - PHAs. Responding to Recommendations. The cost, primarily capital costs and expenses, to implement improvements to address recommendations from PHAs, MOC hazard reviews and incident investigations. As a rule of thumb $22,000 - $25,000 per P&ID might be used as the cost associated for the development and implementation of an OIM system.

Especially promote risk engineering in the conceptual engineering design phase.

Mockingbird ® (EKSS)

Time

Mockingbird ® is an Electronic Knowledge Performance System widely used by the Chemical and Petrochemical industry to support OIM / PSM systems development, implementation and manage compliance.

Research has shown that the following level of compliance can be attained after the implementation of an OIM system: • • • •

Initial implementation

40% level of compliance First Year (Baseline) 50% level of compliance Second Year 100% level of compliance anticipated in Fifth Year Excellence in OIM program anticipated Seventh Year

The initial population of the system takes place during the "Miracle Month". This period is used for training on the system, designing its structure and have all involved take ownership. Mockingbird ® becomes the portal for all OIM elements and its tools.

Time is dependent on how many documented similar elements the business already have in place in the organization that can be brought into OIM or easily adapted.

.incMocKingbirdd

Cost

MocKingbird® lTts^u

The labor cost for developing and implementing an OIM element can be accounted for in one or more of the following categories: Meetings Writing Reviewing Revising Training/Orientation Pilot testing More revising

44

Conclusions •

~ΤΛ, , ·ι · ι ,Λι · r * OIM can be easily implemented m any alumina refinery for which other systems such as ISO, Safety, Health and Environmental already exists.

•

OIM can be a combination of the elements listed in this paper and those defined for PSM

•

The successful implementation and continuity in the use of OIM depends on the organization commitment to support it

•

It is important that a strong safety culture be promoted and nurtured

•

The existence of other systems such as ISO, Safety, Health and Environmental, etc will make the OIM system implementation less costly and shorter in time

•

Refineries that reach an excellence level of compliance with OIM will also show an overall improvement in financial and operating levels

•

Mockingbird ® and its application suites offer a sound and robust platform on which to manage OIM

CCPS, 'Guidelines for Auditing Process Safety Management Systems", -g ' .- β *. „ (μ« 169-0556-8 ' * '

Further Readings Exxon Management Systems h{tp;//w\vw.
Exxon Mobil, Operations Integrity Management System CCPS, "Guidelines for Technical Management of Chemical Process Safety", 1989, ISBN No, Ö-8169-0423-5. CCPS, "Plant Guidelines for Technical Management of Chemical Process Safety", 1992, ISBN No. 0-8169-0499-5. American Petroleum Institute, Recommended Practice 750 "Management of Process Hazards", 1990, reaffirmed 1995. Available via the API at www.api.org Ü.S. Occupational Safety and Health Administration, 29-ŒR-1910.119, "Process Safety Management of Highly Hazardous Chemicals". Canadian Chemical Producers' Association (CCPA), Responsible Careen codes of practice. Available on the CCPÄ website www.ccpa.ca American Institute of Chemical Engineers, "Daw's Fire and Explosion Index Hazard Classification Guide", latest edition, ISBN No. 0 4 1 6 9 4 6 2 3 - & American Iristttute öf Chemical Engineers, "Dow's Chemical Exposure index", latest edition, ISBN No. 0-81694647-5. Kietz, T.Ä., "An Engineer's View of Human Error", London; The Institution of Chemical Engineers, third edition 20Û1, ISBN No. 1 -5603-291Ö-5 (available from CCPS). CCPS, "Guidelines for Preventing Human Error in Process Safety", 1994, ISBN No. Ö4169446V8. Health and Safety Executive (UK), "Human Factors in Industrial Safety", HS(G)48, London HMS0,1989, ISBN No. 0-11-6854864.

45

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INFLUENCE OF SOLIDS CONCENTRATION, PARTICLE SIZE DISTRIBUTION, pH AND TEMPERATURE ON YIELD STRESS OF BAUXITE PULP Carla Napoli Barbato1'2, Silvia Cristina Alves Franca2, Marcio Nele de Souza1 1-Escola de Quimica/ Universidade Federai do Rio de Janeiro- Av. Horacio Macedo, 2030. Edificio do Centro de Tecnologia, Bloco E. Cidade Universitaria. Cep: 21941-909. Rio de Janeiro - RJ 2-Centro de Tecnologia Minerai- Av. Pedro Calmon, 900. Cidade Universitaria. Cep:21941-908. Rio de Janeiro- RJ Keywords: yield stress, zeta potential, van der Waals forces Thus, the purpose of this work is to study the effect of solids concentration, particle size distribution produced by different grinding time, pH and temperature on the yield stress of bauxite pulp.

Abstract In Northern Brazil, bauxite pulp is transported through pipelines to the plant where alumina is produced. In slurry transportation through pipelines, knowing the yield stress value is essential for the pumps and pipeline design. Yield stress is the minimum shear stress and corresponds to the first evidence of flow. This rheological property is influenced by factors, such as: particle form, temperature, particle size distribution and interaction among the particles. Within the context above, the objective of this work is to verify the influence of solids concentration, particle size distribution produced by different grinding time, temperature and pH on the yield stress of bauxite pulp. It was verified that the yield stress of bauxite slurry increases as solids concentration and grinding time go up and decreases with temperature and pH.

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Brazil is a country with large territory and its mining activity usually takes place far from the production or consumption centers. Ore transportation by rail or road is expensive and lacks adequate infrastructure, which makes its transportation in the form of pulps through pipelines an attractive alternative.

J*

*

-

The first bauxite slurry pipeline in the world is located in the North of Brazil and it aims at transporting the aluminum ore from the processing plant, located in Paragominas (PA), to the alumina production plant in the municipality of Barcarema (PA) through a 244 km long pipeline [1]. The mine and the pipeline are located in the State of Para, Brazil, as shown in Figure 1.

φ Figure 1. Map of Brazil and pipeline showing location of Mine and alumina plant.

In slurry transportation through pipelines, it is important to get information on yield stress in order to carry out the pumps and pipeline project. Furthermore, high values of yield stress can retard solid particles from gravity settling for long periods of time [2].

Materials and Methods The bauxite used in this work comes from North of Brazil and it is basically composed of the minerals gibbsite and kaolinite. The dressing process followed the methodology described by [6] and the steps consisted of crushing and grinding.

Yield stress is the minimum value of shear stress and is the first evidence of flow, i.e., the value of shear stress when the velocity gradient tends to zero [3,4].

The particles size distributions studied were based on industrial conditions in pipeline transportation of bauxite located in Para State, in which the particle size distribution is 6% plus 0.208 mm and 40-47% minus 0.043 mm [1]. The particle size distribution of the bauxite sample ground for 30, 35 and 40 min in a rod mill was determined through wet screening with the use of a set of Tyler sieves with openings from 1.2 to 0.037 mm.

Yield stress is related to the internal structure of the suspension. For this structure to begin to flow it is necessary to overcome the minimum shear stress necessary to break the interparticle contact and the unions of aggregates or floccules. This structure is formed by particles whose electric charges are of opposite signs. The interparticle force of attraction of this structure is of the van der Waals type. The variables which influence yield stress are: solids concentration, particles size, shape, pH and the nature of the materials [5].

The zeta potential was determined in a DT 1200 equipment manufactured by Dispersion Technology. The suspensions were prepared with 10% bauxite and 0.01 M KC1. The zeta potential

47

measurements were carried out in a pH range of 2.0 to 12.5. The pH was adjusted with diluted solutions of KOH and HC1

Figure 2 illustrates the size distribution of bauxite ground for 30, 35 and 40 min, obtained through wet screening with a set of Tyler sieves with openings from 1.2 to 0.037 mm. It was verified that the bauxite ground for 30 min presented a size distribution similar to that of the bauxite transported in the pipeline located in the North of Brazil. The bauxite ground for 40 min had a higher amount of fine particles (<0.037 mm) than the bauxite ground for 30 and 35 min, at approximately 13 and 6%, respectively.

In order to study the influence of solids concentration, particles size distribution produced by different grinding times, temperature and pH on the yield stress of bauxite slurry, a matrix of experiments 24"1 around mean values was performed (Table 1). The solids concentration studied was 50 and 60% (w/w), the grinding time of bauxite was 30 and 40 min, the pH studied was 7 and 12 and the temperature was 25 and 45°C.

io(H

Table 1. Experimental conditions following matrix experiments 24"1 around mean values.

90

T GT SC pH (min) (°Q (% w/w) 1 25 30 7 50 2 45 30 60 7 3 45 40 50 7 4 25 40 60 7 12 5 45 30 50 12 6 25 30 60 7 25 40 50 12 12 8 40 60 45 9 35 35 55 9.5 10 35 35 55 9.5 11 35 55 9.5 35 T- Temperature; GT- Grinding Time; SC- Solids Concentration Run

80

60

0.1

K=

*D,

Figure 2. Size distribution of bauxite ground for 30, 35 and 40 min obtained through wet screening. The zeta potential is useful to evaluate the electrical double-layer repulsive forces among particles in suspensions, as charge density and the potential of the particles depend on pH and ionic strength. Figure 3 illustrates the zeta potential of bauxite in relation to pH. It can be observed that the point of zero charge occurs at pH = 10.5. At this pH the degree of particle flocculation is maximum. In the pH range between 7 and 11, the zeta potential is small. It is expected that in this range particle flocculation is high because the density of charge on the particles surface is small, which results in high attraction due to van der Waals forces. At pH 12 the zeta potential is higher, which means that the density of charge on the particles surface is large and the repulsive forces among particles are greater than the attraction forces of van der Waals and, consequently, the degree of flocculation is lower.

(1)

H J_ A,

+

(2)

3

A suspension will be stable when the repulsive forces dominate while the presence of strong attractive forces will cause particle aggregation. Thus, the suspensions prepared at pH 12 are more stable than the suspensions prepared at pH 7 and 9.5.

Tm - measured torque Dv - vane diameter xy = shear stress H= vane height Equation 03 was used to adjust the yield stress values of the bauxite pulps, having as its variables: temperature, grinding time of bauxite, solids concentration and pH. Y=

a0+£aiXi+'£aijXiXj

1

Particle Size (mm)

Measurements of yield stress were obtained from an ARES rheometer, manufactured by TA instruments. A geometry Vane was used. The tests were performed at 1 rpm for 200 s. Torque values were converted to shear stress by means of Equations 1 and 2. Yield stress is the maximum shear stress obtained at low speeds.

-TJK

Bauxite ground for 30 min Bauxite ground for 35 min Bauxite ground for 40 min Bauxite transported along the pipeline

(03)

With Y as the dependent variable (yield stress), Xi the independent variables (temperature, grinding time of bauxite, solids concentration, and pH) and ^ and a^ as the parameters. Results and Discussion

48

H

Figure 3. Zeta potential of bauxite particles.

"

■

The model parameters for yield stress used normalized independent variables [+1,-1] so that the values of the parameter could be associated with the variable effect. Estimates for model parameters were obtained with standard linear regression procedures [7]. Standard statistical tests of significance (t-test of student) were used to allow the evaluation of parameter significance. Whenever parameter significance was smaller than 5%, the parameter and respective effect would be removed from Equation 3. Regression results obtained for the yield stress determination are shown in Table 2.

s;

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7

8

1

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1 10

1 11

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13

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1.964 + 0.064 1.232 + 0.082 0.752 ±0.082 -0.038 + 0.082 0.073 +0.082 0.448 +0.082 0.188 ±0.082 -0.023 ±0.082 0.946

Figure 4. Comparison between (■) observed and (O) predicted yield stress.

aoH aT asc*GT asc*oH asc*T R' SC-Solids Concentration; GT - Grinding Time and T-Temperature

Figure 5 illustrates the values of yield stress from the bauxite slurry obtained under different conditions of solids concentration, grinding time, temperature and pH. It was verified that the experimental conditions that favored a higher value of yield stress of the bauxite pulp was 60% (w/w) solids, bauxite ground for 40 min, 25°C and pH 7.

** Significant effects in bold 5-

It is possible to observe from the correlation coefficient (Table 2) and the model fitting (Figure 4) that the model gave a good fit to the experimental parameters. The value of R2 is larger than a confidence interval of 95% indicating a very good fit. The model relating the yield stress (τ0) and the statically signification preparation variable is:

4-

3CD CO

τ0 = 1.964 +1.232 *SC + 0.752 *GT- 0.448 *SC*GT

(04)

m

Fh

2-

τ

IiI I -

It may be concluded, by observing the parameters values in Table 2, that the yield stress depends mainly on solids concentration and the grinding time of bauxite and these variables show a direct relationship to yield stress. The remaining variables, pH and temperature, are not statistically significant. This means that the effects of temperature and pH are negligible compared to the yield stress measurement error.

ΓΤ 1

m

m

T

m

m

n ..

1

2

3

4

5

6

7

Firn 10

m

11

Experimental Run

Figure 5. Values of yield stress of bauxite pulp obtained under different conditions of solids concentration, grinding time, temperature and pH. It can be stated that yield stress increases when solids concentration and grinding time of bauxite increase and when temperature and pH decrease. This behavior can be explained by the following: 1. The increase of solids concentration results from the decrease of the water layer between particles, so the inter-particle distance is smaller in a denser slurry which produces an increased attractive potential and a larger probability of collisions among particles, resulting in more particles attracting one another and a high close-packing of particles [8,9,10].

2. The increase of bauxite grinding time allows a bigger production of fine particles (< 0.037 mm) which accentuates the strong particle-particle interaction due to the van der Waals forces of attraction, forming floes or aggregates of low packing density and the formation of chain or networks of particles (Figure 6) that reduces the fluidity of the suspension. For the suspension to flow, it is necessary to apply a greater force to break the packing particles network, such as that shown in Figure 7. In this Figure the state of the network of very voluminous particles packing is similar to a "sponge".

The yield stress of the bauxite slurry increases with increase of solids concentration, grinding time of the bauxite and with decreasing temperature and pH. Acknowledgements The authors would like to thank CAPES for their financial support, the School of Chemistry / Federal University of Rio de Janeiro and the Center for Mineral Technology. References

3. When the pH is between 7 and 11, the zeta potential is low which means a high charge surface density and strong van der Waals attraction forces among the particles and the formation of floes. At pH 12, the zeta potential is higher and the yield stress is low because the net attractive force decreases due to the increase of the surface charge density. When this repulsive force exceeds the van der Waals attraction at a pH far from the IEP, dispersed suspensions are usually generated (Figure 7).

1. Gandhi, R. Weston, M. Talavera M. Brittes G. P. & Barbosa E."Design and Operation of the World's First Long Distance Bauxite Slurry Pipeline". Light Metals, 2008, 95-100. 2. Dzuy, N.Q. e Boger, D. V., "Yield Stress Measurement for Concentrated Suspensions", Journal of Rheology, 27 (1983), 321349.

4. The yield stress increases as temperature decreases because the viscosity of dispersant phase (water) increases.

3. Liddell, P. V.e Boger D. V., "Yield Stress Measurements with the Vane", Journal of Non-Newtonian Fluid Mechanic, 63 (1996), 235-261. 4. Stokes, J. R. e Telford, J. H., "Measuring the Yield Behaviour of Structrured Fluids", Journal of Non-Newtonian Fluid Mechanics, 124 (2004), 137-146. 5. Alejo, B. and Barrientos, A., "Model for yield stress of quartz pulps and copper tailing", International Journal of Mineral Processing, 93 (2009), 213-219. 6. Silva F. A. N. G. Medeiros M. E. Sampaio, J. A. Santos R. D. Carneiro M. C. Costa L. S. Garrido F. M. S. "Technological Characterization of Bauxite from Parâ-Brazil". Light Metals, 2009,139-144.

Figure 6. Configuration of maximum packing fraction, particles forming networks of particle-particle contact (floes or aggregates) due to the van der Waals forces of attraction.

7. Box, G. E. P., Hunter, W. G., Hunter, J. S., Statistics for Experiments (New York, NY: Wiley, 1978). 8. Nuntiya, A., Prasanphan, S., "The Rheological behavior of kaolin suspensions", Chiang Mai Journal Sience, 33 (2006), 271281.

OOOOO O OOOOOO

9. He M. Wang Y. Forssberg E. "Parameter Studies on the Rheology of Limestone Slurries" International Journal of Mineral Processing, 78 (2006), 63-77.

OOOOO O OOOOO O

10. He, M., Wang, Y. e Forssberg, E., "Slurry Rheology in Wet Ultrafine Grinding of Industrial minerals: a Review", Powder Technology, 147 (2004), 94-112.

Figure 7.Configuration of minimum packing fraction; disperse particles due to repulsive forces.

11. Krester, R. G., Scales, P. J. "The effect of temperature on the yield stress of mineral suspensions". J. Colloid and Interface Sci., 328 (2008), 187-193.

Conclusion A bauxite slurry prepared at pH 12 is more stable than that prepared at pH 7 because the bauxite particles have a higher surface charge at pH 12. Thus, the repulsive forces between particles are higher at this pH. The parameter that has the greatest effect on the yield stress of the bauxite pulp is solids concentration, followed by grinding time, pH and temperature.

50

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

A NEW METHOD FOR REMOVAL OF ORGANICS IN THE BAYER PROCESS Bai Yingwei1, Shen Mingliang1, Li Junqi2, Zhao Fei2 1. Guiyang Aluminum Magnesium Design & Research Institute, Guiyang, Guizhou, 550081; 2. Materials Science and Metallurgy school,Guizhou University, Guiyang. Guizhou, 550003 Keywords: Bayer process, Organics, Super-concentration, Carbonate, causticization refineries with the Bayer process. The impurities, especially Na20 content in the product will be increased, and this is not good for aluminum hydrate settlement, filtration and classification, which results in generating foam in the settlers, reducing the alumina whiteness, and accelerating the aluminum hydrate scaling in the seed precipitation tanks.

Abstract This article introduces briefly the main organics sources in Bayer process alumina production and the harm due to accumulation during process flows, specifies the organics removal methods applied generally in world alumina refineries, and explains briefly advantages & disadvantages of each method. The article stresses the super-concentrated organics removal method developed successfully by GAMI with Guizhou University together, which not only has good organics removal effects (total removal percent of 57-66%), but also effectively removes the carbonates in the process flow.

(3) It also has effects on the evaporation of mother liquor. The organic impurities can make Na2Oc concentration in the mother liquor increase (make Na2C03 in the liquor supersaturated and concentration increased), the organics content is higher and this phenomenon is more obvious. The organics makes the liquor viscosity increase and the size of the precipitated monohydrate sodium carbonate fine, resulting in difficulties with settlement and filtration & separation. The production practice of Guangxi Branch Company of China Aluminum Company estimates that: the oxalate is easy to be precipitated from the evaporator and flash evaporator when the sodium oxalate concentration in the liquor is high, having the effects on the evaporation and increasing the scaling cleaning workload.

1. Introduction The organics in the Bayer liquor arise mainly from the bauxite. A few of the organics are also brought in from the flocculants, defoaming agents and dewatering agents, but with lesser amounts and smaller effects, according to the literature reports. The organic carbon content in the bauxite is generally 0.1-0.3%, but also can be lower to 0.03% or higher to 0.6% (some surface mines). The organic carbon content in tropical bauxite is higher, generally 0.20.4%; however, that in diasporic bauxite is lower, generally 0.1%. The organics content in the bauxite in Poncho, Africa and Australia is higher and that in Europe, Russia and China is lower [1].

(4) The pitch organics often attach to the surface of filter cloths and heat exchangers, which has effects on the filtration and heat exchange. Because the organics have some serious effects on the alumina production process flow mentioned above, alumina refineries in the world adopt varied methods for the removal of organics from the process flow. Two kinds of methods are applied in the actual production: one is to remove the organics from the liquor, mainly by spent liquor calcination, adsorption, deposition etc; the other is to totally or partly damage the organics in the liquor, mainly by the oxidization which is to partly or totally oxidise the organics to Na2C03. At present the main application methods are as follows [2" 61 : adsorption, oxidization, calcination, aluminum hydrate washing liquor settlement, crystallization etc.

There are three main kinds of organics in the Bayer liquor: natural matter (such as roots), humic acids and pitch [1] formed by precipitation of the organics (plants) and microorganisms with the chemical reaction. These organics undergo some changes during the alumina production flow, and finally become humic acids, benzenecarboxylic acids, oxalic acid etc. The organics content in the production process accumulates step by step along with the progress of production, and harms the Bayer alumina production process when it reaches a certain value. The main effects are as follows:

2. Test principle and test method

(1) It has effects on the digestion and red mud separation & washing [1]. Along with the increase of the organics content, the liquor viscosity will increase, the red mud settlement speed will be reduced and the overflow suspended matters in the settler will increase.

2.1 Test principle The super-concentrated causticization method this article puts forward is different from the traditional organic removal methods. The principle is as follows: according to the characteristic that the solubility of oxalate in the alkaline liquor is reduced, along with the increase of the alkali concentration in a certain alkali concentration range, the liquor required for organics removal is concentrated to a higher concentration (super-concentrated) so as to precipitate the most oxalate from the liquor by crystallization,

(2) It has effects on the crystal seed precipitation[l]. The precipitation speed and the alumina output rate shall be reduced, which results in the alumina size becoming fine and fragile, and especially fragile during calcination. This problem becomes one of the difficulties with producing sandy alumina in the alumina

51

then the precipitated oxalate is translated into the insoluble compound by a causticization reaction so that the compound is removed from the process flow by washing and filtration. Therefore, the effect of factors on the solubility of oxalate in the liquor and the optimum conditions of causticization reaction is required to be found by testing, moreover, the authors have also studied the solution condition of the causticization slag in the liquor so that this method can be better applied to the industry production.

Effect of crystal seed addition on organics removal percent The optimum super-concentrated conditions of organics removal are adopted for concentration and slag making. The concentrated slag is dried and milled to be fine for 1-section causticization test, using the orthogonal method. The chemical compounds in the liquor and solid after causticization and separation, is analyzed and the organics removal percent is calculated so as to determine the optimum 1-section causticization conditions.

In the test: Nk expresses the caustic sodium oxide concentration in the liquor (NaAJ g/1

•

The 2-section causticization treatment with different causticization conditions is done based on the 1-section causticization. The organics removal condition after 2section causticization separation is analyzed so as to determine the optimum 2-section causticization conditions.

•

The anti-solution test of the slag obtained after causticization is done so as to study the anti-solution condition of the organics in the caustic slag during the washing.

Nc expresses the carbonate sodium oxide concentration in the liquor (Na2Oc) g/1 NT expresses the total soda concentration in the liquor (Na2Or) g/1 AO expresses the alumina concentration in the liquor (A1203) g/1 Corganks expresses the organics content in the sodium oxalate (Na2C204) in the liquor

3. Test result and analysis

AC organics expresses the precipitation (settlement) percent (%) of C

3.1 Effect of initial organic carbon concentration on organics removal percent

organics

Crystal seed coefficient of sodium oxalate

When the mother liquor (super-concentrated green liquor) (Nk=153g/1; Corganics =2.46-8.0g/l) is super-concentrated to about Nk 300g/l, table 1 shows that the crystal precipitation percent of sodium oxalate is directly proportional to the concentration of Corganics in the green liquor.

Added Na2C204 Na2C204 in the liquor

(mass ratio)

Crystal seed coefficient of sodium carbonate =

Added Na2C204

Table 1 Relation of initial organics concentration and organics removal percent

(mass ratio)

Na2C204 in the liquor

BSC= Before super concentration ASC= After super concentration

The organics removal percent formula can be expressed as the followings: C% =

Item

cox^o-Qxv/ X l o o % c 0 xv 0

No.

Ντ

Conditions

C0 and V0 respectively express the concentration and the volume of organic carbon (Corganics) in the test liquor.

1

Cj and Vj respectively express the concentration and the volume of organic carbon (Corganics) in the concentrated liquor.

2

2.2 Test method

3

•

Result

Effect of initial organic carbon concentration on organics removal percent: initial organic carbon concentration = 2.468.0 g/1

Ao

Nk (g/i)

Ne (g/i)

m

Q

Ne/Ντ

(%)

BSC

163.8

91.35

153.0

15.3

2.46

9.1

ASC

335.4

191.38

314.0

21.4

1.81

6.4

BSC

168.3

91.35

153.0

15.3

4.0

9.1

ASC

331.2

213.35

309.0

22.2

2.51

6.7

52

(%)

66.89

71.39 BSC

168.3

91.35

153.0

15.3

6.0

9.1

ASC

321.2

181.84

303.0

18.2

3.6

5.7

BSC

168.3

91.35

153.0

15.3

8.0

9.1

ASC

312.6

179.27

298.0

14.6

4.31

4.7

73.36

4

Effect of concentrated Nk on organics removal percent: concentrated Nk= 300 g/1. 320 g/1. 340 g/1. 360 g/1. 380g/l

ACcxgnics

75.97

? 76

I

!.. i

£ 74 J Έ

8

72

i n i t i a l c o n c e n t r a t i o n of o r g a n i c C ( g / I )

concentrated N K v a l u e ( g / l )

Figure 1: Relation of initial organic carbon concentration and organics removal percent

Figure 2: Relation of concentrated Nk value and organics removal percent (1)

The organics removal percent has a direct relationship to the initial organic carbon content, and will increase along with the increase of the organic carbon content in the green liquor. There are main two reasons for this: one is that the more organics shall be precipitated by the crystallization after the initial organics content increases, the first precipitated organics itself can be the crystal seed for the later precipitated organics so as to promote the crystallization precipitation of the organics during the concentration and improve the organics removal percent. The other one is that the solubility of the organics in the sodium aluminate liquor has no big change under certain conditions, the organics removal percent must be increased if the initial organic carbon concentration is increased.

concentrated NKvalue(g/l)

3.2 Effect of concentrated Nk on organics removal percent

Figure 3: Relation of concentrated Nk value and solubility percent of carbonate in the liquor

The mother liquor (super-concentrated green liquor) (Nk=136.8/1; Corgnics=l-44g/l) is super concentrated. The precipitation percent of sodium oxalate is increased along with the increase of the superconcentrated liquor Nk value. The precipitation percent reaches the peak value (refer to table 2 and figure 2) when the super concentrated liquor Nk value reaches about 360g/l, moreover, the solubility of the carbonate in the liquor is continuously reduced along with the increase of the concentrated liquor Nk value, the lowest value is here when the Nk is 360g/l, and men it will be little increased along with the increase of the Nk value (refer to table 2 and figure 2).

The mother liquor (super-concentrated green liquor) (Nk=153g/1; Corganics =6g/l) is super concentrated, the study result shows that the precipitation percent of sodium oxalate is increased along with the increase of Nk value of super concentrated liquor under the condition of higher Corganjcs content in the super-concentrated liquor. The precipitation percent reaches a peak value (refer to Table 3 and Figure 4) when the Nk value of super-concentrated liquor reaches about 360g/l.

Table 2: Relation of concentrated liquor Nk and organics removal percent (1)

Table 3: elation of concentrated Nk value and organics removal percent (2)

\jtem

No.

Ντ

Ao

Nk

Ne

(-Organics

NC/NT

ACotgnics

(g/i)

m

m

(g/0

(g/i)

(%)

(%)

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

333.4

172.04

310.0

23.4

1.06

7.0

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

352.6

101.78

335

17.6

0.98

14.7

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

374.2

200.19

364.0

10.2

0.79

2.7

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

399.0

219.27

386.0

13.0

0.91

3.3

^ResiA Condition

1 2 3 4

No.

I

Ντ

Ao

Nk

Ne

Corgarks

Ν^τ

(g/i)

(g/i)

(g/i)

(g/i)

(g/i)

(%)

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

333.4

172.04

310.0

23.4

1.06

7.0

^ResultX Conditiohs^

1 70.3

2 75.3

3 81.3

4 79.2

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

352.6

101.78

335

17.6

0.98

14.7

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

374.2

200.19

364.0

10.2

0.79

2.7

BSC

160.4

82.9

136.8

23.6

1.44

14.7

ASC

399.0

219.27

386.0

13.0

0.91

3.3

BSC= Before super concentration ASC= After super concentration

BSC= Before super concentration ASC= After super concentration

53

ÂCognics 1

(%) 70.3

75.3

81.3

79.2

Table 5: Sodium carbonate causticization percent and organics removal percent of one-section causticization Factors Test No

Nc

[CaO] [Na2Oc]

(%)

40

350

360

concentrated

I I

I8 |

0.5

ACorganics

ACorganics

%

%

%

1

ACorganics

%

55

2

86.73

62.1

|

10

1.3

75

3

91.21

61.8

I

10

1.4

95

4

94.81

67.8

|

12

1.2

75

4

94.52

69.3

I

12

1.3

95

2

96.12

64.7

I

12

1.4

55

3

93.67

62.9

|

14

1.2

95

3

93.98

76.45

I

14

1.3

55

4

90.64

75.36

|

14

1.4

75

2

93.90

72.68

|

[NC12C2O4]

I|

Table 6: Two-section causticization organics removal percent and total organics removal percent (%)

ACorganics I

%

81.8

82.1

81.5

No.

in table 2 No.4

2

1.2

^s^^organics 81.3

79.2

81.9

82.3

(%)

10

causticization, then we can get the results shown in table 6.

No.3

1

I

(%)

The causticization slurry No.7 to No.9 in table 5 is cooled to 5070°C and added the lime — \ £ α θ ] — = 0 5 ^ 1 0 for two-section

I

1.2

ACorganics

3.4.2 Two-section causticization

(added Na 2 C 2 0 4 liquor / Na 2 C 2 04 in the liquor)

0

9

Time

According to the test results, the maximum difference in organics removal is shown after concentrated slag causticization is carried out. This shows the order of the effect of factors on the organics removal percent are: Nc> time > temperature > [CaO/Na2Oc]; the optimum caustic conditions of the organics removal are: N c -14%, time-3 hours, temperature-95 °C, [CaO/Na2Oc]-1.2. These conditions are also the optimum conditions for carbonate removal.

Effect of crystal seed addition on organic carbon Crystal seed coefficient

5 6

I 7

The crystal seed addition test is carried out under the conditions No.3 and No.4 specified in table 2. The results indicates that the crystal seed addition can increase the settlement percent (ACorganics) of caustic sodium oxalate a littlebit and also is good for improving the settlement and filtration performance. The effect of crystal seed addition on the sodium oxalate precipitation refers to the following table (table 4).

Original conditions

I

2

I

3.3 Effect of crystal seed addition on organic carbon removal percent

No

1

I 34

370

NK value(g/l)

Figure 4: Relation of concentrated Nk value and organics removal percent (3)

Table 4: percent

|

Temperature

Sodium carbonate causticization percent

\. 1

82.1

(%)

N.

Super concentratio n

Onesection

Twosection

Anorganic

ACorganics

Δ^-Organics \ IO)

.(%)

75.3

76.12

(%)

87.6 3

Total ACorganics

(%)

75.3x87.63

=66.0% 1

in table2

The results show that the two-section causticization treatment based on the one-section causticization can increase the organic carbon removal percent in the concentrated slag by about 10%. The total removal percent of the organics in the liquor after super concentration, one-section causticization and two-section causticization is about 66.0%.

3.4 Causticization test 3.4.1 One-section causticization The No. 3 liquor (main components: NT160.4g/l; A1 2 0 3 82.9g/l; NK136.8 g/1; N c 23.6 g/1; Corganics1.44 g/1) in table 2 is concentrated to 364g/l, and then the slurry after concentration is treated by the one-section causticization and the control conditions (such as causticization green liquor Nc, molar ratio [CaO/Na 2 O c ], temperature, reaction time etc.) of the causticization reaction are adjusted. The chemical components of the precipitated liquor and solid are respectively analyzed, and the sodium carbonate casuticization percent and the organics removal percent are shown as follows:

3.5 Anti-solution test of causticization slag The resolution test shall be respectively done for the causticization slag from the test No.7 in table 4 and the causticization slag of this test slurry after two-section causticization. The 2 kinds of test results are almost consistent with the resolution law: the alumina concentration during washing has only a small effect on the resolution percent of Corganics, and the temperature has a big effect

54

on it. The resolution test result of organic carbon in causticization slag is shown in table 7. Table 7: Resolution of organic carbon in causticization

Temperature (D

carbon

Resolution

No.

Nk(g/L)

Ao(q/L)

1

4.0

4.8

70

2.9

4.0

4.8

70

3.1

4.0

4.8

80

5.1

4.0

4.8

95

7.5

5

6.0

7.2

60

3.4

6

6.0

7.2

70

4.2

7

6.0

7.2

80

7.6

8

6.0

7.2

95

10.8

9

8.0

9.6

60

3.6

10

8.0

9.6

70

7.1

11

8.0

9.6

80

11.5

I 12

8.0

9.6

95

15.2

I

2

I

4

3

)

the salt settlers for concentration and to the downstream separators for separation. After separation, the solid after solution is sent to one-section causticization tank with lime milk addition, the onesection causticization slurry is obtained by heating the reaction and is sent to a two-section causticization tank after it is cooled. T h e two-section causticization slurry is obtained and sent to separation washing machines, and then the washing liquor is sent to the alumina production system, the slag is sent to the alumina red m u d washing system and is discharged after washing along with the red mud.

Corganics( /o) 1

|

Crystal seed

1 -

I Liquor *

:T

I

'

Lime nilk* JLJL

Super concentration*·'

Test liquor proportion

4.

Mother liquor

Separation ♦>

Causticization*·'

I Slag.

> Causiticization liquor

,

t Related systems of alumina* production*'

? Solid ^

Conclusion

f

The organic carbon removal percent has the direct relation with the initial organic carbon content; it is increased along with the increase of the initial organic carbon content.

Discharge**

When the initial organic carbon content reaches a certain level, the organics removal percent is increased first by the linearity along with the increase of the concentrated caustic soda concentration; however, the organic carbon removal percent is reduced along with the increase of the high concentrated caustic soda concentration after the concentrated caustic soda concentration reaches the maximum value 360g/l.

References 1. Center-south University, Guiyang Aluminum Magnesium Design & Research Institute. "Study report on organics removal of alumina production with bauxite by Bayer process". 2009.

The crystal seed addition can slightly improve the settlement percent of the causticization sodium oxalate product.

2.

The reaction percent of the oxalate crystallization after onesection causticization is 62.1%-72.68%.

Chen Wengu, Tang Jiaming, Zhang Li. "Study on sodium oxalate in the Bayer Process". Light Metal 2005, 5: 11-15.

3.

B Gnyra and G Lever. "Review of Byer Organics-Oxalate Control Processes". Light Metals 1979,151-161.

The two-section causticization treatment based on one-section causticization can increase the total organic carbon removal percent in the concentrated slag by about 10%.

4. W Arnswald et al. "Removal of Organic Carbon from Liquor by Wet Oxidation in Tjbe Digesters". Light Metals 1991, 2327.

The alumina concentration has a small effect on the resolution percent of CorganiCS, and the temperature has a big effect on it. This method can not only effectively remove the oxalate, but also has a relatively good carbonate removal effect. 5. Process flow of super concentrated causticization The mother liquor in the alumina system is sent to the general evaporators for preliminary concentration, and the concentrated mother liquor is sent to the super-concentrated evaporators for super concentration, then the super-concentrated slurry is sent to

55

5.

G M Bell. "Osidation of Organic Substances in the Bayer Process". Light Metals 1981,117-128.

6.

P Atkins and S C Grocott. "The Liquid Anion Exchange Process for Organics RemovaF. Light Metals 1993, 151-157.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ALUNORTE EXPANSION 3 - THE NEW LINES ADDED TO REACH 6.3 MILLION TONS PER YEAR Daryush Albuquerque Khoshneviss1, Luiz Gustavo Correa2, Joaquim Ribeiro Alves Filho1, Hans Marius Berntsen3, Ricardo Rodrigues de Carvalho2 alunorte - Alumina do Norte do Brasil S.A., Rodovia PA 481, km 12. Distrito de Murucupi, CEP: 68.447-000, Barcarena-PA, Brasil 2 Vale S.A., Av. Graça Aranha, 26 - 10° andar, CEP: 20.030-900, Rio de Janeiro-RJ, Brasil 3 Hydro Aluminium AS, N-0240 Oslo, Norway Keywords: Expansion Project, Alunorte Performance, Plant Management Abstract Alunorte started operation in 1995 with a design production capacity of 1.1 Mtpy. Since then the plant was expanded three times and consists today of seven production lines with a total production capacity of close to 6.3 Mtpy. Expansion 3 included the process lines 6 and 7. The lines were commissioned in 2008 and reached their nominal production capacity after a short startup phase. In Expansion 3 a number of new technology developments were consolidated, such as the bauxite transport by pipeline, an improved precipitation concept and others. The impact of the process modifications applied in lines 6 and 7 are discussed in comparison to the previous lines with regard to effects on production volume, productivity, alumina quality and availability.

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However, a number of challenges are associated with the operation of a plant of the size of Alunorte which has grown very fast: the management of seven process lines, the operation of the plant with a very young organization and strong demand for training.

ceeded well its design production capacity. From Expansion 2 onwards the actual production remains slightly behind the design production capacity. However, all lines individually have demonstrated their design capacity. But the plant has become more and more complex with time. A number of issues arose which are linked to the size of the plant and its fast growth. On the other hand operations, maintenance, logistics can be made very efficient and a number of benefits can be achieved.

Introduction Alunorte started operation in July 1995. The plant had two production lines and a nominal production capacity of 1.1 Mtpa. This production rate was reached in 1997 and exceeded in the following years. A number of process improvements were implemented so that the production capacity could be increased to 1.6 Mtpa in 2000. In eight years only from 2000 to 2008 the plant production capacity increased three times. Work on Expansion 1 started in 2000 and in the first quarter 2003 the third process line was commissioned. The plant design production capacity was increased to 2.325 Mtpa. This production rate could be reached and even exceeded during the first year after the expansion. In 2006 process lines 4 and 5 were commissioned as part of Expansion 2. The production capacity increased further to close to 4.4 Mtpa. This number includes afirm-upof the precipitation area of lines 1-3. Finally, in the fourth quarter of 2008, lines 6 and 7 started operation as part of Expansion 3. The plant has now a nominal production capacity of close to 6.3 Mtpa.

In the following, the Expansion 3 project is described followed by a review of Alunorte's current performance. Then the current main challenges are discussed - technically and organizationally. Some of the projects started to improve the situation are presented as well as an outlook about Alunorte's expected development and performance. Scope of the Project Expansion 3 consists of new filters for bauxite dewatering, two new process lines (lines 6 and 7), two 3,300 t/d calciners, a port expansion and the construction of a new high-pressure circulating fluidized bed (CFB) boiler. This paper focuses on the implementation of the new process lines from bauxite dewatering to calcination and discusses port expansion and boiler erection briefly.

In Fig. 1 the development of production capacity and actual production are shown. The graph makes clear that historically Alunorte exceeded the expectations with regard to production volumes. For Expansion 1 and 2 the actual production before the expansion - which were higher than design - was the design basis onwards. In this way additional production volume is included in today's nameplate production capacity of close to 6.3 Mtpa when compared to the production volume of the original plant and the expansions. Before Expansion 2 was commissioned Alunorte ex-

The design of the new process lines 6 and 7 is partly based on the original design of Alunorte of Alcan International from 1995, partly it includes improvements which have been implemented in the previous expansions and some new elements were added. The most obvious difference between the expansions is their size. Each of the new lines has a nominal production capacity of 930 ktpa or 1.86 Mtpa in total for Expansion 3. The original design capacity of the plant was 1.1 Mtpa, 550 ktpa per line. Thus, the

57

design capacity of lines 6 and 7 is about 70% larger than the original design capacity of the first two lines. The dimensioning of the new lines is larger than that of the old ones but also the process performance considerably increased. In 2000, after the firm-up, lines 1 and 2 produced 1.6 Mtpa and operated at a precipitation productivity of close to 84 g/1. With lines 6 and 7 a precipitation productivity of 89 g/1 can be achieved. This is a result of the continuous improvement of the precipitation concept. Lines 1/2/3, 4/5 and 6/7 apply three different concepts. In lines 6/7 coarse seed and fine seed filtration are installed, lines 4/5 use seed cyclones and lines 1/2/3 coarse seed filtration. The newest lines are operated with the highest solids concentration and achieve the highest precipitation yield. Furthermore, the cooling capacity is well dimensioned. In addition to traditional in-tank plate heat exchangers external plate heat exchangers (Barriquand coolers) are installed. The number of tanks has been increased as well as their size in order to increase residence time for higher precipitation productivity.

The performance of the calcination facilities was increased again with Expansion 3. The energy efficiency of the new two 3,300 t/d circulating fluidized bed (CFB) calciners is higher than that of the old ones. A high degree of automatic control is installed [4]. In total seven CFB calciners of Outotec (formerly Lurgi) are installed. Alunorte received the energy efficiency award 2010 from the German Energy Agency for the two newest calciners [5]. A part of the electrical power used by Alunorte is co-generated at site while the rest is received from the national grid. As part of Expansion 3 a circulating fluidized bed boiler was installed. The utilities system for steam and electrical power generation of Alunorte has a high overall efficiency and good operational flexibility. This is important in order to be able translate the low specific consumption of steam and electrical power into a low specific consumption of boiler fuel. The environmental concept is continued with Expansion 3 in order to guarantee a generally high standard. Emissions and effluents are within the legislative limits. Of high importance is Alunorte's bauxite residue disposal concept - dry-stacking which is state of the art. The mud in filtered in drum filters and then disposed at high solids concentrations. Water from the bauxite residue area (RDA) is collected, cleaned and then send to a nearby river.

The bauxite for lines 6 and 7 is supplied from Mineraçao Bauxita Paragominas (MBP) through the world's first long-distance bauxite slurry pipeline [1]. The bauxite slurry is pumped at a solids content of 50 % and is then dewatered at Alunorte in hyperbaric filters [2,3]. A residual moisture content of about 15 % in the dewatered bauxite is achieved. This is slightly higher than that of Trombetas bauxite which arrives by ship. To be pumpable through the pipeline the Paragominas bauxite is finely ground at MBP. Its particle size distribution is sufficiently fine to be charged to the Bayer plant without additional milling. Different to the other lines no mills are installed in lines 6 and 7 and the bauxite is sent from filtration by conveyor belts directly to the bauxite slurry re-suspension tanks and heating units.

Project Execution The installation of process lines 6 and 7 as part of Expansion 3 was finished on schedule, on budget and the design performance of the two new lines could be demonstrated during the first two months after start-up. The project had a duration of 32 months. The master schedule of the project for the refinery is shown in Fig. 2.

The clarification area has undergone a number of changes compared to the original design. Initially conventional decanters and washers were installed. As part of the expansions deep thickeners were added. Alunorte made good experience with their operation so that in Expansion 3 only deep thickeners are installed and conventional decanters or washers are completely eliminated. As part of Expansion 3 Alunorte has decided to install Diastar filters for liquor filtration. The filters are operated fully automatically. The operational performance of the liquor filtration area could be improved in comparison to the previous process lines.

A number of factors contributed to the fast execution of the project. One of them was the strategy to use the latest expansion as design basis for the new lines. Improvements in specific plant areas were included. However, the concept was not fundamentally changed in most areas. The biggest differences are found in precipitation. The project was executed with an extended owners' team with no EPCM contractor. Alunorte Board of Directors acted as steering committee for the project. An advantage of the organization of the project was the ability to take decisions very fast and avoid decision-related delays of the project. Furthermore, it was possible to work with a very experienced team. Most members of the Alunorte expansion team have already been involved in the previous expansions. Finally, it was possible to perform the start-up of the lines earlier than originally scheduled. Less time than planned was required to reach the nominal production capacity of lines 6 and 7.

In digestion, evaporation and vacuum flash cooling bottom-entry flash tanks are installed. The carry-over of caustic soda to the condensate is lower than for conventional side-entry flash tanks. The regenerative condensate has better quality, the scaling in the vapor transfer lines is reduced and longer cleaning cycles are achieved. Originally side-entry flash tanks were installed in digestion lines 1 and 2. These were modified to bottom entry and in all expansions only bottom entry tanks were used.

The project completion system ProCoSys™ was used for construction and commissioning of the new lines. The system was initially developed by Hydro and mainly used for oil and gas projects. The system provides a systematic approach to ensure a fast project completion. It is ensured that all equipment is available and checked at the time it is required. The system became a well recognized tool at Alunorte and helped to achieve fast commissioning and a start-up of the new lines without major delays.

A modern fully digitalfieldbusprocess control system is installed. At Expansion 2 it was decided to complete the move from an analog system to a digital system. This system is implemented as part of Expansion 2 and Expansion 3. The degree of automatic control increased with each expansion. The high degree of automatic control of the new lines helps to achieve higher overall performance in the new lines.

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Alunorte's shareholders are very satisfied with the performance of the project organization, management and execution in general. An exception, however, is the unacceptable safety performance of Expansion 3. A fatality occurred during the construction caused by a crane fell on a passing construction worker. A TRI rate of 3.4 was reached for the overall Expansion 3 project. Although the implementation of the process lines went according to schedule and the production capacity was achieved faster than initially planned there were some delays in the Expansion 3 project. Both the new boiler and the port expansion were delayed. The delays caused some operational challenges. Solutions were found to operate Alunorte at close to full capacity without having these projects finished. Expansion 3 included a high-pressure circulating fluidized bed boiler for steam and electrical power generation. As the boiler was delayed the required steam for the new process lines was provided by already existing low pressure boilers. For more than one year Alunorte had to operate without spare boilers. Delays in the early phase of the port expansion could not be recovered so that finally the installation of a new alumina shiploader could not befinishedon time. Therefore some logistic challenges needed to be overcome at the port. The situation was challenging but Alunorte succeeded to ship alumina at a rhythm exceeding to 5 million t/yr with only one ship loader without loosing production due to logistical problems at the port.

With Expansion 3 new personnel was contracted by Alunorte. The hiring started enough time before the planned start-up of the new lines. Training for the new operators started about a year before commissioning. In this way it was possible to perform the start-up of the new lines with a team of sufficient experience. Overall, the new lines performed as expected. No major negative surprises were experienced when the lines were started up. Although a new precipitation concept was installed for which no operational experience existed no major problems turned up. Lines 6 and 7 are fully based on Paragominas bauxite and some minor differences in process behavior were observed caused by the bauxite. These differences were mostly found in the clarification area. The fine particle size distribution of the bauxite might be the cause of a somewhat slower settling mud as compared to Trombetas bauxite. Measures were taken to correct the situation in the clarification area.

Start-up and Operational Performance

The energy utilization of Alunorte is among the best in the world with about 8 GJ per ton of alumina. As there are few plants with a similarly low energy utilization this defines the world-wide benchmark for alumina production. A detailed review of Alunorte's energy efficiency is presented in [6]. Alunorte's specific caustic consumption is mainly determined by the noncontrollable losses which are driven by the content of reactive silica in the bauxite. Controllable losses are small with less than 8 kg per ton alumina. Although the content of reactive silica is somewhat high in both bauxites, Trombetas and Paragominas, their quality is generally very good. The content of available alumina is high and the content of organic material very low. In combination this leads to a low specific consumption of bauxite and very clean liquor with low concentrations of impurities. This

23 25 27 29 31 33 35 37 39 41 43 45 47 49 51 53 week(2008)

Figure 3. Ramp-up of Alunorte's lines 6 and 7.

59

Table I. Typical alumina quality parameters of Alunorte. 5.0% 0.015 % Si0 2 cc-alumina LOI (300-1000 °C) 0.8% Fe 2 0 3 0.015 % 0.005 % LOI (110-300 °C) 0.7% Ti0 2 Na 2 0 0.4% BET 75 m2 3.0% ZnO 0.001 % + 149 μπι (+100#) + 74 μπι (+200#) 70.0 % CaO 0.01 % -44μm(-325#) 7.0% v 2 o 5 0.002 % MnO 0.001 % app. density 0.97 g/cm3 0.001 % attrition index 17% p2o5

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the previous alumina quality characteristics could be maintained. However, some minor differences can be observed when the product from lines 6 and 7 is compared to that of the older lines. The content of occluded soda is slightly higher and the lines are operated with slightly coarser crystal size distribution in precipitation. No difference is observed for Attrition Index. The product of the new lines is blended with the product of the other lines before shipped or sent to the nearby Albras smelter. In Table I the typical alumina quality parameters of Alunorte are shown. As result of improvements done in lines 6 and 7 compared to the older lines variations in alumina quality parameters became smaller. As the variations in all lines are generally small and the product of all lines is blended this effect is not observed in the final product, in particular not for alumina which is shipped. The storage in silos at Alunorte, transport by ship and storage at the smelters lead to additional homogenization of the product. However, the target is to operate the process as stable as possible in order to achieve the best possible performance.

Alunorte has grown as fast as no other alumina refinery before and has a size which exceeds the second-largest alumina refinery by about 50 %. In this way Alunorte stands out in comparison to all other alumina refineries in the world. In the following section a critical review is presented about Alunorte's current performance, actual challenges and expected development in the coming years. Safety There is a high focus on safety at Alunorte. The YTD TRI rate in October 2010 was 1.95 (based on one million working hour), after having achieved a TRI rate of 2.37 in 2009. Alunorte has developed and continuously improved a strategy to manage health and safety during the past 15 years of operation [7]. Alunorte is ambitious to further improve this performance in the coming years. The decision to install new technologies at Alunorte always includes some new risks for accidents which need to be carefully addressed.

Project Budget

Process & Technology

The project cost of Expansion 3 was 629 USD per annual ton of alumina. The project budget was set up in both currencies US Dollar and Brasilian Real (BRL). The project budget was met in BRL but there was a budget overrun in US Dollar. The main cause was the development of the Brasilian currency during the project. At the beginning of the project the exchange rate was 2.2 while it dropped to about 1.6 at the end of the project. In Fig. 4 the development of the Brasilian Real in comparison to the US Dollar is shown for the period from 2000 until the end of 2008. The situation of Expansion 3 is compared to the Expansions 1 and 2. A large amount of the material and equipments was purchased in Brasil so that project cost in BRL were not significantly affected by the development of the currencies.

The general technical performance of Alunorte is described above. Operation is characterized by a high productivity, low controllable caustic consumption, low energy utilization, a stateof-the art bauxite residue disposal concept and low plant emissions. For good plant performance and good product quality it is important to operate the refinery as stable as possible. Operational variations must be kept small and the degree of non-planned maintenance must be kept as small as possible. One of the challenges to achieve a good performance in these aspects is the fact that Alunorte can be considered as one huge alumina production complex which can be divided into three major units. The first unit are lines 1/2/3, the second lines 4/5 and the third 6/7. Since the equipment installed in these units is not identical and different process concepts exist, e.g. in precipitation, specific operation and maintenance procedures are required. The three units are with small exceptions process-wise fully independent

Expansion 3 was finished close to the beginning of the worldwide economic downturn end of 2008. The additional alumina volume from Expansion 3 was added to the market at a time of falling LME and alumina prices. The sudden development of the alumina market made clear how important it is to operate a plant

60

from each other. The plant liquor of one unit is usually not mixed with liquor of another unit so that individual characteristics for each of the units exist. One example was given above. The settling behavior of Paragominas bauxite is different from that of Trombetas. The advantage of the separation of the lines is that operational disturbances of the process in one line do usually not affect the other lines. Even a stop of a line is possible without affecting the production of the lines in the other operational units.

Organization & Management In early days when Alunorte consisted of two lines only 450 people were employed. Nowadays, Alunorte employs 1,600 people and about 1100 more workers as permanent contractors. Due to the very fast growth of the plant Alunorte's team is very young and many of the employees have only few years of work experience. The plant has not only grown very fast but also new technologies were employed. The dewatering of the bauxite received through the pipeline from MBP is a new technology not only for Alunorte but also for the whole alumina industry. Alunorte could not profit from the experience of other alumina refineries. The first circulating fluidized bed boilers were installed as part of Expansion 2. Although the technology is not new as such it was new for Alunorte. Furthermore, there is just a small number of circulating fluidized bed boilers installed in South America. Both, the young organization and the application of new technologies at Alunorte result in a strong demand for training for operators and engineers.

The three production units are, however, not fully independent from each other. They share utilities such as the electrical system, steam generators, bauxite supply and port. Operational problems in one of these plant areas can affect the whole process and will in this case lead to significant losses in production. Alunorte recognized that this is a risk for the operation of the whole refinery and started a program to minimize the interconnections between the units as far as possible. This task became more and more important as the plant grew. The complexity of the plant increases exponentially with its size and the possibilities for equipment failure which can affect larger parts of the plant get bigger. The availability of the plant is exposed in a similar way. The more complex the refinery got the more efforts were required to maintain a high global availability.

As explained above the plant consists of seven lines which can be divided into three units consisting of lines 1/2/3, 4/5 and 6/7. The operational issues are not the same in the different lines. In some plant areas, such as precipitation, different process concepts are chosen. Furthermore, the first lines have an age of 15 years while the newest lines are just two years old. All this has to be considered in daily operation, for maintenance, training of the employees, etc. The principal management organization of Alunorte is still very similar to that in place during the early years of operation. It can be questioned if today's organization is still ideal or if other concepts are more suitable. Alunorte works actively to evaluate the current issues and challenges to manage the plant. Alunorte's is generally open to new ideas, technologies and concepts and works also on a continuous improvement of the plant organization.

Logistics are another important issue. Some of the challenges arose from the delay of the port expansion. This issue could be handled as described above. However, the tonnages of alumina, bauxite, caustic and boiler fuels are huge and require a careful planning of port operations. Unsatisfactory performance in this part of the plant could result in an increase of transport and handling cost or in the worst case in production losses. A similar challenge exists for the transport of bauxite residue to the residue disposal area (RDA). It is transported by trucks from the filtration area to the disposal area. In case of bad weather conditions as it canfrequentlyhappen during the rainy season from December to April the disposal to the RDA can become a bottle neck. As the plant has the potential to creep in production in the coming years and the bauxite quality slowly deteriorates the amount of bauxite residue which needs to be disposed will increase in the future and the challenge will gets bigger. Alunorte studies options to improve this situation in order to improve the logistics and also eliminate a potential safety risk caused by the large number of trucks in service. In addition to logistics of materials, transport logistics for Alunorte's employees needs to be mentioned as well. Alunorte is located at about 40 km from Belém, the capital of the state Para, in straight line. The road, however, is much longer about 100 km. Many of Alunorte's employees live in Belém and travel forth and back day by day. To improve the situation Alunorte launches a ferry connection between the plant and Belém city.

Conclusion With Expansion 3 Alunorte has become the largest alumina refinery in the world with a production capacity of close to 6.3 Mtpa. Alunorte has grown as fast as no other refinery in the world. The development of Alunorte from its start-up in 1995 is presented and Expansion 3 is reviewed in detail. The design of Expansion 3 is based on that of the previous lines. However, a number of new technologies are applied, such as a new precipitation concept, different filters for pregnant liquor or the use of deep thickeners only for mud wash. The Expansion 3 project had a duration of 32 month with a short start-up phase of two month for both new lines to reach design production capacity. The project cost of Expansion 3 was USD 629 per annual ton of alumina. The installation of the two new process lines could be performed on schedule and on budget. The generally good performance of Alunorte, such as energy efficiency or alumina quality parameters, could be maintained also after the addition of the two new lines. The challenges of managing a plant of the size of Alunorte with a young team with on average few years of work experience are discussed. Alunorte has not only grown very fast but has implemented a number of new technologies, such as bauxite transport as slurry in a pipeline or the installation of CFB boilers. Training of Alunorte's operators and engineers is therefore important. Nowadays, Alunorte can be considered as a big alumina production complex with seven process lines which can be divided into three units, namely Lines 1/2/3, 4/5 and 6/7.

After many years with a high focus on growth Alunorte has entered into a phase of optimizing the operation of the plant. There is potential for some increase in production. Furthermore, Alunorte works constantly to improve energy utilization and specific consumption figures. Some counteracting factors have to be considered such as the aging of the equipment or a slow deterioration of the bauxite quality in the long-term perspective. As part of the current work on process optimization new technologies are reviewed and tested to guarantee good operational performance also in the long term.

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This offers technologically and organizationally possibilities to operate the units independently of each other with the aim to achieve a high global availability of the plant and to minimize the risk of failures in one line affecting the rest of the refinery. Alunorte has the goal to further improve the operational performance of the plant in order stay among the best plants in the world also in the long term. References 1. Ramesh Gandhi, Mike Weston, Mani Talavera, Geraldo Pereira Brittes and Eder Barbosa, "Design and Operation of the World's First Long Distance Bauxite Slurry Pipeline", In Light Metals 2008, De Young, D.H., Ed., TMS (The Minerals, Metals & Materials Society), 95-100. 2. Ayana Oliveira, Juarez Dutra and Jorge Aldi, "Alunorte Bauxite Dewatering Station - A Unique Experience", In Light Metals 2008, De Young, D.H., Ed., TMS (The Minerals, Metals & Materials Society), 85-87. 3. R. Bott, T. Langeloh and J. Hahn, "Filtration of Bauxite after Pipeline Transport: Big Challenges - Proper Solutions", In Proceedings of 8th International Alumina Quality Workshop, Armstrong, L, Ed., AQW Inc.,2008, 319-323. 4. Michael Missalla, Jan Jarzembowski, Roger Bligh, Hans Werner Schmidt, "Increased Availability and Optimization of Calciner Performance due to Automation", Light Metals 2009, Bearne, G., Ed., TMS (The Minerals, Metals & Materials Society). 5. Michael Missalla, Hans Werner Schmidt, Joaquim Ribeiro Alves Filho and Reiner Wischnewski, "Significant Improvement of Energy Efficiency at Alunorte's Calcination Facility", In Light Metals 2011, Lindsay, S., Ed., TMS (The Minerals, Metals & Materials Society). 6. Reiner Wischnewski, Cleto Maues de Azevedo Jr, Emerson L.S. Moraes, Arthur Barros Monteiro, "Alunorte Global Energy Efficiency", In Light Metals 2011, Lindsay, S., Ed., TMS (The Minerals, Metals & Materials Society). 7. Jorge Aldi Dias Lima, "Safety Performance at Alunorte", In Light Metals 2008, De Young, D.H., Ed., TMS (The Minerals, Metals & Materials Society), 5-11.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

One Green Field Megaton Grade Large Alumina Refinery with Successful Engineering & Operation Experience Luo Xianqing

Yang Xiaoping

(Guiyang Aluminium Magnesium Design & Research Institute, Guiyang, Guizhou, 550081) Keywords: diaspore bauxite, alumina production, Bayer process, comprehensive energy consumption accumulated bauxite and the ancient weathering crust bauxite, in which the accumulated diasporic bauxite is 83.5% and the ancient weathering crust diasporic bauxite is 16.5%. From the ore bed scale it is known that all ore beds in 10 large-size ore beds are accumulated diasporic bauxite and 3 ores in medium-size ore beds are accumulated diasporic bauxite. The accumulated bauxite mainly is distributed in Pingguo, Debao, Jingxi etc. in the west of Guangxi. This bauxite is characterized by medium aluminum content, high ferrous content, high alumina silica ratio, low sulfur content etc. and consists of diasporic bauxite, goethite, hematite and kaolinite being 85% content ,with lower content of other mineral, is the few high-quality bauxite resource in China can be used for producing the aluminum with Bayer process. The ore grade: A1203 is 45%-65%, generally 54%-60%; Si02 is 3.4%-14% &, generally 5%-9%; alumina silica ratio is 4-11; Fe 2 0 3 is 5%-25%, generally 10%-20%.

Abstract: The Phase I alumina project (l.ómillion tons/a metallurgical grade alumina) of the Guangxi Huayin Aluminum Corporation Limited is a greenfield construction project with to-date, the largest disposable investment in the Chinese alumina industry. The project is based on a diasporic bauxite which is difficult to grind and dissolve, with a high aluminum & low silicon content. Two years of operational experience shows that the comprehensive energy consumption per ton alumina production is only 10.51GJ/t. This article introduces the process scheme & technical measures adopted in the engineering and operational stages of this project.

1. Introduction Guangxi Huayin Aluminum Corporation Limited is a partnership of the Guangxi Investment (group) Corporation Limited, China Minmetals Non-ferrous Metals Corporation Limited and China Aluminum Stock Corporation Limited. The major investment is located in Mayi Town, Debao County, Guangxi Chuang Municipality, Guangxi. The construction of its phase I alumina project (hereinafter to be referred as Huayin phase I alumina project) has been targeted to produce 1.6 million tons/a of metallurgical grade alumina and is based on a total budgetary estimated investment of RMB 9.12 billion..

The bauxite in Debao and Jingxi County is the raw material for the Huayin phase I alumina project, and there is a mining division and an ore wash division respectively built in these two counties. The parameters of the mined bauxite in the first mining area are as follows:

The project was designed by the Guiyang Aluminium Magnesium Design & Research Institute, constructed and started up by China Aluminum International Engineering Corporation Limited as the general contractor. Construction was started in June 2005; the first production line was put into production in Dec. 2007, with all the production lines in operation in June 2008 and production & operational standards achieved in June 2009.

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£5.75

£9.84

£7.0

£55.80

£5.94

>9.39

£7.0

|

>56.22

£5.84

£9.62

£^0

|

|

The ore processing performance test [ ] result indicates that the bauxite in Debao and Jingxi is a diaspore dominated bauxite, characterized by high aluminum & low silicon content, and its performance is as follows:

2. Bauxite resource and ore processing performance Guangxi Chuang Municipality is abundant in high-quality bauxite resource reserves. The total explored bauxite reserves are 0.569 billion ton and the proven ones are 0.375 billion ton up to the end of 2005. As far as known from the ore bed investigation degree in Guangxi, the resource reserves reaching exploration degree are 28%, reaching detailed investigation are 22% and reaching general investigation are 48%, thus the bauxite investigation degree in Guangxi is high. Guangxi bauxite contains 3 types of ore, including ancient weathering crust bauxite, accumulated bauxite and laterite bauxite. The proven resource reserves are mainly the

63

•

This bauxite is difficult to grind and its work index is large. The Bond rod mill work index is 20 mesh, Wi=16.02; the Bond ball mill work index is 120 mesh, Wi=22.55.

•

Under the test conditions with a solid content of 320g/l in the slurry, test liquor caustic soda (Na20) being 245g/l and lime addition being 7%, the predesilication rate of this bauxite is close to 80% when the predesilication temperature is 90 to 100°C.

•

Under the test conditions with digestion temperature being 260°C, digestion time being 40 min and lime addition

being 7%, the actual digestion rate reaches 86% and the relative digestion rate is 96%, however, with digestion time being 30 min, the actual digestion rate is 83% and the relative digestion rate is 92%. With a diluted slurry retention of 3 hours, the ratio of alumina content and silicon oxide content in the sodium aluminate liquor is always above 230. The bauxite in Debao and Jingxi is difficult to dissolve, so it requires higher digestion temperatures, higher caustic liquor concentration and finer bauxite granularity. •

belt weigher & roller conveyor. The test liquor required for bauxite grinding is taken directly from the evaporation station, and is added in proportion respectively from the outlet & inlet of the rod mill and the outlet & inlet of the ball mill. The slurry ground by the rod mill and the ball mill together enter the slurry tank, and are then sent to the hydraulic cyclone for classification by the slurry pump. The underflow is sent back to the ball mill and the overflow enters the slurry storage tank through the rotary screen and becomes the qualified slurry, then sent to the high-pressure digestion by slurry pump.

After some flocculant is added to the slurry after digestion, the red mud settling speed can satisfy the industry production requirements, thus the red mud settlement performance is acceptable.

This grinding scheme effectively solves the problem that the bauxite from Debao and Jingxi is difficult to grind and satisfies the digestion requirements to the bauxite granularity, which reduces the power consumption for the bauxite grinding by 6.22kw.h/t compared with the general bauxite grinding process and makes the bauxite granularity more uniform.

3. Process scheme and main operation indexes The alumina refinery uses the Bayer process, based on the bauxite supply conditions. It consists of raw material plant (including bauxite storage & transportation, bauxite grinding, lime storage & transportation, lime slaking etc.), digestion shop (including high-pressure pump house, sleeve-tube preheating digestion & dilution etc.), red mud settling & washing plant (including red mud settling & washing, red mud filtration, red mud transportation, flocculant preparation, hot water station, security filtration etc.), precipitation shop (including pregnant heat exchange, seed precipitation, hydrate classification, intermediate cooling, seed filtration etc.), hydrate calcination (including product filtration, hydrate silo & hydrate conveying, calcination, alumina conveying & packing etc.), evaporation plant (including evaporation station, green liquor storage tank & water washing, test liquor storage tank, desalt, causticization etc.), circulating water shop etc.

3.2 Predesilication Considering the bauxite from Debao and Jingxi possesses good predesilication performance, the predesilication section is not here in the design and just requires more heating in the slurry storage tank. Fast slurry predesilication is achieved in the slurry storage tank, with 2 hours predesilication (the scaling in the single sleeve-tube preheater is very slow and the acid washing cycle of the single sleeve tube preheater is above 80 days during actual production, indicating that the predesilication effect is good and this measure is very successful). 3.3 Digestion An eleven-stage preheating and ten-stage flash evaporation digestion device has been adopted for this plant.

The process flow of the whole plant is as follows: the bauxite is made into slurry after it is ground with the lime and qualified test liquor by the ball mill. The slurry after predesilication enters the digestion unit consisting of the sleeve-tube preheater and the autoclave, and the digested slurry after multi-stage self-evaporation enters the dilution tank. The diluted slurry enters the red mud settling & washing system consisting of large flat-bottomed settlers and drum filters. The red mud after washing is sent to the red mud stockyard and the settler overflow after security filtration and heat-exchange cooling, enters the seed precipitation system. The aluminum hydrate generated from the seed precipitation in converted to alumina through gas suspension calcination and the spent liquor after precipitation is fed back to bauxite grinding through the evaporation.

The slurry from the raw material grinding section is sent to the high-pressure pump house, and to six-stage sleeve-tube preheater through a high-pressure diaphragm pump, and heated to 170 °C, with the secondary vapour generated by the self -evaporator. It then enters 4 autoclaves heated to 220 °C with secondary vapour, and is heated to the digestion temperature being 260 °C in a follow-up reaction autoclave with fresh 6.1MPa steam. It then enters an insulated autoclave for insulated retention for 40 min. The slurry after high temperature digestion, is sent to the dilution tank after the temperature reduces to below 126 °C through a ten-stage self-evaporation, and sent to the after tank by the pump, after it is diluted with the primary washing liquor from the red mud settling & washing and remains for some time. It is then sent to the red mud settler by pump. The feeding pump of this digestion device is a triple cylinder single action diaphragm pump.

The technical scheme of each process section is as follows: 3.1 Bauxite grinding

There are many improvements in the utilization of digestion fresh steam condensate secondary vapour, the pore plate structure of the flash evaporator, heating pipes of the autoclave, continuous emission of non-condensing gas from autoclave, utilization of high-pressure digestion secondary condensate etc. in this engineering design compared with before. These make the steam consumption of the digestion system more reasonable.

A two-step bauxite grinding process is used, with the first step a rod mill open circuit and the second step a ball mill & hydraulic cyclone closed circuit. The bauxite, being no more than 15mm from the homogenization stackpile, is sent to the bauxite silo of the bauxite grinding plant, and the lime from the lime warehouse is sent to the lime silo of the plant.

3.4 Red mud settling & washing

The bauxite and the lime are sent to the rod mill in proportion respectively by plant-type feeder & belt conveyor and electronic

64

A reverse washing system is adopted, consisting of red mud settlement by settler, triple settling & washing and primary washing by drum filter.

precipitation output rate. So intermediate cooling unit is added for higher precipitation output rate (the output of pregnant liquor can reach more than 88kg/m3 ).

The solid-liquor separation of digestion slurry (concentration of caustic soda Na20 being 172g/l ) after dilution in the settler, the overflow (coarse liquor) is sent to security filtration and the underflow is sent to red mud washing system by pump. In this washing system, the red mud is treated using hot water addition through triple settling & washing. The washing water flows along the reverse routing of red mud and its concentration is gradually increased, finally being used for diluting the slurry after digestion.

Moreover, some crystallizing agent is added so as to intensify the precipitation of aluminum hydrate and also increase its granularity. 3.8 Seed filtration This section has the function of seed supply for seed precipitation. The aluminum hydrate in the slurry from the last precipitator of the precipitation system is subjected to liquor-solid separation so as to obtain the aluminum hydrate seed; the spent liquor (caustic alkali Na20 concentration being 181g/l) generated after filtration is sent to evaporation section for carrying out mother liquor evaporation, so as to obtain the test liquor

The settler is a large flat-bottomed settler of 40m diameter. Some flocculant is added during operation so as to ensure the solid content of the settlement overflow is controlled within the specified limit. The red mud slurry after triple washing is sent to drum vacuum filter by the pump so it is washed and filtered with water injection for the last time. The red mud after washing is sent to mixing reaction tank; the filtrate is back to red mud washing settler through the filtrate tank by pump.

3.9 Recycle of aluminum hydrate in spent liquor In the spent liquor obtained after seed filtration and product filtration, the aluminum hydrate content is about 2g/l. The existence of aluminum hydrate in the spent liquor increases the pipe blocking risk index of the evaporator heating pipe and reduces the circulating efficiency of the caustic alkali. So this project is equipped with an aluminum hydrate recycling section and a vertical leaf filter is the main device for this section.

The filtration area of the drum filter is 100m2, being equipped with water injection device. It can simultaneously carry out red mud filtration and washing. The alkali content in the residual liquor of the red mud after filtration, is reduced to 0.32% (dry basis) and the water content is adjusted to about 40% (wet basis). Finally the red mud is transported by a diaphragm pump to the red mud stockyard being 2 km away for dry storage.

The underflow from the spent liquor conical tank in the seed filtration section is sent to the vertical leaf filter for filtering out the aluminum hydrate in the liquor. The filtrate is sent back to the spent liquor tank in the seed filtration section, and the filter cake (aluminum hydrate) is sent to the overflow tank in seed filtration and is finally sent to the precipitation section as seed.

3.5 Security filtration The large double-star vertical leaf filter of the filtration area being 308m2 is adopted to filter the coarse liquor from the red mud settling & washing. It is characterized by light weight, less occupation area, higher automatization, being convenient for maintenance and high capacity. The solid content of filtrate is no more than 15mg/l.

3.10 Evaporation and desalt The super-concentrated desalting technology with a 6-effect falling film evaporator and forced circulating evaporator has been adopted. It is characterized with maturity & reliability, low energy consumption (steam ratio <0.27) and high efficiency desalting.

The refined liquor(pregnant) from security filtration is sent to the pregnant heat exchange section by pump.

The spent liquor being 80 °C from the heat exchange is respectively sent to the 6th and 4th evaporator unit.

3.6 Pregnant heat exchange

The liquor entering the 6th evaporator unit reaches the requirements after it is concentrated through the 6th effect evaporator and the 5th effect evaporator.

This section is for simultaneously satisfying the seed precipitation temperature requirements and recycling the heat resource. The two mediums of heat exchange are respectively the pregnant liquor (inlet temperature being 100-105 °C and outlet temperature being 60-62 °C) from the security filtration and the spent liquor (inlet temperature being 45-48 °C and outlet temperature being 80-82 °C) from the product filtration.

The liquor entering the 4th evaporator unit is concentrated through the 4th evaporator, the 3th evaporator, the 2nd evaporator and the 1st evaporator, and enters 2-stage flash evaporation, then respectively enters the 3th stage flash evaporation and super-concentrated evaporator in proportion. The concentration of the evaporation liquor after 3-stage flash evaporation reaches the plant requirements; the concentration of caustic alkali Na20 in the liquor after concentration through super-concentrated evaporator reaches 320g/l. The sodium carbonate crystals are removed from the slurry with salt grain discharged from the super-concentrated evaporator through the settler and filter, the overflow from the settler and the filtrate from the filter as well as the liquor from 5-effect evaporator and 3-stage flash evaporation converge to become the mother liquor.

The equipment for this section is a plate-type heat exchanger. 3.7 Seed precipitation The seed precipitation system consists of large mechanical agitation precipitators of diameter 14m, intermediate cooling unit and classifier (hydraulic cyclone). The sodium aluminate pregnant liquor after cooling is sent to the precipitation system where hydrate seed is added, and then the hydrate is generated by the precipitation over 45-50 hours. Guangxi is in semi-tropical area in south of China, with a hot climate and high humidity. This limits the natural drop in temperature during precipitation and the increase in

Mother liquor, liquor after causticization, less evaporation green liquor and liquid caustic soda with alkali addition are mixed to be the test liquor with caustic alkali concentration being 245g/l,

65

and then it is sent to the bauxite grinding by pump after circulation through the liquor storage tank.

The four production lines ran under full load in the last month of 2009 along with the recovery of the alumina market. The technical & economic indexes were obviously improved and the production cost is reduced, being below RMB 1100, thus the profit was achieved in July to December 2009.

3.11 Productfiltration& washing The underflow from the hydraulic cyclone classifier flows to an aluminum hydrate slurry tank and the secondary condensate from the evaporation section enters a hot water tank, and the hydrate slurry and the hot water respectively are sent to a horizontal pan filter for hydrate separation & washing.

Production statistical indexes of Guangxi Huayin Aluminum Corporation Limited in July to December 2009 are indicated below.

The spent liquor obtained through filtration is sent to the spent tank in seed filtration by pump. The washing liquor is sent to the red mud washing settler by pump and the filer cake is sent to the calciner or hydrate silo through a belt conveyor.

No

This section is mainly equipped with a large horizontal pan filter. Separation and 2-stage reverse washing of the aluminum hydrate is carried out in this filter. Some dehydrating agent is added so as the water content in aluminum hydrate is below 3%. 3.12 Calcination

Item

Unit

1

Bauxite component

2

Output

3

Product One-class percent

kt/a

product

Comprehensive energy consumption

The aluminum hydrate calcination uses a gas suspension calciner. The aluminum hydrate from the product filtration & washing or the aluminum hydrate silo, is sent to a small silo before the calciner, through the belt conveyor.

4.1 4.2 4.3

The aluminum hydrate in the small silo is sent to a gas suspension calciner through the feeding weigher and screw feeder, being burned to the required alumina specification.

6 7 8 9 10 11

Steam Electrical power Coal gas Alkali (converted into Na2C03) Bauxite Lime Fresh water Compressed air Relative digestion rate Pregnant liquor output Alumina recycling rate

There are two heat exchange stages in calcination: heat exchange between the aluminum hydrate and the exhausted gas generated through coal gas burning and between alumina and air required for coal gas burning.

I 12

The exhausted gas, with dust generated during calcinations, is discharged after passage through electrostatic filters to meet discharge standards.

827.3 ( converted into 1654.7 for whole year)

|

Sandy AI2O3

4

5

Production statistical I indexes in July to December 2009 AI2O3 53.86% A/S 8.504

%

99.78

GJ/t(kg standard coal/t)

10.51 (358.8)

t/t kWh/t Nma/t

2.438 246.3 617.1

kg/t

118.7

t/t

2.26 243.8 2.97 39.93 93.4 87.63 81Ό1

kg/t

t/t ma

%

kg/ma

%

|

5 Conclusions Huayin phase I alumina project is a green field project with the largest disposable investment and construction scale in the Chinese alumina industry. It achieved the target of full commissioning in three years, and production & standards achieved after 1 year's commission, and thus establishes a new record in China alumina construction history.

The utilization of calciner waste heat has been greatly improved in later stage of this project, so RMB 10 million economic benefits per year can be gained. 4 Actual production conditions

Guangxi Huayin Aluminum Corporation Limited as a new plant can get away from the difficult position in the international finance crisis, because this plant possesses the right technical guidelines, excellent engineering design, better construction quality and strong production management, thus various indexes are at international advanced levels, especially comprehensive energy consumption per ton alumina production, being only 10.51GJ reaches the international leading standard.

The first production line of this project was put into production in December 2007; it took only 33 hours to go through the whole alumina production process flow. The second and third production line were successively finished in March and April 2008, and the commissioning of the fourth production line in May 2008 indicated that Huayin phase I alumina project was fully finished. However, due to the international finance crisis coming, the China aluminum product was overstocked heavily and the alumina price declined significantly and continuously. So Guangxi Huayin Aluminum Corporation Limited had to reduce the output to limit loss, and ran at only 50% production before May 2009.

References 1. Tian Fengming Fan Na, Potentional and availability analysis on bauxite resources in Guangxi, China Mining, 19th Volume, 5th Period, May 2010.

Guangxi Huayin Aluminum Corporation Limited checked the performance of Huayin phase I alumina project with the technical guidance of China Aluminum International Engineering Corporation Limited in June 2009. The checking results indicate that various indexes reached or exceeded the designed value, thus the performance checking was qualified.

2. Deng Feiyue, Huang Manxiang, Processing technology performance study of bauxite ore in Debao Guangxi, Mining study and development, 27th volume, 1st period, Feb. 2007. 3. Tang Shitai, Development on grinding process technology

66

with two-section grinding and hydro-cyclone classification in alumina production. 11th volume, 5th period, Oct. 2001.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ADVANCED PROCESS CONTROL IN THE EVAPORATION UNIT C.Satish Kumar1, Tonmoy Banerjee1, Uttam Giri1, Rosalin Pradhan1, Ramu Saha1, Pratichi Pattnaik1 Vedanta Aluminium Limited; Lanjigarh; Orissa - 766027; India Keywords: Advanced Process Control, RMPCT, steam economy Abstract

Process Description

Energy consumption is a significant constituent of production cost in an Alumina refinery. Indian refineries are striving to reach global benchmarks and have made conscious efforts to minimize energy consumption. Evaporation plays a significant role in conserving energy and is responsible for 30% of the thermal energy consumption. In the Evaporation unit, due to a large number of interacting processes, frequent disturbances and the presence of single line equipment, it is necessary to maximize the utilization of assets to ensure the production volumes at targeted efficiencies. The need for reduction in process variability resulting in optimizing operations, challenged by plant constraints, was derived by adopting Advanced Process Control (APC) application in the Evaporation unit.

The evaporation unit removes water that has entered to the liquor circuit from bauxite moisture, mud washing, hydrate washing and miscellaneous dilution streams. The evaporation section consists of two identical parallel batteries of six backward feed, falling film evaporator units in each battery. The weak Spent liquor is collected and received in evaporator feed tanks. The weak liquor is evaporated in the six backward feed falling film evaporators and strong liquor is stored in the evaporator test tanks. In each evaporator train six nos. online calandrias are provided as tubular heat exchangers according to the falling film principle. Each calandria is equipped with a detachable cover and liquor distribution system (on top) for even distribution of spent liquor into the tubes.

APC incorporates a matrix of modeled dynamic process responses with an optimizer to maintain the unit at the most profitable vertex of the allowable operating envelope. This paper will discuss the area selection criteria, activities carried out during the project tenure and post benefit analysis.

Vapor separators and ducts are provided next to the calandrias for efficient separation of the vapor from the product. Low pressure (LP) steam provides the energy required for evaporation. LP steam is fed to the 1st effect from main header through flow control loop. A desuper heater is provided in the steam main header in order to extract superheat from the steam coming from main header and thereby preventing high temperature steam entering the evaporator body. Spent liquor is pumped from the feed tanks to feed flash vessel connected to 5th effect unit where it gets flashed and evaporated and then transferred successively to feed flash vessels connected to the 6th unit and mixing condenser for further concentration. After flashing the feed in three flash vessels, the liquor is pumped to the 6th unit. From the 6th unit, the liquor is pumped successively to the 5th, 4th, 3rd, 2nd and 1st units through transfer pumps. Finally the concentrate coming out of the 1st unit is flashed successively in product flash vessels connected to 3rd unit followed by 4th unit and thus final concentration is achieved. Each body is provided with a recirculation pump and corresponding piping. The recirculation pump circulates the required flow of liquor ensuring the proper wetting rates (to avoid any dry spots on the evaporating surface) to the top of the calandria on liquid distributor. The steam condensate from the 1st effect and 2nd effect is extracted withoutflashing.Balance process condensate is flashed three times before being extracted at 60 °C. Non-condensable gases from 1st unit to 6th unit are connected to a common header and in turn are connected to a condenser. The vaporsfromthe 6th unit & last feed flash vessel are condensed in a direct mixing type condenser (Barometric Condenser) with cooling water from an alkaline cooling tower entering condenser at 35°C and returning at 50°C. The vacuum in the evaporator is achieved & maintained through a two stage steam ejector system. LP Steam is being used in the ejectors to maintain the desired vacuum in the system.

Introduction The challenges for any alumina refinery are to minimize the cost of production per tonne of alumina, while meeting safety and environmental considerations, along with maximum production volumes. In view of the recent step increase in the cost of energy generated from fossil fuels, Alumina refinery operators are constantly looking for various ways and means to reduce specific energy consumption per tonne of alumina production. Utilization of new control technologies using existing infrastructure with a reduced support team has emerged as a new dimension for significant economic saving in alumina production. Multivariable predictive control technology becomes one of the main tools to optimize the capital investment. M/s Vedanta Aluminium Limited is located at Lanjigarh, Orissa state, India, having a capacity of 1MMTPA smelter grade alumina production, employing the Bayer process with low temperature and pressure digestion. Evaporation consumes about 30% of the thermal energy and is an ideal unit in Alumina refinery for APC implementation to generate significant potential savings. Apart from this, other benefits accrue, such as uniform test liquor caustic concentration, which is a key parameter in the Digestion unit for improving productivity. In classic process control, a controller brings a single value close to a set point using a single manipulated variable, or it can involve a reactive process based on currently measured conditions. Classic process control can do things automatically, but it can't respond to other kinds of variables, such as cost constraints. Process variability reduction resulting in optimizing operations challenged by plant constraints was derived by deploying an effective APC application in the Evaporation unit.

Area Selection criteria Digestion, Calcination and Evaporation are the three major areas for APC implementation to generate profits in terms of yield improvement, increase in throughput, minimization of standard

69

deviations of process parameters and reduced energy consumption. The digestion process has the most potential for APC benefits because it is a key production unit, critical for sustainability of the process and a good state of digestion control. Due to frequent changes in the quality of the bauxite feed to the refinery, it is difficult to control the process parameters in the digestion unit. So the option of APC implementation in the digestion unit was kept on hold tempora-rily. However, auto control of alumina to caustic ratio with test liquor and bauxite slurry has been implemented.

future response with respect to control. The prediction model will send communication to optimizer for taking moves on targets of Controlled Variables (CVs) and Manipulated Variables (MVs) and final control element will take action accordingly. The basic idea of the control strategy is to take action before a disturbance reaches the CVs. Disturbances which enter the process are detected using the predictive model and appropriate changes are made in the manipulated variables such that the controlled variables are held constant. Figure 2. shows the behavior of variables with and without control process.

The existing Calcination unit has been planned for an upgrade in throughput by 20%. After completion of the upgrade and achieving the rated throughput, APC implementation will be done in Calcination. The two Evaporation units are designed for 300 Tonnes/hr each. Due to lower bauxite quality than design (-37% Trihydrate alumina (THA) compared with a design of 41% THA), mud generation is on the higher side which results in more wash water demand per tonne of red mud to minimize residual caustic losses. So both the Evaporation units are operated with 65%-70% of their capacity. To maintain consistent product quality, minimizing the effect of disturbance and enhancing performance of single line equipment, raised a challenge to improve the steam economy, so as to reduce the specific energy consumption per tonne of alumina production.

MV Past Move* DV P»*t Vahtei CV Unforced CV Current Velue CV Unforced +Bi»s SP Funnel Error MV Move PUn CV ConlroDeil

Figure 2. RMPCT controlling a CV inside limits APC brings several values close to optimal targets simultaneously using several manipulated variables simultaneously. Also, APC includes a predictive process based on a process model that foresees how variables will behave in the future, and the controllers in an APC system find optimal trade-offs in case of conflicts among the goals. Because several actuators act simultaneously to achieve the best trade-offs, the application can achieve its goals in the fastest or most efficient fashion. In addition, APC's predictive character allows early recognition of potential violations and timely implementation of remedies, which means improved stability and reduced defects. It also lets applications work nearer to process constraints, which further reduces costs. A major objective of APC is to minimize the dotted area (Error) shown in Fig.2

Robust Multivariable Predictive Control Technology (RMPCT) Control designs in the process industry are almost exclusively based on PID controllers these days. Even though they are simple to implement and easily integrated into the control system, they quickly reach their limits when more complexity is involved. The development of advanced control theory has received considerable attention and applications in the process industry are gradually increasing. APC opens new opportunities, as even complex situations can be mathematically described with process parameters or variables and used for automatic, flexible plant operation. APC employs robust control techniques and can handle several constraints simultaneously by handling multiple inputs and outputs, and maximizes the profits derived by pushing the operations closer to limits.

APC Advantages ■ Standard deviations are minimized and faster adjustment can be done ■ Improvement in throughput, yield and consistent product quality ■ Manual intervention will become significantly less, which leads to a reduction in operator stress ■ Lifetime of the plant with the existing automation increases with less maintenance work ■ Significant reduction in energy consumption

Multivariable controller Predicted Steady Optimizer state values

H

Prediction

' Targets, WV targets

Process

Predicted! Controller values

Project Implementation Steps

•"H Model

The scope of the project was the implementation of APC in both evaporation trains of the alumina refinery. A detailed project planning schedule was prepared in collaboration with M/s Honeywell. The following activities were carried out over five months.

Prediction error

Figure 1. Advanced process control components Advanced process control consists of optimizer and prediction model. The function of the prediction model is to predict the

70

Data Collection and Analysis Base case data was collected to understand the operation of evaporation unit and evaluate the variability in process parameters. Six minute average data was considered for variability analysis. Minimum, maximum, average and standard deviation of all process parameters were collected for 15 days to understand the existing operation, and major process upsets during this period were discarded.

Functional Design The functional design specification was based on preliminary plant test and earlier data analysis, discussion with plant operating personnel and engineers and information gathered on operating constraints and future operating strategy. Since the operating variables in the evaporation unit are interacting with each other, RMPCT was selected for advanced control. Two identical RMPCT were selected for evaporation train land 2 to provide reliable operational control and optimization objectives, while maintaining the constraints within specifications. To achieve this, RMPCT manipulates a set of independent variables MVs to maintain a set of dependent variables CVs at target or within constraints. The controller also uses disturbance variables (DVs) as feed forward control to reject the disturbances. A linear programmed optimizer and dynamic control algorithm coordinate the movement of MVs. The optimizer functions were decided to maximize steam economy, improve product density control and to maintain process condensate conductivity within acceptable limit.

Preliminary Plant Test The main objective of Preliminary plant test was as follows: Estimate step size, dead time and settling time of the variables, Evaluate the tuning of the DCS controllers and changes accordingly wherever is required, Check the existing instrumentation. Preliminary plant test was performed for one week. During preliminary plant test, the operation of a few controllers was improved from earlier Manual mode to Auto mode by making the adjustment in tuning factors of the controllers. Fig.3 shows the controllability of steam flow in AUTO mode. This valve was operated manually earlier. After taking the controller in AUTO the fluctuation in steam flow has come down significantly.

^

7

^

Step Test A step test plan was prepared based on preliminary plant tests, functional design and earlier operational experience. The step test was required to build the process model, which is the basis of predictive control. The step test plan was designed to build up a good model without interrupting normal plant operation. A test was carried out for 15 days in both evaporation trains one after another. Eleven numbers of manipulated variables were considered for the step test plan. Those variables were Steam flow controller, Feed low controller, Bypass feed flow controller, Barometric condenser pressure controller, Barometric condenser cooling water flow controller and each effect's level controller. During the step test, one variable was moved at a time and all other variables were kept in auto mode. The test was carried out round the clock and a normal range of operation of the evaporation unit was ensured. The results of each step were monitored and analyzed continuously. Based on this analysis step size and numbers of steps were increased and or decreased in a few cases.

^

Figure 3. Steam Flow control before and after tuning

A few snapshots of the correlation between the MVs and DVs with CVs obtained after the step test are shown below. FIR Order 200 Lap Order 2 Settle T = 2 0 0 TfSettie = 1Q3 FIR Form = Pos Rank = 3Tria! 7 G[sl= .000171

■i \ I *.„

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„***"·♦,'', „

S

,

50

V^yT^s^yC^V 150 2£ 100 -20$

193**.2 + 27 8s * 1

Figure 5. shows Correlation between Product Density and Feed Flow

»p •x*

.*,'.*<

Figure 4. shows Product Flash Vessel#3 level before and after tuning

71

FIR Order 1 2 0 L a p Order 1 Settle T - 1 2 0 TfSettle « 115 FIR Form « Pos Rank = 2 Trial 5 Q 1 Gis) - 0 6 6 26.7* + 1

After the detail validation of the model, the APC Server configuration was done in the existing Experion Process Knowledge System (EPKS) network. Ole Process Control (OPC) client was installed in the APC server for the communication to the OPC server. Redirection manager was installed in APC server machine to communicate to the redundant server in case the primary server failed. A Distributed Control System (DCS) resident watchdog timer was maintained to allow mode shedding of MVs in case of any communication loss between DCS & the profit controller. The APC controller plan moves for all the MVs, and those moves, will be written to DCS MVs. Finally, the APC model was successfully commissioned in both evaporation units, 15 weeks after the start of the project.

120

Figure 6. shows Correlation between Vacuum and 1st effect temperature difference Training Two types of training were given: 1). Engineer Training before Step test: 25 engineers attended engineer training. The objective of the training was to understand the basic function, design and implementation of APC. Also, identification of control models and tuning of the controller were covered in the training. 2). Operator Training after Commissioning: All concerned operators were trained under the operators training in which procedures for taking APC online, offline, and emergency control were covered.

Benefit Analysis Reduction of Variability in CVs Significant reduction of the variability of all the CVs has been achieved as measured by standard deviation and shown in normal distribution graphs. For example, Product density is the important controlled variable in the evaporation circuit for product quality control. Before the implementation of APC the product density variability had a range from 1.321 to 1.350 whereas, after APC implementation the range narrowed to between 1.302 and 1.310. Product density standard deviation has reduced from 0.006 to 0.001. The following figure shows the variability reduction before and after APC implementation.

Detail Design and Commissioning The APC detailed design was developed based on the information generated during the Plant step test. To maximize the steam economy, the optimizer was designed to minimize the steam flow, maximize the vacuum, minimize the product density and minimize the temperature difference across each effect. In order to achieve the minimum temperature difference across each effect, optimizer was designed to maximize the feed flow through the evaporator and minimize the bypass feed. During the step test it was observed that the step change in level controller of each effect didn't give any correlation with CVs. So the basic controller was designed by taking five MVs which are Feed Flow Controller , Bypass feed flow controller, Steam Flow controller, Barometric condenser pressure controller, Barometric condenser cooling water flow controller and fifteen numbers of CVs out of which the important variables were Product Density, Process Condensate Conductivity, Temperature difference across each effect, first effect jacket temperature, Feed density, feed temperature, inlet water temperature to barometric condenser, deviation of feed flow and barometric condenser pressure from set point has been taken as DVs in the detail design. Preference has been given for MVs control in the following order: Feed flow Controller, Direct condenser pressure Controller, Steam Flow Controller, Bypass feed flow controller, Direct condenser cooling water flow controller.

Figure 7. shows Normal distribution curve for the Product density before implementation of APC

Two types of limits have been set for all MVs and CVs, one is the engineering limit and other one is the operating limit. The engineering limit has been set to ensure safe operating range of equipment. Operating limits are set to serve the process control and optimization. The operating limit is accessible to the control room engineer. The high limit and low limit of operating range are set based on forward unit's requirement and overall condition of the plant. After the detail design, the model was run in offline mode at M/s Honeywell, Pune. The validations with respect to achieving the objective were carried out at different input conditions. The direction of movement of each CV with respect to the step change in each of the MVs was checked and validated.

Figure 8. shows Normal distribution curve for the Product density after implementation of APC.

72

Maximization of Steam Economy: Steam Economy has increased by an average of 2% after implementation of APC. This was achieved mainly by minimizing the direct condenser pressure and optimizing process parameters continuously. Before APC implementation, the average direct condenser pressure was -648 rnmHg(g), which improved to -655 mmHg(g) after implementation of APC. Previously, maintaining pressure below -655 mmHg (g) was a constraint for operation as it was leading to an increase in process condensate conductivity intermittently, because of disturbances created by other process parameters. But under present conditions, due to continuous optimization by APC, lower pressures up to -662 mmHg were achieved without deterioration in the process condensate, as well as the direct condenser cooling water quality. Barometric Cond. Pr.Actual kmujiv2}rjrtwNe>HMirtN»TH«ujr>.cnH«inr«.o>t -645 - — — B i—,.y,i

-650

-655 -660 *0*¥4Φ*Ε*Τ*

m

; -665

I -670 Figure 9. shows the direction of movement of condenser pressure towards high limit after giving a step change in the limits. Conclusion RMPCT implementation in the Evaporation unit at Lanjigarh was completed within the planned project tenure and has proven successful benefits. RMPTC provided an opportunity to operate the plant closer to the operational constraints which helped plant debottlenecking without adding assets. Project outcomes such as minimizing standard deviation of process parameters, improvement in throughput, plant uptime with less breakdown maintenance and finally improvement in steam economy were successfully delivered. The way forward involves Digestion and Calcination as the next key units for APC implementation to maximize the production volume and throughout.

References 1. Ayana Oliveira, Jefferson Batista, Jedson Santos, Marcia Ribero, Jorge Charr, Rafael Lopes "Advanced process control in alumina digestion unit" Page no: 91-96, Light Metals 2009 2. Robert K.Jonas, "Application and benefits of advanced control to alumina refining", page no: 43-49, Light Metals 2004 3. Freeman N, Pandya M and De Roos R, "Simulation applications for alumina refining", page no: 52-55, Alumina quality workshop 2008 4. Oliveria A, Batista J, Ribero M, Charr J and Lopes R, "Robust multivariable predictive control technology implementation in an alumina digestion unit", page no: 47-51, Alumina quality workshop 2008

Acknowledgements The authors wish to gratefully acknowledge Mr. Jaydeep Bhattacharya and Mr. Vinoth Ramachandran from M/s Honeywell, Pune for providing assistance during the project.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

IMPROVEMENTS IN SMELTER GRADE ALUMINA QUALITY AT CLARENDON ALUMINA WORKS Ruth Shaw1, Ajamu Duncan2, Marlon Crosdale3 Clarendon Alumina Works; Halse Hall; Maypen, Clarendon, Jamaica Keywords: Diastar, Process Management, Hydro Cyclone Abstract Alcoa's Clarendon Alumina Works recorded a significant improvement in its smelter grade alumina, SGA, quality. This marked improvement came about as a result of changes in green liquor filtration technology and process management in precipitation. This paper discusses the impact that these changes had on iron, calcia, titanium, silica and soda in SGA. Introduction Alcoa's Clarendon Alumina Works, CAW, located in South Central Jamaica, is a low temperature digestion plant with an annual production capacity of approximately 1.4 million tonnes SGA. The refinery is ISO 9001:2008 certified.

Parameter

2006

2007

2008

2009

%Fe 2 0 3

0.019

0.010

0.008

0.009

%Si0 2

0.017

0.017

0.016

0.014

%NazO

0.45

0.45

0.40

0.43

%CaO

0.057

0.054

0.051

0.050

%Ti02

0.002

0.002

0.001

0.001

Table 1: CAW Alumina Quality 2006 - 2009

Since 2007 the quality of the alumina produced at CAW has made progressive improvement as measured by the Customer Satisfaction Index (CSI). The CSI is the metric that measures the refinery's ability to meet critical parameters required by the smelter customers. For Clarendon these parameters include iron oxide (Fe203), silica (Si02), sodium oxide (Na20) and calcia (CaO). These parameters are affected by the quality of the liquor sent to precipitation and operating practices. At the beginning of 2007 the CSI averaged 55%. Following the commissioning of improved technologies and improvement in process management systems, the CSI increased to an average 92% in 2009. See Figure 1. Table 1 show the trend of the alumina quality parameters in SGA for the period 2006 - 2009.

Iron oxide (Fe203), titanium (Ti02) and calcia (CaO) all showed a step reduction which was directly related to the commissioning of the high pressure diastar filters. This resulted in a significant drop in the particulates in the liquor sent to the precipitation circuit. Further, improvement in thefiltrationprocess control and monitoring minimized deviations and sustained gains. In precipitation, strategic changes in process management and hydrate classification using hydro-cyclones were the main contributors to the improvement in silica and soda levels in SGA. Improvement in Fe203, CaO and Ti0 2

Customer Satifaction index

The Effects of High Pressure Diastar Filters grioo

Since the construction of the refinery in the early 1970s the plant has employed the use of sand filtration technology to remove suspended particles from the liquor being fed to the precipitation circuit. In this process the liquor is filtered by gravity through a bed of suitably sized sand and the filtrate fed to the precipitation stage.

E 80 c

t

40

ä

20

There were several issues associated with sand filters that resulted in high particulates in filtrate. These included:

m** 1

y**1

10*

„fi*

y**

<**

* *

Frequent screen ruptures and sand leakages Changes in the particle size distribution of the sand bed over the life cycle of the filter

0ec^

Figure 1: Customer Satisfaction Index 2007 - 2009

75

No efficient way of flushing out the mud entrained in the sand bed during the filtration cycle.

Operations Management: Re-clothing practices Process Control to prevent Cloth Degradation Inventory Management of Filter parts Cloth Design

As a consequence of aforementioned, the operations experienced high variability in the control of iron oxide, titanium oxide and calcia in SGA.

Process Management: Changes to dosing of flocculants to thickeners A visual management system Spigot sample to reduce time taken to identify defective filter elements.

Iron in SGA and Filtrate Solids — -Iron m SGA

——Filtrate Solids

0.04 0.035 0.03 g

0.025

S c

0<02

|

"I 4"AJ

ITP Calcia and Calcia in SGA — -Calcia to Precip

io I

0.015

»10

i^>->^.r^·.

0.01

90

o 1

/V

œ

70

16°

-20

(*>«r°1

^ * &

a#*

i9itt&

g 0ec"^

40

,5

Figure 2: Iron in SGA and Filtrate solids 2006 - 2009 Titanium in SGA and Filtrate Solids — »Titanium in SGA

/^^

·§ 50

-30 i^·01

0.09 0.08 0.07

<:■■■: M - k k

Λ -0M

a!

0.05 * | \

/

, V v - ^

\

■■ ' \

•

t

χ

X

^

•

* ' Λ * * ' * ,

0.04 O

30 20

0.03 0.02

10

0.01

yvtfv06

— F i l t r a t e Soiids

k f\

; :

80

0.005 0 y,***

— C a l c i a in SGA

oe*06

itf**1

oec-0 1

ypSfi

o e

c ^

&*&

o e

c-^

Figure 4: Calcia in SGA and LTP Calcia solids 2007 - 2009 Improvement in Na 2 0 and Si0 2

y****

v***

ju*°1

ç**

1

jtfv**

o e c ^ „g***

In 2006 CAW s precipitation department advanced its classification system by moving from gravity classifiers to hydrocyclone classifiers. These changes required modification of the hydrate management and first precipitator control strategies so as to assure that the soda and silica levels in product meets the customer specification. Figures 5 and 6 illustrate the improvement in both the variation and absolute value of the Soda and Silica in product.

oe c-«

Soda in SGA ofi

Figure 3: Titanium in SGA and Filtrate solids 2007 - 2009

0.55

During the 2007 plant expansion, the refinery replaced the sand filters with Diastar filters which are widely regarded as the most advanced fine filtration technology.

0.5

The change to Diastar filters resulted in the step reduction in iron oxide and titanium oxide in SGA. Refer to Figures 2 and 3.

* 1

cu

* 0.35 0.3

Process Improvements - Diastar Filters and Thickener Operations

0.25 -

Following the commissioning, the refinery was faced with the challenge of adapting to the Diastar operations. Significant improvements were made in the following areas with the gains in calcia control being realized during 2008. See Figure 4.

**<*

#**

tfH*1

yfi*

ç**

***

„**

*#*

Figure 5: Soda in SGA 2006 - 2009

16

o*&

«*'*

CAW experienced high variability in its first precipitator soda control immediately following the introduction of the hydrocyclone classifiers in 2007. Changes in head tank control with the improved classification system positively influenced the fixed soda in SGA. These changes allowed the sustained reduction in the average value of soda in product from 0.45% in the second half of 2008 to 0.41% in 2009 and a reduction in variability from 0.09% in 2008 to 0.03% in 2009. Silica in product is CAW's highest-weighted component capturing over 35 % of the CSI. Silica in SGA is affected by inter alia hydrate seed in contact with liquor for extended periods of time. Vessel failures in precipitation are one of the leading causes that present this condition. Silica in SGA - Box and Whisker

T

s

-E 3-

H2 - 08

Π.

H1 - 09

H2- 09

Figure 6: Silica - Box and Whisker plot Introduction of a new hydrate management system and the revision to the procedure for returning a failed vessel to service has eliminated occurrences of silica excursions. In 2007 silica averaged 0.17 + 0.02 while in 2009 the silica averaged 0.14 ±0.02 Conclusion CAW has significantly reduced iron oxide, calcia, titanium oxide, soda and silica in SGA through: Switching from gravity sand filter to short cycle Diastar pressure filters Improved process management and control of the Diastar and precipitation operations CAW has seen marked improvement in its product quality and continues to drive for improvement through the process of best practice transfer. References 1 Diastars filters, manufactured by Filters Gaudfrin (France), are cylindrical shaped pressure vessels with a cone shaped bottom and a domed shape removable top. They utilize a filter cloth as the filter medium, similar to a Kelly filter.

77

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINA and BAUXITE

Red Mud SESSION CHAIR

Linda-Therese Kristiansen Hydro Aluminium Oslo, Norway

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

APPLICATION OF NANOFILTRATION TECHNOLOGY TO IMPROVE SEA WATER NEUTRALIZATION OF BAYER PROCESS RESIDUE Kelvin Taylor1}, Mark Mullett^, Lee Fergusson2), Helen Adamson1* & Juerg Wehrli1} υ Hatch Associates Pty Ltd, 144 Stirling Street, Perth, WA 6000, Australia 2) Virotec Global Solutions Pty Ltd, Coomera Waters, QLD 4209, Australia Keywords: Nanofiltration, Membranes, Red Mud, Bayer Process, Sea Water Neutralization plants. Typical sea water NF plant operating pressures are around 18 bar compared to around 60 bar for SWRO plants [2]. The capacity for concentrating calcium (Ca) and magnesium (Mg) ions is therefore likely to be limited by precipitation rather than osmotic pressure.

Abstract Sea water neutralization of alkaline Bayer process solid waste residue is a sustainable solution to turn a hazardous waste material into a benign, non-hazardous material that can be re-used in some applications. The concentrations of the active neutralizing agents in sea water, calcium and magnesium, are low and large volumes of sea water are required to neutralize the alkaline residue.

The production of alumina by the Bayer process results in the generation of large quantities of "red mud". Between one and two tonnes of residue are typically generated for every one tonne of alumina produced and, due to its high alkalinity and pH, the red mud generated from alumina refineries is classified internationally under the Basel Convention (Article 1, paragraph la) as a "hazardous waste" (classification #B2110) and its trans-boundary movement is tightly controlled [3].

Nanofiltration of sea water can produce a concentrate containing up to four times the calcium and magnesium levels in sea water, while the mono-valent ions, like sodium and chloride, are only marginally concentrated. This significantly reduces the volume of sea water required for the neutralization process and hence the size of neutralization equipment. Other advantages include improved reaction kinetics and reduced salinity of the neutralized residue, making it more suitable for certain applications. The nanofiltration permeate, with lower scaling potential, is also an improved feed for sea water reverse osmosis plants to produce potable water.

The alumina industry has for many years sought ways to minimize and reduce the environmental and social impacts of red mud and alumina refineries face two basic choices in managing their red mud residue, either:

Introduction

i)

store the residue indefinitely at considerable cost; or

ii)

attempt to reuse it in some form [4],[5].

Long-term storage costs and environmental liabilities are high, and it is estimated that management of residue in Australia alone costs more than AUD80 million per year [6].

Nanofiltration (NF) is a membrane-based separation method that is similar to reverse osmosis (RO) but the NF membranes are more selective in retaining chemical species than RO membranes. RO membranes retain >99.0% of all chemical species in solution, except for some small, polar organic molecules such as acetone and acetaldehyde. NF membranes on the other hand will allow mono-valent ions such as sodium and chloride to pass through to permeate but will retain divalent species, such as calcium and magnesium, in the retentate.

Refinery waste reuse requires some form of conversion from its highly caustic state to a more environmentally benign and "userfriendly" state for both transport and reuse applications. Thus, it is desirable to neutralize refinery residues in a way that they are no longer highly caustic, irrespective of their potential or actual reusability. Most neutralization options that attempt to address red mud residue storage and management problems do not make full use of the inherent acid neutralizing, metals and phosphate binding, and odour destroying properties of the neutralized red mud residue. One method for neutralizing alumina refinery residues, called Basecon Technology can be used to economically convert the residue and thereby imbues it with environmental and commercial value [7]-[ll]. Such a technology can reduce the long-term environmental, financial and social liabilities of storing caustic residues, and has the added advantage of producing a reusable solid material, called Bauxsol Raw Material (BRM), which has favorable environmental and industrial reuse characteristics. Different reactants, including sea water, brines, concentrated brines and metal salts can be used to neutralize red mud residue.

In general, the species that are retained by an NF membrane are multi-valent inorganic ions and large organic molecules. Different NF membranes have different rejection characteristics depending on their surface chemistry and surface charge density and are normally characterized by measuring the rejection of magnesium sulfate, typically in a range between 95% and 99%. Typical applications for NF membranes are concentration of dissolved metals in acid solutions, concentration of dyes and sugars, removal of chloride from sulfate solution and removal of calcium from solution to stop scaling. Nanofiltration technology has also been applied to pre-treat sea water for desalination plants, where the NF membranes have been used to remove calcium, magnesium and sulfate prior to reverse osmosis treatment. It has been shown that the RO membranes, or thermal equipment in the case of distillation plants, operate more efficiently if NF pretreatment is applied [1].

Sea water neutralization is the most widely implemented neutralization process, due to its cost-efficiencies and readily available feedstock [12] for alumina refineries located close to the sea, and having environmental licenses that provide access to sea water. When sea water is used to neutralize red mud, the volume of sea water added is between three and ten times the volume of

The concentration of species retained by the NF membranes is low and hence the operating pressure is also low when compared to the operating pressures of sea water reverse osmosis (SWRO)

81

red mud slurry being treated (also refer Table 4). Nanofiltration now offers an opportunity to significantly reduce the volumes required. The expected advantages of using NF concentrate to neutralize red mud are: •

Reduction in reacting volume and therefore tanks, pumps, piping and the capital and operating costs associated with this equipment.

•

Reduction in salinity of the neutralized product, thus making it more useable for soil remediation and other similar applications.

•

MEMBRANE FILTRATION

m

Production of an NF permeate stream that will perform better in SWRO plants than untreated sea water.

PERMEATE

Figure 1: Schematic Diagram of Test Arrangement

An R&D program was initiated to determine the extent to which sea water could be concentrated using NF membranes before scaling commenced, or before the osmotic pressure became excessive. In addition the comparative reactivity of the sea water NF concentrate and untreated sea water with red mud was investigated in order to gain an understanding of the equipment capacities required and capital reduction benefits that could reasonably be anticipated by alumina refiners. Experimental Membrane tests were conducted in two phases: laboratory bench tests (2L start volume), and pilot scale (lkL start volume). Cross flow filtration over a membrane surface under pressure was used in all tests to effect the chemical separations, the aim being to concentrate magnesium and calcium by rejecting it to the retentate, while passing sodium chloride to the permeate. All tests were run using a sample of sea water collected from the Western Australian coast near Perth, and pre-filtered using 5 μπι cartridge filters.

Figure 2: Pilot Test Rig

Two types of tests were conducted. Perturbation tests (P-tests) determined the relationship between rejection, pressure and flux for a given membrane and solution type, while recovery tests (Rtests) provided the relationship between operating pressure and flux as the volume recovery and feed analyte concentration increased. The results were used to:

The first pilot scale test run was performed at a constant pressure of 3200 kPa. The flux was allowed to decline from an initial flux rate of 45 LMH as the osmotic pressure increased and conductivity, temperature and pH measurements were taken at increments of 10% recovery up to a maximum recovery level. The maximum recovery level was characterized by a significant rate of flux decline.

i. Determine flux behavior over a range of operating pressures for a given solution matrix and assist in determining which membrane type is most suitable for a given application. ii.

Feed and permeate samples were collected at various recovery points and analyzed for Ca, Mg and S by ICP-OES. The concentrated feed sample produced at the maximum recovery level of 80% during the first pilot run was then used for comparisons of neutralization characteristics with untreated sea water by titration of the free alkalinity of a generic Bayer residue water (refer Figure 3). Analytical results from this first pilot run were also used as the input to modeling of the neutralization using Basecon Technology (refer Table 4).

At pilot scale, provide quantitative data that facilitates an engineering design process.

During R-testing, the retentate is recycled to the feed tank such that the concentration of rejected species increases, causing the osmotic pressure to increase toward a maximum value. A schematic diagram showing the layout of the equipment for both the laboratory and pilot scale tests is shown in Figure 1.

A repeat of the previous R-test was performed at a reduced starting pressure of between 1000 and 2400 kPa and then gradually increased to achieve a constant flux rate of 35 LMH. As recovery increased beyond 60%, samples were collected at recovery increments of 5%. P-tests were conducted concurrently during the test at a range of recovery levels to better determine the ideal recovery recommended for a full scale process.

Laboratory scale tests were conducted using 21.5 cm2 flat sheet membranes that represented three commercially available spiral wound NF membranes. The capacity of each membrane to concentrate Ca and Mg into the retentate at suitable NF operating pressures (<4000 kPa) and flux rates (>20 liters per square meter per hour - LMH) were compared and the best performing membrane wasfinallytested in the pilot scale program.

82

magnesium and calcium hydroxides and reactions with carbonate ions also start to occur.

Results and Discussions Table 1 shows that, of the three membranes P-tested in the laboratory, the Dow DK membrane demonstrated superior magnesium rejection performance at relatively low operating pressures and it was selected for use in the pilot scale tests. The TS80 membrane achieved similar magnesium concentrations, however the flux decline was more significant and the applied pressures required were much higher. The TS50 achieved the lowest magnesium rejections. Applied

Recovery

Mg Rejection

Initial Flux

Final Flux

[kPa]

[LMH]

[LMH]

[%]

TS80

2820

35

9.5

60

98

TS50

1700

35.4

22.5

75

92

DK

1425

35.4

18.6

78

97

Type

Pressure

achieved

[%]

(3)

Ca2+ + C032- ->CaC0 3

(4)

Mg2+ + C032" ->MgC0 3

(5)

Once the hydroxide concentration has sufficiently reduced, the hydrotalcite-like compounds precipitate at suitable nucleation sites, typically on the previous precipitates. These compounds are layered double hydroxide compounds with the general formula [M(II)]1.z[M(III)]z(OH)2[A]z.mH20 where M(II) and M(III) are divalent and trivalent metal ions and A is an anion e.g. C0 3 2 \ S0 4 2 \ Cl-, OH\ Typically hydrotalcite, hydrocalumite and paraaluminohydrocalcite can be produced as shown in the reactions below:

Table 1 : Membrane Screening Results Summary I Membrane

Al(OH)4- -> Al(OH)3 + OH-

1

6Mg2+ + 2Al(OH)4- + 80H" + C032" + 4H20 -> Mg6Al2C03(OH)16.4H20 (6)

The sea water neutralization reaction was simulated by titration of 250 ml of a generic Bayer residue water with both untreated sea water, and sea water concentrate (NF retentate). The aim of the titration experiments was to confirm that the increase in magnesium and calcium concentration did not change the reactions from those typically observed during sea water neutralization and the pH was allowed to equilibrate after each addition. Samples were also collected at the start, at the perceived inflection point (-11.5 pH for sea water and -10.3 pH for NF concentrate) and at the end of the titration (~pH 9.4) and details of analytical data are shown in Table 2.

2Ca2+ + Al(OH)4' + 30H" + 3H20 -* Ca2Al(OH)7.3H20 (7) Ca2+ + 2Al(OH)4- + 2C032- + 3H20 -* CaAl2(C03)2(OH)4.3H20 + 40H (8) Determination of the exact stoichiometry is impossible due to the number and complexity of the reactions taking place.

Table 2: Chemical Analysis of Titration Solutions Bicarbonate Carbonate Hydroxide

A|k°..

a

.

[mg Na2C03/l]

Aluminium

Calcium

Magnesium

[mgAI/l]

[mg Ca/I]

[mg Mg/I]

3100

0.5

<0.1

0.34

410

1200

Bayer Residue Water

<1

Sea Water

.

Residue & Sea Water at Inflection

<1

5300

160

5460

620

4

8.4

2800

480

<1

3280

2.4

110

220

0.55

760

6600

6700

4500

<1

11200

40

12

18

9000

720

<1

9720

0.71

240

1900

Residue & Sea Water at end Sea Water NF Concentrate Residue & Concentrate at Inflection Residue & Concentrate at end

8200

46000

54200

.

|

|

Figure 3: Mass Mg Added vs pH for Concentrated Sea Water in Residue Water Inspection of the neutralization curves (Figure 3) shows that there are two areas where a small inflection is observed: around pH 11.5 and pH 10.3. Thefirstinflection point corresponds to the pKsp of magnesium hydroxide of 11.3. This is expected because magnesium is present at three times the concentration of calcium in sea water and in sea water NF concentrate. The second inflection corresponds to the carbonate/bicarbonate equilibrium as the hydrotalcite-type reactions (equations 6, 7 and 8) remove carbonate from solution. Analysis showed that while some hydroxide and aluminate was still present at the upper inflection point they had all been consumed by the lower inflection point. The alkalinity remaining at the lower inflection and the end point was due to the presence of carbonate and bicarbonate ions. These observations were similar, regardless whether sea water or sea water NF concentrate was used for neutralization.

The titration curves (Figure 3) show similar symmetry with two inflection points when the added reagents are normalized to magnesium addition, suggesting that similar reactions are taking place. The causticity in a Bayer residue water can exist as both free hydroxide (OH) and as the aluminate ion (Al(OH)4). Further alkalinity is provided by the carbonate ion (C032~). When this Bayer residue water is reacted with sea water, or sea water concentrate, there are a multitude of reactions that can take place. The initial reactions precipitate magnesium and calcium hydroxides (brucite and portlandite) thereby lowering the pH. 3Mg2+ + 60H- -» Mg3(OH)6 2+

Ca + 20H" -* Ca(OH)2

(1) (2)

Two points of difference between the sea water and sea water NF concentrate neutralization curves are evident: i) at the end point

As the hydroxide concentration reduces, the aluminate ion precipitates, releasing hydroxide ions which then form more

83

where a lower pH was reached using sea water NF concentrate (9.0 cf 9.4 pH), and ii) that the endpoint was achieved much sooner. This is explained by the higher concentration of reactants in the sea water NF concentrate pushing the reaction equilibrium further to the right (equations 1 to 5), thus removing more hydroxide, carbonate and aluminate ions from solution thereby reducing the pH. There is also less dilution if the sea water NF concentrate is applied, which reduces the total amount of unreacted calcium and magnesium in solution and hence improves the magnesium and calcium reaction efficiency.

The suitability of using sea water NF concentrate with Virotec's Basecon Technology was also considered. Untreated sea water and NF concentrate were compared as the feedstocks for neutralization of three different alumina refinery red mud residues in different parts of the world. The sources were: i)

a refinery in Australia, processing bauxite from North Queensland using the Bayer Process; ii) a refinery in North America, processing bauxite from Jamaica using the Bayer Process; and iii) a refinery in China, processing diasporic bauxite from China using the Lime Sinter Process.

Although a faster reaction with concentrate was observed, in good agreement with Virotec's experience, this was not investigated further. Alumina refinery residues have different properties and reaction kinetics need to be specifically evaluated with on-site pilot tests which will determine actual reaction rates and equipment sizing.

The chemical composition of neutralizing streams used in the kinetic model was based on the world average sea water, and the chemical composition of NF concentrate at 70% recovery as analyzed in the first pilot trial. This was the point at which scaling became apparent and therefore the maximum condition for operation. Details of the chemical composition of the two neutralization reagents are shown in Table 3.

The results for the first pilot scale test are shown in Figure 4. In addition to observing the deportment of magnesium in this test, the behavior of calcium was also considered of high importance as the propensity of calcium sulfate to form a scale on the surface of the membrane will lead to rapid flux decline. A highly concentrated layer of calcium and sulfate ions at the membrane surface can form as a function of increasing bulk concentration of each ion in the feed stream and as a function of flux. The formation of this layer is called concentration polarization and was identified in the tests by a sharp decline in flux for a given pressure (Figure 4). The determination of the conditions at which this significant flux reduction occurs is of considerable importance in determining the achievable recovery level.

Table 3: Chemical Analysis of Reagents used for Basecon Technology Computer Model of Sea Water Neutralization Kinetics 1 Chemical Analysis 2+

Sea Water

NF Concentrate |

[mg/1]

412

940

Mg2+ [mg/1

1290

4000

SO/' [mg/1]

910

9400

K+ [mg/1]

400

645

Na+ [mg/1]

10400

13000

Cf [mg/1]

19400

27100

pH

8.3

8.0

Ca

|

1

1

The Basecon Technology neutralization kinetics for each of these three refinery residues was modeled using Virotec's computer model. Table 4 presents the pre- and post-neutralization properties and data of the three alumina refinery residues investigated. Table 4: Basecon Technology Kinetic Model Outputs for Three Different Refinery Residues 1 |

Figure 4: Pilot Test 1 - Analytical Trends

Refinery

North America

Australia

China

Bauxite Source

Jamaica

Nth Qld (Aus)

Central China

1

Process

Bayer

Bayer

Lime Sinter

|

1 1

Properties :

Figure 4 also shows the concentration of each analyte in the feed at each recovery level and the flux at each recovery level. The rejection of chemical species was very high and resembled those achieved in the screening tests recorded in Table 1. At 80% recovery the magnesium rejection was 95% and the sulfur rejection was up to 100% (the limit of detection for these analyses was 100 mg/L). The calcium concentration at 80% recovery was less than that at 70% recovery. This implied that some calcium precipitation occurred at about the 70% recovery level. It can also be seen that the decrease in calcium concentration coincides with a significant increase in the rate of decline in flux and this supports the hypothesis that concentration polarization of calcium sulfate was occurring. As a result recovery levels were reducing because of the resultant rapid rate of flux decline at the applied pressure of 3200 kPa.

1

pH of residue before neutralization

12

12.4

pH of residue after neutralization

9.0-9.5

9.0-9.5

11.6 9.0-9.5

Total alkalinity of residue before neutralization [mg Na 2 C03/L]

19,000

11,370

44,000

<500

<500

<500

5:1

3:1

10:1

2: 1

0.9:1

3: 1

2.5

3.3

3.3

[Total alkalinity of residue after neutralization [mg Na2C03/L] 1 Treatment ratio of Seawater to refinery residue [kL/kL] 1 Treatment ratio of NF Cone, to refinery residue [kL/kL] |

Ratio of Sea Water to NF Concentrate

|

The model supported pilot testing results showing the end point is reached sooner and the volume of neutralizing reagents can be reduced approximately three fold and hence there is significant potential to reduce pumping, piping and size of major equipment if an NF concentrate is applied.

84

The impact of concentration polarization can potentially be reduced by lowering the operating pressure and therefore flux rate, leading to greater recovery levels. This hypothesis was assessed in the second pilot scale test.

reduced to 7.8% at 80% recovery. This is another indication of calcium precipitation on the membrane surface and in the feed. At 85% recovery it was observed that precipitation was occurring in the feed vessel as well.

The applied pressure was varied as required to maintain the flux at a lower constant target flux rate of 35 LMH. The test incorporated a number of P-tests at various recovery points. The operational data and the species transmissions are summarized in Figure 5, and the relationship between flux, pressure and recovery is shown in Figure 6. By comparing with Figure 5 it can be seen that controlling flux and minimizing concentration polarization improves permeate recovery, and therefore calcium and magnesium concentration. It is apparent in Figure 6 that the pressure / flux curve at 75% recovery is almost identical to those at lower recoveries, and the onset of scaling has been delayed by reducing flux.

For these reasons 75% recovery was chosen as the practical concentration limit for a sea water NF plant. Pilot Test No 2 (P-Tests) : Flux vs Applied Presure at Different Recovery Rates

Pilot Test No 2 : Concentrations of Ca, Mg & S in Retenate vs Recovery 1600

2000

2400

Applied Pressure [kPa]

Figure 6: The Relationship between Pressure and Flux at a Range of Recoveries Anti-scalents are commercially available to reduce the rate of nucleation and these products achieve an induction period during which scaling can be avoided. However, they do not indefinitely prevent the precipitation of minerals such as calcium sulfate under supersaturation conditions and the application of anti-scalents was therefore not considered in this investigation. Another advantage of sea water NF is the fact that the NF permeate is largely free of the scaling species and it is therefore an ideal feed to a sea water RO (SWRO) plant producing potable water.

Figure 5: Pilot Test 2 Analytical Trends Predicting the point at which calcium sulfate nucleation occurs is very important in managing a sea water NF process. Precipitation of calcium sulfate is suppressed by the total ionic strength of sea water (in excess of 45 g/L TDS). Residence time in a membrane element is also short (<0.2 sec). As a result, the conditions under which precipitation occurs cannot be readily determined from standard thermodynamic data. The point at which calcium sulfate precipitation occurs on the membrane surface was therefore determined empirically based on the observation of flux versus pressure relationships.

Simulations were performed using Dow's Reverse Osmosis System Analysis (ROSA) simulation software to compare untreated sea water feed with the NF permeate feed. The results are summarized below in Table 5.

Table 5: Comparison of SWRO and Sea Water NF and RO Sea Water RO

Figure 6 shows a linear relationship between flux and pressure at low flux rates. A breakpoint is observed at a flux rate of about 45 LMH, the critical flux. This is the point where the membrane was operated in pilot test 1, explaining why the recovery results were poor. Poor membrane performance in pilot test 2 also occurred around 75% recovery, evidenced by the sideways shift in the 80% and 85% flux / pressure curves as shown in Figure 6

Feed Row Basis NF Permeate Recovery RO Plant Feed RO Permeate Recovery

For each recovery level up to 75% the relationship between applied pressure and flux was very similar and the applied pressure to achieve these recoveries was approximately 2200 kPa. Increasing the pressure beyond 2200 kPa caused a significant rate of flux decline suggesting that resistance at the membrane surface increased significantly with the increase in applied pressure causing a concentration polarization effect. The maximum achievable recovery, while maintaining a flux of 35 LMH, appears to be 75%. Beyond this point, the pressures required to achieve an equivalent flux increased significantly. The transmission of calcium at 75% recovery was 8.3%, however this

Potable Water Production

Sea Water NF &| RO

1000 m3/hr

1000m3/hr

|

N/A

75%

|

1000 m3/hr

750 m3/hr

|

48%

64%

1

3

3

|

480 m /hr

480 m /hr

The precipitation of calcium sulphate is performance limiting in both traditional SWRO and multi-effect distillation (MED) plants. When using sea water NF permeate as a feed to the RO unit, the limiting factor is instead osmotic pressure and the results of this comparison indicate that permeate recovery will increase from 48% to 64%. Hence, a significant reduction in the RO unit capacity can be achieved for equivalent potable water production,

85

reducing capital and operating costs and significantly reducing waste retentate discharge from the process.

and chloride concentrations are maintained close to the original concentrations of sea water.

Meeting fresh and potable water requirements will continue to be a global challenge for alumina refineries. For those refineries in proximity to the sea and with environmental approval to access sea water, the opportunity to combine sea water NF and RO for potable water and neutralization of refinery red mud solid residue using NF concentrate (retentate) offers great benefits.

This NF concentrate improves the reactivity and kinetics of sea water neutralization of alumina refinery residue and the required volumes of NF concentrate are about three times lower than for neutralization with untreated sea water. The size of neutralization equipment like pumps, piping, neutralization tanks, settlers and décantation and drying ponds are significantly reduced, leading to reduced capital costs and footprint for new operations or enabling expansion of capacity for existing alumina refineries using sea water neutralization.

The capital and operating costs for a sea water NF & RO plant will be higher than for an SWRO plant due to the extra membrane installations and pumping costs respectively. However the extra costs for the sea water NF plant need to be compared to the significant cost reductions from the smaller neutralization equipment required, where the volumes of NF concentrate added are about three times lower than for traditional sea water neutralization (refer Table 4). Pumps, piping, neutralization tanks and settler size are significantly reduced for an equivalent amount of neutralized red mud and these changes can be seen in the process schematic below in Figure 7

The low salinity of neutralized residue improves options for reuse in agriculture and soil beneficiation The nanofiltration permeate is also a valuable and improved feed to sea water reverse osmosis (SWRO) and multi-effect distillation (MED) plants. With most of the scaling species removed by the NF membrane, RO process efficiency is much improved, leading to a significant reduction in the RO unit capacity for equivalent potable water production, reducing capital and operating costs and significantly reducing waste retentate discharge from the process. Alumina refinery residues have different properties and reaction kinetics need to be specifically evaluated with on-site pilot tests which will determine actual reaction rates and equipment sizing. References

► Refinery

Le

Eend:

[

Refiner^— '

I Scope Additions I Scope Reductions

A. M. Hassan et al, "Conversion and Operation of the Commercial Umm Lujj SWRO Plant from a Single SWRO Desalination Process to the New Dual NFSWRO Desalination Process", IDA Conference, Manama Bahrain, March 2002.

[2]

A. M. Hassan et al, "A Demonstration Plant Based on the New NF-SWRO Process", Desalination 131, 2000, 157-171.

[3]

L. Fergusson, "The conversion and sustainable use of alumina refinery residues: Global solution examples", Light Metals, TMS, Orlando, 2007, 105-111.

[4]

G. Graham, and R. Fawkes, "Red mud disposal management at QAL". International Bauxite Tailings Workshop, Perth, Australia, 1992, 188-195.

[5]

D. McConchie, P. Saenger, and R. Fawkes, "An environmental assessment of the use of sea water to neutralise bauxite refinery wastes". The Minerals, Metals & Materials Society, 2nd International Symposium on Extraction and Processing for the Treatment and Minimisation of Wastes, Scottsdale, Arizona, October 1996,407-416.

[6]

C. Madden, "Mud, glorious mud?", Science Network Western Australia, 6 October 2006.

[7]

L. Fergusson, "Virotec: A ten-year story of success in environmental remediation". Prana World Publishing, ISBN 978-0-646-53324-7, 2010, 182.

[8]

D. McConchie, M. Clark, and F. Davies-McConchie, "New strategies for the management of bauxite refinery

■ | No Scope Change

Figure 7: Schematic Showing Sea Water NF and RO Process Impact on Bayer Residue Treatment The cost comparisons are dependent on the particular situation but order of magnitude costing estimates have shown that using a sea water NF concentrate to neutralize alumina refinery residue has significant economic potential in situations where: •

[1] Residue Storage Area

An existing refinery using sea water neutralization is looking to expand without spending capital on neutralization equipment and especially if there is limited land area available for décantation and drying ponds.

•

The refinery is not adjacent to the sea but still wants to neutralize its residue with a sea water based reagent.

•

The refinery wants to reduce salinity of neutralized residue to improve options for re-use, and

•

There is a local sea water desalination plant (e.g. SWRO, MED), where the combined benefits of nanofiltration to reduce scaling of desalination equipment and provide a sea water NF concentrate for refinery residue neutralization can be achieved Conclusions

Nanofiltration of sea water has been shown to be a viable method to significantly concentrate calcium and magnesium while sodium

86

resides (red mud)", 6th International Alumina Quality Workshop, Brisbane, Australia, September, 2003. [9]

D. McConchie, F. Davies-McConchie, M. Clark, and L. Fergusson, "The use of industrial mineral additives to enhance the performance of Bauxsol-based environmental remediation reagents", Australian Industrial Minerals Conference, Brisbane, 21 March, 2003.

[10]

D. McConchie, M. Clark, F. Davies-McConchie, L. Fergusson, T. Prowse, and S. Pope, "The use of industrial mineral additives to enhance the performance of Bauxsol-based environmental remediation reagents", Australian Institute of Geoscientists Bulletin, No. 38, 2003, 27.

[11]

M. W. Clark, J. Berry, and D. McConchie, "The longterm stability of a metal-laden Bauxsol reagent under different geochemical conditions", Geochemistry: Exploration, Environment, Analysis, 9, 2009, 101-112.

[12]

C. Hanahan, D. McConchie, J. Pohl, H. Creelhman, M. Clark, and C. Stocksiek, "Chemistry of sea water neutralisation of bauxite refinery residues (red mud)", Environmental Engineering Science, 21(2), 2004, 125138.

87

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CAUSTIC AND ALUMINA RECOVERY FROM BAYER RESIDUES Gu Songqing Science & Technology Division of Chalco, No.62 Xizhimenbei Street, Beijing China Keywords: Bayer residue, Sintering process, Floatation-Bayer process, Lime-Bayer process, Hydro-process, DSP, Hydrogamet Abstract

Even though China is rich in bauxite resources, only about 20 % of bauxite reserves are of sufficient quality for Bayer processing as shown in Table 1, while bauxites of A/S 4-6 make up about 60% of total reserves.

The Bayer process is not suitable for alumina production from high silica bauxites due to the high caustic and alumina loss into the Bayer residue. In this study, caustic and alumina recovery from the Bayer residue by various processes, such as sintering, lime treatment and hydro-processing etc, is analyzed and compared theoretically. The most important target to treat Bayer residue for the recovery of these products is to find an efficient Desilication Product (DSP) containing less alumina and caustic soda by some suitable and efficient processes.

Most of Chinese bauxite is diasporic with both high alumina & silica content and has very complex silicate mineral composition. However the bauxite reserves per capita are only 10 % of the world average. Furthermore the bauxite mines are usually small in size, and there are lots of underground bauxite reserves in China, which brings about bauxite mining difficulties and higher mining cost. See Table 2.

Introduction

Based on the situation mentioned above, the Chinese alumina industry has to face serious raw materials issues and is forced to put great effort into R&D of new efficient processes to deal with lower grade bauxites.

The biggest challenge for the Chinese alumina industry is the bauxite supply shortage and the reduction of the grade of the available resource, including the ratio of alumina to silica (A/S) and the alumina content in these bauxites. The bauxite A/S supplied to some Chinese refineries has continuously been reducing from 12 to less than 6 in the past 8 years and even to 5.3 last year, which brings a major negative impact on alumina production cost. The bauxite grade change trend is shown in Fig. 1.

Too low a grade of bauxite means higher silica and reduced alumina content, which greatly impacts the cycling efficiency and productivity. The DSP produced in the Bayer process will carry more caustic and alumina to the residues increasing red mud disposal. Consequently, caustic consumption and alumina loss will increase and the operational cost grows quickly. Comparison of various processes to treat low grade bauxites There are many studies looking for new, high efficient processes for the low grade bauxite. The invention of the Rotation - Bayer process in the last decade, tries to pretreat low grade bauxite for silica minerals removal before introduction to the Bayer process. It was a success and applied in some Chinese refineries. But the alumina recovery is too low due to two stage losses in both the flotation and Bayer processes, and the organic substances entering the Bayer process might lead to some behavior changes of desilication and precipitation.

2001 2002 2003 2004 2005 2006 2007 2008 2009

Years Fig. 1 Bauxite grade reduction for some Chinese refineries Table 1 The Grade Distribution of Chinese Bauxites low grade

middle grade

A/S

3-4

4-6

6-7

7-9

9-10

>10 |

%

~8

-49

-11

-15

-12

-5 |

Bauxite grade

high grade

The Lime-Bayer process was developed to reduce caustic consumption by adding more lime to substitute some caustic from the residues. The Lime-Bayer process is simple and easy to apply.

Table 2 The Characteristics of Chinese Bauxite Reserves

Bauxite

1 World China

Reserve

Reserve Base

Reserve/ Population

A/S

23 B t

35 B t

4t

>10

0.54 B t

0.72 B t

0.41

4-7

* B t—billion tons

89

Mine Size

Bauxite Types

Large >0.1Bt Small <0.03Bt

Gibbsite Boehmite Diaspore

Silica Minerals Types Simple Kaolinite, Quartz Complex Kaolinite, Wite etc.

1 1 |

But too low an alumina recovery and high lime additions makes the production cost increase and the amount of residue increase, as well due to its lower desilication efficiency.

DSP analysis for the existing processes to treat low grade bauxite

The "Bayer-Sintering in Series" process is being tested in some Chinese refineries. The core concept for this process is to recover caustic and alumina from Bayer residue by sintering to form calcium silicates as the DSP. It seems that only a small part of the energy can be recovered and the energy consumption is still very high in the "Bayer-Sintering in Series" process compared with the original "Bayer-Sintering Combined" process.

To improve the desilication efficiency in caustic alumina production, the most important factor is to get suitable DSP with low A/S, N/S and C/S, that is, the less A1203 and Na20 content in the residues and the less lime addition, the more efficient for the process.

Fig 2 and Table 3 show the major DSP in the residues for existing processes.

This is the goal of looking for a targeted DSP. The DSP from the sintering process contains no AI2O3 and Na20 so that it is high efficient for silica removal and is energy intensive as well. The DSPfromthe Bayer process has the biggest N/S so it could not be used for high silica bauxites. The DSP from lime-Bayer process contains much A1203 and CaO, but no Na20 at all, so the process can only be used in the refineries with access to cheap lime and bauxite.

The high pressure hydro-process invented by former Russia scientists is a complete hydrometallurgical process without any pyro-process, and offers high alumina recovery and less caustic loss. The major concept is to recover caustic and alumina from Bayer residue by hydro-processing to form calcium sodium silicates as DSP. But this process is too complex and the evaporation is too large for production of a high concentration liquor and sodium aluminate precipitation, and again this process features high energy consumption.

Key technical concepts to develop high efficient alumina production

According to the process review above there is no highly efficient and energy saving process at present to deal with low grade bauxites. A comprehensive summary of the advantages and disadvantages of all the processes and theoretical investigation of the process reactions and products will provide the opportunity to develop new and efficient processes. Caustic Processes

For more economical recovery of caustic and alumina from Bayer residues, it is essential to find an efficient DSP and to apply energy saving processes. The alumina production efficiency can be improved by higher desilication efficiency.

Desilication Elements Ca and Na

DSP 2CaO· SiOv xH 2 0

Sintering

Na 2 0 A1203- 2Si02- 2H 2 0

Bayer Lime-Bayer

3CaOAl203Si02-4H20

High Pressure Hydro

Others

Fig 2 DSP analysis for the residue from different processes Table 3 The chemical compositions of the various DSP |

Processes

DSP in residues

A/S

N/S

c/s

|

Sintering

2CaOSi0 2 xH 2 0

0

0

1.87

|

Bayer

N a 2 0 AI2O3 -2Si02xH 2 0

0.85

0.52

0

|

Lime-Bayer

3CaO · AI2O3· nSiP 2 ·χΗ2Ο(η:0.2-1)

8-1.7

0

14-3

|

A/S—ratio of A1203 to Si0 2 content; N/S—ratio of Na20 to Si0 2 content; C/S-- ratio of CaO to Si0 2 content

90

The only way to greatly reduce energy consumption in the sintering process is by recovering as much energy from sintering as possible, and to enhance the process efficiency so that the total cost of sintering process can be competitive with the Bayer process.

The new highly efficient process to be found must be producing the DSP with relatively lower ratios of A/S, N/S and C/S so as to greatly reduce consumptions of bauxite, caustic soda and lime addition. Another criteria for any suitable process is that it should be as, energy efficient as possible, by adopting hydrometallurgical processing instead of pyrometallurgical processing to save energy.

Conclusions (1) The bauxite A/S for some Chinese refineries has continuously been reducing, which greatly impacts on the technical and economical performance and production cost. (2) The advantages and disadvantages of various existing processes to treat the low grade bauxite are analyzed according to their desilication efficiency and energy consumption. (3) A theoretical study has been carried out by comparison of the chemical compositions of the DSP from the different existing processes. (4) The key technical concepts to develop new and efficient processes for low grade bauxite are summarized. The most important is to produce a DSP with low A/S, N/S and C/S in a hydro-process. (5) It is possible to develop a process to form hydrogarnet structured compounds, which may indeed be the most efficient DSP.

Calcium compounds are much cheaper than other possible elements to replace Na20. But the multi elements compounds containing CaO are very complex and form at quite variety of reaction conditions. So the key solution to get the most efficient DSP will be based on the tests results at various conditions to find suitable compounds. There are numerous compounds produced from A1203, Si02, Fe203 and Ti02 (existing in bauxite) and Na20, CaO and MgO (added into the processes) etc, which can be found from X-ray diffraction tables and the phase diagrams. The most possible DSP in the processes to treat Bayer residues can be seen in Table 4. In this Table, the desilication efficiency increases from top to bottom. Table 4 Possible DSP in Bayer residue treatment

|

DSP in residues

A/S

N/S

c/s

Na 2 02Ca02SiO r xH 2 0 CaO - 2A1203 · 2Si0 2 H 2 0 3CaO · A1203· 2Si02 ·2Η20 3CaO Fe203· 2Si02 ·2Η20 3CaO · 2Si02 · mH20 CaOSiQ2xH20

0 1.7 0.85 0 0 0

0.52

0.94 1

o

0.47 1.4 1.4 | 1.4 0.94

0 0 0 0

| References [1] Gu Songqing, Wu Lichun, Liu Fengqin et al., Progresses of Nonferrous Metals in China (in Chinese), (Changsha, China, Central South University Publishing House. 2007) [2] Bi Shiwen, Yu Haiyan, Alumina Production Processes (in Chinese), (Beijing China, Chemical Industry Press, 2006) [3] Gu Songqing." Chinese bauxite and its influences on alumina production in China". Light Metals 2008, 79-83 [4] Claudia Brunori et al., "Reuse of a treated red mud bauxite waste", Journal of Hazardous Materials, B117, (2005), 55-63 [5] Ma Shuhua, et al., "Recovery of soda and alumina from red mud", Multipurpose Utilization of Mineral Resources (in Chinese), No.l (2008), 27-30

It is found from Table 4 that the DSP from CaO and Si0 2 is the most efficient because there are no A1203 and Na20 losses. Also, the higher the Si0 2 coefficient in the DSP, the higher the desilication efficiency. Laboratory tests show that the DSP from both CaO and Si0 2 usually form at very high temperatures, or in the pyrometallurgical processes. Hydrogarnet is a series of compounds containing CaO, Si0 2 and the oxides of other elements, such as A1203, MgO, Fe203 and Ti0 2 etc. Usually CaO in the hydrogarnet like structured DSP can be replaced by Na20, MgO and K20, while A1203 in the DSP can be replaced by Fe203 and Ti02. The minerals containing Fe203 and Ti0 2 are commonly found in bauxites and will take part in the reactions during the residue treatment so the study on the replacement of Fe 2 0 3 and Ti0 2 to A1203 should be carefully monitored to reduce A1203 content in the final hydrogarnet like structured products. It seems that the hydrometallurgical process is the first choice for getting an efficient DSP since the pyrometallurgical processes such as the sintering process will consume too much energy without much better waste heat recovery.

91

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INVESTIGATION OF ALUMINA DISCHARGE INTO THE RED MUD POND AT NALCO'S ALUMINA REFINERY, DAMANJODI, ORISSA, INDIA B. K. Mohapatra1, B.K.Mishra2 and C.R.Mishra3 institute of Minerals and Materials Technology (CSIR), Bhubaneswar-751013, Orissa, India institute of Minerals & Materials Technology (CSIR), Bhubaneswar-751013, Orissa, India 3 National Aluminium Company Ltd., Bhubaneswar-751013, Orissa, India Keywords: Bauxite, Alumina Discharge, Red Mud electron microscopy and electron probe micro-analysis with a view to finding out the reason for alumina discharge into the red mud pond during the alumina refinery process.

Abstract Around 14% of alumina gets discharged into the red mud pond from NALCO's alumina refinery. The minerals hosting alumina in red mud have been investigated using Optical microscope, Scanning Electron Microscope (SEM), Electron Probe Micro Analyzer (EPMA) and XRD. The minerals in red mud sample, as identified through integrated techniques, are: gibbsite, boehmite, goethite, silimanite, muscovite, garnet and kaolinite besides minor hematite, sodalite and rutile. The average particle size of red mud is ~8μιη except isolated grains of gibbsite, goethite and garnet which are greater than 20μπι in size. The gibbsite invariably contains up to 60 mole % of boehmite. The nodular goethite contains over 30-mol % of ΆΓ in its lattice and termed as alumo-goethite. Around 34% of the A1203 is present in muscovite/silimanite/garnet. Studies reveal that most of the minerals in red mud have either alumina in their lattices or are undigested aluminium silicates that do not get dissolved during alumina refining, thus a considerable amount of alumina gets discharged through these phases into the red mud.

Materials and Methods For this study, representative red mud sludge was collected from the red mud pond of NALCO alumina refinery. After washing and oven drying, the sample was sprinkled on double sided adhesive conducting tape, placed over the sample holder and studied under SEM (JSM 35CF) equipped with WDS & EDS systems. Compositional x-ray analysis was undertaken for selected minerals. Chemical compositions of various phases were determined by an ARL-SEMQ-II electron microprobe in the Geochemisches Institute, Universität Gottingen, Germany, which is equipped with six spectrometer and four different analyzing crystals. The operating conditions were 15 kV accelerating voltage and 15 nA sample current. ZAF - corrected representative mineral analyses are presented in Table -1.

Introduction Morphology and Mineralogy

Around 2 tons of caustic insoluble waste residue known as 'Red Mud' is generated from NALCO aluminium refinery at Damanjodi, Koraput district, Orissa for every ton of alumina produced. Aluminium metal is extracted from aluminium oxide phases that constitute only 38 to 60% of the original bauxite ore. The rest is made up of Fe203, Si02, Ti02 and many other oxide phases. After dissolution of the bauxite in caustic soda, these impurities remain in suspended form which is separated out after being washed and then pumped in a slurry to the nearby pond called the Red Mud Pond. India accounts for about 2 Million tons of red mud per year of which is more than half the quantity is generated at the NALCO refinery. This industrial waste material poses tremendous environmental and disposal problems. Reduction in the quantity of red mud is possible only through their utilization in one form or another. However, the inherent complexity of red mud poses problem in its bulk utilization. Prior to going for bulk utilization of red mud rejects from the NALCO refinery, it needs in-depth characterization. Rao et al [1] have reported the characteristics of undigested sand rejects from the NALCO refinery. Only limited attempts have been made on the characterization of red mud sludge from this refinery [2-5].

In the red mud sample, about 35% by weight of solids contain less than 5-micron particles and the average size is found to be around 8 micron. However, a few independent grains > 20μπι size (about 10%) are also present. Due to the very fine size, some of the particles appear coagulated. Broad mineral species present in the red mud sample were identified through XRD (Philips). As can be seen from XRD pattern illustrated in Fig.l, the red mud sample contains gibbsite, boehmite and sodalite as alumina bearing; goethite and hematite as iron bearing and rutile as titanium bearing phases. Goethite, hematite, rutile, kaolinite and mica were recognized under optical microscope. Some of these alumina bearing minerals were subsequently analyzed under electron micro-probe. Gibbsite and goethite are observed to be dominating phases in the red mud sample. Gibbsites occur in pseudo hexagonal or prismatic habit. Goethite is either flaky or nodular; the later showing bunches of rotund units. Presence of fine grains of rutile, ilmenite, and zircon are recorded in subordinate amount. In addition, flakes of kaolinite, mica, needles of sillimanite (as litho relics) and sub-rounded grains of garnet are also noticed.

Fine particles of red mud can be characterized only through selected instrumental techniques. The present paper describes the characteristics of NALCO red mud through scanning

93

H 6

H

Table 2. Electron Probe Microanalysis results of Kaolinite and Goethite 2 4 1 Compounds, % 3 43.834 45.859 0.795 0.70 Si0 2

Gb

Cû

GD

Os"

J V

fel v** ^χ,ν

A rv Γ

A1203 FeO MgO MnO Ti02 CaO

UL τ τ τ

ΓΤΓΐτιτιτΓη^ττ ΐΓτΡτ ι ττ ΐΊ y

CuK α

<- 2Θ

Fig.l: XRD pattern of Red-mud. [ H: Hematite, Go: Goethite, S: Sodalite, Gb: Gibbsite, R: Rutile, B: Boehmite]

κ2ο

Na 2 0 Cr 2 0 3 BaO NiO CaO

Mineral Chemistry A few selected mineral grains viz. gibbsite, goethite and kaolinite, muscovite (precursor minerals not completely replaced by alumina) from red mud were exposed under electron probe microanalyser (EPMA) to know the extent of alumina in them (Table 1-3). Most of the idiomorphic crystals appearing as gibbsite under scanning electron microscope are found to have undergone dehydroxylation containing up to 60 mole % of boehmite in solid solution (Table 1).

H20Calc

SI Totali Al Fé

Mg Mn Co Total 2

The nodular type of goethite contains alumina in its lattice (Fig. 2). Such alumina-rich goethite is termed as alumo-goethite by Jonas and Solymar [6] Since goethite grains in the NALCO red mud are found to contain 30 mol % of alumina, it may be termed as alumo-goethite. 30 to 39% of A1203 are recorded in litho relict minerals like kaolinite and muscovite flakes (Table 2 & 3). Compositional maps of kaolinite, silimanite and garnet grains in red mud (Figs. 3-5) indicate that these minerals are not converted to gibbsite.

Na K Ca Ti | Total 3

A1 2 0 3 FeO Si0 2 MgO MnO CaO

κ2ο Na 2 0 Ti0 2

AlOOH Al(OH)3 FeOOH Other Oxides

2

1

3

4

78.30 0.26 0.04 0.00 0.00 0.49 0.04 0.00 0.00

71.82 0.00 0.00 0.00 0.00 0.22 0.00 0.00 0.00

69.94 0.00

65.60 0.00

0.36 0.07 0.00 0.02 0.00 0.00 0.00

0.10 0.00 0.00 0.00 0.00 0.04 0.01

69.44 29.63 0.32 0.93

49.90 49.87 0.00 0.23

35.65 73.69 0.00 0.66

2.35 97.31 0.00 0.34

-

—

38.009 0.049 0.058 0.018

0.054 0.025 0.203 0.076 0.000 0.009 0.236

0.134 0.033 0.064 0.127 0.017 0.000 0.211

--

~

--

--

Based on 08 cation 3.95 4.05 3.95 4.05

0.957 85.012

---

0.013

0.015 0.974

0.30 0.62

--

21.44 12.20 Based on 1 cation 0.01 0.01

0.02 4.02

0.01 3.93

0.03

--

0.01 0.01

-

---

0.03

0.02

~

--

0.01

~

6.12 0.42

~ ----

3.90

--

~

---

3.98 0.01 0.01

--

18.48 52.63

-

-~

0.01 0.06

~

Note: Column 1 & 2: Kaolinite; 3: Goethite; 4: Alumo-goethite

Table 1. Electron Probe Microanalysis Results of Selected Alumina-rich Grains from Red Mud Compound, %

39.228 0.145 0.076 0.000

1

Reasons for Discharge of Alumina Phases into Red Mud After the soluble aluminium containing oxides and hydroxides are recovered in the Bayer process, other metal oxides present in the bauxite are disposed as a red mud material. However, discharge of alumina rich phases into the red mud pond, resulting loss in alumina recovery, is of great concern. Discharge of alumina to an appreciable extent (Av. A1203: 14%) into the red mud pond may be attributed to gibbsite containing boehmite in solid solution, alumo-goethite having 30 mol% of alumina in its lattice, needles of silimanite, specks of mica and garnet (litho-relict minerals) and minute kaolinite platelets recorded in the red mud. All such mineral phases do not get dissolved during the industrial treatment of bauxite with caustic soda and ultimately get into the red mud pond. However, an in-depth study is necessary to improve the efficiency of digestion of all these phases during the Bayer process, so as to recover the alumina value from at least some of them. Volkov et al [7] reported that by increasing the amount of lime at 240° C, it is possible to completely break down aluminogoethite.

Note: Column 1: Boehmite dominating; 2: Mixed boehmitegibbsite; 3: Mixed gibbsite-boehmite; 4: Gibbsite dominating

94

Fig.4. Compositional Map of Si & Al in Silimanite Grain Present in Red Mud. Presence of Silica and Alumina indicates its poor leaching during Bauxitisation Process

Fig. 2. Compositional X-Ray Map of Fe & Al in Alumogoethite found in Red Mud. Image Map of Iron, Alumina confirms the Presence of significant Alumina in the Goethite Lattice

Table 3. Electron Probe Micro-analysis Results of Muscovite Si0 2 A1203 FeO MgO MnO CaO

κ2ο Na 2 0 Ti0 2 BaO

Si Al Totall Al Fe Mg Mn Ti Cr Total 2 Na K Total 3

Fig. 3. Compositional X-Ray Map of K, Al & Si in a Litho Relict Observed in Red Mud. The Silica Rich Phase Still Occupies the Core Region and hence escaped leaching

95

02 03 01 44.529 46.526 46.566 34.074 34.566 33.120 1.651 1.725 1.711 0.482 0.498 0.455 0.000 0.102 0.045 0.012 0.000 0.015 10.734 11.697 10.270 0.441 0.400 0.455 0.214 0.350 0.355 0.000 0.00 0.00 Based on 14 cations 6.22 1.78 8.00 3.60 0.19 0.10

-

0.02

-

3.91 0.10 1.99 2.09

6.28 1.72 8.00 3.78 0.19 0.09 0.01 0.03

6.20 1.80 8.00 3.63 0.20 0.10 0.01 0.04

4.10 0.12 1.78 1.90

3.98 0.12 1.90 2.02

~

~

04 44.407 30.493 1.177 0.375 0.00 0.00 10.304 0.426 0.248 0.076 6.47 1.53 8.00 3.72 0.14 0.08

»

0.03

-

3.97 0.12 1.91 2.03

The mineral chemistry of some of the constituent phases in red mud reveals the presence of lattice bound alumina which escapes the digestion stage in the Bayer's process, resulting in loss of alumina during the refining process at NALCO.

References 1.

M.B.S. Rao, B. K. Mohapatra, B. Das and A. K. Paul, "Characteristics of sand rejects- A case study from an Alumina Refinery Plant," Vistas in Geological Research, UU Spi. Publication in Geology, 2 (1997), 164-172.

2.

. S. Thakur and S. N. Das, "Red Mud Analysis and Utilization," Pubi. & Inform. Directorate & Wiley Eastern Ltd., New Delhi, (1994), 291p.

3.

J. Pradhan, S. N. Das, J. Das, S.B. Rao and R. S. Thakur, "characterization of Indian red mud and recovery of their metal values," Light Metals: Proceedings of Sessions, TMS Annual Meeting (Warrendale, Pennsylvania), Minerals, Metals & Materials Soc, USA., (1996), 87-92.

4.

M. G. Sujana, R. S. Thakur, B. C. Acharya, S. N. Das and S. B. Rao, "Effect of calcination and physico-chemical properties," Light Metals: Proceedings of Sessions, TMS Annual Meeting (Warrendale, Pennsylvania), Minerals, Metals & Materials Soc (TMS), USA, (1986), 93-98.

5.

B.K.Mohapatra, M.B.S.Rao, R.Bhima Rao and A.K.Paul. Characteristics of red mud generated at NALCO refinery, Damanjodi, India. Light Metal.(Warrensdale, Pa), (2000) 161-165

6.

K. Jonas and K. Solymar, "Preparation, X-ray, derivitographic and infrared study of aluminium substituted goethites," Ada Chim. Acad. Sci. Hung., Tomus. v.66 (4), (1970), 383-383.

7.

V. V. Volkov, M.I. Soboi, N. I. Eremin, "Lime addition effect upon Alumina extraction from Alumino-goethite," kv. V.u.z. Tsvetn. Metal, 5 (1984), 22-24.

Fig. 5. Compositional Map of Si & Al in Relict Garnet Present in the in Red Mud Conclusions The red mud sludge from the NALCO alumina refinery at Damanjodi, Orissa, has been characterized using SEM-WDS and EPMA techniques. These are very fine grained muds, the average size being 8 micron. Different alumina containing phases such as alumo-goethite, sillimanite, kaolinite, mica and garnet were identified from their micro-morphology and in-situ chemical analysis, which contribute to the loss of alumina into the red mud. The presence of dehydroxylated grains of gibbsite probably indicates the higher temperature (than that for pure gibbsite) required for their dissolution.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PRODUCTION OF ORDINARY PORTLAND CEMENT (OPC) FROM NALCO RED MUD Chitta Ranjan Mishra1, Devendra Yadav2, P.S.Sharma2& M.M.Alli2 former Deputy General Manager (R&D), National Aluminium Company Limited (NALCO), NALCO Bhawan, P-1, Nayapalli, Bhubaneswar-751013, Orissa, India Scientist, National Council for Cement & Building Materials (NCB), Ballavgarh, 34-KM Stone, Delhi-Mathura Road, PIN-121004, Haryana, India Keywords: Red Mud, Limestone, Shale, Coal Ash, Clinker, Gypsum, OPC Abstract A process for production of Ordinary Portland Cement (OPC) from NALCO Red Mud has been successfully developed from a raw mix containing limestone, red mud, shale and fine coal. The raw materials are ground to the required fineness and then blended to prepare the raw mix. The raw mix is fed in to a kiln and fired to a temperature of 1400-1450 degree centigrade to obtain clinker. Clinker was cooled and gypsum was added in to it to obtain OPC. 3-3.5% of NALCO red mud was used for production of OPC.OPC prepared from this clinker conformed to the requirements of three Indian Standard Specifications for 33, 43 & 53 grade of OPC. The process is efficient, cost-economic and effectively addresses the environmental problems associated with the waste red mud generated during the refining of bauxite for alumina manufacture.

Sample Collection & Sampling Five shift samples of red mud, bauxite and fly ash were collected. Red mud samples were designated as RMD-1 to RMD-5. Red Mud: Five shift samples of red mud weighing about 25kgs each were dried in electric oven at temperature of 105±5°C. The dried lumps in the samples were crushed. The samples were reduced to ~3kg by coning and quartering procedure and ground to pass through 100 mesh sieve. The laboratory samples were prepared by drawing ~200gms from each shift samples by coning and quartering procedure and were subjected to chemical and mineralogical characterization. The results of chemical and mineralogical characteristics of these laboratory samples are discussed in the following section.

Introduction National Aluminium Company Limited (NALCO), a Navaratna Company and a Govt. of India Enterprise under Ministry of Mines has set up Asia's largest integrated alumina and aluminium complex in the state of Orissa, utilizing state of the art technology from Aluminium Pechiney, France and is producing 4,8,00,000 tpa of bauxite, 1,5,75,000 tpa of alumina, 3, 45,000 tpa of aluminium metal and 960MW of captive power. During the refining of bauxite for production of alumina employing the Bayer process at NALCO's alumina refinery, at Damanjodi in the district of Koraput, Orissa, on a dry basis ,about 2 million tones /annum of red mud is generated as a by-product. The red mud contains more than 55% water and is disposed off in the form of slurry in to the nearby red mud pond. The accumulated quantity of red mud in the ponds is estimated to be around 18.5 million tones. The water overflowing the pond after the mud is settled, is recycled to the alumina plant for partial recovery of soda. Therefore, the soda content in red mud is minimized. It was thought prudent to explore the possibilities for utilization of this uncausticised red mud dried in air/sunlight for production of value added items. Development of Ordinary Portland Cement (OPC) was one of the most positive efforts undertaken in this direction [1]. NALCO in collaboration with National Council for Cement & Building Materials (NCB), Ballavgarh, and Haryana, India has developed this process.

Limestone and Other Raw Materials: Limestone and other raw materials viz. shale, coal and gypsum samples were ground to pass through 100 mesh sieve. The laboratory samples of these raw materials were prepared by drawing ~200gms from each of the ground by coning and quartering procedure. The results of chemical and mineralogical characterization of these samples are discussed in the following section. Characterization and Evaluation of Samples Red Mud Samples: (a) Quality and quality variation: Five ground samples of red mud (RMD-1 to RMD-5) were subjected to chemical analysis for their major constituents' viz. LOI, Si02, Fe203, AI2O3 and alkalies for assessing their quality and the variation therein. The results are presented in Table-2 indicated the variation in the composition of red mud shift samples with respect to major oxides to be low. In view of this, all the five shift samples of red mud were mixed together to make a composite sample. The composite sample was designated as RMD-C. Further investigations were carried out on the composite sample of red mud, RMD-C only. Table 2. Quality and Quality Variation in Shift Samples of Red Mud.

Table 1. Chemical Composition of NALCO Red Mud. Constituents % By Weight 16.13 A1203 53.92 Fe 2 0 3 Ti0 2 04.82 06.29 Si0 2 Na 2 0 03.45 CaO 01.88 P205 00.11 00.11 v2o5 LOI 11.78

97

SI No

Constituent Determined

1 2 3 4 5 6

LOI Si0 2 FeA A1203 Na20

(%)

κ2ο

Shift I RMD1 11.32 6.32 50.10 20.51 4.07 0.07

Shift II RMD2 11.86 6.35 49.70 20.70 4.11 0.08

Shift III RMD3 11.03 6.48 50.15 21.07 3.40 0.06

Shift IV RMD4 11.13 6.07 50.52 20.14 3.50 0.06

Shift V RMD5 11.64 5.90 49.57 20.55 3.78 0.06

mineral and muscovite, yavapaiite and anorthite as minor minerals.

Red Mud Composite Sample (RMD-C):

Gypsum:

(a) Chemical Characterization: The composite red mud sample, RMD-C was subjected to chemical analysis and the results are presented in Table-3.

(a) Chemical Characterization: The gypsum sample was subjected to chemical analysis and the results are presented in Table-4. The sample is considered suitable as a set retarder in the manufacture of cement. (b) Mineralogical Characterization: X-ray diffraction of the gypsum sample indicated the presence of gypsum as a major mineral and α-quartz and albite as minor minerals.

Table 3. Chemical Analysis of Composite Red Mud Sample (RMD-C) SI No Constituents Results Determined (%) 1 11.45 LOI 2 6.32 Si0 2 49.62 3 Fe 2 0 3 4 20.74 A1203 5 CaO 2.45 6 0.15 MgO 7 0.08 so3 8 Na 2 0 3.66 9 0.07 κ2ο 10 0.33 p2o5 11 4.76 Ti0 2 12 Mn 2 0 3 0.11

Coal: (a) Chemical Characterization: The results of proximate analysis of the coal sample are presented in Table-5. The ash prepared from the above coal was subjected to chemical analysis and the results are presented in Table-4 Table 5. Proximate Analysis of Coal Sample. Characteristics Sl.No. Results 35.34 1 Ash Content (%) 2 Moisture Content (%) 1.82 Volatile Matter (%) 30.04 3 4 Fixed Carbon (%) 32.80 5 Calorific Value(kcal/kg) 5140

(b) Mineralogical Characterization: X-ray diffraction analysis of the composite red mud sample, RMD-C indicated the presence of gibbsite and hematite as major minerals and goethite as minor mineral.

The above results of chemico-minerological characteristics of red mud, limestone and other raw materials indicate that their matrix is compatible with the cement matrix and can be used in the manufacture of cement.

Other Raw Materials: Limestone: (a) Chemical Characterization: The limestone sample was subjected to chemical analysis and the results are presented in Table-4. The sample can be classified as cement grade limestone.

Sl.No. 1 2 3 4 5 6 7 8 9 10 11

Table 4.Chemical Analysis of Raw Materials Constituents Limestone Shale Gypsum Determined

(%)

LOI Si0 2 Fe 2 0 3 A1203 CaO MgO

so3 Na20

κ2ο Ti0 2 cr

37.25 10.78 0.81 1.95 44.90 3.06 0.00 0.13 0.73

-

0.008

2.72 85.93 1.14 7.22 0.58 0.21 0.13 0.18 1.76 0.27 0.01

17.10CW 11.44(+IR) 0.26 NIL 28.85 0.66 39.90 0.24 0.04

-

Technical Suitability of NALCO Red Mud for the Manufacture of Cement Chemical and mineralogical analysis of the NALCO red mud sample indicated that its composition is quite compatible with the cement matrix. Therefore, the NALCO red mud sample was considered prima-facie suitable for its gainful utilization in the manufacture of Ordinary Portland Cement (OPC)clinker in combination of other raw materials viz. limestone, shale and fine coal [2 -5]. The process of making OPC is discussed below.

Coal Ash 0.87 58.64 6.25 25.85 3.19 0.67 1.72 0.08 0.83 1.56 0.017

Raw Mix Design & Its Optimization: The Approach: While designing the raw mixes, attempts were made to maximize the use of red mud, RMD-C as raw mix component in the raw mix design for the manufacture of OPC. Keeping this in mind a large number of different raw mixes were designed to achieve good quality clinker. The proportions of raw material in the raw mixes designed (RM-1 to RM-5) are shown in Table-6. The design parameters optimized in the designed raw mixes are shown in Table-7-11.

(b) Mineralogical Characterization: X-ray diffraction analysis of the limestone sample indicated the presence of calcite as a major mineral and α-quartz, dolomite and muscovite as minor minerals.

Raw Mix No. RM-1 RM-2 RM-3 RM-4 RM-5

Shale: (a) Chemical Characterization: The shale sample was subjected to chemical analysis and the results are presented in Table-4. (b) Mineralogical Characterization: X-ray diffraction analysis of the shale sample indicated the presence of α-quartz as major

98

Table-6. Various Raw Mix Designs Prepared Red Mud Shale Limestone Coal Ash Absorption(%) (%) (%) (%) 3.00 94.70 2.30 4.75 2.14 3.50 94.36 4.75 94.02 4.00 1.98 4.75 4.30 93.81 1.89 4.75 93.34 5.00 1.66 4.75

Raw Materials Proportion

Table-7. Design Parameters of Raw Mix RM-1 Coal Ash 1 Red Mud Limestone Shale Absorption 4.75 3.00 94.70 2.30

(%)

Composition

(%)

Raw Mix

S 35.68

Clinker

MgO: 2.89 Na 2 0: 0.27 K 2 0: 0.49

4.39

5.37

62.69

MgO: 4.30 Na 2 O:0.40 K 2 0: 0.76

SM

AM

1.08

2.17

1.08

Liquid Content(%)

0.93 C3S

2.12

1.22 C3A

C4AF

55.94

17.04

6.80

13.35

2.63

42.61

MgO: 2.91 Na 2 O:0.24 K 2 O:0.73

Raw Mix

21.11

3.68

5.13

63.25

MgO: 4.34 Na 2 O:0.35 K 2 0: 1.13

Clinker

Liquid Content(%)

Modulii Values RawMix

LSF

Clinker 1 Potential Phase Compo1 sition(%)

2.52

1.15

Clinker 1 Potential Phase Compo1 sition(%)

0.93

2.40 C2S

1.40 C3A

C4AF

57.32

17.32

7.37

11.19

c3s

(%)

29.25

35.53

12.09

20.65

PU

£5

Pu

12.23

2.53

2.72

42.47

MgO: 2.90 Na 2 O:0.25 K 2 O:0.73

Raw Mix

20.08

4.03

5.25

62.97

MgO: 4.32 Na 2 0: 0.38 K 2 0: 0.76

Clinker

Liquid Content(%)

Modulii Values RawMix

LSF

Clinker Potential Phase Composition^)

S

9

κ2ο

Modulii Values RawMix

LSF

SM

AM

1.08

2.43

1.08

Clinker Potential Phase Composition^)

0.93 C3S

2.25

1.30 C3A

C4AF

56.64

17.17

7.08

12.27

c2s

c2s

(%)

Composition

MgO Na20

d <

d

d

d

κ2ο

31.17

Table-10. Design Parameters of Raw Mix RM-4 Raw Red Mud Limestone Shale Coal Ash Absorption Materials 4.75 4.30 93.81 1.89 Proportion

Table-8. Design Parameters of Raw Mix RM-2 Red Mud Limestone Shale Coal Ash Absorption 3.50 2.14 94.36 4.75

(%)

Clinker

42.32

2.28

κ2ο

1.08

35.61

2.80

12.37

(S

AM

Raw Mix

2.77

PU

SM

(%)

9

S

d

MgO Na20

d <

9

LSF

Composition

(%)

Composition

MgO Na20

1 Modulii Values 1 RawMix

Raw Materials Proportion

Table-9. Design Parameters of Raw Mix RM-3 Limestone Shale Coal Ash Red Mud Absorption 94.02 1.98 4.75 4.00

d <

O

d

Raw Materials Proportion

(%)

30.03

99

en

MgO Na20 K20

q

9u

2.92

2.85

42.24

MgO: 2.88 Na 2 O:0.28 K2O:0.72

4.62

5.44

62.51

MgO: 4.29 Na 2 O:0.42 K 2 0: 0.75

SM

AM

1.08

2.11

0.98

Liquid Content(%)

0.93 C3S

2.04 C2S

1.18 C3A

C4AF

55.44

17.03

6.63

14.00

S

O

35.49

12.01

20.52

PU

31.86

Raw Materials Proportion

Red Mud

Limestone

Shale

5.00

93.34

1.66

(%)

Composition

(%)

Raw Mix

up to 5% in the raw mixes designed, the raw mix RM-2 was selected to avoid process problems attributed to the presence of a high proportion of fluxing agents (A1203 and Fe203) in such raw mixes. From Table-7 to 11, it is found that on increasing the red mud content from 3.0% to 5.0%, the liquid content was found to be increased continuously. The presence of increased liquid content known to cause process implications in the cement manufacture and especially beyond 30% liquid content is detrimental to the life of refractory lining of the kiln. Generally, the liquid content has been maintained in the range of 27 to 30% in cement manufacture. 3.5% has been selected as the optimal red mud content keeping in mind the maximum utilization of red mud in cement manufacture. Its liquid content is 30.03% and is on border line but is still in an acceptable range. On going from red mud content of 3.5 to 5%, the liquid content increases from 30.03% to 33.45% and the resultant clinkers were found to be sticky to the refractory linings and will reduce the life of the refractory linings and have many process implications. Also, the presence of high liquid content results in the formation of boulders instead of normal sized clinkers which further makes problems in the process. The role of fluxing agents like Fe203, A1203 etc is very critical in the formation of clinker. The general range for A1203 content is 4 to 7% and for Fe203 is 2 to 4% in the clinker. The details of the optimized raw mix RM-2 is given in Table-8. The raw mix RM-2 is capable of yielding good quality clinker at 1450 °C with a retention time of 20 minutes.

Table-11. Design Parameters of Raw Mix RM-5

d

S

cJ5

35.39

Clinker

d Pu

d <

9

Coal Ash Absoφtion 4.75 MgO Na20 K20

11.80

3.26

2.97

Aim

MgO: 2.87 Na 2 O:0.31 K2O:0.48

20.19

5.10

5.61

62.13

MgO: 4.26 Na 2 O:0.46 K 2 0 : 0.75 Liquid Content(%)

Modulii Values RawMix

LSF

SM

AM

1.09

1.90

0.91

Clinker 1 Potential Phase Compo| sition(%)

0.93 C3S

1.89 C2S

1.10 C3A

C4AF

54.53

16.78

6.24

15.51

33.45

Table 12.Burnability Studies of Cement Raw Mixes with Red Mud as a Component (Retention Time 20 Minutes)

Raw Mix Preparation: Raw mixes, RM-1 to RM-5 were prepared by taking weighed quantities of raw materials as per the designs, blending them thoroughly and grinding the mixes to fineness of 10% residue on 90μ(170 mesh) sieve. Nodules of about 1 cm in diameter were prepared by mixing about 12% water and were dried in an electric oven at 105±5° C for about 2 hrs before subjecting them to burnability studies. Burnabilities Studies: Burnability studies were carried out on all the raw mixes. The dry nodules were introduced in to a laboratory furnace at ambient temperature, which was gradually raised to 1450°C. The raw mixes were fired at 1300, 1350, 1400 and 1450°C with a retention time of 20 minutes. The clinkers, CL-1 to CL-5, prepared from the raw mixes, RM-1 to RM-5 respectively were room cooled and their free lime content determined. The results of free lime determination are presented in Table-12 which indicate that all the raw mixes have good burning characteristics and are capable of yielding quality clinkers even at 1450°C with a retention time of 20 minutes. The free lime content in all the clinker samples was found to be < 0.50% at 1450°C.

SI. No. 1

Raw Mix RM-1

2

RM-2

3

RM-3

4

RM-4

5

RM-5

Temperature(°C) 1300 1350 1400 1450 1300 1350 1400 1450 1300 1350 1400 1450 1300 1350 1400 1450 1300 1350 1400 1450

Free Lime % 1.66 0.82 0.32 0.06 1.73 0.88 0.34 0.06 2.44 1.57 0.55 ! 0.13 2.35 1.18 0.44 0.09 1.73 0.61 0.31 0.15

Preparation and Evaluation of Bulk Clinker

Optimization of Raw Mix Design:

10 kg sample of optimized raw mix RM-2, selected for detailed investigations was prepared by taking weighed quantities of raw materials viz. red mud, limestone, shale and coal ash as per its raw mix design, blended them in a ball mill and ground to a fineness of 10% residue on 90μ(170 mesh) sieve. The nodules were prepared in a pan nodulizer and dried in an electric oven at 105±5 °C for 2 hrs before introducing in an electric furnace at ambient temperature and eventually firing it at 1450 °C for 20 minutes. The resultant clinker CL-2 from raw mix RM-2 was

While designing the raw mixes, every effort was made to keep the level of utilization of red mud to the maximum extent possible. All the raw mixes prepared viz. RM-1 to RM-5 yielded good quality clinkers when test fired during burnability studies at 1450 °C with retention time of 20 minutes. But, keeping in view the maximum utilization of red mud, design parameters and process implications there from, raw mix RM-2 was taken as the optimized raw mix design. Although the red mud could be utilized

100

13KWH/T. As this value of Bond Index lies on the middle of the range, the clinker appears to be hard enough in nature.

studied for chemical, mineralogical and grinding characteristics and the OPC prepared there from was evaluated for performance as per relevant Indian Standard Specifications. Three batches of such bulk clinkers, lOkgs each, were prepared by adopting the same procedure to evaluate the reproducibility of the results.

Reproducibility of the Clinker In order to ascertain the reproducibility of the product i.e. clinker, two more batches of the clinker were prepared in the laboratory from the optimized raw mix, keeping the process parameters same as maintained during preparation of bulk clinker sample, CL-2.

Chemical Analysis: The chemical analysis of the bulk clinker, CL-2, carried out as per Indian Standard Specification, IS: 4032-1985 and is presented in Table-13. These results indicated that the quality of the clinker was good and capable of yielding good quality cement.

Preparation and Evaluation of Two Additional Batches of the Clinker: Two additional batches of the clinkers CL-2-A and CL-2-B were prepared by taking 10 kgs of raw mix RM-2 adopting the same procedure as described under Preparation and Evaluation of Bulk Clinker. The resultant clinkers designated as CL-2-A and CL-2-B were studies for chemical and mineralogical characteristics. The results are discussed below.

Table 13. Chemical Analysis of Bulk Clinker

SI. No. 1 2 3 4 5 6 7 8 9 10 11 12

Results Obtained

Constituents Determined LOI Si0 2 Fe 2 0 3 A1203 CaO MgO

(%)

0.35 20.60 4.02 5.30 63.27 4.22 0.27 0.38 0.78 0.003 0.38 0.25

so3 Na20 κ2ο cr

Insoluble Residue Free Lime

Chemical Analysis: The chemical analysis of the two additional batches of clinkers, CL-2-A and CL-2-B, were carried out as per Indian Standard Specification. IS: 4032-1985 and the results are presented in Table-15. Results indicated that the chemical composition of both the clinkers, CL-2-A and CL-2-B, are comparable to the composition of thefirstbatch of the clinker, CL-2. Table 15. Compositional Variation in Three Batches of Clinkers

SI. No.

Mineralogical Analysis: The clinker sample CL-2 from raw mix RM-2 was evaluated for its mineralogical composition by optical microscopy and the results along with the granulometry of the clinker phases are presented in Table-14. Clinker phases are moderately developed and are homogenously distributed. Majority of alite grains are subhedral in shape. Few pseudo hexagonal alite grains are also developed in the clinker. Transformation of belite into alite has also been observed. Some belite clusters were also presented in the sample. Few crystals of free lime were also present. It indicated that the quality of the bulk clinker sample so prepared was good. Table 14. Mineral Phase Analysis of Bulk Clinker Sample by Optical Microscopy SI. No.

Clinker Sample

Phases

Quantity

1

CL-2

C3S C2S C3Al C 4 AFj CaOf

48 33 18

(%)

1

1 2 3 4 5 6 7 8 9 10 11 12

Granulometry (μπι) Min. Avg. Max. Size 12 3 18 3 17 10

-

-

-

1

9

4

Constituents Determined LOI Si0 2 Fe 2 0 3 A1203 CaO MgO

so3 Na20 K20 cr

Insoluble Residue Free Lime

Results Obtained (%) CL-2-A CL-2 CL-2-B 0.35 0.45 0.40 20.60 20.48 20.43 4.02 4.12 4.18 5.30 5.29 5.26 63.24 63.27 63.35 4.22 4.23 4.26 0.34 0.27 0.31 0.37 0.38 0.31 0.78 0.69 0.77 0.004 0.003 0.006 0.34 0.38 0.36 0.25

0.22

0.24

Mineralogical Analysis: (a) Optical Microscopic Analysis: Samples from the two additional batches of clinkers, CL-2-A and CL-2-B, prepared in the laboratory ascertaining the reproducibility were evaluated for their mineralogical composition by optical microscopy. The results along with the granulometry of clinker phases are presented in Table-16. The alite grains were subhedral in shape. Transformation of belite into alite was observed in both the samples. Most of the belite grains were sub-rounded in shape with corroded margins. Few crystals of free lime were also observed in these two samples. The results have revealed that the granulometry and mineralogical characteristics of the two clinkers of additional batches, CL-2-A and CL-2-B, are comparable to the first batch the clinker i.e. CL-2.

The X-ray diffractogram of the bulk clinker sample CL-2 indicated the presence of homogenously distributed and well developed phases. Grindabilitv Studies: In order to determine the energy required for grinding clinker, laboratory grindability test was conducted using standard bond ball mill. The Bond Index for clinker, CL-2, has been found to be 10.9KWH/T which lies in between the range of Bond Index 9-

101

Table 16. Variation in Minerai Phase Analysis of Three Batches ___^ of Clinker Samples by Optical Microscopy

SI. No.

Clinker Sample

Phases

Quantity

1

CL-2

C3S C2S C3Al C4AFJ CaOf C3S C2S

48 33 18

2

3

CL-2-A

CL-2-B

Granulometry^m) Min. Avg. Max. Size 12 18 3 17 10 3

-

-

-

C3A!

1 45 34 17

1 3 4

9 19 17

4 11 11

C3A!

1 46 34 18

1 3 4

9 18 18

4 11 12

1

1

9

4

C4AFJ CaOf C3S C2S C4AEJ CaOf

Table 17.Performance of Ordinary Portland Cement (OPC)

SI. No.

-

-

-

-

1 2 3

-

4

-

Property

Fineness(M2/kg) Setting time (Min.) Initial Final Compressive Strength(N/mm2) 3 Days 7 Days 28 Days Soundness Le-chatelier(mm) Autoclave (%)

Results 305

Requirement of IS: 12269-1987 (53 Grade OPC) Not less than 225

128 254

Not less than 30 Not more than 600

29.0 42.5 60.8

Not less than 27 Not less than 37 Not less than 53

1 0.073

Not more than 10 Not more than 0.8

The above results indicated that good quality clinker can be prepared by using 3.50% red mud, 94.36% lime stone and 2.14% shale. The Ordinary Portland Cement (OPC) prepared from this clinker conformed to all the requirements of three Indian Standard Specifications for 33, 43 and 53 grade of OPC. In addition, OPC made from 3.0% of NALCO red mud, 94.70% of lime stone and 2.30% of shale also confirmed to the above specifications. Complete performance evaluation of different OPC samples prepared from 3.0 and 3.5% of NALCO red mud was conducted.

(b)X- Ray Diffraction Analysis: Samples from the two additional batches of clinkers CL-2-A and CL-2-B, were also subjected to X- Ray diffraction analysis. It indicated the presence of homogenously distributed and well developed clinker phases in these two samples, CL-2-A and CL-2-B, similar to the first batch of the sample, CL-2. The chemico-minerological evaluation of the results of the two additional batches of clinker samples, CL-2-A and CL-2-B, indicated their close resemblance to the clinker, CL-2, prepared in the first batch. Results have clearly established the reproducibility of the product i.e. clinker made from NALCO red mud.

Conclusions Chemical and mineralogical analysis of NALCO red mud indicated that its composition is quite compatible with the cement matrix. The NALCO red mud therefore was considered primafacie suitable as a raw mix component in the manufacture of OPC. While designing the raw mixes, every effort was made to keep the level of utilization of NALCO red mud to the maximum possible extent. Upto 5% red mud could be gainfully utilized in the raw mixes. However, to overcome process problems attributed to the presence of high proportion of fluxing agents (A1203 and Fe203) in other raw mixes, the raw mix RM-2 with 3.5% red mud was considered most suitable for the preparation of bulk clinker, CL-2. Physical performance of the OPC made from the bulk clinker, CL-2, utilizing 3.5% of NALCO red mud conformed to all the requirements of all the three Indian Standard Specifications for 33, 43 and 53 grades of OPC viz. IS.269-1989, IS:8112-1989 and IS: 12269-1987 respectively. Reproducibility of the clinker made from NALCO red mud is also established.

Preparation and Evaluation of Ordinary Portland Cement (OPC) Ordinary Portland Cement ,OPC-2 was prepared by grinding the bulk clinker CL-2 so prepared with 5.0% gypsum to a fineness of ~300m2/kg and tested for setting time, compressive strength, Lechatelier and autoclave expansion tests as per IS :4031-1988 and the results are presented in Table-17. Setting Time: The initial and final setting times of Ordinary Portland Cement prepared in laboratory from bulk clinker were determined as per IS: 4031-1988 and the results are presented in Table-17. The results indicated that the cement sample OPC-2 conformed to the requirements of the standard.

References 1. H.Z. Xu, Technological and Economic Feasibility Study on Producing Building Materials with Red Mud, Gold 17(1996), pp. 17-21(in Chinese). 2. Maneesh Singh, S. N. Upadhyaya and P. M. Prakash, Preparation of Special Cements from Red Mud, Waste Management, Vol. 16, Issue 8, (1996), pp.665-670. 3. M Singh, et al., Preparation of Iron Rich Cements using Red Mud, Cement Concrete Research, 27(7)(1997), pp.1037-1046. 4. P.E.Tsakiridis, S.Agatzini-Leonardou, P.Oustadakis, Red Mud Addition in the Raw Meal for the Production of Portland Cement Clinker, Journal of Hazardous Materials, B 116(2004), pp.103-110. 5. C.Brunori, C.Cremisini, P.Massanisso, V.Pinto, L.Torricelli, Resuse of a Treated Red Mud Bauxite Waste, Studies on Environmental Compatibility, Journal of Hazardous Materials, 117(l),(2005),pp.55-63.

Compressive Strength: The compressive strength of OPC-2 prepared above was determined as per IS: 4031-1988 and the results are presented in Table-17. The results indicated that the OPC-2 sample conformed to all requirements of all the three Indian Standard Specifications for 33, 43 and 53 grades of OPC viz. IS:269-1989, IS:8112-1989 and IS: 12269-1987. Soundness: Autoclave and Le - chatelier expansion tests on OPC-2 prepared above were carried out as per the procedures laid down in IS: 4031-1988. The results indicated high volume stability of the cement sample and conformed to all the requirements of all the three Indian Standard Specifications for 33, 43 and 53 grades of Ordinary Portland Cements viz. IS:269-1989, IS:8112-1989 and IS: 12269-1987 respectively.

102

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

RECOVERY OF METAL VALUES FROM RED MUD P.K.N.Raghavan1, N K. Kshatnya1 and Katarzyna Wawrynink2 Îharat Aluminium Company Limited,Korba (CG) India 2 Warsaw University of Technology,Poland Keywords: red mud, roasting, alumina, iron, titania values. However, there is a very strong argument against this in the countries with large availability of Iron ores, Titanium ores and Aluminum ores. However, the time is not very far off when depletion of mineral resources will make it viable to convert these wastes into useful products. The growing awareness of environmental pollution and concern for protecting ecological balance has created alternative ways of its utilization converting the waste into wealth. The successful use of red mud as a raw material source for metal production depends on the recovery of each metal value present in the particular red mud.

Abstract In processing bauxite for production of alumina by the conventional Bayer process, huge quantities of so called red mud are generated which are disposed of as a waste product. The red mud is normally contains caustic soda and therefore poses pollution hazardous and requires storage in specially made large size ponds. This waste material at present does not find any use, hence, in addition to the pollution hazard, considerable expenditure and wastage of land is involved in disposal of this waste material. Therefore, in view of the increasing problem of waste disposal, it is desirable to recover the oxides of iron, titanium and aluminum from red mud.

Besides the undissloved alumina, red mud contains the alkali insoluble oxides like Fe 2 0 3 and Ti02.The alkali aluminosilicate is formed due to the reaction between combined silica of bauxite with caustic soda. Various formulae have been proposed to represent this desilication compound. Bayer gave the formula for this desilication product as Al203.Na20.3Si02.9H2O.Though the composition of these silicates has not been definitely ascertained, it is certain that it depends upon the content of the reactive silica and digestion conditions adopted.

The process applies to the recovery of alumina, titania and ferric oxide from red mud which comprises the following steps (i) Roasting of red mud in presence of alkali (ii) Extracting the salts produced in step (i), (iii) Precipitation of hydroxide salt and (iv) Converting the hydroxide salt into oxides. Balco's red mud contains 45-47% Fe 2 0 3 , 1820% Ti0 2 and 16-25% of undigested alumina .Many methods have been tried by others to recover the metal values and it is proposed in the laboratory on bench scale. The present papers provide an alternative and more environmentally acceptable process for the recovery of iron, titanium and aluminum from Balco's red mud.

It has been indicated earlier that the alumina is present as an alkali aluminosilicate with minor amounts of undissloved (diasporic) aluminium oxyhydroxide. Experimental The possible method for treating red mud for the recovery of metal values comprises of the following steps: (i) Roasting the red mud with soda ash and lime (ii) Extracting the fused mass (iii) Digestion and precipitation of hydroxide salt (iv) Conversion into appropriate oxide The red mud sample of Bharat Aluminum Company Limited, Korba (India) was in the form of cakes having pale reddish color. This was easily friable and could be crushed with ease. The major constituents of red mud are analyzed by XRF (PW-2440 Philips, Netherland). The sample analyzed as follows as shown in Table 1 :

Introduction To produce every ton of the aluminum metal 5 to 5.5 tonnes of bauxite ore has to be treated by the well known Bayer process. On digestion with caustic soda, bauxite leaves behind about 50% of its own weight of an insoluble residue known as Red Mud. Disposal of such huge quantities of this solid waste by their impoundment in mud lakes causes problems of increasing land cost, storage and pollution. The application of any red mud depends on its Physico-chemical properties. The red mud is highly complex material containing largely the following six oxides namely ferric oxide, Aluminum oxide, Titanium dioxides, Silica, Sodium Oxide and Calcium Oxide in varying quantities besides minor amount of traces of other elements as oxides.

Tablel: The Major Constituents of Red Mud Weight % Constituents 10.48 LOI 5.45 Si0 2 37.33 Fe 2 0 3 Ti0 2 18.88 18.38 A1203 Na 2 0 5.89 CaO 2.71

Furthermore, it is alkaline, fine sized and poor in settling. Dumping of red mud is certainly an environmentally unfriendly activity. However the solid waste, red mud is a store house of wealth because it contains the metal values of Iron, Aluminum and Titanium. Perhaps the best way would be to convert this waste into the corresponding metal

103

Results

XRD analysis was carried out using PANalytical X'Pert Cubix Pro Series diffractometer equipped with a copper target tube, X'celerator detector and operated at 40kV and 30mA.The dried sample were scanned within 2Θ range from 10-70°. Diffraction data were analyzed using PANalytical X'Pert High Score Plus Version 2.1.The main crystalline phase of red mud were Hematite (Fe 2 0 3 ) »Calcite (CaC0 3 ), Goethite (FeOOH), Gibbsite (Al(OH)3), Boehmite (AlOOH), Cancrinite (3NaAlSi04.NaOH), Quartz(Si02), Anatse (Ti0 2 ). The X-ray diffractogram of red mud is shown in Fig.l.

In the roasting experiments burnt lime was used to immoblize the Si0 2 of the red mud. A burnt lime containing 85% of total CaO was used. The effect of the different variables on the recovery of A1203 was studied.

RED MUD SODIUM CARBONATE

Roasting (1100 Of

The method adopted for the recovery of A1203, Ti0 2 and Fe 2 0 3 from red mud is presented in Fig.2. The process for the recovery of metal values from red mud is briefly outlined below: Crushed red mud was thoroughly mixed with (i) burnt lime and soda ash in the different weight ratios and then roasted at the desired temperature, ranging from 900-1100°C.The roasted mass was crushed and extracted with hot water and filtered. The Filtrate, sodium aluminate can be (ii) accommodated by recycling into the aluminate solution in the Bayer process. (iii) The residue is light and can be treated for the recovery of Iron and Titanium. The residue is digested with concentrated sulfuric acid. (iv) The digested slurry was dissolved in water and filtered. The residue containing high silica was separated. Titanium hydroxide was precipitated from the (v) filtrate by hydrolysis. The Titanium hydroxide was filtered and calcined to get pure Ti0 2 . The Filtrate was evaporated to obtain FeS0 4 . (vi)

SODIUM ALUMINATE Digest with H2S04, Fijte$_

Hydrolysis

FERROUS SULPHATE

Fig. 2. Schematic diagram for recovery of metal values from Red Mud.

Effect of Soda ash on Alumina Recovery: The red mud mixture containing burnt lime and soda ash was roasted at 1100°C for 3 hours and subsequently leached. Lime was added to the mixture so that the mass contained CaO:Si0 2 in the molar ratio of 1.5:1.The Na 2 C0 3 weight ratio was varied from 0.3 to 0.7 by the addition of varying quantities of soda ash and the alumina recoveries are indicated in Table 2 and Fig. 3.

Red Mud-14 21 08 10

It is observed that on addition of increased quantities of soda ash, the sinter becomes hard and has a tendency to absorb moisture thereby making dry grinding difficult and wet grinding has to be employed. I,

*VW**»

mm ÌW w

%M

i p

Table 2. Effect of varying sodium carbonate ratio % A1203 Sinter Weight Ratio Extracted Characteristics ( Mud:Na2C03:CaO) 64.98 1:0.3:0.1 Easily extract 69.09 1:0.4:0.1 Easily extract 76.17 1:0.5:0.1 Slightly hard 76.32 1:0.6:0.1 Slightly hard 1:0.7:0.1 76.47 Slightly hard

■ ■ ■ i ■

Position [*2Theta]

Fig.l. XRD Pattern of Balco's Red Mud

104

1 1

Fig.3.Bfect of Na2C03 ratio on alumina

Fig.4. Bfect of CaO ratio on alumia recovery

recovery

Io 80

« 80 O|70

2

1 75 -

<5 60 X

M

o

S 50

M

0.3

0.4

0.5

0.6

0.7

< 70 * 0.1

0.8

Weight ratio of Na 2 C0 3

0.12

0.14

0.16

0.18

Weight ratio of CaO

From Table 2 and fig.3 it is observed that the alumina recovery increases from 64.98% to 76.17% as the Na 2 C0 3 weight ratio increases from 0.3 to 0.5 and no appreciable increase is observed thereafter. This may be attributed to the side reactions taking place during the roasting operation itself, such as the formation of sodium ferrites and sodium titanate. The presence of Fe 2 0 3 & Ti0 2 affects the alumina recoveries. It was observed that use of Na 2 C0 3 and red mud in the weight ratio of 1:0.5, 76.17% A1203 could be recovered and further roasting tests were carried out using this quantity and the effect of other variables.

Table 4. Effect of roasting temperature %Mole ratio of 1 Temperature %A1203 Extracted Si0 2 :Al 2 0 3 dissolved in of roasting,°C solution 900 1000 1100 1150

71.12 74.34 76.30 78.09

1.01 1.09 1.11 1.23

It is observed in figure 6 that on increasing the roasting temperature from 1 hour to 3 hour, the A1203 recovery is increased from 70.25% to 75.95%. Further roasting for 4 hours and 5 hours decreases the A1203 recovery to 74.14%. A retention time of 2 to 3 hours is sufficient at this roasting temperature.

Effect of CaO on alumina recovery: Varying quantities of burnt lime were used for roasting red mud with sodium carbonate in the range of 1:0.5 and the alumina recoveries are indicated in Table 3 and Fig.4 were studied. It is observed from the Table 3 and Fig.4 that as the CaO content is increased, the alumina extraction is decreased and this can be attributed to the formation of a complex compound of lime and alumina.

Discussion When the red mud is roasted with sodium carbonate and lime, the alumina is converted into a sodium aluminate while silica is converted to an insoluble dicalcium silicate as follows: Na 2 C0 3 + A1203 = 2 NaA102 +C0 2 Si0 2 + 2CaC0 3 = (CaO) 2Si0 2 + 2C0 2

Effect of temperature of roasting on recovery of alumina: In order to determine the effect of roasting temperature mixtures containing red mud, sodium carbonate and burnt lime were heated in the weight ratio of 1:0.5:0.1 for 3 hours at temperature ranging from 900°C to 1150°C.

The soda besides reacting with A1203 reacts with Fe 2 0 3 .Ti0 2 .Si0 2 as follows: Na 2 C0 3 + Fe 2 0 3 = Na 2 O.Fe 2 0 3 + C0 2 Na 2 C0 3 + Ti0 2 = Na 2 O.Ti0 2 +C0 2 Na 2 C0 3 + Si0 2 = Na 2 O.Si0 2 + C0 2

Table 3. Effect of varying lime quantity on alumina recovery %A1203 Weight Ratio % Mole Mud:Na2C03:CaO extracted ratio of Si0 2 :Al 2 0 3 (inclusive of lime in mud) dissolved in solution 76.03 1:0.5:0.10 1.08 75.12 1:0.5:0.12 0.97 1:0.5:0.14 74.05 0.89 0.83 1:0.5:0.16 72.66

During leaching, the sodium ferrite and part of the sodium titanate decompose to Fe 2 0 3 and Ti0 2 . Na 2 O.Fe 2 0 3 + H 2 0 = 2NaOH + Fe 2 0 3 Na 2 O.Ti0 2 +H 2 0 = 2NaOH+Ti02 The sodium silicate dissolves in water and thus enters the solution.

105

Acknowledgement

Fig.6.Bfect of time of roasting on alumina recovery * * 80 Ί e« g Z 70 <

-t—

-f—

*

^^

♦

-♦

3

4

5

The authors acknowledge the constant encouragement from our Chief Operating Officer,Mr.Bibhu Prasad Mishra and Chief Executive Officer and Whole Time Director,Mr. Gujan Gupta during the progress of this work.The authors would like to thank the Management of Vedanta Resources for allowing us to publish this paper.

X*50 2

References

6

Time,hours

1. J. D. Edwards, F.C. Frary, The Aluminium Industry, Vol.1, Mcgraw Hill Book Co, p. 127. 2. Edwards, Trans, A.I.M.E., Vol.182, 1949. 3. F.R. Archibald, Trans.A.I.M.E. Vol.159, 1944. 4. R.S. Thakur and S.N. Das, Red Mud Analysis and utilization, Wiley Eastern, Limited, New Delhi, 1994.

The residue resulting from the lime soda roasting process is very light, fine and can be treated for the recovery of Iron and titanium values. The following observations were found on roasting of red mud with Na 2 C0 3 and lime in the weight ratio of 1:0.5:0.1 atll00°Cfor3hours: (i) The alumina extraction efficiency was 76%. (ii) Recovery of Iron as Fe 2 0 3 was 74.81 % (iii) Recovery of Titania as Ti0 2 was 72.13% Conclusion It was observed that 76% of the A1203 could be extracted by roasting the red mud with Na 2 C0 3 and CaO (weight ratio 1:0.5:0.1) at 1100°C for 3 hours. About 0.8-1.1% mole ratio of Si0 2 dissolved simultaneously in solution. The aluminate liquor obtained is unsuitable for precipitation of desired grade of aluminium hydroxide. It will thus be necessary to adopt a separate desilication step. It is suggested that if this silica containing solution from the lime soda process is added, the resulting pregnant solution will contain low Si0 2 concentrations suitable for subsequent precipitation of aluminium hydroxide. The proposed method provides more environmentally acceptable process for the extraction of Aluminium oxidejron Sulphate and Titanium dioxide from red mud.

106

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

RED MUD FLOCCULANTS USED IN THE BAYER PROCESS F. Ballentine, M. E. Lewellyn, S. A. Moffatt Cytec Industries Inc.; 1937 West Main Street; Stamford, CT 06904-0060 USA Keywords: Bayer process, flocculants, flocculation, red mud characteristics of each type of flocculant can ultimately be attributed to the molecular composition and also to the molecular structure. This paper examines chemistry of each type of flocculant, and correlates these differences to the flocculation performance.

Abstract Flocculation and separation of red mud is an integral part of the Bayer process. Over the latter half of the 20th century, flocculant technology dramatically changed from natural starches to use of "rationally designed" polymers. Many of these advancements were due to the introduction of liquid or emulsion based flocculants which enabled elaborate post-reaction chemistry to be done on the polymer backbone. This paper presents a historical overview of milestones of flocculant technology used in the Bayer process up to present day. Discussion of flocculants is based on inventions in the published literature that have gained widespread use throughout the industry and will included the benefits/advantages of different flocculant technology for settling red mud.

Background It is worth noting that following his discoveries, Dr. Bayer received contracts to build alumina refineries in France, Germany, Italy, the United Kingdom, and the United States [1]. Thus, it can be conjectured that the global alumina industry likely originated from the plants that Dr. Bayer helped design and build. In all likelihood the initial design of these plants did not utilize flocculants in the separation of red mud, and mud was removed by direct filtration of digested slurry. Undoubtedly, efforts at process improvement commenced in parallel among the different geographical regions where Bayer helped build refineries. However, for the purpose of this paper and the role of red mud flocculants, discussion will be limited to developments within the North American and Jamaican refineries.

Introduction The Bayer process for the production of alumina has changed little since it was discovered and developed by Dr. Karl Joseph Bayer. The initial discovery (1888) that aluminum hydroxide could be precipitated by seeding a supersaturated sodium aluminate solution, and the subsequent discovery (1892) that aluminum hydroxides and oxides in bauxite could be selectively dissolved in sodium hydroxide under pressure to produce a sodium aluminate solution remain the fundamental steps in the production of alumina today [1]. Obviously, there have been tremendous advances in terms of design, engineering, and control of the process since then, but the basic operations are still based on Dr. Bayer's discoveries.

In the early 20th century, there are numerous references to the use of starch as a flocculant [2-4]. The coal industry was among the first to utilize starch, in the 1930s, as a flocculant for the removal of fine coal particles from washery water [2]. Around this time, 1939, starch was also reportedly used as a flocculant to aid in the filtration of red mud [3]. It is further reported that in 1940, due to the use of lower grade bauxites in the United States, the first cone bottom settlers were installed in the process. This led to the development of counter current décantation (CCD) circuits comprised of primary settlers and washers. At this point it seems that the alumina industry was by far the largest consumer of starch based flocculants [4].

One of the most important advances in the Bayer process has been the introduction of red mud flocculants. Interestingly, it is reported that Dr. Bayer further described the advantages of his process in terms of the purity of the alumina produced, relative to precipitation by C0 2 , and the crystalline form which made it easy to wash and filter [1], However, he also cited the main difficulty in the process was the separation and removal of the red mud; fine, insoluble, impurities that needed to be removed from the sodium aluminate solution. This is where flocculants have played an indispensable role in the process. Considering the history of flocculant use for alumina production by the Bayer process there are arguably three distinct milestones: starch, synthetic polymers, and modified synthetic polymers. Each of these flocculants have unique characteristics which is why the industry gradually moved from starch to synthetic polymers, and then to modified synthetic polymers. The benefits and performance

A significant advance in flocculant technology occurred in the mid-1950s with the introduction of synthetic flocculants for the mining industry [5,6]. The technical literature is rich with information on the development and use of synthetic flocculants of this era as several large chemical manufactures, in addition to governmental and academic researchers, were active in this area. These flocculants, based on acrylamide and/or acrylate offered several advantages over starch. Among these were: significantly lower dosages, higher settling rates, better filtration, improved overflow clarity, and higher underflow density. Within the North American region of the alumina industry, these synthetic flocculants were first used continuously on an industrial scale in Jamaica in the early 1970s. Synthetic

107

flocculants either replaced starch or were used with significantly reduced dosages of starch.

CH 2 OH

CH 2 OH

0-K°»

Over the years significant advancements have been made to synthetic flocculants. Many of these have been described in technical literature and include product form (solution, gel, dry powder, and oil based emulsion), molecular weight, variation in anionicity, structure, and composition. However, the most significant advancement has been the development of hydroxamated polyacrylamide (HXPAM) polymers. Development of these polymers as a flocculant was first announced in 1988 [7], and over the last 20 years they have become the primary flocculant used in red mud settlers around the world.

H

OH

H

,/L< OH

H

OH

Figure 1. Structure of amylose in starch based flocculants

With above as a background, the following sections of this paper explain why the industry moved from starch to polyacrylate/polyacrylamide and then onto hydroxamated polyacrylamides. The argument presented is based on the compositional and structural differences among these polymers and the benefits they offer.

Figure 2. Structure of amylopectin found in starch.

The composition and structure of amylose and amylopectin provide insight into how starch interacts with red mud particles to form floes. The hydroxyl groups extending from the cyclic rings are likely to adsorb on the red mud particle surface. However, because of the sterically hindered structure and the limited charge density, the starch is likely to be poorly adsorbed on the surface. More importantly, the effective molecular weight (< 1 million for amylose) is relatively low for a flocculant, and therefore there is minimal extension of the molecule from the mud particle surface to allow contact with other particles and form floes. Floe or aggregate formation is believed to occur via a a physisorption mechanism, which includes electrostatic, hydrogen and hydrophobic bonding. Although hydrogen bonding can form strong attachment of the starch to the mud particle surface, limited attachment points and molecular extension from the mud surface likely explain why such high dosages of starch are needed to effectively flocculate red mud. Dosages on the order of approximately 16 kg of starch per ton of mud solids have been reported for muds produced from some bauxites [9].

Starch Starch is a set of natural polymeric materials primarily composed of amylose and amylopectin, and belongs to the general class of polysaccharides. It has been commonly derived from sources such as flour, corn, and potatoes. In terms of composition, amylopectin composes the outer shell of the starch grain while amylose is contained within the grain. Before starch can be used as a flocculant, a solution of it must first be made. This is accomplished through a process described as gelatinization whereby the starch grain is heated in water. During this process, the outer shell of the grain swells or expands allowing the amylose to be released. Once in solution, the amylose has flocculating properties. Amylose has an appreciable solubility in water, more so than amylopectin. Literature on the use of starch states that the addition of caustic to the water used to create starch solutions enhanced the flocculation performance. It was conjectured that the caustic was responsible for hydrolyzing phosphoric esters, present in the amylopectin, into amylose and phosphoric acid [2]. Thus, the concentration of amylose in solution increased resulting in enhanced flocculation performance.

Furthermore, this would also seem to explain the significant contribution made by starch towards liquor organics concentration. Starch is very soluble in Bayer process liquor, and because of the relatively weak adsorption mechanisms; a significant portion of the dosage stays in the liquor and reports to the settler overflow. In one study, it was reported that approximately 40-50% of the total organics were due to the use of starch as a flocculant [10].

Structures of both amylose and amylopectin are shown in Figures 1 and 2. Structurally, both of these molecules are composed of joined glucose units. Amylose is described as more linear in its linkage configuration, a more flexible molecule, and to have molecular weights below 1 million daltons [8]. Amylopectin is a more branched molecule, and can reach higher molecular weights on the order of 10-100 million daltons [8].

Despite the drawbacks listed above, starch is still used in a few refineries today. One positive attribute it does have is that due to its structure and size it generally does not blind red mud filter presses that are found immediately downstream of the primary red mud settlers. Furthermore, because of the high amount of residual starch in the Bayer liquor, unflocculated mud that may escape from a settler

108

feedwell can be flocculated in the more quiescent region of the settler before reporting to the overflow. However, despite these features, and a significant economical cost advantage, starch has essentially been replaced by synthetic polymers as the primary flocculant. Synthetic flocculants Synthetic polymers used as flocculants in the Bayer process are derived from acrylamide and/or acrylate monomers. Other noteworthy flocculants have been invented over the years; however the most widely used by the alumina industry have been the aforementioned. The basic structure for polyacrylamide and polyacrylate are shown in Figures 3 and 4, respectively. Furthermore, Figure 5 shows the structure of a copolymer synthesized by copolymerizing monomers of both acrylamide and acrylate. While the characteristics of these polymers have been well described in technical literature [11-15], a common feature to the structures is the flexible linear "backbone" with amide and/or carboxyl groups extending off the backbone. The linear structure enhances the aqueous solubility of the polymer, relative to highly branched polymers. This conformation is also optimal for adsorption of the polymer onto mud particles, thereby improving the flocculation performance of the polymer.

(—CH2—CH—)n C=0 H2 Figure 3. Basic monomer unit of a polyacrylamide.

-^CH2-ÇH^

Oe N a®

Figure 4. Basic monomer unit of polyacrylate. As shown, the sodium exists as a counter ion on the negatively charged carboxyl group. -4-CH 2 —CH4x

fCH2-CH^

C=0

C=0

NH 2

O^Na®

Conventional primary flocculant, x = 5-0%, y = 95-100% Conventional washer flocculant, x = 50-5%, y = 50-95% Figure 5. Copolymer of acrylamide and acrylate. Respective ratios illustrate composition differences used for primary settlers and red mud washer flocculants.

As opposed to natural polymers, such as starch, there is a great deal of flexibility in manufacturing synthetic polymers. It is possible to customize the molecular weight, structure, and the amount and type of charge for a particular application. In terms of flocculation, it is quite possible to create linear polymers with molecular weights on the order of 20 million, much higher than starch or naturally occurring polymers. On the other hand, if a filter aid or dispersant is needed it is possible to reduce the molecular weight to the order of 100,000 or < 100,000, respectively. In terms of charge it is well known that primary red mud settlers require flocculants with a high degree of anionicity, such as found in polyacrylate polymers. However in middle to lower stage red mud washers, moderate (90-70%) to lower (70-50%) charged copolymers provide superior performance. Additionally, synthetic flocculants are made in a variety of product forms. Introduced initially as aqueous solutions, gel logs, or dry powders; these were the only forms until the mid to late 1970s. At this time, the invention of water-in-oil emulsion polymerizations enabled much higher molecular weight polymers to be produced in a much easier-to-use form. While dry powders are still used and offer certain benefits, the vast majority of flocculants used in the alumina industry today are inverse emulsions. While the above information provides the compositional and structural characteristics of synthetic flocculants it does not tell the full story about why they offered such superior performance relative to starch. This is attributable to the high molecular weight and high charge density which affect how, and how much, polymer is adsorbed on mud particle surfaces. In primary red mud settlers polyacrylate is by far the most effective flocculant of the polymers discussed so far. Polyacrylate contains 100% carboxyl groups extending from the backbone of the polymer, and therefore has the highest charge density possible. The repulsive forces between adjacent carboxyl groups help extend, or open up, the conformation of the polymer in solution. This serves to increase the probability that part of the polymer will adsorb to the surface of a mud particle when a collision occurs. Adsorption of polyacrylate and copolymers of acrylamide/acrylate to the mud surface is widely believed to be through a multi-valent ion bridging mechanism whereby multi-valent ions act as a bridge between the carboxyl group on the polymer and the mud particle surface [16]. This type of mechanism would fall under a broad category of chemisorption mechanism. It is conjectured that only part of the molecule is attached to the mud particle surface and that a significant portion of the molecule extends out from the surface. The unattached portion can exist as "tails" or "loops" and serves to attach to other mud particles. Once this happens, a floe or aggregate is formed. The higher molecular weight and linear structure of synthetic flocculants is a distinct advantage in this regard. Larger floes can be formed that have faster settling rates, and provide better mud compaction. Perhaps the most prominent attribute of synthetic flocculants, relative to starch, is the higher settling rates achieved with significantly lower dosage. Settling rates

with synthetic flocculants were reported to be an order of magnitude higher than those achievable with starch; easily exceeding 10 m/hr. Furthermore, effective dosages were reportedly an order of magnitude less with the synthetic flocculants. The higher settling rates produced other benefits, such as increased settler throughput, and increased underflow density. Higher throughput was often an artifact of attaining higher settling rates in large settlers designed with starch as the flocculant. Increased underflow density led to improved soda and alumina recovery in the washer circuit. Interestingly, although synthetic flocculants produced higher settling rates, the floes were more fragile than those made with starch [9, 17]. Floes formed with synthetic flocculants were easily sheared through the underflow pumps so that additional flocculant was required in subsequent washing stages. However, the fragility was also perceived as a positive in that the floes were more compressible and easier to dewater. Much of the "robustness" of starch formed floes was due to small (an artifact of the low molecular weight of starch) but strong floes, and perhaps residual starch in the entrained liquor, which acted to reflocculate the mud after the underflow pump. The benefits of synthetic flocculants listed above far outweighed deficiencies in supernatant clarity, and filtration. While clarity was "good" with polyacrylate or various copolymers, starch was often used in conjunction with these flocculants to achieve better clarity [18]. A general performance characteristic with these flocculants was that as dosage was increased, and consequently settling rate, overflow clarity degraded [19]. This represented an operational issue in that if more flocculant was required to control mud level or increase underflow density, overflow clarity could be compromised. The solution was to supplement the synthetic flocculant dose with some starch to control clarity. One definite drawback to synthetic flocculants that was not an issue with starch was the risk of blinding red mud pressure filters. The extremely high molecular weight of the synthetic polymers would easily blind filters if nonadsorbed, or residual, flocculant escaped the feedwell and reported to the settler overflow. A common indication of this was an immediate decrease in filter throughput, a shiny or glossy appearance of the filter cake, and a "sticky" consistency to the mud on the individual leafs of the press. Again, this difference between starch and synthetic flocculants can be attributed to not only molecular weight, but also structural differences in the molecule. Obviously, the alumina industry has embraced the use of synthetic flocculants in the separation of red mud. This was despite a large "learning curve" associated with determining makeup and application requirements. Something normally anticipated with the introduction of new technology. As reported, the time invested in developing and converting one (of two) Jamaican refineries to synthetic flocculants easily incurred enough savings in raw materials in one year to finance conversion of both plants to synthetic flocculants [9].

Hydroxamated polyacrylamides The latest advancement in synthetic flocculants, that gained widespread use in the alumina industry, occurred in the late 1980s with the introduction of hydroxamated polyacrylamides (HXPAMs) [20]. These polymers, as shown below in Figure 6, are an example of a terpolymer in that there are three different functional groups on the polymer backbone: amide, carboxyl, and hydroxamate. This type of polymer is only available in liquid (i.e. inverse emulsion) form and was made possible by the development of emulsion polymerization technology. It is essentially made by carrying out a post-reaction on polyacrylamide to incorporate the hydroxamate and acrylate groups on the backbone. A post reaction, meaning a separate chemical reaction after the main polymerization reaction to create the backbone, can only be carried out in any practical measure in a liquid or emulsion form. Dry powders do not lend themselves easily to additional chemical reactions, and aqueous solutions are too dilute.

-eC^-CIftr-fCH^H)^— (CH2-ÇH)—

ç=o

NH2

ç=o

0"Na

+

ç=o _

NHO Na+

Figure 6. Hydroxamated polyacrylamide flocculant which contains three separate functional groups: amide, carboxyl, and hydroxamate. These polymers have unique performance characteristics in that they have a positive dosage response in terms of settling rate and clarity. This means that as the dosage of an HXPAM flocculant is increased settling will increase, and clarity will also improve. This is a departure from the performance of traditional synthetic flocculants in that an increase in dosage would produce an increase in settling rate, but at the compromise of clarity. Laboratory data illustrating the different response characteristics is shown in Figure 7. In terms of settler operations, this is of tremendous value. When situations arise where higher settling rates, or higher dosage is needed the operator can make the necessary adjustment without being concerned about degrading the quality of the overflow of the settler. Development of these types of polymers also represent the current era of flocculant research and development. The incorporation of other functional groups into traditional polymer backbones is intentional, by design, to give the polymer improved flocculation capabilities [16]. This could be improved dosage response, or the ability to flocculate selective mineralogy, such as DSP, in red mud [21]. Since the first industrial application [22] of HXPAMs in the early 1990s, these flocculants have slowly become the main flocculant used in primary red mud settlers. Initially, several refineries used these flocculants not only on the red mud settlers, but also on the washer circuit. The reason for this was the exceptional performance, and positive control

formation in the settler, thereby extending the life of the settler [25].

that these flocculants offered relative to conventional polyacrylate and copolymers. Since then, use of HXPAMs in washer circuits has been scaled back significantly, mainly due to economics, and the fact that copolymers deliver better mud compaction. However, a number of plants use HXPAMs on the upper stages (first or second) because of superior clarity (when the overflow goes to precipitation) and reduced auto-precipitation. Occasionally, the use of HXPAMs is extended down the washer train to deal with difficult settling mud.

A

• : t ^

-

B

C

D

E

F

G

H

I

J

K

tlTT

L

M

N

I

&ENETTS Ac I ligner Prwiuctica 8: Improved F$krit»n C: latfrov«i Clarity D: Staxh Elimînaâcn E: Handling Diöicu t Bauxite F: Higher Uff Density G: Lesa Thkkener Scaie H: Una Reduction I: Reduce Na Scttk-s J: Bern* Mud l£vé Control K: Incseased A/C L: Lowei Floe. Costs M: Improved Wasbg- Opciarion Η'. fiter By-hss

Figure 8. Summary of benefits provided by hydroxamated polyacrylamide flocculants on primary red mud settlers. ♦ Hydroxamated polyacrylamide

Conclusions

50

100

150

This paper has summarized the main types of flocculants used by the alumina industry, since the process was first discovered by Bayer. Flocculant technology was shown to have progressed from the introduction of natural polymers, such as starch, to the use of synthetic polymers based on acrylamide and/or acrylate. This lead to modified synthetic polymers such as hydroxamated polyacrylamide, which are used throughout the industry on primary red mud settlers.

200

Floccu Ian t dosage - grams of f loccu lant / ton of mud

Figure 7. Supernatant clarity from lab tests showing the different responses achieved with hydroxamated polyacrylamide and traditional polyacrylate flocculants.

So what does the future likely hold for red mud flocculants? It is unlikely that there will be a change from synthetic flocculants based on acrylamide and/or acrylate in the near future. These types of polymers have clearly demonstrated the ability to provide suitable performance in the separation of red mud for approximately 40 years. Furthermore, it is difficult to compete with the economics of the raw materials used to produce these flocculants, and therefore they should continue to be the primary flocculants in the industry for the foreseeable future.

The unique interaction of HXPAMs with red mud particles is due to the hydroxamate functional group. This group has a very strong affinity for iron bearing minerals and tends to adsorb through a chemisorption (chelation) mechanism. Studies using atomic force microscopy (AFM) have confirmed that surface adsorption of hydroxamated polymers is stronger relative to polyacrylate [23]. Furthermore, the presence of this functional group provides faster adsorption rates for these polymers relative to polyacrylates or starch. This is particularly true in certain liquors with a high organic load. Additionally, it is what also has enabled HXPAMs to flocculate a wide range of red muds, even those considered to be difficult to settle.

Research efforts on flocculants will likely focus on "specialty" or "rationally designed" polymers that have unique capabilities to (selectively) flocculate certain solid phase impurities, such as DSP for example. New flocculants may be utilized as a supplement to the primary acrylamide/acrylate flocculants outlined in this paper, and be used on an occasional basis. Efforts to develop this type of flocculant technology would seem to fit well with the direction that the industry is moving; processing bauxite reserves with deteriorating quality. The ability to separate and remove mud produced from these bauxites, and maintain industry specifications for alumina quality, may rely heavily on the use of processing aids, such as new flocculants.

As described earlier in this paper, the chief benefit of starch was the ability to settle red mud and provide a significant reduction in mud load to the filters. Synthetic flocculants brought significant increases in settling rate, settler throughput, flocculant dosage, and underflow density. Performance improvements associated with the use of HXPAMs have also been reported [24]. These are summarized in Figure 8 and include: settling of difficult mud, increased settler throughput/production, improved filtration, improved overflow clarity, and elimination of starch. Many of these improved characteristics are due to the stronger, faster polymer adsorption on mud particle surfaces. This leads to better scavenging of fine mud particles, and less residual polymer in the liquor that can blind filters. Lower suspended solids and a more distinct mud interface reduces seed sites for gibbsite scale

Ill

References 1. 2.

3. 4. 5. 6.

7.

8. 9. 10.

11.

12.

13.

14.

15.

16.

Fathi Habashi, "A Hundred Years of the Bayer Process for Alumina Production", Light Metals (1988), pp. 3-11. George R. Gardner and Kenneth B. Ray, "Flocculation and Clarification of Slimes with Organic Flocculants", AIME Transactions, (1*939), pp. 146-168. L. Keith Hudson, "Evolution of Bayer Process Practice in the United States", Light Metals (1988), pp. 31-36. Donald A. Dahlstrom, "Liquid-Solid Separation Challenges in the Fast Lane", Challenges in Mineral Processing, (1989) SME, pp. 467-475. Merrill F. McCart and Robert S. Olson, "Polyacrylamides for the Mining Industry", Mining Engineering, January (1959), pp. 61-65. William F. Linke and Robert B. Booth, "Physical Chemical Aspects of Flocculation by Polymers", AIME Transactions, Vol. 217, March 12, (1959), pp. 364-371. D.P. Spitzer, A.S. Rothenberg, H.I. Heitner, and M.E. Lewellyn, "New polymers for the Bayer process", Alumina Quality Workshop, Gladstone, QLD, September (1988), pp. 221-230. Marguerite Rinaudo, "Polysaccharides", KirkOthmer EncycL Chem. TechnoL, 5th Ed., J. Wiley & Sons, New York (2006), 20, pp. 549-586. J.L. Chandler, "Advances in the use of synthetic flocculants", Light Metals, (1976), Vol. 2, pp. 163-171. M.J. Pearse and Z. Sartowski, "Application of special chemicals (Flocculants and Dewatering Aids) for red mud separation and hydrate filtration", Bauxite - Chapter 38, (1984), pp. 788810. L. J. Connelly, D. O. Owen, and P. F. Richardson, "Synthetic flocculant technology in the Bayer process", Light Metals, Vol. 2, (1986), pp. 61-68. M. J. Pearse, "Synthetic flocculants in the mineral processing industry- types available, their uses and advantages", Light Metals, (1984), pp. 101-106. S. Weir and G. M. Moody, "Trends in the development of flocculants as aids to solid/liquid separation", Journal of The Filtration Society, Vol. 1(4), (2001), pp. 11-12. M. J. Pearse, S. Weir, S. J. Adkins, and G. M. Moody, "Advances in mineral flocculation", Minerals Engineering, Vol. 14, No. 11, (2001), pp. 1505-1511. Richard E. Ellwanger, "Use and application of synthetic organic flocculants in the mining industry", AIME Transactions, Vol. 270, (1980), pp. 1812-1815. H. T. Chen, S. A. Ravishankar, and R. S. Farinato, "Rational polymer design for solidliquid separations in mineral processing applications", International Journal of Mineral Processing, Vol. 72, (2003), pp. 75-86.

17. Ronald J. Allain, "Water soluble polymers - the future?", Reagents for a better metallurgy, Chapter 26, SME, (1994), pp. 263-267. 18. T. K. Hunter, G. M. Moody, S. E. Sankey, and C. A. Tran, "Advances with chemical additives for the alumina industry", Light Metals, (1991), pp. 159-165. 19. David O. Owen, Lawrence J. Connelly, and Peter A. Dimas, "Evaluation of downstream effects of specialty chemicals in the bayer process", Light Metals, (1991), pp. 173-176. 20. D. P. Spitzer, A.S. Rothenberg, H.I. Heitner, M.E. Lewellyn, L.H. Laviolette, T. Foster, and P.V. Avotins, "Development of new Bayer process flocculants", Light Metals, (1991), pp. 167-171. 21. M. Davis, Q. Dai, H.-L. Chen, and M. Taylor, "New polymers for improved flocculation of high DSP-containing muds", Light Metals, (2010), pp. 57-61. 22. Maritza Faneitte, Alfredo Galarraga, and Terry Foster, "Utilization of new polymer in Interalumina", Light Metals, (1994), pp. 129-131. K. E. Bremmell and P. J. Scales, "Adhesive forces between adsorbed anionic polyelectrolyte layers in high ionic strength solutions", Colloids and Surfaces A: Physicochemical and Engineering Aspects, Vol. 247, (2004), pp. 19-25. Roderick G. Ryles and Peter V. Avotins, "Superflue HX, a new technology for the alumina industry", 4th International Alumina Quality Workshop, Darwin, Northern Territory, Australia, June (1996), pp. 205-215. Peter V. Avotins, Larry L. Laviolette, Emiliano F. Repetto, and Araujo M. Eli, "The effects of flocculants on thickener scaling", 5th International Alumina Quality Workshop, Bunbury, WA, March (1999), pp. 448-455. © 2010 Cytec Industries Inc. All rights reserved.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

REDUCTIVE SMELTING OF GREEK BAUXITE RESIDUES FOR IRON PRODUCTION A. Xenidis1, C. Zografidis2,1. Kotsis1, D. Boufounos3 national Technical University of Athens, School of Mining and Metallurgical Engineering, Greece 2 LARCO General Mining & Metallurgical Company; Larymna Lokridos; Fthiotida, GR350 12, Greece 3 Aluminium of Greece SA; Agios Nikolaos, Viotia, GR 32003, Greece Keywords: bauxite, residue, smelting, reduction, iron have been reported [4, 5, 6]. The most common reducing agent utilized is coke while the iron recovery finally obtained is always over 90%.

Abstract The reductive smelting of Greek bauxite residues was investigated for the production of an iron product that meets some crucial industrial requirements as a blast furnace feed. Fine-grained Greek bauxite residue - either as is or in the form of pellets - and solid fuel reducing agents -lignite and coke-, were used as raw materials. The effect of parameters such as the smelting temperature, the amount of the reducing agent in the mixture, the retention time and the addition of fluxes on the quality of the metallic product, as well as the basicity and the desulfurization capability of the slag, was investigated. The results obtained regarding the chemical properties of the metallic product were very promising, providing input for further research on the optimization of the proposed pyrometallurgical method for the production of an attractive ferrous raw material for the iron ore industry.

Experimental The sample of bauxite residue utilized in the present study was provided by the Greek alumina refinery owned by Aluminum of Greece S.A. It is a fine grained material, given that a percentage of 40% by weight is less than 2 urn. Lignite and coke were used as reducing agents. Chemical analysis of the raw materials is given in Table I. The total iron (Fetot) content in the bauxite residue is about 33.5%, which is indicative of the fact that it constitutes an important potential iron source for the iron ore industry. Its main mineral constituents, as determined by mineralogical analysis conducted by X-Ray diffraction, were hematite, gibbsite, diaspore, calcite and hydroxysodalite. Table I. Chemical Analysis of Greek Bauxite Residue and Reducing Agents Lignite Coke Bauxite Component residue

Introduction The application of various processing methods for the effective recovery of major metallic constituents of bauxite residues, has been a matter of scientific interest by numerous researchers [1]. The current work constitutes a part of an integrated research effort which aims to investigate alternative ways of utilizing bauxite residue as a raw material in the iron ore industry, by applying pyrometallurgical treatment methods. Within this framework, roasting reduction, magnetic separation for iron enrichment and pelletization of the Greek bauxite residue have already been investigated [2, 3]. The present study is focused on the reductive smelting of the same raw material for iron production.

Fe 2 0 3 Na20 Si0 2 CaO MgO A1203 K20 Ti0 2 C0 2 S0 2 L.O.I.

The crucial technical requirements that determine the usability of an iron-bearing raw material for the iron ore industry -either as a feed for a rotary kiln or a blast furnace-, are the following: iron content, "gangue constituents" (Si02, CaC03, MgO, A1203), "deleterious impurities" (S, P, Na 2 0, K20 and Ti02), physical properties (porosity, durability, ore structure), reducibility and its amenability to concentration. Given that about 1.61 tonnes of iron ore with an iron content of 60% are required for the production of each tonne of cast iron, and also considering the fact that iron ore prices have almost quadrupled since 2004, it is easy to understand how crucial it is for the industry to make use of an iron bearing raw material with a lower iron content (20-30% Fe) which can substitute part of the iron ore feed. The iron content of bauxite residue, mainly in the form of iron (III) oxide, ranges from 30 to 50%, while at the same time the typical content of gangue material -such as CaO and A1203 - is such that this does not seriously affect its melting point. Thus, much attention has been focussed towards iron recovery in these residues.

C fix

Ash Volatile Matter

(%)

48 3.26 6.96 14.84 0.24 15.85 0.07 7.06 2.24 0.78 11.37

(%)

(%)

4.44 2.83 53.81 4.91 3.39 19.11 0.69 0.79 3.78

29.09 2.67

30.1 20.1 49.75

81.7 5.0 13.3

0.5 2.9

Moreover, a single experiment was conducted regarding smelting reduction of bauxite residue in form of pellets with lignite (bauxite residue/lignite: 1/3 by weight and bentonite: 0.6% by weight as a binding agent). The melting temperature of the bauxite residue was initially determined with a LECO AF 600 apparatus, by employing the Seger cones technique, based on the determination of four characteristic temperature values: the start of softening, end of softening, start of melting and end of melting points. The melting point of the bauxite residue was determined to be 1400°C, as a mean of five values. The determined value is in agreement with the theoretically calculated value from the ternary phase diagram of FeO - CaO - A1203 (the Fe203 content of the bauxite residue is stoichiometrically converted to FeO content).

Reduction smelting constitutes the most significant pyrometallurgical process for the production of cast iron. It is mainly conducted in blast furnaces or electrical furnaces. Many efforts in the smelting reduction of bauxite residue material

The smelting reduction experiments were conducted in two types of laboratory furnaces:

113

The sample fed in both furnaces was 70-80 g. It is noted that two types of furnaces were used for the conduction of smelting reduction of experiments, due to the fact that each one has several advantages and disadvantages. It is noted that the determination of temperature was more precise in the experiments conducted in the Tamman furnace, however, the duration of each experiment was more than 4 hours in Tamman and only half an hour in the induction smelting furnace. Nevertheless, due to the intense eddying of the material resulting from the Foucault current, there was loss of fine material and therefore, it was difficult to conduct mass balance calculations.

i) Electric resistance heating furnace Tamman (Figure 1) (60 KVA, Tmax: 2500°C): its main part is a cylindrical heating tube made of carbon material, surrounded by a thermal insulating material, so that the inside of the heating tube acts as a heat treatment chamber, in which the graphite sample holder is placed. The temperature was recorded by a Raytek optical sensor pyrometer (laser). ii) Induction smelting furnace (Figure 2): This consists of an induction heating system, a water cooled induction coil, a supporting body having a suitable circumference for supporting a graphite crucible. The temperature during the experimental procedure was measured by the same Raytek optical sensor pyrometer.

1. Crucible 2. Heating tube 3. Graphite protecting tube 4. Silimanite mantle 5. Amianthus 6. Water connection 7. Cooling water connecting tube 8. Grains offillermaterial 9. Graphite grains 10. Graphite grains 11. Cooling water 12. Steel mantle

Figure 1. Electric resistance heating furnace Tamman

Figure 2. Heating chamber of the induction smelting furnace

In Figure 4 the effect of the type of the reducing agent and reaction time on carbon content in the pig iron is given. It is seen that increase of the retention time of the bauxite residue melt results in a significant decrease of the carbon content. Moreover, after 15 minutes of smelting reduction, the carbon content of the metallic product fluctuates between 4.6-5.1%, while after 45 minutes the respective values are 2.04-3.01, in compliance with the requirements of commercial cast iron. The use of a solid fuel like coke, which is much more reactive at higher temperatures (>1000°C) favors the reduction, contrary to lignite, which is much more reactive and evolves its thermal energy at lower temperatures.

Results and Discussion The effect of the following parameters on the final result of smelting reduction was investigated: temperature, retention time of the smelted material, bauxite residue/lignite ratio and the addition of fluxes. The determination of the bauxite residue/lignite ratio was based on the stoichiometrically required quantity of carbon required for the complete reduction of Fe203 to metallic iron. Moreover, the addition of fluxes (CaO and Si0 2 reagents), was based on the theoretical determination of the desirable lower melting point, according to the ternary phase diagram of Al203-CaO-Si02, taking into consideration that the aforementioned are the three basic constituents of the bauxite residue slag. Thus, the addition of three different ratios of CaO and Si0 2 was investigated per 100 g of the mixture bauxite residue/lignite (4/1 by weight): i)Si02-CaO: 20-0 g, ii)Si02-CaO: 35-5 g, iii)Si02-CaO: 62-38 g.

1800 1750

|;..;,;; ..:,;;..:Ί

1700

u 1650 e g, 1600 oo

The experimental conditions of reduction smelting as well as the chemical analysis of the metallic products, are presented in Table Π. The most important conclusion deduced is that regardless of the conditions employed, a metallic product with iron and carbon content ranging from 91.3-95.8% and 2.4-5.4 % respectively, was produced. The quality of iron produced can be characterized as satisfactory, since according to a typical analysis of cast iron, the iron content is approximately 94% and carbon content fluctuates between 3.5-4.5%. In Figure 3, the effect of the ratio of bauxite residue/lignite on mixture melting point is presented. It is noted that increase of the aforementioned ratio results in a considerable increase of the melting temperature.

I 1550

I

Ì::-:::'::: ;·■>;:; ■>::;■:■)

1500 1450 1400

4:1 3:1 1.5:1 Bauxite residue/lignite ratio (wt/wt)

Figure 3. Effect of the ratio bauxite residue/lignite on the melting point

114

Table IL Experimental Conditions of Reduction Smelting Experiments and Chemical Analysis of the Metallic Product (%) Content in the Metal | Fluxes BR/ Retention Smelting Raw Reducing Reducing Time Rest CaO Si0 2 s/n Material Agent Agent Fe S C (min) Admixtures (g) (g) Ratio Pellets B.R. * 1 B.R. 2 B.R. 3 B.R. 4 B.R. 5 B.R. 6 B.R. 7 B.R. 8 B.R. 9 B.R. 10 B.R. 11 * BR: Bauxite Residue

0

Lignite Lignite Coke Lignite Lignite Lignite Lignite Coke Coke Lignite Lignite

4/1 4/1 11/1 4/1 4/1 4/1 4/1 11/1 11/1 3/1 1.5/1

35

5

20 62

38

15 15 15 15 15 30 45 30 45 15 15

94.8 91.3 94.2 91.7 91.9 94.5 95.3 95.0 95.8 95.6 95.8

0.1 0.7 0.2 0.4 0.4 0.3 0.2 0.1 0.5 0.3 0.4

3.4 5.4 4.6 5.3 5.2 3.5 3.0 3.2 2.4 2.7 2.8

1.7 2.6 1.0 2.6 2.5 1.7 1.5 1.6 1.2 1.3 1.0

1450°C. On the contrary, the melting temperature value measured the same way of the mixture of bauxite residueAignite (4/1 by weight) was 1560°C. The aforementioned values are in agreement with the theoretically calculated melting temperature values from the ternary phase diagrams in Figures 6 and 7. 4.5

60 40 Time (min) Figure 4. Effect of the reducing agent and retention time on carbon content of the metallic product (reducing agent in 50% excess of the stoichiometrically required quantity) 20

1

2

4

5

6 7 8 9 10 Smelting S/N Figure 5. Basicity index of the slag samples produced Moreover, the slag produced by the addition of 35 g Si0 2 and 5 g CaO (Point B) per 100 g of mixture (bauxite residue/lignite), is in the eutectic region of the slag produced by the blast furnace operation.

The addition of fluxes results in the alteration of the basicity index of the slag. The basicity index is determined within the framework of the current study as the ratio (bases/acids). The oxides FeO, MgO and CaO are considered as bases while Si0 2 is considered as acid. A1203 is considered as an amphoteric oxide, which means that an A1203 (%) content higher than 18% is considered as basic, and lower than 15% is considered as acidic. On the contrary, for A1203 (%) content 15-18, it does not participate in the slag basicity calculations.

The smelting reduction of pellets of Greek bauxite residue with lignite (diameter of pellet: -9.5+6.3 mm, bauxite residue /lignite: 4/1 by weight) resulted in the production of a metallic product of good quality (94.8% iron and 3.1% carbon), something which is ascribed to the better contact of the bauxite residue with the solid reducing agent, favoring in such a way, the iron oxide reduction. This constitutes an essential conclusion taking into consideration that pelletization is a very common agglomeration process used in the iron ore industry.

More precisely, the addition of fluxes results in the alteration of the basicity index of the slag from 4.5 to 0.6, at an addition rate of 35 g Si0 2 and 5 g CaO per 100 g of bauxite residue/lignite mixture (Figure 5). Due to the reduction of slag basicity, the (%) sulfur content in the metallic product increases to approximately 0.7%. It is noted that the sulfur content of a typical cast iron should not exceed 0.04% and this constitutes a crucial parameter of its quality.

A metallographical section of grains of the metallic product as well as the slag samples from the smelting reduction of bauxite residue with lignite (4/1 by weight) with the addition of 35 g Si0 2 and 5 g CaO per 100 g of the mixture, was studied by SEM/EDS (JEOL® JSM-6380LV). Back-scattered electron image of the aforementioned samples, are presented in Figures 8 and 9.

The addition of fluxes in the current work can be characterized as successful, since the melting temperature value measured by the optical pyrometer during the three experiments conducted with the addition of CaO and Si0 2 , fluctuated between 1400-

115

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**l / «£»*>

*g$#*

Ce02Ä%5, WLXfr

CaOt

Jfe

Figure 6. Melting temperature value of the slag sample produced by the mixture bauxite residue /lignite (4/1 by weight)

~ * W

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CaO mfa

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Figure 7: Melting temperature values of the slag samples (A,B,Q produced by the addition of: A) Si02-CaO: 20-0 g, B) Si02-CaO: 35-5 g, C) Si02-CaO: 62-38 g per 100 g of the mixture bauxite residue /lignite (4/1 by weight)..

116

Figure 8 shows a typical feature, where the metal matrix and the lamellar structure of iron carbide flakes can be seen. In Figure 9 a typical matrix of an iron ore slag is seen, where the black area corresponds to the slag oxides as determined by the EDS analysis. We have also focused on metal grain 'trapped' in the slag (white area), which is probably ascribed to the intense eddying of the molten bath due to the Foucault current, resulting in non satisfactory phase segregation (metallic product - slag).

reduction smelting with solid reducing agents resulted in the production of a metallic product with an iron content fluctuating between 91-95.8%, which can be characterized as satisfactory considering that a typical cast iron has an approximate analysis of 91-96 % in iron. The addition of solid reducing agents results in a significant increase of the melting point of the bauxite residue, from 1400°C up to 1740°C for a great excess of the reducing agent, due to the formation of the refractory iron carbides. The addition of CaO and Si0 2 reagents as fluxes favored the decrease of the melting point, in agreement with the respective ternary phase diagrams. The type of the solid reducing agent used, proved to be a critical parameter affecting the quality of the metallic product, in terms of the desired sulfur content, the basicity of the slag and the kinetics of the reductive procedure. The results render the future investigation for optimization of the proposed processing method very promising. Apart from optimization of the iron recovery, intensification of the future research is necessary concerning the extraction of impurities, such as S, P and Ti from the metallic product. Acknowledgements The authors would like to acknowledge the financial support of the General Secretariat for Research and Technology of Greece, Programme PAVET 2005, No. 05PAB102. References

Figure 8. Back-scattered electron image of the metallic product from the smelting reduction of Greek bauxite residue with lignite.

1. R. K. Paramguru, P. C. Rath, and V. N. Misra, "Trends in Red Mud Utilization - A Review" Mineral Processing and Extractive Metallurgy, 26 (2005), 1-29. 2. A. Xenidis et al., "Reductive Roasting and Magnetic Separation of Greek Bauxite Residue for its Utilization in Iron Ore Industry" (Paper presented at the Light Metals as held at the 138th TMS Annual Meeting, San Francisco, California, 2009), 63-67. 3. A. Xenidis et al., "Pelletization and Reductive Smelting of Greek Bauxite Residues for Iron Production" (Paper presented at the 3rd International Conference AMIREG: Assessing the Footprint of Resource Utilization and Hazardous Waste Management, Athens, September 2009), 433-439. 4. L. Visnyovszky, "Complex utilization of Hungarian bauxites", Femip. Kut. Intez. Kozlem, 9 (1971). 5. N. I. Eremin, "Complex processing of bauxites" (Paper presented at the Proceedings of the 2nd International Symposium ICSOBA Bauxites, Alumina, Aluminium, Leningrad, 1969).

Figure 9. Back-scattered electron image of the slag phase from the smelting reduction of Greek bauxite residue with lignite. Conclusions Smelting reduction of Greek bauxite residue revealed that it can be used successfully for direct cast iron production. The

6. A. I. Zazubin et al. "Complex processing of red mud", Tr. Inst. Met. Obogashch., Akad. Nauk. Kaz. SSR, (1967), 25.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINA and BAUXITE

Precipitation, Calcination and Properties SESSION CHAIR

Hans-Werner Schimdt Outotec GmbH Oberursel, Germany

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECT OF TECHNOLOGICAL PARAMETERS ON PSD OF ALUMINUM TRI-HYDROXIDE FROM SEED PRECIPTATION IN SEEDED SODIUM ALUMINATE SOLUTIONS Yusheng Wu, Mingchun Li Yanping Qu School of Materials Science and Engineering, Shenyang University of Technology, Liaoning 110178, P. R. China Keywords: Bayer process, Sodium aluminate, Particle size distribution, Aluminum tri-hydroxide extremely slow growth rate. Agglomeration consumes smaller particles and creates larger particles, representing the additive volume of the smaller particles consumed. Nucleation is the formation of new, minute particles which replenish the particle count of the system. Particles are continuously lost through agglomeration or product removal. Thus for efficient operation, nucleation must provide enough particles to maintain the total particle count, and to provide surface area for continuing growth. Particle size distribution of products mainly depends on all of the above four precipitation steps.

Abstract Periodic attenuation of particles, which interferes seriously with normal alumina production, is a characteristic of the Bayer process. In order to construct a mathematical model of total number of particles and particle size distribution (PSD) which is helpful to predict and control the PSD of aluminum tri-hydroxide for an alumina refinery, the PSD of aluminum tri-hydroxide from the seeded precipitation process with different technological parameters, has been investigated under industrial conditions modeled in the laboratory. The results show a move of PSD to smaller particle sizes with decreasing initial precipitation temperature and increasing molar ratio of Na20/Al203. The volume fraction of particles below 44um first increased and then decreased with increasing initial concentration of Na20. The PSD has the opposite variation tendency with different initial seed amounts, compared with the initial concentration of Na20.

Most alumina plants nowadays have already developed, or are in the process of developing, their own precipitation model for the simulation of the aluminium hydroxide precipitation mechanisms. In order to construct a mathematical model of total number of particles and particle size distribution (PSD) which is helpful to predict and control the PSD of aluminum tri-hydroxide for alumina refinery, the PSD of aluminum tri-hydroxide from a seeded precipitation process under different technological parameters has been investigated under industrial conditions, modeled in the laboratory.

Introduction The most common industrial process for alumina production from bauxite is the Bayer process. Seed precipitation is the key stage, significantly affecting the yield and specification of product [1-3]. Lots of work has been done on the precipitation process for enhancing precipitation ratio in the sodium aluminate liquor, and improving product specifications [5-8]. However, the periodic attenuation of grain size, that mass percent of particle size above 44um and attenuating periodic change from 50% to 90% and 3 to 5 months respectively, inherently exists in the Bayer process in China. During attenuation, the functioning of filters deteriorates and product quality drops. Meanwhile, the electrical consumption and the dust density in the exit of electrofilter increase [9-12]. Although some work has been done to solve the problem, such as using sandy alumina production technology invented by France and investigating the particle size distribution (PSD) of production alumina in some branches of China Aluminum Co., Ltd, the explosive attenuation of alumina particles can't be avoided.

Experimental The main apparatus for seed precipitation study included a bladepaddle mixer tank and a Model LB-801 super constant temperature bath. The supersaturated sodium aluminate solutions, with industrial concentration and initial gibbsite seed for all experiments, were provided by the Shandong branch of China Aluminum Co., Ltd. Precipitation time is 40 h. After adding seed, the precipitation commences by stirring at 260 rpm and the temperature decreased uniformly over the desired ranges. Solution samples (precipitate suspension) were analysed by acidbase titration and complexometric titration methods. The precipitated aluminum tri-hydroxide samples obtained were washed with hot deionised water and dried at 60° C for 24h before observations with Scanning Electron Microscopy (S-3400N). The PSD and average size were measured by laser diffraction (Malvern Mastersizer 2000).

In general, the precipitation of aluminium hydroxide from seeded sodium aluminate solutions (SAS) is comprised of crystal growth, nucleation, agglomeration, and attrition [13-15]. By the mechanism of crystal growth, dissolved alumina hydrate transfers from the aqueous phase to the solid phase, depositing on the surface of existing particles. Thus the growth rate determines the overall liquor/solid mass balance. Growth also impacts the particle population balance because particles will increase in diameter. The crystal growth rate in gibbsite precipitation has been shown to depend on three factors: the supersaturation of sodium aluminate in solution, the temperature, and the solid surface available for reaction. Agglomeration is essential in gibbsite precipitation because it is the only economically feasible method of producing particles with the required size, given the

Results and discussion Effect of technological parameters on particle size distribution In our study, the question was to understand how all the experimental parameters usually considered in the Bayer process influenced PSD. In order to focus our attention on the effect of other parameters, only one parameter was varied in the experiments. The PSD of the products with different parameters are displayed in Fig. 3. It is found that all particle size distribution curves have a similar shape, but show different trends.

121

Figure 3(a) shows that PSD moves to smaller particle sizes with decreasing initial precipitation temperature. Fig. 3(b) shows that 100 0s-

a

PSD moves to coarser particle sizes with decreasing initial molar ratio Na20/Al203. The results we obtained are in agreement with 100

80 60 40

o > 20

0 50 100 Particle size, μπι

50 100 Particle size, μΐΏ

150

50

100

150

50 100 Particle size, μηι

150

Particle size, μπι

150

Fig. 1. Variety of Particle size with (a) temperature; (b) molecule ratio of Na20/Al203; (c) concentration of Na 2 0; (d) fine seed content. three fine seed concentration loading tests have the same degree of supersaturation. That is, have the same driving force for seeded precipitation. Therefore, the secondary nucleation takes place with the fine seed content at 10%, and the agglomeration degree decreases with the fine seed content at 30%.

of those of Yamada [15]. Yamada indicates that the agglomeration of Al(OH)3 particles, for similar supersaturations, is more liable to take place at higher temperatures. He supposes that, at high temperature, the dispersion of the growth species across the particle surface, is faster, and provides a higher probability for the successful cementation of the particles. At the same temperatures, higher initial supersaturations also favour agglomeration, due to the higher concentration of adsorbed growth species. We have worked with fine seeds which are more liable to participate in the agglomeration process. At the same concentration of Na20, as the supersaturation of sodium aluminate solution increases, this favours agglomeration as the initial molar ratio of Na20/Al203 decreases.

Table 1 shows the effects of technological parameters on crystal growth. It is found that when the initial concentration of Na20 was varied from 140g/L to 160g/L. The volume percentage of product particles below 44μπι is increased by 6.05% and then decreases by 3.88%. The surface area is increased by 0.015 m2/g and then decreased by 0.009 m2/g. However, the average volume size is decreased by 3.77μπι and then increased by 1.828μηι. Compared with the initial concentration of Na20, the volume percentage of particles below 44μιη, surface area, and the volume average size have opposite variation tendencies with different initial fine seed content.

Fig. 3(a) shows that the PSD moves first in the direction of coarser and then to finer particle sizes with increasing initial concentration of Na20. Compared with the initial concentration of Na20, the PSD has the opposite tendency with different initial fine seed content. In a seeded precipitation process, the fine particles need more aluminium tri-hydroxide deposition from SAS to transform into coarse particles. However, the seed surface area must respond to the decreasing degree of supersaturation. If not, this is harmful to the agglomeration process. In the tests, the supersaturation of SAS favours agglomeration with Na20 at 140g/L and that can't match agglomeration with Na20 at 160g/L. However, with further increase in concentration, the precipitation solution has a higher viscosity, which increases the chance of collision between fine particles, thereby increasing the agglomeration efficiency and reducing fine particle levels. The sodium aluminate liquors in

The volume percentage of particles below 44um and the surface area, decreased with increasing initial temperature. This varied from 32.17% to 12.41% and 0.148 m2/g to 0.107 m2/g respectively. That is, higher precipitation temperatures favour particle size coarsening. For example, the average volume increased by 11.51 lpm with initial precipitation temperature increase from 60°C to 75°C. The product characteristics therefore show the opposite variation tendency with different o^ compared with initial precipitation temperature.

122

140g/L

12.03

78.171

surface area /m2/g 0.104

150 g/L

18.08

75.401

0.119

160 g/L

14.20

77.229

0.110

1.35

13.65

78.047

0.111

1.45

15.72

76.223

0.115

Technological parameter Nk

Table 1. Characteristics of product volume fraction volume average of<45um/% size /urn

ak

1.55

18.27

75.740

0.120

60°C

32.17

66.890

0.148

65°C

20.92

74.453

0.120

70°C

15.54

77.016

0.110

75°C

12.41

78.401

0.107

10%

23.24

64.683

0.111

20%

18.38

74.611

0.115

30%

29.65

62.130

0.120

Fine

Note: Nk - concentration of Na20; ak - molecule ratio of Na20/Al203; T- temperature; Fine - fine seed content. 8.1. Seyssiecq, S. Veesler, D. Mangin, J. P. Klein, and R. Boistelle, "Modeling gibbsite agglomeration in a constant supersaturation crystallizer," Chemical Engineering Science, 55(2000) , 5565-5578. 9. Y. S. WU, D. Zhang, M. C. Li, S. W. Bi, and Y. H. YANG, "Periodical attenuation of Al(OH)3 particles from seed precipitation in seeded sodium aluminate solution," Transactions of Nonferrous Metals Society of China, 20(2010), 528-532 10. C. B. Zhang, and P. Zhao, "Studied on polarizing particle size of aluminate hydroxide in Bayer process," Light Metals, (1999), 17-19. (In Chinese) 11. G. Zhang, and J. N. Yang, "Studied on particle size fluctuating of aluminate hydroxide in Bayer process," (2002), 9-12. (In Chinese) 12. B. Garner, B. Cristol, and A. Soirat, "Precipitation particle size control," Light Metals, (1999), 71-76. 13. B.K. Satapathy, and T. Padhi, "Determination of grain-size distribution of sandy alumina using electron sensing zone method," Light Metals, (1990), 185-191. 14. T. K. Hunter, G. M. Moody, and S. E. Sankey, "Advances with chemical additives for the alumina industry," Light Metals, ( 1991 ), 159-162. 15. K. Yamada, "Nucleation and agglomeration during crystallization of aluminium trihydroxide in sodium aluminate solution," Light Metals, 32(1980), 720-726.

Conclusions A systematic study has been made on the PSD of aluminum trihydroxide from a seeded precipitation process with different parameters. The PSD moved to smaller particle sizes with decreasing initial precipitation temperature, and increasing molar ratio of Na20/Al203 . The volume fraction of particles below 44um first increased, and then decreased with increasing initial concentration of Na20. The PSD has the opposite tendency with different initial seed amount compared with the initial concentration of Na20. Acknowledgments The authors are grateful to appreciate the financial support of National Natural Science Foundation of China (NO. 50804031) and the educational department of Liaoning Province of China (NO.L2010395). References 1. R. Chester, F. Jones, M. Loan, A. Oliveira, and W. R. Richmond, "The dissolution behavior of titanium oxide phases in synthetic Bayer liquors at 90 °C," Hydrometallurgy, 96(20093), 215-222. 2. H. Watling, J. Loh, and H. Gatter, "Gibbsite crystallization inhibition 1. Effects of sodium gluconate on nucleation, agglomeration and growth," Hydrometallurgy, 55(2000), 275-288. 3. F. Farhadi, and M. B. Babaheidary, "Mechanism and estimation of Al(OH)3 crystal growth," Journal of Crystal Growth, 234(2002), 721-730. 4. N. Brown, "A quantitative study of new crystal formation in seeded caustic aluminate solution," Journal of Crystal Growth, 29(1975), 309-315. 5. Z. Wang, S. W. Bi, Y. H. Yang, and Z. F. Yuan, "Evolution of particle size and strength of hydragillite from carbonization in seeded sodium aluminate liquors," Journal of Crystal Growth, 274(2005), 218-225. 6. A. M. Paulaime, I. Seyssiecq, and S. Veesler, "The influence of organic additives on the crystallization and agglomeration of gibbsite," Powder Technology, 130(2003), 345-351 7. H. LI, J. Addai-Mensah, J. C Thomas, and A. R. Gerson, "The crystallization mechanism of Al(OH)3 from sodium aluminate solutions," Journal of Crystal Growth, 279(2005), 508-520.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Methods to Reduce Operating Costs in Circulating Fluidized Bed Calcination Cornells Klett, Bernd Reeb, Michael Missalla, Hans-Werner Schmidt Outotec GmbH, Ludwig-Erhard-Str. 21, 61440 Oberursel, Germany Keywords: Circulating Fluidized Bed, Calcination, Energy Efficiency, Product Quality, Particle Breakage, Dry Hydrate the free moisture of the hydrate, with typical values of 6 to 8% moisture, is evaporated and the hydrate is dried.

Abstract The calcination of gibbsite or hydrate to alumina is one of the most energy consuming process steps in every alumina refinery based on the Bayer process. Approximately 30% of the thermal energy input is used for the calcination process step.

The dried hydrate is then separated from the gas in an electrostatic precipitator and fed to the second preheating stage, where it is mixed with the 950°C hot off-gas of the furnace. Here a significant step of the reaction is already taking place with approx. two thirds of the hydrate water released. The solids are separated again from the gas and transported to the CFB furnace.

Over the years new technologies such as Circulating Fluidized Bed (CFB) Calciners by Outotec (formerly Lurgi) have reduced the energy consumption for the calcination step significantly (see also recent Energy Efficiency Award, given in 2010 by the German Energy Agency). The CFB technology was introduced as early as 1961. Up till then rotary kilns were the standard technology for the calcination of alumina. This innovation reduced the consumption of fuel by up to 30%. Since then, CFB calciners have been constantly improved, and methods have been developed to reduce fuel consumption even further.

Finally under the high solids density and homogeneous mixing and temperature field in the CFB furnace, the precalcined material is calcined to the specified alumina quality. The good mixing and high solids density, with a sufficient residence time of several minutes, allows very homogeneous calcination of the material and the optimum product quality, with very low alpha alumina and no residual gibbsite, as well as low particle breakage.

Not all, but some, of the methods to reduce fuel consumption have resulted in increasing process complexity and operator and maintenance demand. However, a number of measures have also been introduced to mitigate the negative effects of the increased process complexity and to even improve on operability and maintainability. In this paper the different options and methods introduced for reduction of fuel consumption, but also to increase of operability and maintainability are compared and evaluated with regard to their effects on installation and operating costs.

Cooüng Sfege U

Cooling Wafer

Introduction Calcination of hydrate to alumina is the last step in the production of smelter grade alumina from bauxite in an alumina refinery. The energy consumption for the production of alumina varies between 6.5 and 18 GJ/t alumina for the whole refinery. Depending on the technology, approx. 30% of this is consumed in the calcination step.

Additional Atr Secondary Air Primary Air Cooling Wafer

Figure 1: CFB Calciner Flowsheet of recently installed calcination plants

Until stationary calciners, such as flash and CFB calciners were introduced, rotary kilns were used. Rotary kilns have a thermal energy consumption of around 4.5 to 5.5 GJ/t alumina based on lower calorific heating value. With the introduction of stationary calciners and the first CFB calciner by Outotec (at that time known as Lurgi) in 1961, the consumption dropped instantly to approx. 3.3 GJ/t. Thus operating costs dropped significantly, which gave a huge incentive to invest in this new technology.

The final product is discharged from the CFB calcining stage and then cooled down in two direct gas solids contact cooling stages and one indirect cooling stage, which is the fluid bed cooler. In the first two cooling stages the hot alumina is mixed with air and the air heated up as the alumina is consequently cooled down. In each stage the alumina is separated from the heated air by means of a gas cyclone and transported further to the next stage, whereas the air is flowing in direction of the CFB furnace to serve as heated combustion air.

Since then the technology has matured and improved with current energy consumptions of 2.79 GJ/t for the state of the art flowsheet as shown infigure1.

At last in the fluid bed cooler all residual heat, which is not easily recovered in the calcining process, is extracted from the alumina to enable discharge with safe alumina temperatures for handling. The fluid bed cooler consist of several sequential compartments, where primary fluidizing air is preheated for the nozzle grate of the CFB furnace and finally cooling water lowers the alumina temperature to safe handling levels.

The CFB calcination process utilizes maximum heat recovery by having air and furnace off-gas flowing in a countercurrent scheme to the hydrate feed or alumina product, respectively. In the current flowsheet the hydrate is preheated in two stages with the off-gas of the CFB furnace. In the first preheating stage

125

thus defining the general temperature level in the cooling stages and consequently has an influence as described on the performance of the cooling cyclones.

Key Consideration for Fuel Efficient Calciner Design Essential for the success of the CFB calcination flowsheet for optimized performance is not only the design of the Fluid Beds (CFB and Fluid Bed Cooler [1]), but also the cyclones in the system. In the presented calcination process heat recovery is done by counter current solid gas heat transfer. The means of separating solids and gas are the cyclones. Any amount of solids in the cyclones off-gas is counteracting the counter current heat recovery scheme. Further, the fines in the off-gas of one cyclone will be fed to the cyclone in the upstream recovery stage and subsequently reduce the separation efficiency as well.

The process calculation results show that in fact with a reduced furnace temperature, the fuel consumption of the calciner is less than with a higher furnace temperature. This can then only be compensated by more cooling stages. Not surprisingly reduced cyclone separation efficiencies also lead to increased fuel consumption. Measures for Further Reduced Fuel Consumption

The cyclone design is not only critical for best separation performance and thus for the process efficiency, but also for the amount of fines in the system. The fines are both fed to the process with the hydrate feed, and also generated in the process by particle breakage. A major part of the particle breakage is generated in the cyclones. This needs to be taken into account in the process and specifically cyclone design [2]. Cyclone design, e.g. geometry, velocity profile, etc. needs to be carefully optimized to have max. separation efficiency not only for one cyclone, but also for the overall process as breakage and fines influences all of the cyclones at the same time. This phenomenon is not only known for alumina calcination [3]

In the above sections the basic flowsheet of Outotec's CFB calcination technology is described and assessed. As the technology is now well proven and lot's of research and development has been undertaken in recent years, there are several developments and options available in this flowsheet, which further improve energy consumption, availability and operability. Which option is best suitable depends on client's requirements, fuel prices and project structure such as investment budget and predicted plant life time and life cycle. In the following sections, the different options available for reduced fuel consumption and improved plant performance are described and the impact on investment and operating costs are discussed. The discussions further distinguish between options for greenfield plants or revamps and upgrades of existing units.

■ 100 m90

Cyclone Separation Efficinecyin%

Di

Based on the above basic flowsheet several further options and variations of the flowsheet have been developed to reduce fuel consumption and hence reduce operating costs. Naturally these options will have an impact on the investment costs as well. There are several options available to reduce mei consumption of the calcination process. The available measures are targeting the optimization of the heat recovery and energy utilization within the calcination process inside the battery limits of the calciner. The two main options possible are discussed in this publication.

2 Cooling Cyclones 2 Cooling Cyclones 3CoolingCyclones 3CoolingCyclones with 1100°C with 950°C Furnace with 1100eC with 950°C Furnace Furnace Temperture Temperture Furnace Temperture Temperture

The available options are: • Hydrate bypass, where preheated material is bypassed around the CFB furnace and calcined with heat from CFB discharge alumina • Pre-drying of hydrate feed with waste heat from fluid bed cooler

Figure 2: Fuel Consumption depending on Cyclone Efficiency and Calcination Temperature Besides velocity and solids concentration in the cyclone inlet, the operating temperature also has a strong influence on the performance [4]. With increase of gas temperature, the gas viscosity increases and the drag on the particles inside the cyclones increases. The higher wall friction also causes reduced tangential velocities for the centrifugal field. As a result the separation efficiency decreases for a cyclone with constant inlet velocity. In order to compensate the inlet velocity needs to be increased, which will then increase breakage and generation of fines. As described above this again will have an impact on the overall process performance.

In figure 3 the first option the hydrate bypass is shown. This option has been developed in the 1990s by Outotec [6] and [7]. In the bypass flow sheet a part of the dried hydrate from the first preheating stage is taken and directed to the discharge of the CFB furnace. The bypassed solids are mixed with the hot alumina from the CFB discharge and allowed to react in a so-called mixing pot, which will provide sufficient residence time for the reaction. With this method, heat from the CFB alumina discharge is utilized for the calcination reaction of the bypassed hydrate, rather than for heating up of combustion air. Hence the heat is utilized more efficiently and secondly the amount of hydrate or alumina respectively, which needs to be heated to the reaction temperature

In figure 2 the influence of cyclone performance on the fuel consumption of a stationary calciner is shown. For comparison two calciners one with 2 and one with 3 cooling cyclones and two different calcination temperatures are assessed. CFB calciners are typically built with two cooling cyclones and operated at 950°C in the furnace. Stationary calciners are often equipped with 3 cooling cyclones and operated at 1100°C [5]. The furnace temperature is

126

P< «ïïmïng

ïvige !

Secondary Air Primary Air CBnaWst*r ■«·■*■

Figure 3: Flowsheet with the Hydrate Bypass implemented. of the CFB of more than 950°C is reduced. This reduces the required amount of fuel significantly. The Bypass was developed in the 1990s and has already been installed in more than 6 plants worldwide since then.

Figure 4: CircoCal flowsheet with bypass, hydrate dryer and 3rc preheating stage As a result, less heat is utilized in the preheating stage 1fromthe calciner off-gas. Consequently the off-gas temperature would rise. To utilize this heat, a third preheating stage can be installed, which then recovers the heat. Finally this means that less heat is rejected with the cooling water to atmosphere and there is also reduced heat loss with the calciner off-gas.

With the hydrate bypass in operation, the temperature profile in the cooling stages changes, as the solids enter the first cooling stage with a lower temperature than before. Furthermore, due to the reduced fuel consumption, less air is required and also less combustion off-gases are produced. This allows the design of the preheating stages, CFB stage and cooling stages to be smaller, with less construction material for the same nameplate capacity.

The flowsheet for this process is shown in figure 4 and forms the CircoCal process by Outotec. Advantages are that all extra equipment does not need to be refractory lined. Also there is only one more cyclone installed on the preheating side and no further cooling cyclone is needed to increase heat recovery. Cyclones are one of the major sources for particle breakage [2] and the most sensitive is the alumina on the product side (cf. section above). Therefore the impact of product quality is minimal for this approach, compared to the classical approach to expand the counter current heat recovery scheme.

To further reduce the fuel consumption of the calcination process, a closer analysis of the process is required. A thermodynamic heat and mass balance analysis of the calcination flowsheet shows that CFB with its fixed and homogenous temperature profile is a "pinch point" for the heat flow and exchange within the process. This means that with the existing counter current heat exchange and recovery scheme it is very difficult to transport heat from the back end of the process through the CFB stage into the front end of the process.

Alternatively the dried hydrate can also be transported to a dry hydrate storage silo instead of feeding to the calciner if dry hydrate is desired as an extra product from the refinery. However this option is not investigated further in this publication.

One way could be to install more preheating and more cooling stages to increase heat recovery at both ends. However this would mean more investment costs for very big and refractory lined vessels.

3.00

Another possibility is to collect the heat in one end and transport it around the "pinch point" to the other end to utilize it there. A method for this approach has been developed by Outotec in the past [8].

.2 O 2.95 i £> 2.90

Heat can be recovered from the fluid bed cooler by means of a heat carrier, such as water under pressure or thermo oil. The heat carrier is then pumped to the front end of the calciner. There a part of the moist hydrate feed is taken and fed into a fluid bed hydrate dryer (for details see the section below). In the hydrate dryer the heat from the heat carrier is used to dry the hydrate before it is fed to the preheating stage I. Then the heat carrier is pumped back to the fluid bed cooler. Hence the waste heat from the cooler is used to evaporate the free surface moisture of the hydrate.

Base Case

Hydrate Bypass

CircoCal™

Figure 6: Specific energy consumption of different flowsheet variations

127

Base Case

Hydrate Bypass

CircoCal™

Figure 7: Impact on investment costs Figure 8: Schematic of fluid bed hydrate dryer design

Evaluation of Measures for Reduced Fuel Consumptions

costs / benefit ratio can possibly be further improved if carbon projects are being included.

In the previous section, different options are presented to further reduce the fuel consumption of CFB calciners. The base calcination flowsheet has a specific energy consumption of-2.79 GJ/t with a feed hydrate moisture of around 7%. The results of the fuel consumption comparison are shown infigure6.

Drying Hydrate in a Fluidized Bed Key equipment of the CircoCal flowsheet is the hydrate dryer. There are a lot of refineries, which also produce dried hydrate. However in most cases direct heat transfer with hot off-gas from combustion is used to dry the hydrate. Here the waste heat from the fluid bed cooler is used. As the heat carrier, water or oil is used in a closed circuit (cf. figure 4). To transfer the heat to the wet hydrate from the hydrate filtration, indirect heat transfer via heat transfer surface is required.

In the comparison the hydrate bypass reduces the specific energy consumption by - 3 % and the CircoCal flowsheet has the lowest specific fuel consumption of all the options discussed. The reduction compared to the base case is about 6%. For all options, the reduction in fuel consumption can be seen as direct reductions in operating costs. There will be an increase in labor and maintenance costs for the extra equipment in the plant. However these can be seen as marginal. None of the options include maintenance intensive equipment such as extra rotating equipment or huge refractory lined vessels. The extra fluidizing air requirements can be covered with the existing blower concepts and Outotec's technology for solids mass flow control does not require any moving parts inside the solids streams.

Outotec has developed for this a hydrate dryer based on a bubbling fluidized bed similar to the fluid bed cooler. In the fluid bed dryer, heat transfer bundles are submerged in the fluidized solids. A schematic of the fluid bed dryer is shown in figure 8. In this fluid bed, large transfer areas can be installed in very small vessel volumes. For optimum bundle design the same design criteria are applied as for the fluid bed cooler with same success as reported in 2007 [1]. To ensure easy maintenance the bundles can be removed from the side and cleaned.

Investment costs of the plant will be impacted as presented in figure 7. It shows that an improvement in energy efficiency does not necessarily mean also an increase in investment costs.

The implementation philosophy of the dryer into the overall calcinerflowsheetis shown infigure9.

Figure 7 shows the changes of plant weights normalized to the base case. Again here the boundary condition is that all plants have the same nameplate capacity. The changes in weight here are considering vessel plate work, refractory and steel structure. Instrumentation, piping and other equipment such as blowers, screws, etc. do not change significantly.

Hydrate from Filtration

Preheating Stage I Offgas to ESP

I

Dryer Off gas to ESP

Due to the changed temperature profile in the hydrate bypass option the vessels can be reduced significantly in size as described above. This then leads to reduced overall plant weight and does in fact reduce investment cost for the plant when nameplate capacity is maintained asfigure7 demonstrates.

A

Heat Carrier from Fluid Bed Cooler

The reduced fuel consumption of the CircoCal flow sheet is offset by a potential increase of investment cost of about 15% if based on weight. This is the highest increase of investment of all options. On the other hand the savings in fuel can give a return of investment in 2 years based on an oil price of 80US$/barrel. The

Fluidizing Air for Dryer

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Hot Offgas from Preheating Stage II

Figure 9: Schematic of hydrate dryer implementation in preheating stage I

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The base case includes 5 main control loops, which are: • Hydrate feed • Furnace temperature • Furnace inventory • Fluid bed cooler level • Emergency water sprays

The wet hydrate is fed into the calciner feed bin. The feed bin has two separate bottom discharges, which are connected to screw feeders. One of the screw feeders is feeding wet hydrate directly into the venturi of the preheating stage I. The second screw feeder is feeding into the fluid bed hydrate dryer. The dryer screw feeder is controlled by a temperature measurement inside the dryer fluid bed and ensures that wet hydrate is only fed into the dryer, when the temperature is above 100°C. This guarantees that the fluid bed is always dry and easily fluidizable. Secondly, the speed of the screw feeder is controlled to maintain stable dryer temperature and ensures that all the heat available from the heat carrier is used.

All these loops are tied together in a governing overall control scheme with smart feed forward and troubleshooting mechanisms. In the recent years the improvements of process control and automation have led to very easy operation and fast ramp up performance [10]. As result there is little difference anymore between the different calcination technologies available and the limitations for fast process changes, ramp ups and shut downs are only in the speed in which process equilibrium can be achieved. Figure 10 demonstrates the comparability in ramp up performance for different calcination technologies.

Furthermore the inventory of the fluid bed dryer needs to be kept stable to keep the bundles submerged during normal operation. This is achieved by means of a lifting seal pot in the same way as it is used to maintain the level of solids in the fluid bed cooler. The solids level in the fluid bed dryer is detected by means of a differential pressure measurement across the fluid bed height. The differential pressure is then controlling the fluidizing air to the fluid bed dryer seal pot, which discharges the dry hydrate directly into the venturi of the preheating stage I.

In 2008 the influences on availability and consequences from unplanned outages of CFB calciners have been discussed [11]. Although not the major reason for unplanned outages, the consequences of refractory failures are very time consuming and costly. With the latest installations of CFB calciners the impact of plant trips and subsequent cooling and reheating cycles on the refractory could be explored. The stable operation, as reported in [2] and [12], proved to have very positive impacts on refractory life, thus not only decreasing operating costs by reduction of fuel consumption, but also reducing maintenance costs and loss of production by increasing reliability and plant integrity.

Finally in the preheating stage I both hydrate streams (wet and dry) are mixed with the hot off-gas of the preheating stage II. With the mixture of wet and dry hydrate, the utilization of the available heat in the heat carrier and thus in the fluid bed cooler is maximized. The vapor from the hydrate dryer is then added to the preheating stage I off-gas and leaves the calciner together with the combustion off-gas through the ESP and stack. Alternatively the vapor can also used to recover water if required. The off-gas from the hydrate dryer will have very high amounts of water with small amounts of air.

The above mechanisms and approaches will also be applied to the options described above, to further reduce fuel consumption on CFB calcination. For the bypass option only one more control loop is required which is the mixing pot temperature. For CircoCal the following control loops are needed: • Hydrate feed • Furnace temperature • Furnace inventory • Fluid bed cooler level • Emergency water sprays • Mixing pot temperature • Hydrate Dryer temperature • Hydrate Dryer level

Measures for Improved Operating Stability and Availability Besides the energy optimized flow sheet and plant layout, stable operation of the process is critical for the achievement of low operating costs. It is not only important for the optimized performance with minimum consumption of utilities, but also for maximum lifetime of plant equipment and low maintenance requirements. For example the typical it experience has been that stable process operation and low frequencies of unplanned plant downtime is one key for long refractory life.

With the base case, a thermal efficiency of -88% is achieved. In order to get closer to the theoretical fuel consumption of 2.45 GJ/t more initiatives need to be taken. Therefore the complexity of the plant will likely also increase. For the operator, the complexity is only visible in the amount of information he has to process and the attention he has to pay to each individual unit.

A further advantage for the refractory lifetime of the CFB system is the fact that it operates at significantly lower calciner reactor temperatures, compared to other stationary calciners while achieving similar product quality [9]. In the recent past, Outotec hasfrequentlyreported on the progress in plant operation and stability. With new technologies available, and with the experience of more than 60 plants built, the controls and automation has been developed to an extent that constant operation is achieved under nearly all conditions. This leads to a minimized number of unplanned plant downtimes as the available control philosophies now prevent plant drop outs by automated reaction of the system. Further the controls now assist in automated start-ups, change of plant loads and shut downs [9].

In this case it can be argued that the complexity grows significantly from the base case from the Bypass options to the CircoCal process, when one looks at the number of control loops to supervise. However the extra control loops are be integrated in the overall control philosophy to achieve same temperature stability as in the base case. The relationships between temperature changes and solids mass flows to different areas in the plant are well known, which makes the controls easily manageable.

129

consumption, but also on product quality such as fines in the final product. E.g. precipitation and calcination need to be looked at together in order to optimize investment and operating costs to achieve the required product specification. Acknowledgement Outotec herewith thanks their customers Alunorte S.A. and AOS Stade GmbH for their support and cooperation. The support from Alunorte led to the recently given energy efficiency award 2010 from the German Energy Agency.

—CFB Calciner 1 —CFB Calciner 2 - -GS Calciner

References

time, min Figure 10: Ramp Up of stationary calciners. Data for GS Calciner performance adapted from [5]

[I] M.Missalla, C.Klett, R.Bligh, "Design Developments for Fast Ramp-up Easy Operation of New Large Calciners" TMS Light Metals, (2007)

Therefore it can be suggested that the increase of complexity lies only in the increased demand for instrumentation and control maintenance, which can be handled via the general refinery maintenance scheme and relieves the operator.

[2] C. Klett, M. Missalla, R. Bligh, "Improvement of Product Quality in Circulating Fluidized Bed Calcination", TMS Light Metals, (2010)

Retrofits to existing CFB Calciners

[3] J. Reppenhagen, A. Schetzschen, J. Werther, "Find the optimum cyclone size with respect to fines in pneumatic conveying systems", Powder Technology 112, (2000)

The measures presented for reduced operating costs and increased availability, can also be applied to existing units. The improved control schemes are implemented easily with only limited costs involved, depending on the existing instrumentation infrastructure.

[4] M. Missalla, "Calculation Method for Highly Loaded Cyclones", Ad. libri Hamburg, Dissertation Technische Universität Braunschweig (2009)

Implementing a hydrate bypass or turning an existing unit into CircoCal operation is also feasible, but will involve installation of extra hardware. However this would not only reduce specific energy consumption and improve operability, etc. it also leads to an increase of plant production capacity. This is in contrast to the cases assessed above, where the plant capacity is deliberately kept constant. In a retrofit situation, the blower capacities and existing vessel and cyclone geometries are kept unchanged. This means that with changed temperature profiles, the throughput can be increased to maintain the velocity profile in the plant before and after the modifications. This also means that particle breakage can be kept constant or even improved depending on the existing unit.

[5] S. Wind, B.E. Raahauge, "Energy Efficiency in Gas Suspension Calciners (GSC)", TMS Light Metals, (2009) [6] H.-W. Schmidt, W. Stockhausen, A. N. Silberberg, "Alumina calcination with the advanced circulating fluid bed technology", TMS Light Metals (1996) [7] D.J. Brodie, H.W. Schmidt, "Custom Design Fluid Bed Calciner for Nabalco Pty Ltd." Proceeding of 5th International Alumina Quality Workshop, (1999) [8] H.-W. Schmidt, W. Stockhausen, "Latest developments in circulating fluid bed calcination based on operating experience of large calciners", Proceeding of 6th International Alumina Quality Workshop, (2002)

Conclusions In the above sections different options have been presented as to how to make CFB calciners more fuel efficient, how to reduce the operating costs and the impact on investment and process complexity and control.

[9] A. Saatci, H.W. Schmidt, W. Stockhausen, M. Ströder, P. Sturm, „Attrition Behaviour of Laboratory calcined Alumina from various Hydrates and ist influence on SG Alumina Quality and Calcination Design", TMS Light Metals, pp 81-86 (2004)

With these different options, thermal efficiencies of up to 92% can be achieved and fuel consumptions as low as 2.65 GJ/t. The risks involved for particle breakage, product quality and process complexity is minimal and can easily be mitigated.

[10] M. Missalla, J. Jarzembowski, R. Bligh, H.-W. Schmidt, "Increased availability and optimization of calciner performance due to automation", TMS Light Metals , (2009)

The options presented are not only of interest for new installations, but also for older ones and even viable for units of the very first generation of CFBs.

[II] P. Hiltunen, R. Bligh, C. Klett, M. Missalla, H.-W. Schmidt, "How to achieve high availability with large calciners and avoid unforeseen downtime", TMS Light Metals, (2008)

However for some options, the integration of the calciner in the overall refinery energy balance can give extra benefits. This needs to be the focus of future investigation on how to reduce overall operating costs in the production of smelter grade alumina from bauxite. These investigations shall not only focus on fuel

[12] M. Missalla, H.-W. Schmidt, J. R. A. Filhio, R. Wischnewski, "Significant Improvement of Energy Efficiency at Alunorte's Calcination Facility", TMS Light Metals, (2011) to be published

130

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PRESSURE CALCINATION REVISITED F. S. Williams, CMISCORP, Longmont, CO 80503 and C. Misra, AluminaTech, Pittsburgh, PA 15239 Bayer Process, Calcination, Pressure Alcoa Pressure Calcination Process

Abstract tth

The pressure calcination process developed at Alcoa Laboratories addresses both the questions of energy use and fines generation. It enables a substantial recovery of calcination energy together with the production of a stronger alumina.

Twenty-five years ago at the TMS 100 anniversary meeting of the Hall-Heroult Process, two Alcoa scientists, S. W. Sucech and the co-author of this paper, C. Misra presented a paper on an improvement to the alumina calcination process - pressure calcination.

In this process, alumina hydrate is calcined in two stages. In the first stage alumina hydrate is heated indirectly to 500°C in a decomposer vessel under a steam pressure of about 8 arm. About 85% of the combined water is released and is recovered as process steam. There is also a partial transformation of gibbsite to boehmite during decomposition involving alteration of the crystal morphology. The product from the decomposer has about 5% LOI and is calcined by direct heating to 750-850°C to obtain smelting grade alumina. The alumina product has high attrition resistance.

The improved process offered an opportunity for a net energy reduction in fluid flash calciners of 1.6 GJ/ton alumina (and the subsequent green house gas reductions). This could be retrofitted into existing fluid flash calciners and produce an alumina meeting smelting requirements with the added advantage of high attrition resistance and thus low dust generation. The question is: Why hasn't this improvement been incorporated into today's alumina plants?

Heat Effects in Hydrate Decomposition

Technical data in the previous paper will be reviewed and updated in the present paper.

The process of thermal decomposition of alumina hydrate can be studied from X-ray diffraction, DTA and TGA data [27]. The TGA plot for a typical American Bayer hydrate is shown in Figure 1 for atmospheric pressure calcination.

Introduction Energy requirements for alumina calcination have decreased from 4.5 GJ/t A1203 to less than 3.0 GJ/t A1203 following the replacement of rotary kilns by stationary calciners. Since the commercial introduction fluid flash calciners in the 1960s, there have been numerous papers on the development and improvement in design as well as increases in installed throughput capacities of this class of calcination equipment [3, 5, 6, 7, 8, 9, 12, 15, 18, 19, 20, 28, 30]. One major drawback of the new calciners was that the type of alumina trihydrate particles being produced in alumina refineries using the new calcination technology tended to break down and create a significant amount of fine dust during calcination and that dust, in addition to being an environmental problem, led to a range of feeding problems for alumina smelters using the flash calcined aluminas [1].

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This problem of dust creation has been mitigated over the years by the alumina refineries changing precipitation practices to create "sandy" "mosaic" crystal habit alumina trihydrate particles [22, 23, 24, 26, 31], reduction in internal velocities and sharpness of bends in calciner design [8, 28], and methods of measuring or characterizing the potential attrition resistance of alumina trihydrate particles [2, 4, 10,13, 14, 16,17, 21,22, 23, 28].

Figure 1, TGA Analysis of Hydrate As can be seen, more than 85% of the water is released below about 520°C. Some boehmite is formed due to hydrothermal conversion within the crystals. This is shown by the exothermicity at 270°C and endothermicity at 520°C in the DTA plot (Figure 2).

Today, the majority of the calcined alumina produced has been calcined in some form of fluid flash calciners as opposed to rotary kilns. The problem of alumina fines, while reduced, remains a continuing problem for handling and smelter cell feeding systems.

Heat requirements have been estimated from DSC measurements andare: For water removed between 200-350°C = 3660 J/g H 2 0 For water removed between 350-1000°C = 1686 J/g H 2 0 Overall = 1128 J/g hydrate

In 1986, an Alcoa paper at the TMS meeting introduced the concept of pressure calcination [25]. The present paper will reintroduce the data presented in that paper.

Similar tests carried out under eight atmospheres (120 psig) steam pressure show, as expected, larger amounts of boehmite formation and, unexpectedly, an 11% decrease in heat requirements to reach thefinalLOI value of about 1%.

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Figure 2. DTA Analysis of Hydrate Characteristics of Pressure Decomposer Product Bench scale decomposition experiments were carried out in a modified autoclave system with provision to inject dry alumina hydrate. In a typical run the autoclave was brought to pressure and temperature by starting with a small amount of water. Hydrate was then charged from the injection vessel by pressurized gas. A venting arrangement maintained constant steam pressure in the autoclave during the test. Range of conditions studied were: Pressure: 4-30 atmospheres Temperature: 250-650°C Residence Time: 30-120 minutes At the end of the test the autoclave was vented fully and cooled. The product was removed and examined for properties.

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The decomposed products were then subjected to the second stage atmospheric calcination in an electric furnace for one hour at the desired temperature. The products were again analyzed for LOI, surface area, attrition resistance and crystalline phase. Morphological features were examined by SEM and also by sectioning the alumina particles.

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132

Table I. Sieve Analysis( ' Comparing Hydrate Particle Size distribution to Pressure Calcined Alumina(2) Particle Size Distribution Sieve Size +100 +150 +170 +200 +270 +325

Hydrate 1 Hydrate Alumina 9 32 44 61 85 93

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Notes: (1) Values represent cumulative weight percent retained on indicated sieve size. (2) Average of all aluminas pressure calcined at 8 atmosphere 1st stage and 2nd stage at 1 atmosphere and 850°C. Results of these tests, which have a direct bearing on the development of a practical pressure calcination process, are summarized as follows: a. More than 50% transformation of gibbsite to boehmite occurs at 120 psig pressure and 500°C for a residence time of 60 minutes. A part of the boehmite subsequently converts to Y-A1203. b. This transformation affects the development of surface area. There is no rapid rise in surface area; the surface area increases to about 70 m2/g under the above conditions. This surface area value remains practically unchanged after second stage calcination at 750-850°C. c. The above decomposer conditions result in a product having a LOI of about 5%. A final LOI of <1% (corresponding to smelting grade alumina) can be attained by carrying out the second stage calcination at the relatively low temperature of 750-850°C. The final product consists largely of γ- A1203 with the complete absence of a - A1203. d. The formation of boehmite has considerable impact on the internal structure of the alumina and consequently on the attrition behavior. Figures 6 and 7 show SEM pictures of internal structures of atmospheric and pressure calcined (after second stage) aluminas. The absence of cracks and fractures in the pressure calcined material is remarkable.

Figure 7. Section of Pressure Calcined Alumina Attrition test results confirm the above observations. First, there is little particle breakdown during the pressure decomposition stage. Results actually show a slight coarsening effect. Second, the product from the second stage has very high attrition resistance as measured by the usual modified Forsythe-Hertwig test. Some results are given in Table Π. Table II. Sieve Analysis(1) and Attrition Indices, of Bench Pressure Calcined Aluminas(2) Umos. 8 atmospherei$ Hydrate 1 Atmos Sieve 300°C 400°C 500°C Mesh 12 17 19 17 17 =100 48 56 60 55 54 +150 64 74 74 78 76 +170 78 87 91 90 89 +200 90 98 96 99 99 +270 99 93 98 100 100 +325 Notes: (1) Numbers represent cumulative weight percent retained on indicated sieve size. (2) Second Stage Calcination at 850°C for 1 hour. (3) Atmospheric calcination performed in laboratory flash calciner for 30 minutes at 850°C.

Figure 6. Section of Atmospherically Calcined Alumina

Results also suggest that the residence time of the hydrate in the decomposer, and hence the extent of transformation to boehmite,

133

strongly influences later attrition behavior. Higher attrition resistance corresponded with longer residence time. Pilot Scale Continuous Decomposer

In general, pilot plant findings fully confirmed bench scale results with respect to product properties. Some results are displayed in Table III. Heat transfer results showed the gas side coefficient to be limiting with overall heat transfer coefficients in the range of 8-10BTU/ft2/hr/°F.

A pilot continuous decomposer was designed and operated to test some concepts of indirect heating of the hydrate and compare product quality results from batch operated bench tests with continuous decomposition tests. The aim was to obtain information useful for the final design and construction of a fullscale pressure calcination system. This must include geometry to incorporate a large amount of surface area for indirect heat transfer from hot combustion gases, practical feed and discharge schemes for the decomposer, and recovery of clean steam.

Table III. Sieve Analysis(1) and Attrition Indices of Pilot Pressure Calcined Aluminas Sieve Testi Test 2 Test 3 Mesh Feed Prod. Feed Prod. Feed Prod. +100 3 3 2 3 8 9 +150 27 28 22 26 36 39 +170 42 44 38 42 50 52 +200 73 75 71 74 73 76 +270 94 96 94 95 93 97 +325 97 98 97 98 96 99 A. I. 3 2 4 3 3 4 Notes: ( 1 ) Numbers represent cumulativeweight percent retained on indicated sieve size

Several practical decomposer designs were considered. Observations in an externally heated- transparent quartz tube showed that the decomposition of the hydrate bed is selffluidizing due to release of steam. This fluidizing effect can be utilized to improve heat transfer to the hydrate bed. On this basis, the most economical design concept for the decomposer resembles a vertical shell and tube heat exchanger with a selffluidized bed of hydrate flowing down the tubes and hot combustion gases flowing counter currently upwards in the shell.

Modeling of the fluidizing behavior inside the tube showed that the bed behaves as a bubbling fluidized bed in the top 80% of the tube and as a packed bed in the bottom section.

A picture of the pilot pressure decomposer is shown in Figure 8. The design is based around a single tube identical to what would be used in the plant calciner. It resembles a double pipe heat exchanger, with solid hydrate flowing down the tube and hot combustion gases flowing upwards in the annulus. Continuous feed to and discharge from the system were through pressurized lock-hopper arrangements. System steam pressure was maintained through a backpressure regulator. A gas-fired combustion system supplied hot combustion gases. A computerized data acquisition system connected to sensors provided flow, temperature and pressure data and computed mass and energy balances, heat transfer coefficients and pressure drop measurements along the column.

Heat Balance for Pressure Calcination Process A simplified heat and mass balance for the pressure calcination process is displayed in Figure 9.

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Figure 8. Pilot Pressure Calciner

134

effective fuel consumption for the process is only 1.55 GJ/t A1 2 0 3 . Design Of Industrial Pressure Calciner

3. D. J. Brodie and H. W. Schmidt, Custom designed Fluid Bed Calciner for Nabateo ΡΊΎ LTD, Presented at 5 th International Alumina Quality Workshop, Bunbury, Aus., 1999.

At the time of the original paper the preliminary design of a 76 t/hr AI2O3 capacity pressure calciner was worked out based on bench and pilot plant experience. The design allows retrofitting of an existing Alcoa fluid-flash calcination system to operate in the pressure calcination mode. The design utilized a pressurized lockhopper system for feeding dry hydrate to the decomposer vessel. Further development work and implementation of this technology, however, was stopped at this point.

4. P. Clerin and V. Laurent, Alumina Particle Breakdown in Attrition Test, TMS Light Metals, 2001, 41-47.

Discussion

7. E. Guhl, and R. Arpe, Nearly 30 Years of Experience with Lurgi Calciners and Influence Concerning Particle Breakage, TMS Light Metals, 2002, 141-144

5. J. Fenger et al., Experience with 3x4500 TPD Gas Suspension Calciners (GSC) for Alumina, TMS Light Metals 2005. 6. W. M. Fish, Alumina Calcination in the Fluid-Flash Calciner, TMS Light Metals, 1974, vol. 3, 673-682.

An even more energy efficient unit, not discussed in the previous paper, would be to build a stand-alone pressure calciner. Wet filter cake would be fed into a pressurized dryer/decomposer unit where the alumina would exit at 850°C. A tubular design with self-fluidized hydrate/alumina on the inside of the tubes and POC on the outside, similar to the retrofit unit discussed above, is visualized. The hot alumina product could than go to a two-stage atmospheric indirect cooling section using air and water to cool the alumina to handling temperature and preheat air for fuel combustion. In this case all of the steam latent heat from hydrate moisture and water of hydration would be recovered as usable steam. Dust collector flow volume and particle load sizing would be significantly less than in current flash calciners.

8. Vladimir Hartmann, et al., Upgrade of Existing Circulating Fluidized Bed Calciners at CVG Bauxilum without Compromising Product Quality, TMS Light Metals 2006, 125-130. 9. J. B. Henin, Flash Calcination by Fives-Cail Babcock, TMS Light Metals, 1984, 1669-1696. 10. H.P. Hseih, Measurement of Flowability and Dustiness of Alumina, TMS Light Metals 1987, 139-149. 11. Cornells Klett, et al., Improvement of Product Quality in Circulating Fluidized Bed Calcination, TMS Light Metals 2010, 33-38.

Conclusions By recovering the water released during alumina calcination as process steam, the Alcoa pressure calcination system decreases effective heat requirement for alumina calcination to less than 1.6 GJ/tAl 2 0 3 .

12. E. W. Lussky, Experience With Operation of the Alcoa Fluid Flash Calciner, TMS Light Metals, 1980, 69-80. 13. Valerie Martinent-Catalot et al., A New Method for Smelting Grade Alumina (SGA) Characterization, TMS Light Metals 2004, 87-92.

The large gibbsite - boehmite transformation occurring during pressure decomposition results in a strong attrition resistant product and enables the production of < 1 % LOI smelting grade alumina at the low calcining temperature of 750-850°C. The surface area of the product is around 70 m2/g suitable for effective F recovery.

14. James Metson et al., Evolution of Microstructure and Properties of SGA with Calcination of Bayer Gibbsite, TMS Light Metals, 2006, 89-93. 15. Michael Missalla, et al., Increased Availability and Optimization of Calciner Performance Due to Automation, TMS Light Metals, 2009, 241-245.

The process was studied in bench and continuous pilot units and an initial design for a 76 t/hr alumina pressure calciner was developed but not implemented.

16. Dag Olsen, Alumina Dustiness Related to Physical Quality Parameters - User Experience and R&D in Hydro Aluminum, Presented at 5 th International Alumina Quality Workshop, Bunbury, Aus, 1999.

With the energy saving, ability to retrofit into existing fluid flash calciners and the ability to produce a better SGA for smelters, it is very surprising that this technology has languished without being implemented! References

17. S. Perra, Measurement of Sandy Alumina Dustiness, TMS Light Metals 1984, 269-286.

1. Evan W. Andrews and Barry J. Welch, Alumina Quality Requirements for Improved Cell Feeding, Presented at 5th International Alumina Quality Workshop, Bunbury, Aus., 1999.

18. Andre' Pinoncely, and Karim Tsouria, FCB Flash Calciner Technology: Ten Year Performances, TMS Light Metals, 1995, 113-120.

2. T. R. Barton et al., Flash Calcination of Alumina: An NMR Perspective, TMS Light Metals 1995, 71-74.

19. B. E. Raahauqe et al, Energy Saving Production of Alumina With Gas Suspension Calciner, TMS Light Metals, 1982.

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20. L. Reh and H. W. Schmidt, Application of Circulating Fluid Bed Calciners in Large Size Alumina Plants, TMS Light Metals, 1973, vol. 2, 519-532. 21. G. I. D. Roach, and J. B. Cornell, Dust and Dustiness Testing of Smelter Grade Alumina, Presented at 2 nd International Alumina Quality Workshop, 1990. 22. G. I. D. Roach et al., Charge Contrast Imaging -Applications to the Bayer Process, Presented at 5 th International Alumina Quality Workshop, Bunbury, Aus., 1999. 23. Alpaydin Saatci et al., Attrition Behaviour of Laboratory Calcined Alumina from Various Hydrates and its Influence of SG Alumina Quality and Calcination Design, TMS Light Metals 2004, 81-86 24. J. V. Sang, Factors Affecting the Attrition Strength of Alumina Products, TMS Light Metals, 1987, 121-127. 25. S. W. Sucech and C. Misra, Alcoa Pressure Calcination Process for Alumina, TMS Light Metals, 1986, 119-124. 26. O. Tsamper, Improvements by the New Alusuisse Process for Producing Coarse Aluminium Hydrate in the Bayer Process, TMS Light Metals, 1981,103-115. 27. K. Wefers, and C. Misra, Oxides and Hydroxides of Aluminium, Alcoa Tech. Paper No. 19, 1987, Alcoa Technical Center. 28. Suzanne Wind and Benny E. Raahauge, Energy Efficiency in Gas Suspension Calciners (GSC), TMS Light Metals, 2009, 235240. 29. Suzanne Wind et al., Development of Particle breakdown and Alumina Strength During Calcination, TMS Light Metals, 2010, 17-24. 30. K. Yamada et al., Development of a Fluid Calciner With Suspension Preheaters, TMS Light Metals, 1983, 159-172. 31. J. D. Zwicker, The Generation of Fines Due to Heating of Aluminium Trihydrate, TMS Light Metals, 1985, 373-395.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DYNAMIC SIMULATION OF GAS SUSPENSION CALCINER (GSC) FOR ALUMINA Susanne Wind1, Mengzhe Wu2, Torsten Vagn Jensen2 and Benny E. Raahauge1 F.L.Smidth A/S Minerals, automation Denmark Vigerslev Allé 77, DK-2500, Denmark Keywords: Dynamic Simulation, Operator Training, Alumina Calcination This is especially so, if the Bayer circuit can process more bauxite, liquor and red mud as illustrated in Table 1 below, assuming:

Abstract Training of plant operational personnel is becoming more important today than ever, to sustain high availability and productivity of high capacity equipment. The Gas Suspension Calcination process for production of Smelter Grade Alumina is very easy to operate and control regardless of calcining capacity. But increasing calcining capacity of single GSC units exceeding 4500 tpd of SGA makes it increasingly costly to lose operating time. The process dynamics of GSC units are very fast with some true response times in fraction of seconds. To train GSC operators, FLSmidth has developed a dynamic Calciner Simulator which serves the primary purpose of training both new operators, as well as maintaining the skills of experienced ones. The dynamic Calciner Simulator has been developed from supply of more than 75 Pyro process simulators by FLSmidth to the global cement industry. The first Calciner Simulator for Alumina will be commissioned in Australia later this year.

1) 2) 3)

SGA production of 1,125,000 tonne per year; SGA Sales Value: 275 US$/ton; SGA production cost: 230 US$/ton; Calcination Capacity 3,500

Capacity

ton/day S&%

3,087

3,190

3,259

3,430

3,602

1 Availability - o r [Operating Factor

55%

2,993

3,092

3,159

3,325

3,491

|

9B%

2,930

3,027

3,092

3,255

3,418

1

2,867

2,962

3,026

3,185

3,344

1

|

|profrt/{Uïs)

(236)

4,394

7,481

15,199

22,916

(4,4891

0

2,993

10,474

17,955

|

1 Operating Factor

93%

(7,324)

(2,930)

0

7,324

14,648

1

91% US$/yr

(10,159)

(5,859)

(2,993)

4,174

11,340

]

(84,507)

1,571323

2,676,043

5.436,593

347

(1.556.474)

0

1,037,649

3,631,773

6,225,896 |

33$

{2,486,047}

(594,419)

0

2,486,047

4,972,094 J

332

(3374,229] (1,946,067)

(993359)

1,386,311

3,766,581 |

1 Availability in 1 Days per Year

If the Operating Factor decreases from 95% to 93% (at 93% Capacity Utilization) due to lack of operator training, the loss of just 8 days production, will incur a loss of US$ 994,419 to the refinery per year.

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WÊL |MF^'

ÈÊ 2Α|Γ^^ΡΒ

Bfë* *

8,197,143|

Table 1: Gross Profit and Loss Opportunities.

BP^-V

JE

US$/d 3 y |

95%

I Profit /(Loss)

As demonstrated over many years the Gas Suspension Calcination (GSC) process for production of Smelter Grade Alumina is very easy to operate and control regardless of calcining capacity.

1

1 Availability - or

1

Introduction

Capacity Utilisation (CU)

1 Nominal Design

jt ·

Or, if due to sufficient operator training, the Operating Factor can be increased from 93% to 95% and the Capacity Utilization can be increased to 100% based on improved market conditions as well, the gross profit will increase with US$ 3,631,733 per year.

1 - *JÎ

The above examples illustrates that to sustain high availability and productivity of high capacity equipment, training of plant operational personnel is more important today than ever before.

~^*?

Figure 1: 3 x 4500 tpd GSC Units at QAL, Australia

Furthermore, the training needs to be done without exposing the refinery to undue risk which can be done by using a Dynamic Process Model Simulator.

However, when the GSC operating factor exceeds 98% as is the case in some GSC units, very little opportunity is left for the operator to gain actual hands on shut-down and start-up experience.

In summary improved operator training can provide the following benefits:

With calcining capacity of single GSC units now exceeding more than 4500 tpd of SGA a day, it also becomes more and more costly to lose operating time owing to unnecessary calciner down time.

> > > >

137

Reduced and improved handling of incidents such as plugging; Improved Safety of operation with less stops; Maximize Capacity Utilization to meet demand; Maximize the Operating Factor at all times;

> >

develop a model that reliably predicts the quality of the calcined alumina in response to the operating conditions imposed. The FLS calcination model is built on the experience and know how gathered since the first GSC was put into operation in 1986, and incorporates modeling of the degree of calcination as well.

Improved quality by more stable operation; Minimize down time and cost;

Needles to say, in addition to operator training - proper plant and equipment maintenance is also required in order to maximize return on the assets installed.

The cooling section (COI, C02, C03 and C04) is used for recovery of heat from the hot calcined alumina. This is efficiently done by using four cyclones. The counter-current flow obtained with four (4) cyclone stages of co-current flow in series is providing a high thermal efficiency with respect to cooling the alumina and simultaneously pre-heating the combustion air. Again modelling heat transfer as well as solids separation becomes paramount.

To train GSC operators, FLSmidth has developed a dynamic Calciner Simulator which serves the primary purpose of training both new operators, as well as maintaining the skills of experienced operators. The dynamic Calciner Simulator described in this paper has been developed from operational experience, theoretical principles and the supply of more than 75 dynamic calciner process simulators by FLSmidth to the global cement industry.

The last section of the GSC system comprises a fluid bed cooler, which will reduce the alumina temperature from approx. 180 °C to a temperature low enough for the alumina to be transported with a belt conveyor. However, this process will not have any significant influence on the overall heat consumption of the GSC system.

CEMulator Unit Sale Volume

DUST RECOVERY

HEAT RECOVERY

Figure 2. CEMulator unit sales volume: 1986 to 2009.

jj~

As seen in Figure 2 the sales of such simulators to the cement industry have increased over the past decade. FLSmidth did experience a sales volume reduction in 2008 and 2009 due to the global economic conditions, but sales are expected to be at the 2007 level in the near future.

*""** — i

fis »MM

»*V Figure 3. FLS Gas Suspension Calciner (GSC) flow sheet.

The first dynamic Calciner Simulator to be described in the remainder of this paper, will be delivered to Australia for a new 3500 tpd GSC unit before it is started-up and commissioned later this year.

The Gas Suspension Calciner is straightforward to control with the following control loops: > > > >

Gas Suspension Calciner Process Flow Sheet Many processes are involved in the calcination of alumina hydrate. When the moist hydrate enter the GSC system it encounters hot gases from combustion of fuel and the calcination reactions. The purpose of the venturi and the first two cyclone stages is to utilize this heat for drying, pre-heating and precalculation of the hydrate. The foremost processes are heat- and mass-transfer and solids separation.

>

Production Rate controlled by Hydrate Feed Rate; Alumina Quality controlled by Calcination Temperature; Calcination Temperature controlled by Fuel Flow Rate, Alumina Discharge Temperature controlled by Cooling Water Flow Rate to Fluid-bed Cooler; Excess Oxygen controlled by Exhaust Gas ID-fan;

The control parameters for the closed-loop process simulator can be altered as the parameters in the plant DCS. Tuning these loops will enhance the closed-loop dynamic response of the process even further in the direction of truthful plant dynamic behavior.

The hydrate is partly calcined in the riser duct to the second preheater cyclone, but will continue to react in the next section, which is the calciner furnace and holding vessel. The fuel is injected into the furnace and combusted to achieve the necessary furnace temperature. The hot combustion gases entrain the partially calcined alumina into the Holding vessel, where sufficient retention time is provided to reach the final degree of calcination and alumina quality (LOI, SSA and alpha phase). Again heat transfer is important, but in combination with a calcination model and a combustion model that can handle both natural gas and heavy fuel oil. The central point is however to

The control strategy is to keep the Hydrate Feed Rate and thus Production as stable as possible and then modulate the Fuel Flow to maintain the Calciner Furnace temperature. This offers the advantage of a rapid control response as the manipulated variable (Fuel Flow) and the measured control variable (Furnace temperature) is situated in the same vessel. Dynamic Process Modeling of the GSC

138

reaches its steady state velocity (= gas velocity less its terminal velocity) in less than 1 second.

The aim of the simulator was to create a dynamic model that also corresponded with our steady state model when solving the timedependence mass, energy and momentum equations. Building the model thus allowed us to simulate starting and stopping the equipment, but also to introduce changes in any variable to model different operation scenarios.

Finally the operator observed process dynamics are further slowed down by the sampling frequency of the DCS. While the dynamic response of thermocouples is measured in minutes, the alumina quality like SSA is measured in hours, owing to the turnaround time of alumina samples through the laboratory.

In general constructing a model can be approached in different ways. It can be either: • • •

Fully theoretical, derived from first principle and theoretical properties Semi- empirical, using first principles and derived model parameters by comparison with plant data Fully empirical, derived by correlating input and output data.

Still other parts of the plant never reaches a true steady-state such as the outer part of the refractory lining being exposed to the ambient temperature cycle covering 24 hours per day and seasonal cycles over the entire year. In all, the FLS model is a mathematical analogy of the process generated by applying the principle of conservation of mass, energy and momentum. It can be used in real time simulations, thus improving the predictability and flexibility of plant operation.

The FLS model is semi empirical, solving the dynamic balance equations to reach steady state, but utilizes the considerable FLS know-how about process and operation of the GSC to obtain the right model response in comparison with observed behavior.

Model equations - An Example The mass and heat balance models are derived from first principle:

Some of the process dynamics of Gas Suspension Calciners are very fast with some true response times in fraction of seconds, such as accelerating the particles discharged from a cyclone into the down-stream riser duct as described below. Once particles are dispersed into the flowing gas stream in the Cyclone Riser Duct, each particle in the gas suspension is accelerated by the flowing gas until each particle reaches its steady state velocity which depends on the gas velocity and the particles terminal velocity. The time needed for each particle to reach their steady state velocity can be estimated by solving the equation of motion for each single sized particle. The vertical upwards flowing particle velocity relative to the gas velocity, at time t>0, assuming that the Stokes drag law applies, is:

Accumulation = Input - Output + Generation

In other cases a force balance is used as in the example of the pressure drop over a cyclone stage including its up-stream riser duct (heat exchange & separation stage): 0 =

Pressure drop due to acceleration Δρ,

Particle Velocity

v /// /

/

__

'

/

/

100 um/

/

/

(FgM + Fdust + Fs)-v2

{FgM +

FdustyVl

(4)

Pressure drop in the riser pipe bend

^~^~

/""

/

(3)

Pout " 4 P

drop components associated in the separation vessel:

The calculated particle velocity (please see list of symbols below) is shown in Figure 4.

/

Pin'

The pressure drop Δ/7 is determined by summing five pressure

Uo = [dp2 (pp - pg) g /18μ] (1 - exp[-18t/dp2pp]) + Uio exp[-18t/dp2 pp] (1)

45 ym

(2)

/ /ί

Δρ φ6 =0.5/^(1+%±£-)Λ 2 ν 2 2

/

/ 150 pm

//

'" m

(5)

Pressure drop due to rise pipe area change Apm = 0 . 5 ( 1 + ^ ^ - ) / ^ (v32 -vl )

/

(6)

Pressure drop due to friction

χ έ £ ^

Apcyc=0.5kcycpg^

Time, sec

(7)

Pressure drop due to gravity

Figure 4. Particle Acceleration in a Vertical Riser Duct.

(8)

From the above figure it can be seen that a 45 micron particles reaches its steady state velocity (= gas velocity less its terminal velocity) in less than 0.1 second, while a 150 micron particle

139

The total pressure drop will thus be expressed as the summation of the contributors:

4 -

Δ

0 -

Ρ = 4 P , « + Δ / V + APrPa + APcyc + 4 P , ™

2

<9>

-2 -

Below the dynamic response to a step change in the ID-fan speed can be seen on the outlet suction pressure from the dedusting filter and cyclone POI respectively. 95 94 93 92 91 90 89 88 87 86 85

• ID-Fan »POI press. » BH press.

1n

-7.5 -8.5

0

100

200

300

400 500 Time (s)

-10 -10.5 r

i

600

700

»

..Lez

■ pppi 1

''"ZZZZZZZZZZZzzy

However, matching the stationary steady-state response shown above is a necessary criteria, but insufficient to qualify the dynamic model as applicable for training operators!

-9.5

—i

1 1 1 1 1

Figure 7. Pressure Profile at 50% Turn Down ratio with 0,2% deviation relative to absolute pressure of 101,325 kPa.

-8 -9

-I

1 I

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.8 γ~~~~

-6.5 -7

^_

H^J

"6 T

-6

T

«Simulation «SteadyState

T

Needless to say, if a dynamic simulator model does not respond dynamically in a way that the operator can recognize from the real plant dynamic response observed from the DCS, the dynamic model will surely be abandoned and not used.

-11

800

Figure 5. Dynamic response in relative suction pressure to change in ID fan speed.

This raises the very important issue of proper model validation. We are of the opinion that the response of the dynamic process model must be tested and verified by our experienced GSC commissioning engineers in order to be qualified as a valid training tool - and thus applicable in the Training Simulator.

Implementation of Dynamic Model Results and Validation The overall pressure profile model prediction at steady-state is compared below with the resulting pressure profile calculated by the independent steady-state design model, which in turn has been verified against practical GSC plant operation:

Training Simulator System Structure To cover the needs for conducting simulation based training the FLSmidth training simulator consists of 4 program modules. The modules that are illustrated in Figure 8 covers: • • • •

Figure 6. Pressure Profile at Nominal Production with 0% deviation relative to absolute pressure of 101,325 kPa. As can be observed from Figure 6, a reasonable accurate convergence is obtained from the model when applied at nominal production of the GSC. But to be of use to the operator, the model must also provide a reasonably true response over the entire GSC capacity range as demonstrated below in Figure 7 at 50% Turn Down ratio.

140

Dynamic model and solver, simulating the dynamic process behavior; Soft PLC, simulating control loop actions and equipment unit/sequence interlocking behavior; Simulator, control of the training communication between instructor and trainee, etc. DCS (Distributed Control System), enablesriskfree trainee operator interaction;

Simulator Dynamic solver Process simulation

CERTIFICATE

Session control

Session evaluation

ACTION REPORT

ACTION REPORT

DCS Operator Access

Unit simulation

Soft PLC

ACTION REPORT

Automated tutorials

Data transfer

Figure 8. Simulator program structure split in program modules (dashed lined boxes can be standard FLSmidth setup or Honeywell).

Figure 9. Example of action reports and session plots representing the offline session evaluation.

In the Dynamic solver module the process dynamics are simulated. This includes the heat, mass and momentum balances and will represent the fast as well as the slow dynamic behavior in the GSC system based on mass, component and energy holdup. At FLSmidth we use the gPROMS, solver which is a state of the art dynamic solver delivered by PSE based in UK.

Next in the Simulator module is the automated Tutorials. Tutorials are prerecorded sessions including session control events like the action described in online session control. These can be used to uniform the training as events will have exactly the same timing for all Trainees, which results in that evaluation of two Trainees can be compared. The messages included in Tutorials are used to support the training and can therefore be used to ensure that all Trainees learn the same procedures covering startup, hot standby and shutdown.

The Soft PLC module is where unit simulation is hosted. As in a true plant control system setup the PLC holds the process and plant information defining the operating state of the process. For this purpose PLC programs are uploaded in the Soft PLC. These programs are modified to simulate the physical equipment responses such as a starting motor. Besides the FLSmidth standard Rockwell Soft PLC setup a solution supporting Honeywell Soft PLC s are also supplied.

Finally the Simulator module includes the data transfer between the other modules. The main data transfer is synchronizing the Soft-PLC and Dynamic solver so that all process data is updated in the PLC as well as all motor states and set-points in the dynamic solver. This module also ensures that the Soft PLC and Dynamic solver are aligned when restoring a saved session.

In the Simulator module all the programs used to set up the simulation is hosted.

In the Distributed Control System (DCS) objects are presented to the Trainee just as in a real DCS. Here object trending, alarm handling, set-point changes and the remaining features covered by the DCS will be available.

The first area covered by this module is the Session control, where the Instructor can interact with the simulation session allowing control of the following: • • • • • •

Start/Stop/Pause session Simulation speed - accelerate process responses if needed Disturbances - offset on demand Input parameters - external inputs/definitions On line evaluation Text messages

In the FLSmidth standard solutions the in-house developed DCS named ECS (Expert Control and Supervision) is used, but the simulation setup also supports a solution using Honeywell Experion DCS as operator interface. Common for all setups are the need for interaction from the Trainee and Instructor through their different stations. Figure 10 and 11 illustrates the accessibility needed for these two stations.

Secondly the Simulator module includes online and offline Session evaluation. In the online Session evaluation the Instructor can monitor the current performance of the Trainee throughout the session. In offline evaluation the completed sessions are presented as reports representing evaluation objective trends and action reports. These reports can be used to evaluate the performance of the Trainee and dig into what learning areas that needs focus of the instructor.

141

In addition to the increased use of alternative fuels and focus on emissions, the current situation also raises demand for the operators to get the right skills by simulator training. Subsequently, this training results in more continuous and optimal plant operation with economical benefits to follow.

Trainee Station

In the cement industry we have successfully developed an array of standard simulation solutions, which can be purchased as off the shelf applications for fast delivery. The trend is that these are sufficient to cover the needs for training in most organizations, while specific plants tend to need minor adjustments for the process to look more like the specific plant. Back in the pioneers days in the 1980's the simulator was sold to specific plants. This trend has changed today where we still deliver half of our simulators to specific plants, but an increased interest from global companies is to purchase simulators for corporate training centers.

Figure 10. The program module access from the Trainee station. From the trainee station the Trainee can make some of the basic session controls such as start/stop/pause and send messages. To ensure focused training the Trainee must be able to monitor the evaluation throughout the session. As the natural environment for an operator is in the DCS the Trainee's main access to the simulation will go through this interface to do control, monitoring and error tracking.

Apart from the CEMulator units sold externally there are 70 internal licenses spread across FLSmidth where the use varies from training to assistance in development of high level control standard solutions. Conclusion

Instructor Station

Increase in calcining capacity of single GSC units, sometimes exceeding 4500 tpd of SGA, makes it costly for the refinery to lose operating time. When at the same time the GSC operating factor exceeds 98%, as experienced in some GSC units, very little opportunity is left for the operator to gain actual hands on shutdown and start-up experience. The Dynamic Calciner Simulator for SGA production developed by FLSmidth, will provide the Alumina refinery with the option of a risk free training of its new and experienced operators. In the cement industry the trend is that sufficient and documented training of operators is becoming a growing demand from local authorities as more and more focus is placed on sustainable production and ensuring environmentally constrained emissions.

Figure 11. The program module access from the Instructor station.

Consequently FLSmidth has experienced a surge in supply of training CEMulator units to the global Cement Industry, a demand that may come from the global Alumina Industry in the not so distant future.

The instructor station will act as a base for the Instructor. Here the full palette of session control must be available so that the Instructor in addition to the Trainee control can introduce disturbances, change input parameters and evaluation objectives on line. During training preparation the Instructor can use the tutorial editor to tailor make training sessions to the upcoming training needs. While undertaking the training the Instructor will also need to access the Operator DCS to fine tune the tutorials. The DCS can also be used during the session to monitor the current Trainees sessions during training.

List of Symbols Symbol Fs ^g.in Fdust

υι υ2 υ3

FLS Experience with Training Simulator In the cement industry the trend is that sufficient and documented training of operators is becoming a growing demand from local authorities. A recent example is the requirement for documented training in cement industry in USA. The reason for these demands is not only for the purpose of plant and staff safety, but also a requirement for implementation of a sustainable production.

Ai Août Krpb Kcyc Pg,2 Pg,3

142

Definition Solid Mass flowrate Gas Mass flowrate Solid entrained in gas flow Gas velocity at inlet of riser pipe Gas velocity at outlet of riser pipe Gas velocity in cyclone inlet area of riser pipe outlet area of cyclone Friction constant Friction constant Gas density at outlet of riser pipe Gas density in cyclone

Unit kg/s kg/s kg/s m/s m/s m/s m2 m2

-

kg/m kg/m

M Δρ Pin Pout

H g

Δρ 8 1 1 8 Aprpa Aprpb Apcyc Apgrav

u0 dp

PP Pg

t μ

ui0

Solid hold up Pressure drop Inlet pressure Outlet pressure Height of cyclone Gravity Pressure drop of acceleration Pressure drop of changed area Pressure drop in the riser pipe bend Pressure drop in the cyclone Pressure drop through gravity lift Rei. particle velocity at t>0 Particle diameter Particle density Gas density Time Gas viscosity Rei. particle velocity at t=0

kg kPa kPa kPa m kg/m2 kPa kPa kPa kPa kPa m/s mm kg/m3 kg/m3 s kg/s m m/s

References 1.

2. 3. 4. 5. 6.

B.E. Raahauge,et al.: "Application of Gas Suspension Calciner in Relation to Bayer Hydrate Properties", The Australasian Institute of Mining and Metallurgy, 1981 Annual Conference, Sydney NSW, 1981. B.E. Raahauge,et al.: "Energy Saving Production of Alumina with Gas Suspension Calciner", 111th AIME Annual Meeting, Dallas,US,1982. T.A.Venugopalan,"Experience with Gas Suspension Calciner for Alumina", Proceedeings 1st International Alumina Quality Workshop, pp 53-66, (1988). S.Wind and B.E.Raahauge: "Energy Efficiency in Gas Suspension Calciners (GSC)",TMS Light Metals, pp 235- 40, (2009). J.Ilkjaer, L.Bastue and B.E.Raahauge:"Alumina Calcination with the Multi Purpose Calcine", TMS Light Metals, pp 107111,(1997). R.K. Jonas:"Design and Training Improvements through the use of Dynamic Simulation", TMS Light Metals, pp 63-67, (2004).

143

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Physical Simulation and Numerical Simulation of Mixing Performance in the Seed Precipitation Tank with a Improved Intermig Impeller Zhang Ting-an1, Liuyan1, Wang shuchan1, Zhao hongliang1, Zhang chao2, Zhao qiuyue1, Lv guozhi1, Dou zhihe1 (1 School of Materials and Metallurgy of Northeastern University, Key Laboratory of Ecological Utilization of Multi-metal Intergrown Ores of Education Ministry, Shenyang, 110004; 2 Shenyang Aluminum & Magnesium Engineering & Research Institute, Shenyang, 110001)) Keywords: Seed precipitation tank; Mechanical stirring; Numerical simulation; Solid-liquid phase flow Abstract A PC6D fiber optic reflection probe was used to measure local solid concentrations distributions in the seed precipitation tank with a new-style Improved Intermig impeller, and the commercial software FLUENT 12.0 on parallel computing graphic workstation was used to simulate numerically its flow fields. The physical simulation results show that at high rotation and viscosity, high-speed particles are more conducive to uniform distribution. When //=3.50cp, «=172rpm, the multiple relationships of the average value of the experiment and theory is 1.011 times. The numerical simulation results show that the too long or too short impeller off-bottom clearance height(C) is not conducive to the suspension of Al(OH)3 particles. Enlarging the blade diameter is good for suspending of Al(OH)3 particles. The blade diameter has a big influence on the stirring power, and the longer blade diameter needs more power consumption, but the C has little effect on the stirring power, which can be ignored. The physical simulation results compare well with numerical simulation.

could reduce the sediment significantly, was devised by the Shenyang Aluminum & Magnesium Engineering & Research Institute. Power consumption also declined from 200kW to 75kW. The difference between the traditional Intermig impeller and the improved impeller has been reported in the literature141.

At first, we want to find the best impeller off-bottom clearance height and impeller diameter just in single impeller experiments. And in next step, we will study the multiple impellers' performance based on the results of single impeller experiments.

Cold water experiments The profile of local solid concentrations on the water-glass bead particles in liquid-solid system, were measured by using a PC6D fiber optic reflection probe and analyzed under different conditions. The PC6D particle concentration measuring instrument is manufactured by the Institute of Process Engineering, Chinese Academy of Sciences, the measuring range is 0~800g/L. The instrument uses an optical fiber as the measuring probe. The reflected light of materials at the fiber end comes back to the photo detector in the instrument through the same fiber and is converted into a voltage signal proportional to the concentration of materials. The average concentration of solid particles is obtained from the voltage signal.

In order to fully understand the impact of the structure of the blade on the solid suspension characteristics and power consumption, the seed precipitation tank with Intermig impeller was studied in our laboratory delegated by the Shenyang Aluminum & Magnesium Engineering & Research Institute. The study includes mixing uniformity, suspension of Al(OH)3 particles and power consumption, etc. By this research we intended to improve the internal structure of the seed precipitation tank, solve the problem of sedimentation, and improve the yield and quality of alumina. This has important theoretical and practical significance for the alumina industry. Research Method

Introduction Technological revolutions have recently hit the alumina industry. As one of the key steps in the Bayer process for the production of alumina, the seed precipitation step has a great influence on alumina product output and quality, and also has an indirect effect on other processes111. So it is necessary to optimize and improve this step. Currently, the mechanically agitated precipitation tank has become the major equipment of seed precipitation step, because it has many advantages which include low power consumption, better mixing effect and less sediment; avoiding "Short-circuiting" of slurry; high reliability, etc[2]. But the application of mechanically agitated precipitation tanks was late in China, and we have less information on the flow field in the tank.

The cold water experiment is based on the similarity principle. The ratio between water model and industrial seed tank is 1:33, and the experimental equipment was made with plexiglass 043.5cm x 52cm. Liquid height was 35cm and material system was glass beads-water (with sugar added to change the viscosity). Table 1 shows the comparisons of physical parameters. Glass beads were used instead of Aluminum hydroxide and an amount of 3800g added in the cold water experiment. The syrup viscosity in the cold water experiment was consistent with the viscosity of the NaA102 solution in the seed precipitation process.

By comparing the power consumption of a traditional Intermig impeller and a pitched blade impeller in industrial tests, the former is much lower than the latter, but using the traditional Intermig impeller produced a large amount sediment133. So a new-style impeller, - an improved Intermig impeller - which

145

n=172rpm, it was 1.340 times; when r|=3.50cp% n=172rpm, it was 1.011 times. The results showed that the Experimental concentration was closer to theoretical concentration under this condition, which was more evenly mixed. Turbulence model: Realizable k-ε; Multiphase model: Eulerian; Interphase drag: Gidaspow; Rotate model: Multi-Reference Frame; Numerical solution: SIMPLE algorithm for pressure velocity coupling; Discretized with the second-order upwind; All Residual converges to 10"3

Table I Comparisons of physical parameters Physical Parameters

Density /kg/m3

Viscosity /cp

Sodium Aluminate Solution

1330

3.5

1000 1149

1.52 3.50

2430

Granularity ΙΟΟμπι

Running Water Syrup Aluminum Hydroxide Glass Beads

2380

Numerical simulation Some research has been done on solid - liquid phases numerical simulation15"71, that may verify the feasibility of study on seed precipitation tank by CFD. (1) Grid Generation In order to reduce computing time and improve the convergence and stability of solution process, we should reduce the grid number and improve the mesh quality. So we took the following methods: A: Ignored the thickness of baffle. B: Changed grid from a tetrahedron into a polyhedron. Figure 1 shows the grid model of the seed precipitation tank. The grid number was about ten million.

Fig.l Grid model of the seed precipitation tank

Results and analysis of numerical simulation

(2) Simulation Strategies From the literature about simulation of solid-liquid phase flow[8"11], we chose the simulation strategies are as follows:

The numerical simulation research is for the seed precipitation process, the liquid-solid system of NaA102 solution and Al(OH)3 particles (for detail about material parameters see attached Table 1). At first, we validated the simulation strategy based on cold water experiments and then we found the best conditions by simulating the seed precipitation process.

Results and analysis of cold water experiment Experimental measurement points were set to 7.5 % 9.5 * 11.5% 13.5% 15.5x 17.5 s 19.5> 21.5cm in radial distance (the center of slot as the 0 point that was indicated as 1); 2x 6^ \0S 14% 18% 22% 26> 30cm in axial distance (the bottom of slot as the origin that was indicated as d).

Simulation validation based on cold water experiments In this section, the liquid-solid system was composed of water and glass beads. By comparing the results of water model experiment and numerical simulation on the solid particles concentration field, we got the following results:

Figure 2 showed the concentrations along the radial and axial distance. A uniform distribution of particles in the tank was assumed to be the theoretical concentration. When r|=1.52cp% n=96rpm, the multiple relationships of the average value of the experiment and theory was 0.962 times; when r|=1.52cp>

90 r

10 12 14 16 18 Radial distance(cm)

10 15 20 25 Axial distance(cm)

22 (a)f/=1.52cp, w=96rpm

146

ÖD

O

8

10 12 14 16 18 20 22 Radial distance(cm)

10 15 20 25 Axial distance(cm)

(b)»pl.52cp, n=172rpm 90

90

80 l·

80 70l· ^60 3 50 40 30 8

20

10 12 14 16 18 20 22 Radial distance(cm)

10 15 20 25 Axial distance(cm)

30

(c)7=3.50cp, «=172rpm Fig.2 Concentrations along radial and axial distance by experiment

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147

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Radial distettero») Radial dìstaiiceCcii*) (a)Experiment results (b)Numerical simulation results Fig.4 Radial distribution of solid particle concentration (z is the distance between measuring points and tank bottom in radial direction) Figure 3 and Figure 4 showed that particle concentration decreased from the bottom to top surface in the axial direction (z), the particle concentration was higher closer to the tank wall in the radial direction (x). By comparing the results of the cold water experiment and numerical simulation, the latter value of particle concentration was slightly higher, but both of them had

the same trend on particle concentration gradient, so the simulation strategy and results were credible. Keeping other conditions the same, we increased the viscosity of the solution, and then we got the following results: 210

180l· 5* 150l·

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consistent with the experiment results. The effect of distance between the impeller and the bottom(Ci on particle suspension and power consumption

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Figure 7 shows the distribution of particle volume fraction with different C. (y= 0 section). The results showed that too long or too short a value of C may lead to aggregation of Al(OH)3 *»·** S1sr**t

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particles at the tank bottom. Figure 8 shows the average of particle volume fraction with high and low stirring speed.

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The results showed that C had little effect on the particle suspension at low stirring speed, while a large C may make the particle distribution more uniform in the whole tank at high stirring speeds.

0.132 h

10.129

The power consumption with different C at 150rpm can be seen in Figure 9. With the increase of C, the stirring power at first increased, and then began to stabilize, but the C had little effect on the stirring power. In this section the best value of C was 111mm.

M C

0.126 h

on 0.123h 40 80 120 Distance between impeller and bottom(nim)

The effect of impeller diameter (D) on particle suspension and power consumption Figure 10 shows the distribution of particle volume fraction with different D/T (T was the tank diameter, y= 0 section), and Table 2 showed the power consumption under the same conditions.

160

Fig.9 Power consumption with different C by numerical simulation, 150rpm

149

(a)D/T=0.5 (b)D/T=0.6 (c)D/T=0.7 Fig. 10 Concentration field with different D/T by numerical simulation, 150rpm 2. Zhang Chunyang, Fu Yan. Alumina Plant Design, (Beijing: Metallurgical Industry Press, 2008), 115-141. 3. Wu Jianqiang, Dong Baocai, "Evaluation of Operating Results of Large-scale Precipitation Tanks with Mechanical Agitators", Linght Metals, 2000(03). 4. Wang Rui, "Numerical Simulation of Multi-Phase Flow in the Alumina Seed Precipitator" PhD Thesis of Northeastern University, Shenyang, China, 2009:29-30. 5. G Montante, "Experiments and CFD Predictions of Solid Particle Distribution in a Vessel Agitated with Four Pitched Blade Turbines", Chemical Engineering Research and Design, 2000,79(8):1006-1010. 6. R. J. Weetman, "Automated Sliding Mesh CFD Computations for Fluidfoil impellers", 9th Euro.Conf.On Mixing, 1997:195-202. 7. I. Naude, C. Xuereb, J. Bertrand, "Direct Prediction of the Flows Induced By a Propeller in an Agitated Vessel Using an Unstructured Mesh", Can Chem. Eng., 1998,76:631-640. 8. Montante Q "Experiments and CFD Predictions of Solid Particle Distribution in a Vessel Agitated with Four Pitched Blade Turbines", Chemical Engineering Research and Design, 2000,79(8):1006-1010. 9. Weetman R J, "Automated Sliding Mesh CFD Computations for Fluidfoil Impellers", 9th Euro. Conf. on Mixing, 1997:195-202. 10. Naude I, Xuereb C, Bertrand J, "Direct Prediction of the Flows Induced By a Propeller in an Agitated Vessel Using an Unstructured Mesh", Can Chem. Eng., 1998,76:631-640. 11. Bakker A, Fasano J B, Myers K J, "Effects of Flow Pattern on the Solids Distribution in a Stirred Tank"(8th European Conference on Mixing, Cambridge, U.K), IChemE Symposium, 1994,136(329):l-8.

Table II Power consumption with different DAT by simulation(unit: W) C=71mm

D/T=0.5

D/T=0.6

D/T=0.7

40rpm

0.0472

0.127

0.258

150rpm

2.025

6.060

13.690

Figure 10 and Table II show that enlarging D may promote the suspension of Al(OH)3 particles, but the power consumption increased substantially. So we should choose moderate D to get a better efficiency in particle suspension, with low power consumption. In this section the best value of D/T was 0.6. Conclusion (1) In terms of overall distribution, the particle concentration with radial distance increases, while the axial distance reduces. The high rotation and viscosity, high-speed particles are more conducive to uniform distribution. When r|=3.50cp> n=172rpm, the multiple relationships of the average value of experiment and theory is 1.011 times. It shows that the experimental concentration is closer to theoretical concentration under this condition, indicating the particles are more evenly mixed. (2) By comparing simulation and cold water experimental results, the feasibility of the CFD method is verified to simulate the flow and solid-liquid mixing characteristics. (3) At a constant stirring speed, there is an optimum value of D and C which can get a better effect of particle suspension with low power consumption. In this study, the best values of C and D/T are 111mm (C/T=0.55) and 0.6 respectively. Acknowledgement The authors acknowledge the support of the National Natural Science Foundation of China (NO.50974035, 51074047). References 1. Yang Zhongyu. Alumina Production Technology, (Beijing: Metallurgical Industry Press, 1993.10),133.

150

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Two Perspectives on the Evolution and Future of Alumina 1 Linus M. Perander, Barnes B. Metson, 2Cornelis Klett ^ight Metals Research Centre, The University of Auckland, New Zealand 2 Outotec GmbH, Frankfurt, Germany

Alumina, Bayer process, Hall Heroult process, Emissions, Energy, Calcination continuous way to feed the alumina resulted in the development of so called Break and Feed systems. The first implementation of this way of feeding was the Bar Breaking method. Using this method, a breaker beam is lowered to break the layer of crust (which covers the cell and anodes) and then a specified amount of alumina (determined by volume) is dumped into the cell. Along with the alumina a large amount of broken crust and cover material also enters the cell. Dissolution is slow and large amounts of undissolved alumina/crust are deposited on the surface of the cathode and gradually dissolved until the next feed cycle. The bar breaking method has given way to Point Feeding. Up to five of these point feeders (shot size 0.5 - 3 kg of alumina determined by volume per feeder) operate every few minutes, changing the operation from a semi-batch type to a more continuous nature [3].

Abstract Over the 125 year history there have been a number of stepchanges in the Hall-Heroult process, despite a remarkable adherence to the original concepts of the invertors. In addition to the steady increment in scale, most noteworthy perhaps have been the introduction and impending disappearance of Soderberg technology, the introduction of magnetically compensated cell design, changes in dynamics of alumina feeding and the introduction of dry-scrubbers for HF control and fluoride recovery. The Bayer process has also seen some significant advances, driven by the demands of energy and environmental imperatives and the steadily narrowing window of product specifications, driven in turn by refinements in the Hall-Heroult process. Demands for coarser particle size distribution, higher specific surface areas and lower attrition index have been accompanied by changes in precipitation strategy and conversion to more energy efficient stationary calcination processes.

As a result, particularly of these feeder changes, the demands on alumina quality have also changed. The so called Floury Alumina (with fine particle sizes, low specific surface areas and high alpha alumina contents) has poor flow and dissolution characteristics and has been displaced by Sandy Alumina (with larger particle sizes, high surface areas and low alpha alumina content). The impact of alumina properties on transport, feeding and dissolution characteristics, the ability to form a stable crust and anode cover as well as the adsorption of HF in the dry-scrubbers, have been of particular interest. Of note is the continued evolution of these properties (for example the inexorable demand for higher surface areas), at times with a limited understanding of the net process impact of such changes.

The properties of a "typical" metallurgical alumina have thus changed. Indeed the term "alumina" is now more indicative of stoichiometry than it is of structure, and even in this, it is less than precise. In this paper we discuss how new scientific tools and insights are changing the way we define (and perhaps should specify) this material. The History of Alumina Calcination The early history and development of the alumina and aluminium processes and industries are summarised elsewhere [1, 2] Fundamentally these processes remain the same, although a number of technological breakthroughs and step-changes have occurred. The past 50-years in particular have fundamentally changed what a smelter demands in terms of properties and performance of the primary raw material, the alumina fed to the reduction cell. This paper addresses the impact of the calcination step in the Bayer process on the development of alumina properties and our understanding of how these properties impact on the smelter operations.

Stationary calciners, as an alternative to the Rotary Kilns were introduced in the 1950s to 60s. The gains, both in increased production volumes, reduced maintenance and in energy savings, resulted in considerable research efforts into these technologies, and as a result variations of thefluidisedbed technology emerged. Today these are known as: Circulating Fluidised Bed (CFB), Gas Suspension (GS) and Fluid Flash calciners. For the main calciner technologies typical energy consumption, residence time, production capacity, alpha alumina content, calcination temperature and free heat transfer surface are presented in table 1. With typical production capacities between 2500 to 4000 tons per day for modern Gas Suspension or Circulating Fluidised Bed calciners, significant effort goes into minimising the downtime of these installations which has also led to other process and control improvements. Most noteworthy, however, is the reduction in energy consumption compared even to best practice Rotary Kilns.

The use of dry scrubbers in the aluminium smelters started in the 1960s as a response to therisingemissions concerns, and potential fluoride losses associated with this process. The dry scrubbers make use of the high specific surface area and the reactivity of partially calcined alumina (LOI < 1 wt-%) to capture the volatile fluorides emitted from the electrolyte in the reduction cell. Around the same time as the dry scrubbers were introduced, another significant change also occurred. Traditionally, the primary raw material, alumina, was fed to the electrolysis cell periodically (often manually) and in relatively large doses. Environmental requirements to enclose the cells as much as possible and the clear process control benefits of a more

The energy reductions can mainly be attributed to more rapid heat transfer into particles, heat recovery in the cooling stages and from the waste gas heat, resulting in common features such as the pre-heating stages and cooling stages with direct gas solid heat transfer as well as indirect heat recovery in the fluid bed coolers for both Gas Suspension and Circulating Fluidised Bed calciners. The energy consumption should however be compared to the

151

theoretical energy for gibbsite calcination which lies around 1.98 to 2.40 MJ kg"1 for dry and moist gibbsite, respectively [4], indicating that there is still margin for improvements. With the calcination stage amounting to approximately a third of the energy consumed in the Bayer process the potential energy savings becomes significant.

stationary calciners will inevitably span the diagram of reaction pathways and challenge the value of any type of quantitative phase analysis. Nevertheless such an analysis is one of the few comparative tools we have in explaining differences in alumina properties and performance. <€| 100 — — i

Table 1. Typical energy consumptions, residence times, production capacities, alpha alumina contents, calcination temperatures and free heat transfer surface for different calciner technologies [4-8].

Specific energy (MJ / kg product) Material residence time in the hot zone in the furnace (s) Typical calcination temperature (°C) Free heat transfer surface (m 2 /g product) Typical alpha alumina range (wt-%) Typical production capacity (tons per day)

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Circulating Fluidised Bed

Rotary Kiln

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180-300

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1100

950

1100

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1000

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500

600

700

800

900

1000

1100

1200

1300

1400

1500 |K

Figure 1. Thermal decomposition pathways of gibbsite, adapted from Wefers et al. [2]. High intensity, high resolution synchrotron (and neutron) diffraction data has been shown to be extremely useful in terms of monitoring gibbsite crystal growth mechanisms from Bayer liquors in-situ [15], for identifying minor crystallographic impurities [16] as well as for examining the transition aluminas and their transformation reactions [17-20]. However, for deducing information on the structures of the poorly crystalline transition alumina phases, even these techniques face some limitations [21], particularly in complex mixed phase environments such as the MGAs.

Differences in the alpha alumina contents for the different technologies are directly tied to the residence times and calcination temperatures, whereas the higher free heat transfer surface for the modern calciners is related to the high gas velocities and large surface areas. It has been proposed that the heating rate during the calcination may influence the reaction pathways [9, 10], and it is therefore hardly surprising that material from rotary kilns often display different properties [11, 12]. Surprisingly limited effort has however gone into understanding how these differences arise and more specifically what impacts differences in the precursor material have on the calcination reactions and resulting product properties. The speed of the modern calciners, together with the variation of particle size and morphology, is expected to produce a structurally more disordered and heterogeneous material than those produced in rotary kilns. The disorder and co-existence of phases, often within a single particle, poses significant challenges for characterisation and understanding behaviour of these materials. It also challenges our conventional definition of what we call alumina.

Due to the short residence times in modern calciners, deviations from the average transition alumina structures are observed. This is caused by local disorder, which results in significant peak broadening which, combined with the overlapping peaks for the transition aluminas, complicates structural analysis using traditional diffraction methods. Despite the improvements in fitting diffraction data on MGA, made possible through the works by Ashida et al. [22] and Paglia et al. [19], conventional diffraction techniques are limited by the need for crystallographic long range order. Indeed from the analysis of a typical fluid bed MGA, Ashida proposes that what we call A1203 is better described as H2Al10O16 [22]. This is consistent with previous views of the structure of gamma alumina [2] and provides a clear insight into the origin of HF generation in the reduction cell as examined by Patterson et al. [23]. The hydrogen (as residual -OH groups) is a consequence of the increasing undercalcination of alumina, necessary to meet the demand for increased surface area. Thus we gain HF adsorption capacity at the scrubber, at the direct expense of an increased HF burden circulating with the cell gas.

The Science of Alumina Calcination Several excellent reviews on transition alumina phase changes during the calcination of gibbsite and boehmite exist; particularly noteworthy are those of Levin et al. [13], and the overview of Wefers et al. [2]. The reactions have been found to be influenced by several parameters (such as temperature, heating rate, residence time, particle size and morphology, crystallinity, impurities and atmospheric conditions) some of which are directly related to Bayer operations [2, 9, 10, 12, 14]. Figure 1 summarizes the possible reaction pathways for gibbsite dehydroxylation. It should be pointed out that, apart from the initial decomposition to chi alumina, rho alumina or boehmite, the reactions proceed slowly; even if the energetic barriers are overcome prolonged heating is required to reach equilibrium for any of the meta-stable transition alumina forms. Thus the particles formed in modern

The best practice in phase analysis is exemplified in the Rietveld refinement in figure 2 for which the diffraction data was obtained at a synchrotron source. Apart from the obvious overlap of several peaks in the diffraction pattern, the broad and diffuse peaks from the transition aluminas, the presence of additional X-ray amorphous components and the incomplete structural models for the transition aluminas, results in the discrepancies between the fitted (black) and observed (red) spectra. In the distinction

152

between the transition alumina phases (or forms) it is generally agreed that theta alumina is more ordered than delta alumina which again is more ordered than the gamma alumina phase [2426], however, exactly how the transformation into the more ordered forms proceeds is still debated [2, 13, 27]. The formation of gamma - gamma' - delta and theta alumina can also be seen as waypoints on the gradual transformation into a fully ordered state (represented by the thermodynamically stable alpha alumina).

5

13

21

29

37

45

53

61

65

77

85

Figure 2. Synchrotron X-ray powder diffraction patter of a CFB calcined MGA sample and resulting Rietveld refinement results. Note that the shape of the background is a result of the amorphous nature of the quartz capillary used for mounting the sample. The move to stationary calciners has apparently swapped the historical problem of the broad distribution of residence times, and thus phase composition for individual particles, for the problem of distributions of phases within single particles due to the heat transfer constraints in these technologies. As reported elsewhere, Environmental SEM can be used to observe alpha alumina directly in cross sectioned alumina grains through the Charge Contrast phenomenon [28]. The different dielectric properties of the structurally more ordered alpha alumina, compared to the transition alumina forms, result in a contrast difference in the ESEM. This has allowed for new insights into the alumina phase distribution and phase transformation mechanisms within single particles. The results indicate that alpha alumina formation is closely tied to growth morphology, and seems to follow the same pattern as the growth rings revealed in gibbsite cross sections (figure 3). This suggests that local structure and impurities are important in the nucleation and transformation reactions and ultimately the formation and location of alpha alumina within particles. The phase inhomogeneities between and even within particles exemplifies a wider challenge in understanding alumina behavior. The properties reported on a alumina specification sheet represent at best a heavily averaged view, where the outlier populations may be of more significance in, for example influencing alumina dissolution and flowability, than the mean value specified. The dominance of over and under calcined material in the fine particle size fraction is a good example of this [29].

Figure 3. Top left: Alpha alumina observed around the edge of a cross sectioned GS calcined alumina particle. Top right: Alpha alumina formation in a soak calcined alumina grain following the gibbsite growth ring pattern. Bottom left: gibbsite growth rings as revealed by Environmental SEM and corresponding Na distribution (bottom right), obtained using a ToF-SIMS instrument. The same consideration applies to specific surface area, LOI, and the rarely reported, but important, pore size distribution. As discussed more extensively elsewhere [30], the low order transition aluminas (gamma, rho, chi alumina) contain more residual hydroxyls (confirmed by LOI measurements) which inevitably results in more HF being generated upon dissolution. At the same time, the fine pore size in these under-calcined components might be expected to influence their ability to capture HF in the dry-scrubber. It seems that the narrow pores restrict access to internal porosity and readily become blocked (when HF reacts to form oxy-fluorides). This further restricts access to internal sites, thus reducing the capacity and rate of HF absorption. An example of a distinctly bi-modal pore-size distribution in a SGA sample is provided in figure 4. and corresponding HF generation and emissions data are presented and discussed in elsewhere [30].

I 0,0052 °~ 0.004-

0.10 0.08

importance and interest [34]. Another critical performance criterion for the alumina is rapid dissolution in the molten cryolite based electrolyte. To achieve this, sufficient dissolution power (or superheat) is needed, but also the method, amount and frequency of the alumina additions and the quality (dissolvability) of the alumina are of importance [35-38]. The operational stability and feed strategy relies on a consistent alumina quality. However, variations between, and even within, alumina shipments is often a reality. Such variations are frequently not reflected in the specifications of the alumina, making it difficult to anticipate and make process adjustments to accommodate the raw materials variations. Typical outcomes are process fluctuations, sludge formation or other feed related instabilities and on occasion, emissions problems.

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The alumina specification sheet currently represents a contractual quality index in terms of a number of (often historically) defined key properties. It is frequently less helpful as a predictive tool in terms of assisting process optimization and informing the smelter in terms of how the alumina is expected to behave.

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Figure 4. A sample taken from a GS calciner shows a bi-modal pore size distribution indicating the presence of under calcined components as well as over calcined material.

This is then exacerbated by an incomplete understanding at a smelter level of how alumina properties impact on smelter performance. Good examples of this are: the impact of microstructure (particularly specific surface area and pore size distribution) or phase distribution between particle size classes on HF generation, dissolution and dry scrubbing.

The Future of Alumina Calcination With the increasing number of green- and brownfield alumina refineries it is no surprise that research today focuses primarily on production gains and energy savings through process and control improvements. Typical energy consumption is around 11.6 GJ per metric ton of A1203. However, the broad range of energy consumption across the industry [31], indicates that there is still room for significant energy savings in a number of operations and process areas. As the higher grade bauxite reserves are being depleted the industry increasingly turns to more energy intense and difficult to process diasporic and boehmitic ores.

These properties are primarily influenced by precipitation and calcination strategies but these relationships are complex. Most critical is the "averaging" impact of numbers reported in the specifications sheet where the outlier populations may be the dominant contributor to process fluctuations and instabilities in the smelter. References 1.

In the past decades alumina production volumes have increased dramatically, with calciner operating capacities up to 4000 metric tonnes per day. As a result of the push towards faster and more energy efficient, but also more abrasive, stationary calcination processes, attrition continues to receive a lot of attention [32, 33]. Many options and improvements have been developed to the calcination technologies to reduce breakage and improve product quality. Not all, but some of these improvements come with the price of higher operating costs. Thus research programmes are increasingly focusing on a better understanding of the role of the precipitation step and resulting particle morphology and gibbsite strength, for the alumina attrition behaviour during calcination and product handling. There is also a wider consensus emerging that new definitions for how attrition is defined, and indeed measured, are needed, as again the specification sheet is not always informative as to smelter experience. The next stages of this discussion should include the processes downstream of calcination in alumina handling, and also at the smelter to get a more holistic view. This informs the optimum compromise between costs, product quality and stable operation.

2. 3. 4. 5. 6.

7.

In the smelter surprisingly few studies systematically examine the impact of alumina quality (or properties) on operations, although the impact of alumina on HF emissions is an area of increasing

8.

154

Edwards, J.D., F.C. Frary, and Z. Jeffries, The Aluminum Industry - Aluminum and Its Production. Chemical Engineering Series. 1930. Wefers, K. and C. Misra, eds. Oxides and Hydroxides of Aluminum. Alcoa Technical Paper No. 19. 1987, Aluminum Company of America: Pittsburgh, PA. Andrews, E.W., A Controllable Continuous Mass-feed System for Aluminium Smelters. PhD Thesis, The University of Auckland, New Zealand, 1998. Hudson, L.K., ed. Alumina Production. Alcoa Research Laboratories. 1982, Aluminum Company of America: Pittsburgh, PA. Jenkins, B. and C. Bertrand, Improvements in the Design and Operation of Alumina Flash Calciners. IFRF Combustion Journal, 2001 (November): p. 1-18. Hiltunen, P., R. Bligh, C. Klett, M. Missalla, and H.-W. Schmidt, How to achieve high availability with large calciners and avoid unforeseen downtime. Light Metals (Warrendale, PA, United States), 2008: p. 63-68. Raahauge, B.E., Advances in gas suspension calcination technology. Aluminium (Isernhagen, Germany), 2007. 83(1/2): p. 40-42. Mclntosh, P., R. Greenhalgh, and P. Mills, Advanced control techniques for alumina calcination rotary kilns.

9.

10.

11.

12.

13.

14. 15.

16.

17.

18.

19.

20.

21.

22.

Light Metals (Warrendale, PA, United States), 1987: p. 59-63. Ingram-Jones, V.J., R.C.T. Slade, T.W. Davies, J.C. Southern, and S. Salvador, Dehydroxylation sequences of gibbsite and boehmite: study of differences between soak and flash calcination and of particle-size effects. Journal of Materials Chemistry, 1996. 6(1): p. 73-9. Whittington, B. and D. Ilievski, Determination of the gibbsite dehydration reaction pathway at conditions relevant to Bayer refineries. Chemical Engineering Journal (Amsterdam, Netherlands), 2004. 98(1-2): p. 89-97. Metson, J., T. Groutso, M. Hyland, and S. Powell, Evolution of microstructure and properties ofSGA with calcination of Bayer gibbsite. Light Metals (Warrendale, PA, United States), 2006: p. 89-93. Yamada, K., T. Harato, S. Hamano, and K. Horinouchi, Dehydration products of gibbsite by rotary kiln and fluid calciner. Light Metals (Warrendale, PA, United States), 1984: p. 157-71. Levin, I. and D. Brandon, Metastable alumina polymorphs: crystal structures and transition sequences. Journal of the American Ceramic Society, 1998. 81(8): p. 1995-2012. Bennett, I. and R. Stevens, Calcination and phase changes in alumina. British Ceramic Transactions, 1998. 97(3): p. 117-125. Loh, J.S.C., A.M. Fogg, H.R. Watling, G.M. Parkinson, and D. O'Hare, A kinetic investigation of gibbsite precipitation using in situ time resolved energy dispersive x-ray diffraction. Physical Chemistry Chemical Physics, 2000. 2(16): p. 3597-3604. Latella, B.A. and B.H. O'Connor, Detection of minor crystalline phases in alumina ceramics using synchrotron radiation diffraction. Journal of the American Ceramic Society, 1997. 80(11): p. 2941-2944. O'Connor, B., D. Li, B.K. Gan, B. Latella, and J. Carter, Time-resolved studies of alumina ceramics processing with neutron and synchrotron radiation data. Advances in X-Ray Analysis, 1999. 41: p. 659-667. Neissendorfer, F., U. Steinike, B.P. Tolochko, and M.A. Sheromov, On the decomposition of hydrargillite investigated by synchrotron x-ray diffraction. Nuclear Instruments & Methods in Physics Research, Section A: Accelerators, Spectrometers, Detectors, and Associated Equipment, 1987. A261(l-2): p. 219-20. Paglia, G., CE. Buckley, A.L. Rohl, R.D. Hart, K. Winter, A.J. Studer, B.A. Hunter, and J.V. Hanna, Boehmite Derived Gamma-Alumina System. I. Structural Evolution with Temperature, with the Identification and Structural Determination of a New Transition Phase, Gamma'-Alumina. Chemistry of Materials, 2004.16(2): p. 220-236. Zhou, R.S. and R.L. Snyder, Structures and transformation mechanisms of the eta, gamma and theta transition aluminas. Acta Crystallographica, Section B: Structural Science, 1991. B47(5): p. 617-30. Billinge, S.J.L. and I. Levin, The Problem with Determining Atomic Structure at the Nanoscale. Science (Washington, DC, United States), 2007. 316(5824): p. 561-565. Ashida, T., J.B. Metson, and M.M. Hyland, New approaches to phase analysis of smelter grade

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aluminas. Light Metals (Warrendale, PA, United States), 2004: p. 93-96. Patterson, E.C., Hydrogen Fluoride Emissions From Aluminium Electrolysis Cells. PhD Thesis, The University of Auckland, New Zealand, 2002. Jayaram, V. and C.G. Levi, The structure of deltaalumina evolved from the melt and the gamma -> delta transformation. Acta Metallurgica, 1989. 37(2): p. 56978. Wilson, S.J., The dehydration of boehmite, γ-ΑΙΟΟΗ, to γ-alumina. Journal of Solid State Chemistry, 1979. 30(2): p. 247-55. Wilson, S.J. and J.D.C. McConnell, A kinetic study of the system boehmite/alumina (γ-ΑΙΟΟΗ/Α1203). Journal of Solid State Chemistry, 1980. 34(3): p. 315-22. Wolverton, C. and K.C. Hass, Phase stability and structure of spinel-based transition aluminas. Physical Review B: Condensed Matter and Materials Physics, 2001. 63(2): p. 024102/1-024102/16. Perander, L., C. Klett, H. Wijayaratne, M. Hyland, M. Stroeder, and J. Metson. Impact of Calciner Technologies on Smelter Grade Alumina Microstructure and Properties, in Proceedings of the 8th International Alumina Quality Workshop. 2008. Darwin, Australia. Perander, L.M., Z.D. Zujovic, T.F. Kemp, M.E. Smith, and J.B. Metson, The Nature and Impacts of Fines in SGA. Journal of Metals, 2009. 61(11): p. 33-39. Perander, L.M., M.A. Stam, M.M. Hyland, and J.B. Metson, Towards Redefining the Alumina Specifications Sheet - The Case of HF Emissions. Light Metals (Warrendale, PA, United States), 2011. Henrickson, L., The need for energy efficiency in bayer refining. Light Metals (Warrendale, PA, United States), 2010: p. 173-178. Klett, C , M. Missalla, and R. Bligh, Improvement of product quality in Circulating Fluidized Bed calcination. Light Metals (Warrendale, PA, United States), 2010: p. 33-38. Wind, S., C. Jensen-Holm, and B.E. Raahauge, Development of particle breakdown and alumina strength during calcination. Light Metals (Warrendale, PA, United States), 2010: p. 17-24. Iffert, M., M. Kuenkel, M. Skyllas-Kazacos, and B. Welch, Reduction of HF emissions from the trimet aluminum smelter (optimizing dry scrubber operations and its impact on process operations). Light Metals (Warrendale, PA, United States), 2006: p. 195-201. Dando, N., X. Wang, J. Sorensen, and W. Xu, Impact of thermal pretreatment on alumina dissolution rate and HF evolution. Light Metals (Warrendale, PA, United States), 2010: p. 541-546. Liu, X., S.F. George, and V.A. Wills, Visualization of alumina dissolution in cryolitic melts. Light Metals (Warrendale, PA, United States), 1994: p. 359-64. 0stb0, P., N, Evolution of Alpha Phase Alumina in Agglomerates upon Addition to Cryolitic Melts. PhD Thesis, Norwegian University of Science and Technology, Trondheim, Norway, 2002. Welch, B.J. and G.I. Kuschel, Crust and alumina powder dissolution in aluminum smelting electrolytes. Journal of Metals, 2007. 59(5): p. 50-54.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Significant Improvement of Energy Efficiency at Alunorte's Calcination Facility Michael Missalla1, Hans-Werner Schmidt1, Joaquim Ribeiro Alves Filhio2, Reiner Wischnewski2 outotec GmbH, Ludwig-Erhard-Str. 21, 61440 Oberursel, Germany 2 Alunorte-Alumina do Norte do Brasil S.A., Rodovia PA 481, km 12, Distrito de Murucupi, CEP 68 447-000 Barcarena, PA Brazil Keywords: Circulating Fluidized Bed, Calcination, Energy Efficiency, Product Quality, Particle Breakage Abstract The Alunorte refinery produces 6.3 million t/a of alumina with seven Circulating Fluid Bed (CFB) Calciners. The calcination facility needs about 3 GJ per ton of alumina of the total energy of 8 GJ per ton for the refinery. Until the introduction of CFB Calciners by Outotec (formerly Lurgi) in 1960, rotary kilns were the standard technology for the calcination of alumina. Since then, stationary calciners such as CFB s are the preferred technology for new installations due to their superior energy efficiency and uniform product quality.

To achieve significant lower energy consumption than rotary kilns, CFB Calciners utilize much more energetic means of gassolids contact and gas solid separators (such as cyclones) with significantly higher velocities and hence strong mechanical stress on the particles. With rising energy prices it became also more favorable to increase the number of heat exchanger stages. The number of heat exchanger stages it is very important to have an efficient particle mixing into the gas stream [9]. Further a high separation efficiency of the cyclones in the heat exchanger stages is necessary to achieve low energy consumption.

The CFB calcination system has implemented several preheating and cooling stages with cyclones for the separation of gas and solids. The inefficiency of these cyclones is leading to a reduction in the heat recovery and increased energy consumption.

The cyclone inlet velocity can be increased to achieve improved separation efficiency. However because of attrition, this leads to higher generation of fines in the process and has the result that the hydrate produced by the alumina refinery must be stronger and contain less fines in order to meet the final product quality specification. While this has certainly been achieved in many refineries, nevertheless, there are additional constraints placed on refinery performance indicators (e.g. yield) [11].

The paper presents a procedure to improve the efficiency of these cyclones and the solid dust load significantly by using a new method of simulation technology. With the optimized process considerably lower energy consumption figures are achieved. The results of these improvements regarding fuel and electrical energy consumption, as well as return on investment are presented. Alunorte has installed the improvements already at two of their seven Calciners and has received the energy efficiency award 2010fromthe German Energy Agency.

Since the introduction of CFB Calciners, a major part of Outotec's calcination research has been dedicated to reducing fines generation without compromising on energy consumption or other aspects of product quality (e.g. attrition index)[2]. While Outotec has been steadily improving its technology to reduce fines generation for many years, the latest installations represent a significant step improvement, with the result that the particle breakage in CFB Calciners has been reduced to only marginally higher than in rotary kilns and much lower than in earlier installations. Also the energy consumption has been reduced significantly and the latest alumina calciner installed in Alunorte has received the energy efficiency recognition prize at the industrial technology exhibition the Hannover Messe 2010 in Germany.

Introduction Product quality is always a major focus for every industrial operation, with only production safety being on the same level or even more important. This applies to the production of alumina calcined from Aluminium Hydroxide with the same intent. The calcination step is the last step in the chain of process steps to produce alumina from bauxite. Until the 1960s the calcination reaction was always performed in rotary kilns. In 1961 Outotec (then known as Lurgi) invented the circulating fluidized bed technology, and applied it to the calcination of alumina in CFB Calciners. With this new technology, energy consumption for alumina calcination was immediately cut by approx. 30% and fluidized bed has been the technology of choice ever since.

The following main improvements have led to the energy reduction in the calcination process: • •

The new technology had quite some impact on product quality. While criteria such as LOI, BET, pick-up of iron and silica kept constant, the production of alpha alumina was significantly decreased while fines generation was increased. Although reduction of alpha alumina is basically considered to be an improvement in quality, increased fines is not. The mass fraction of particles smaller than 45μπι is a common measure of fines content and should be as low as possible but certainly within the limits specified by the smelter customers.

•

Improved Cyclone design to reduced the recirculation of fines and dust in the calcination system Improved Cyclone design to reduce the generation of fines Optimization of process stability for stable operating conditions Improvements at Alunorte

Alunorte S.A. was founded in 1978. Production started in 1995 and since then its alumina production capacity increased continuously [14]. Currently Alunorte S.A. is the world largest alumina refinery with a total production capacity of 6.3 million tones of alumina per year.

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hot gas. In the cyclone, which is arranged downstream of the venturi preheater, the flow of gas and solids are separated. From the cyclone the pre-calcined alumina is delivered into the fluid bed furnace through the material feeding line. The waste gas leaving the cyclone of the second preheating stage is conveyed to the first preheating stage where it assumes the above-mentioned function. The final calcination of the preheated and partly dehydrated hydrate takes place in the fluid bed furnace as shown in figure 3.

Figure 1: Alunorte S.A. 6.3 million tons per year alumina refinery Figure 1 shows the alumina refinery of Alunorte S.A. located in Barcarena, Brazil. With its latest expansion Alunorte S.A. added two alumina calcination plants from Outotec each with a name plate capacity of 3300 tons per day (tpd).

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Figure 3: CFB Calciner Flowsheet of recently installed calcination plants at Alunorte S.A. The required heat for calcination is generated by direct combustion of fuel oil in the fluid bed. Part of the air required for combustion is introduced through the nozzle grate as primary air, and the remaining air is added above the grate as secondary air. Due to the intensive mixing and heat exchange that takes place in the fluidised bed, the furnace temperature adjusts itself as a mixed temperature between the combustion temperature and the solids temperature, and is kept steady at a pre-set level. In the lower furnace zone, between the grate and the secondary air inlet, a fluid bed of high solids concentration is adjusted. It favors the combustion of the fuel and increases the mean retention time in the calcining furnace. In the upper furnace zone the internal recirculation of the solids causes a continuous reduction of solids concentration until a relatively low concentration is reached. With this solids concentration, the hot gases enter the recycling cyclone where they are separated from the solids. The hot alumina, which is separated in the recycling cyclone, passes through the seal pot and re-enters the fluid bed furnace [12],[13].

Figure 2: View of the seven CFB calciners at Alunorte S.A. In figure 2 the seven alumina calcination plants of Alunorte S.A can be seen. The wet aluminium trihydrate is conveyed by belt conveyors to the hydrate feed bin of the calcination plant. From the hydrate feed bin, the hydrate is discharged by a screw feeder, which delivers the material into the venturi preheater of the first preheating stage. There, the solids are mixed with the waste gas, which leaves the cyclone of the second preheating stage. The heat contained in the waste gas evaporates the entire surface moisture of the hydrate.

The recirculation of the solids leads to uniform and practically identical product and gas temperatures in the entire calcination stage, which consists of the fluid bed furnace, recycling cyclone and seal pot. The alumina, which is discharged from the calcining stage is cooled in two direct cooling stages each of which consists of a liftduct and a secondary air cyclone. The third cooling stage is designed as a fluid bed cooler and mainly relies on indirect heat transfer. The alumina coming from the calcining stage is first mixed with pre-heated air in the liftduct and conveyed into the secondary air cyclone 1. At the same time, the air is heated from the temperature of the second cooling stage to the temperature of the first cooling stage. This air is routed from the outlet of the secondary air cyclone 1 to the CFB system as secondary air and

Then the preheated hydrate entrained with the waste gas is conveyed into the two-stage electrostatic precipitator (ESP). The solids precipitated in the ESP are conveyed to the venturi preheater of the second preheating stage. The conveying air is dedusted in the airlift cyclone and delivered via ducts to the secondary air cyclone and used finally as combustion air. The hydrate entering the venturi preheater of the second preheating stage is mixed with the hot waste gas leaving the recycling cyclone and is dehydrated by the heat contained in the

158

To calculate the grade efficiency curve for a cyclone, a model was used which is shown in figure 4. The cyclone is divided into different separate regions which are:

used for combustion. The same applies for the second direct cooling stage. In order to achieve low specific heat consumption during the calcining process, there is further heat recovery in the air-cooled part of the fluid bed cooler. Primary air is pre-heated by indirect heat exchange in a counter-current flow to the alumina. Secondary air is also preheated by direct contact with the alumina. Further heat recovery occurs in the water-cooled section of the fluid bed cooler as secondary air is preheated in the air-cooled section. Any part of the heat content of the alumina, which cannot be utilized for preheating of air or liquid media is removed by means of cooling water.

o o o o o

For the above mentioned regions, differential particle mass balances are solved to calculate the grade efficiency curve for the cyclone. The model considers the different separation efficiencies for the main flow V and the secondary flow Vsek over the top and the wall of the vortex finder. In the model the re-entrainment of already separated particles at the solid outlet is also considered at the re-entrainment region 3.

With the process described above the specific energy consumption at Alunorte has been reduced to 2790 kJ/kg Alumina.

Australian Average

3,900

Previously lowest recorded value m World

I

3,000

1,000

1,500

2,000

2,500

3,000

3,500

4,000

m

kJ/kg Alumina

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Cyclones are firstly named in 1886 in a patent of Knickerbocker Company, USA [5]. Since then the design and calculation of cyclones have been improved significantly. Basically cyclones are used to separate particles from a fluid or gas by centrifugal force. Therefore the dust loaded gas stream is fed into the cyclone where a rotary motion is induced. The solid particles, due to their higher density, are separated to the wall of the cyclone and flow downwards to the solid exit, while the gas leaves the cyclone via the vortex finder. The flow field in a cyclone is highly complex and can be differentiated in:

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m im

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Inlet flow Mainflowwith o Tangential velocity o Redial velocity o Axial velocity Secondary Flows with o Top Secondary flow o Vortex Finder Secondary Flow o Wall Secondary Flow Vortex Finder Flow

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Cyclone improvement

•

w

m

In Alunorte S.A. the outstanding performance with respect to the specific energy consumption was reached by improved design of the cyclones which are used for the separation of the solids from the offgas and the combustion air as described above. Furthermore, the process stability was improved significantly which also contributed to the reduction in specific energy consumption.

•

ir F

if I

Figure 3: Specific energy consumption according to CSIRO National Framework for Energy Efficiency, 2007 including Alunorte S.A. calciner.

• •

Inlet region e Separation region 1 and 11 Secondary Flow regions at the top d and at the vortex finder tr Separation region 2 and 4 Re-entrainment region 3

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159

In CFB calciners as well as in flash calciners the typical solid loading in the preheating and cooling cyclones inlets is high. Due to the very high solids concentration, it dominates the flow field in the cyclone. It has been found that with higher solid loading the tangential velocity decreases significantly in a cyclone [8].

strand formation. Figure 6 shows the effect of the formation of strands in the inlet of a cyclone. It can be seen that increasing the load in the inlet of the cyclone enhances the separation due to the formation of strands. loo i 0.90

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Measured feed to Cyclone

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Figure 7: Calculated and measured particles in the offgas of a cyclone at the Alunorte calciner Figure 5: Relative tangential velocity of a loaded to an unloaded cyclone as function of the solids loading μ* in the inlet stream

Figure 7 shows the measured particle size distribution in the inlet as well as in the offgas of a cyclone in the Alunorte calcination plant. It can be seen that the measured particle size distribution fits quite well with the calculated curve from the model, thus allowing the optimization of the cyclones.

Figure 5 shows that the tangential velocity of a cyclone with high solid loads in the inlet drops significantly. A cyclone with a solid load of μ*=1 in the inlet has just 60 % of the tangential velocity of a cyclone with no solid load (μ*=0) in the inlet. Thus it becomes very difficult to separate fine particles since the tangential velocity is reduced substantially. Muschelknautz [10] showed that the separation in high loaded cyclones is mainly driven by the

Figure 6: Formation of Strands in a cyclone at different inlet loads: a) μ*=0,01 b) μ*=0,1, c) μ*=1, d) μ*=10 [10]

160

The process flowsheet shown in figure 3 has four cyclones of which one is in the preheating stage for the hydrate, one is in the CFB and two are in the preheating stages for the combustion air. The integration of the ESP in the overall CFB process ensures that all dust is collected and returned to the process. In other stationary calciner systems the dust returns to the fluid bed cooler [9]. The disadvantage of the latter method is that it cannot be ensured that all the hydrate is completely calcined.

Operating Stability In order to achieve the best results for the separation in a cyclone the operating stability of the process becomes more and more important. Over the last few years, [3] [4] the stable operation of CFB calciners has been in focus and great improvements have been made up to the point of complete automation of the plants. Figure 8 shows the ability to maintain calcination temperature in the furnace under all circumstances. The modern automation approaches enable the operators to change the load of the plant in a flexible manner and also to keep the furnace temperature within arangeof+/- 10tol5°C.

According to the grade efficiency curve of a cyclone, certain fractions of a specific particle size will not be separated. Assuming two cyclones in series, as in the flowsheet in figure 3 for the secondary air cyclones, the fine particles not separated by the second secondary air cyclone serve as additional solid feed for the first secondary air cyclone. So the first secondary air cyclone is fed with the fresh solid feed from the CFB and with the fine particles from the second secondary air cyclone. Thus the efficiency of the first is reduced.

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Furnace Temperature

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Applying this logic further, then due to the fines in the offgas, the other upstream cyclones are reduced in their separation efficiency. In table 1 a comparison between CFB systems with two and three cyclone cooling stages with different total cyclone separation efficiencies ηιοί is made. A total separation efficiency of r|tot =100 % means that no particles are leaving the cyclone with the offgas while ritot = 90 % means that 10 % of the particles been fed to the cyclones are leaving it with the off gas.

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Table 1 : Specific energy consumption normalized to the system 1 with the highest energy consumption Number of Furnace Temp. °C Normalized nto %t Cooling Energy Stages Consumption 2 950 100 953% 90 96.1% 3 950 92.4% 100 93.4% 90 80 94.8%

The temperature stability in the CFB calciner is very important for the separation efficiency of heat exchanger cyclones in the flowsheet, since the separation efficiency is negatively affected by increasing temperature. Also an increase in particle breakage due to higher furnace temperatures can be observed [2]

* 45 S 40

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The results in table 1 show that cyclone efficiency is of utmost importance to reduce recirculation in the process and thus the energy consumption. Comparing a CFB system with two cooling stages and a ηιοί =100 % with a CFB system of three cooling stages with not optimal total cyclone separation efficiency of r|tot =80 % the result for the specific energy consumption is almost the same.

• ^ 0

200

· 400

600

800

1000

Calcination Temperature, °C

With the model as described in figure 4 the optimum vortex finder length can be found to increase the separation of fine particles. With increasing height H of the cyclone its separation efficiency is improved. Finally also increased recirculation of fines in a calcination process leads an increased amount of alpha alumina in the system. Due to its stickiness, this type of alumina can lead to processfluctuationswhich finally affect the separation efficiency of a cyclone negatively.

Figure 9: Friability of alumina calcined at different calcining temperature. From Figure 9, between 800 and 1000°C the particle breakage changes approx. 1.2% for every 10°C increase. With a value of 40% in the Saatci Test [2] at 1000°C this means a ratio of 3% on the actual breakage. For a plant with a normal particle breakage of 3.5% this would correspond to a variation of 0.1% for each temperature variation of 10°C. To control the particle breakage means to control the stability of furnace temperature.

The energy efficiency of a calcination process does not only depend on the number of cooling stages but also on its total separation efficiency. An investment comparison of the different system can only be done when the weights of the cyclones are known.

However the dependency could also be more than proportional, which would mean a disproportionate increase of breakage with increasing temperatures. In this case the fluctuations of calcination temperature around the set-point would be a

161

demonstrates Outotec's capabilities to design optimum calcination technology for energy efficiency and product quality.

disadvantage to the overall breakage, as every increment of temperature would cause more damage than a decrease could compensate. The effect of increased particle breakage in the calciner and thus the generation of more fines is an increased specific energy consumption due to the negative effect on cyclones performance.

References [I] J. D. Zwicker, "The Generation of Fines due to Heating of Alumina Trihydrate", TMS Light Metals (1985)

Future Developments

[2] S. Saatci, H.W.Schmidt, W. Stockhausen, M.Ströder, P.Sturm, " Attrition behaviour of Laboratory calcined Alumina from various Hydrates and its influence on SG Alumina Quality and Calcination Design", TMS Light Metals pp 81-86 (2004)

Beside designs for low fuel consumption [6], [7] the existing plants can be further improved by adding a Hydrate Bypass as shown infigure9.

[3] P. Hiltunen, R. Bligh, C. Klett, M. Missalla, H.-W. Schmidt, "How to achieve high availability with large calciners and avoid unforeseen downtime", TMS Light Metals pp 63 - 68, (2008) [4] M. Missalla, J. Jarzembowski, R. Bligh, H.-W. Schmidt, "Increased availability and optimization of calciner performance due to automation", TMS Light Metals , (2009) [5] Knickerbocker Company, " Staubsammler", in DRP Nr. 39219 (1886)

Primary Air

**

[6] C. Klett, M. Missalla, R. Bligh, M. Stroeder, P. Hiltunen, M. Graham, "Outokumpu Technology's State-of-theArt CFB Calciners - A Review", Proceedings to EMC (2007)

»

Cooling Wartw - # »

1

Figure 10: CFB Calciner Flowsheet with Hydrate Bypass

[7] C. Klett, M. Missalla, R. Bligh, "Improvement of Product Quality in Circulating Fluidized Bed Calcination", TMS Light Metals pp 33 - 38, (2010)

Figure 10 shows a CFB Calciner Flowsheet as it is installed at Alunorte S.A. but with the addition of a Hydrate Bypass. The upgrade with a hydrate bypass will further decrease the specific energy consumption of the already for its lowest consumption awarded plant. With the hydrate bypass pre-dried hydrate from before preheating stage 2 will be bypassed around the CFB to a mixing pot. In the mixing pot the hydrate is calcined with by the hot alumina leaving the CFB. Beside using the energy from the CFB the efficiency of the cooling cyclones is improved since they are operating with decreased temperature.

[8] M. Missalla, "Calculation Method for Highly Loaded Cyclones", Ad.libri Hamburg, Dissertation Technische Universität Braunschweig (2009) [9] S. Wind, B.E. Raahauge, "Energy Efficiency in Gas Suspension Calciners (GSC)", TMS Light Metals pp 235 - 240, (2009) [10] U. Muschelknautz, „Hart bleiben Verschleißschutztechnik in schüttguttechnischen Anlagen" Chemie Technik Nr. 10 (2005), pp 14-18

However, the by-pass is a major part of the energy saving methodology and can operate to up to 15% of the total dry hydrate flow. As a result, a few effects happen simultaneously. Firstly, only the remaining part of the hydrate flow will actually pass through the CFB and undergoes breakage there. Secondly the recombined flows of alumina originating from the furnace material and from the by-pass material have now been calcined at different temperatures. According to the temperature effect from above, the by-pass material is significantly less brittle and will lead to a reduction of breakage in the cooling stages. This effect is even large enough to compensate for any temperature increase in the furnace to achieve the target LOI. The specific energy consumption is improved by approx 3-5%

[II] E.Guhl, R.Arpe, „Nearly 30 years of experience with Lurgi Calciners and influence concerning particle breakage", TMS light Metals (2002) [ 12] L.Reh," New and Efficient High Temperature Process with Circulating Fluid Bed Reactors", Chem. Eng.Technol. 18 (1995) page 75-89 [13] A. Squires, "Origins of the Fast Fluid Bed, Advances in Chemical Engineering", Vol. 20, Fast Fluidisation (Ed. Kwauk, M.), Academic Press (1994), pp 4-35

Conclusions

[14] Daryush Albuquerque Khoshneviss, Luiz Gustavo Correa, Joaquim Ribeiro Alves Filho, Hans Marius Berntsen, Ricardo Rodrigues de Carvalho, "Alunorte Expansion 3 - The New Lines Added to Reach 6.3 Million Tons per Year", TMS Light Metals 2011.

With the recent research activities and developments Outotec has made a significant effort to understand cyclone design, process stability and particle breakage in CFB Calciners. The recent success with the latest CFB Calciners Outotec has received the energy efficiency award 2010fromthe German Energy Agency. It

162

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ATTRITION OF ALUMINA IN SMELTER HANDLING AND SCRUBBING SYSTEMS Stephen J. Lindsay Alcoa, Inc.; Primary Metals; 300 N. Hall Rd. MS S-29, Alcoa, Tennessee, 37701-2516, USA Keywords: Alumina, SGA, Attrition, Attrition Index, Gas Treatment Center, GTC Abstract

Alumina Attrition Mapping

Customers place various levels of importance upon the Attrition Index of Smelter Grade Alumina. The concerns are generally associated with the content of fines in fluorinated alumina as it arrives to the reduction cell. In this paper the author discusses factors of importance related to the design and operation of alumina handling systems from the refinery to the reduction cells. Examples are given in which the actual attrition of alumina particles has been minimized. Techniques are shared on how to separate the contribution of fine particles of bath evolved by the pots from the attrition of alumina itself.

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Introduction Many producers of smelting grade alumina, SGA, include data on % Attrition Index, or %A.L, in each certificate of analysis. This is one metric for assessing alumina particle "toughness". Although suppliers maintain strict standards to determine values of attrition index, the methods used to determine %A.I. do vary. At this time there is no single industry standard for measurement of % Attrition Index.

Figure 1 - Generic example of Attrition Mapping A sampling regimen was prescribed for the participating smelters. The goal was to be able to gather composite samples that would be as representative as possible of alumina at various locations along its flow path toward the pots. Some of these were straightforward sampling plans that focused on a few key sampling locations. Other sampling plans were more detailed. Complexity beyond the basic sampling points was left to each participating location.

There have been contributions to the literature that help to illustrate the topic of attrition, or particle toughness. Particular significance has been given to particle velocity at the time of impact against an obstacle such as a steel plate [1]. But, what does the %A.I. mean to a smelting customer of SGA? Can it be used with the particle size distribution, PSD, of an alumina source to predict the PSD of secondary alumina at the reduction cell? Is it of greater significance for some screen mesh fractions than others? By definition %A.I. is a measure of change in the %+325 mesh. But, is it meaningful for the %-20 micron fraction, or superfines?

Analysis of the gathered samples was made in cooperation with the Technology Delivery Group of Alcoa World Alumina. This included the determinations of particle size distributions using a Coulter Particle Size Analyzer model LS100Q laser diffraction instrument. Interpretation of superfines content was based upon electroformed sieve calibration samples that are periodically restandardized. Measurements of flowability were also made by use of a standard Alcoa flow funnel with a 6 mm orifice.

The best answer was probably inscribed on the portico of the Temple of Apollo at Delphi, "Know thyself. The particle size distribution at reduction cells appears to have more to do with the design and operation of alumina handling systems and dry scrubbers than with the source of the alumina and its particular morphology or particle toughness. When a smelting customer understands their system quite clearly, then some detailed meaning may be able to be ascribed to the %A.I. of an individual alumina source.

This analytical capability allowed the study to go beyond other studies in the literature that have relied upon dry sieving analytical methods. This enabled many of the conclusions that follow on the nature of pot fume and of the amount of attrition that occurs in various types of alumina handling and processing equipment. Results were summarized according to typical particle size fractions. These were mapped to visually illustrate changes in the percentages of each fraction along the alumina flow path. Particular attention was given to superfine fractions. This approach was used to enable conclusions about alumina attrition. It was understood from the outset that increases in superfine fractions might be confounded by thefineparticles of alumina and bath that are found in reduction cell exhaust.

Experimental Design In mid-2008 more than a dozen Alcoa smelters participated in an "Attrition Mapping" exercise. The exercise is generally described by Figure 1. The activity was designed to allow its participants to understand how much alumina particles break down as they pass through their alumina handling and dry scrubbing systems.

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No effort was made to determine the chemical composition or structure of each sample or fraction. This additional dimension of study is left for future consideration. It was not within the boundaries of the test design for this Attrition Mapping exercise.

Sampling at or after point feeders in pots was not advised due to the impacts of dust collection and gas flow patterns in cells. This may have introduced significant bias in superfine fractions that could not be easily controlled by the experimental design.

The design of the study was to create composite samples from various places along the alumina flow path. The minimum requirement to participate was to sample:

Composite samples were then placed into plastic containers with screw down lids. They were clearly marked prior to shipment for analysis.

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Particle size determinations were all made in the same laboratory, using the same equipment, people, and analytical methodology.

Fresh alumina as received Fresh alumina prior to the inlet of the dry scrubber Secondary alumina after the discharge point of the dry scrubber silo Secondary alumina at, or in close proximity to, the reduction cell

Results Study #1 - The first handling system to be analyzed included a nearly ideal configuration for minimal attrition. Handling systems included; air gravity conveyors, silos, one air lift for fresh alumina and another for secondary alumina, a fluid bed type gas treatment center and air gravity conveyors to pot bins.

As indicated, additional sampling locations were added to this baseline if so desired by participating locations. Sampling

The %A.I. of the alumina source was between 15% and 17% during the period of study. This moderately high %A.I. was determined by the Alcoa method for attrition. See equation 1.

Obtaining representative samples of alumina in an industrial setting can be challenging. Fine and superfine fractions can easily be biased via improper sampling technique. As this study was conducted over an array of smelting locations care was taken to prescribe detailed sampling methods. However, direct oversight of sampling at each location was not conducted by any one person or group.

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Sampling was conducted by taking approximate 100 gm samples over multiple days to build composite samples of approximately 1 kg from a minimum of 10 grab samples. When possible these samples were to be taken directly from a conveyor belt or by using a simple grain thief device. Refer to Figure 2.

Results indicated no significant degradation of particle size diameters at the 10th, 50th or 90th percentiles of the PSD. These are otherwise referred to as dlO, d50 and d90. See Figure 3.

Figure 3 - Attrition mapping results of study location #1 Most surprisingly there was essentially no change in dlO at -40 microns diameter along the flow path to the pots. Calculation of the actual %A.I. in the field yielded a result of only 0.7%. This result was roughly 1/20* of the % A.I. reported for fresh alumina.

Figure 2 - Example of a grain thief sampling device It was advised to avoid taking samples from conveying equipment that was under negative pressure while using any open sampling devices such as scoops or cups. In these circumstances a sampling thief was to be used with a fabric seal around the tube at the sampling access point. An alternative approach was to briefly shut down the system to take a sample. This required shutting off dust collection simultaneously with conveying equipment.

Even though the handling systems and fume controls were known not to be harsh on alumina particles, such low results were not expected. There was no apparent attrition of alumina particles along the flow path through the smelter. However, these results were not in concert with the general observation of reduced

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In Study #1 48%+/-2% of the sub-2(ty material was sub-ΙΟμ in diameter and 30%+/-3% was s u b ^ in diameter. These proportions were consistent over the entire flow path from fresh alumina to the pots. Note that 72% of the sub-2(^ material found in alumina at the pot was originally present in the fresh alumina.

flowability according to Alcoa flow funnel test results for Study #1. See these results in Figure 4. Study #i; Alcoa Flow Funnel Time

Further, the data indicates that 28% of the sub-2(^ material at the pot has a similar PSD to the sub-2(^ material found in the fresh alumina. Note that alumina "at the pot" may be a bit of a misnomer. There were no attempts made to actually capture and analyze alumina that was being delivered to a feed hole in the pot crust. The reason is that some dust generated during feeding is lost to the pot exhaust and is not delivered into the pot per se.

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Study #2 - Attrition mapping Study #2 also had a nearly ideal configuration for minimal attrition. Handling systems with this study included; air gravity conveyors, silos, one air lift for fresh alumina and another for secondary alumina, a fluid bed type gas treatment center and air gravity conveyance to pot bins.

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As with Study #1 the dlO and d50 particle diameters remained essentially the same. However, there was a more pronounced difference at d90. Refer to Figure 7 to see this difference.

Figure 5 - SuperfinefractionsfromStudy #1 A number of interesting observations are associated with this data. There is an apparent increase in superfines content prior to the gas treatment center, or GTC. This may be attributed to factors such as: variation in sampling, segregation in the fresh alumina silo, or perhaps minor attrition in the airlift. Note that these flow path studies did not follow specific lots of alumina along the flow path. Samples were gathered at various points in the flow path all but simultaneously. The patterns of %-20μ, %-10μ and %-5μ in secondary alumina match each other quite well as is shown in Figure 5. This infers that the increases have been driven by the same mechanism. The most likely source of these particles is the fine particulate evolved with pot fume. This is illustrated in a typical PSD for fume particulate that is shown in Figure 6 [2].

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Study #3 - Attrition mapping Study #3 was also performed in a modern pre-bake smelter with various types of dense phase conveying systems in its alumina flow path. This location also uses injection type dry scrubbers. Delivery to the pots is via modern transport systems with low particle velocity.

As with Study #1 the distribution of superfmes in secondary, or reacted alumina, were similar to the PSD in the literature for fume particulate in the range of 2 to 50 microns. Refer to Figure 8 for the cumulative distributions of superfmes in fresh and secondary alumina.

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Calculation of actual %A.I. in thefieldwas 0.6% for Study #2. This is approximately l/25th of the % A.I. reported infreshSGA.

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These first two studies establish a baseline for what might be expected with a nearly ideal alumina handling configuration. "Ideal" implies a system with little pneumatic transport to move particles onward at any substantial velocity. See Figure 9 for the approximation of the PSD for fume particles in Study #2.

Note that the attrition index for this SGA is the lowest of the three studies presented thus far. However, actual attrition was more pronounced across the entire PSD. The primary difference at this location is the use of various dense phase transport systems over both long and short distances. Note the significant decrease in average particle diameter prior to fresh alumina arrival to the GTC. A short path dense transport system is used in conjunction with an airlift to move alumina from unloading into the main silo. This is followed by a dense phase transport system over a moderate distance to the fresh alumina silo of GTC#1. This dense phase system includes multiple 90 degree turns. Sub-4^ was increased by -6% in these systems.

166

in superfine fractions may be attributed primarily to fine bath and alumina particles carried by the pot exhaust gas stream.

Passing through the GTCs there was only a minor increase in sub45μ, -2%, or 1.1% A.I. across the GTC. This was followed by a second dense phase system for fresh alumina from the fresh silo of GTC#1 to the fresh silo of GTC#2. This caused to another increase in 8υ^45μ of -4.5%.

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Unlike Figure 8, note that there is no convergence of the fresh and secondary PSDs above 45μ particle diameters in Figure 11. This implies that attrition of alumina particles is combined with fume particulate in the βυο-45μ fraction. Sorting out the relative contributions of alumina particles broken by attrition and that of pot fume can require chemical analysis under such circumstances. However, it is possible to draw upon the results of Studies # 1 & #2 to estimate thefractionthat originated with pot fume assigning the balance to the attrition of particles.

The shape of the curve for the injection type GTC in Figure 12 is similar to that of Study #3, but with lower overall percentages of superfines in alumina. The implication is that the energy imparted as alumina makes multiple passes through the process gas stream of the dry scrubber fractures particles. This is true even with a low value of 7.4% A.I. for SGA. The increase in the %-45 micron across the injection type GTC was 2.7% more than with the fluid bed type scrubber. While this amount of attrition is not abnormally high it does illustrate that various types of dry scrubbers in various operating conditions can also have a range of impact upon particle attrition.

For study #3 %A.I. in the field was 7.9% after GTC#1 and 12.8% after GTC#2. These are much closer to the value reported by the refinery for fresh alumina, 10.5%, than with the first two studies. The significance of this is that alumina handling system design is important to the actual amount of attrition observed in the field.

Fluid Bed type - Study #1 ~0%A.I. across the GTC Fluid Bed type - Study #2 ~0%A.I. across the GTC Injection type - Study #3 1.1%A.I. across the GTC Fluid Bed type - Study #4 0.6%A.I. across the GTC Injection type - Study #4 3.5%.A.I across the GTC

Study #4 - In all, more than a dozen attrition mapping studies were performed on both pre-baked and S0derberg smelters. A portion of Study #4 is included here to illustrate the impacts of various types of dry scrubbing systems. The location of Study #4 primarily utilizes conveyor belts, airlifts and air gravity conveyors to transport alumina. But, this location has two types of gas treatment centers. One is a fluid bed type and the other is a modern injection type dry scrubber.

Discussion Note that all studies that have been presented here are the product of multiple grab samples taken according to strict procedures and made into composites. In all cases composites were replicated at each sampling site between two and four times. Each composite was then analyzed using the same equipment in the same labs for determinations of particle size distribution and flowability.

The %A.I. of the alumina in this study, a fourth source, was 7.4% during the period of study as determined by the Alcoa method. The %A.I. in the field, across the dry scrubbers only, was 0.6% for the fluid bed type vs. 3.5% for the injection type. This result is consistent with similar comparisons that have been made at Alcoa locations that use two types of dry scrubbers.

There is margin for error in any industrial scale study. When dealing with fine granular material the risk of introduction of bias is high. This may come either from segregation issues in the handling system or with inadequate sampling technique. Precautions have been made in the presentation of this data to validate it against the findings of other sampling points along each alumina flow path and against other locations that utilize similar equipment or handling system configurations.

Refer to Figure 12. Note that increases in superfines in secondary alumina from the fluid bed scrubber closely mimic fresh vs. secondary alumina from Study #2. Studies #1, #2 and #4 all utilized identical fluid bed dry scrubbers. One again, the PSD converges with that of fresh alumina near particle size diameters of 5 and 50 microns. This appears to confirm that such increases

Conclusions

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2. Grjotheim and Welch "Aluminium smelter technology, 2nd edition, 1988, p. 202

Unfortunately, the output of this study cannot draw a specific conclusion about what may happen when the %A.I. of an SGA changes. Changes in the attrition index with any one source of alumina and changes in sources of alumina passed through alumina handling systems were not included in the scope of this study. At this point it is still uncertain that changes in the magnitude of %A.I. for an alumina source may be used to accurately predict changes in the PSD of alumina prior to the reduction cell.

Acknowledgements The author would like to acknowledge the following people for their insights and contributions: - Jack Sorenson, Warrick Operations (retired) for great dedication and perseverance to the study of alumina attrition and flowability. These studies could not have been completed without his invaluable assistance. - Travis Baroni, Alcoa World Alumina Technology Delivery Group for particle size analysis and support for this study. - Participants in Attrition Mapping studies from more than one dozen Alcoa locations worldwide. Their efforts have provided the raw materials to understand and confirm the results of this study. - The support of Alcoa Primary Metals and Alcoa World Alumina to share these findings. - Merino, Dr. Margarita R. (Ph.D. - Florida State University) - for her encouragement, dedication and support.

The major factor that impacts actual attrition of particles in the field is the configuration of an alumina handling system. It has been observed that Actual %A.I. at a smelter may be roughly equal to the %A.I. reported by a refinery. It may also be much lower with Actual %A.I. being only 1/25* of the value reported by the refinery. Of particular concern are pneumatic transport devices that may rapidly move air to move alumina. In situations such as those described in Study #3 there were high rates of attrition at dlO, d50 and d90. In such cases it appears that %A.I. has similar general meaning for a wide range of particle diameters, not only at 45 microns. Differences in particle attrition were also observed between fluid bed and modern injection type dry scrubbers. This was not unexpected, but some systems did demonstrate higher rates of attrition than others, even when the %A.I. of the fresh alumina was relatively low. Some of these differences may be ascribed to system operating conditions and some may be ascribed to GTC design factors. The most common system presented here, fluid bed GTCs with low impact handling systems, did not produce significant amounts of particle attrition. These type systems were observed across a wide range of attrition indexes, from 7.4% to 17%, as measured by the Alcoa method for %A.I. The comparison of fresh and secondary alumina from Study #2, shown in Figure 8, is typical for low impact handling system configurations. The differences between these two distributions produce a sub-distribution shown in Figure 9 that is quite similar to the literature reference provided for fume particulate that is shown in Figure 6. With low impact systems it appears that the primary contributors to changes in PSD are fine bath and alumina particles carried by the pot exhaust gas. Observations indicate that fume particulate has its greatest impact in the sub-40 micron range. Thus, standard methods for measurement of %A.I. will not be easily extrapolated into superfine particle diameters. With the wide range of alumina handling and dry scrubbing systems that exist in our industry it is most important to "Know thyself before taking differences in the %A.I. for various sources of SGA into serious consideration. References 1. Audet, D., Clegg, R.E., "Development of a New Attrition Index Using Single Impact", Proceedings of the 8th International Alumina Quality Workshop, September, 2008, pp. 117-120

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINA and BAUXITE

Energy and Environment SESSION CHAIR

Benny Raahauge F. L. Smidth Copenhagen, Denmark

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PERSPECTIVE ON BAYER PROCESS ENERGY D.J. Donaldson P.E, ChE 1416 Village Center Drive Medford, Oregon 97504 Keywords: Bayer Process, Energy Pintle

Abstract The three most important cost items in the production of alumina are bauxite, caustic soda, and energy. Alumina energy cost will likely rise more than other costs because of energy price increases and energy related environmental issues. As energy cost increases, refineries will need to adjust the process to maintain or improve their position on the world cost curve.

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INTRODUCTION Energy is one of the three largest costs in the production of alumina from bauxite. In the future, energy could be two or three times today's cost (relative to other costs) because of availability and environmental issues. The better the understanding of the factors that make up the usage of energy in the alumina refinery, the better prepared process engineers and designers will be to evaluate refinery design changes in event a significant change in energy supply price or availability.

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This paper on energy in the Bayer Process considers only the interrelated areas of Powerhouse (Boilers), Digestion, and Evaporation. Excluded are Calcination and peripheral energy use areas such as Seed Wash, Liquor Purification, Causticization, etc. As shown in Figure 1, the source of energy to the Bayer Process is the Powerhouse. Energy is distributed essentially as steam to Digestion and Evaporation and electrical power to the refinery. The amount of energy required is determined by four interrelated factors:

Digester temperatures range from about 140 °C for trihydrate alumina (THA) bauxite up to about 280 °C for monohydrate alumina (MHA) bauxite. Because of the flash cooling and feed slurry heat recovery system used in Bayer Process refineries, the Digester temperature itself does not have a significant affect on Boiler Steam usage as long as there are no other streams entering or leaving (e.g., Blow Off). That is, theoretically, it will take nearly the same amount of Boiler Steam for a high temperature digest as for a low temperature digest. This is illustrated in Fig. 2.

1. Type of Bauxite. Trihydrate bauxite (low temperature digest or monohydrate bauxite (high temperature digest).

In practice there are other heat demands that do require additional Boiler Steam to Digestion:

2. Productivity. The amount of alumina produced for each unit volume of liquor circulating (see Fig. 1)

Heat of Reaction (HR). This is the endothermic heat of caustic attacking bauxite alumina hydrate to put the hydrate in solution. The HR is higher for THA than for MHA.

3. Evaporation. The facility that removes the net amount of water added to the process.

Sensible Heat Loss (SHL). This loss will be higher for MHA digestion because there will be over three times as many flash tanks (FT) and heat exchangers (HX) of which 2/3 will be at significantly higher temperatures than for THA Digestion.

4. Refinery Facilities. The physical design of the refinery with respect to heat recovery facilities.

Blow Off Vapor. By design, Blow Off vapor is kept small because it is a direct loss of steam (heat) to the atmosphere.

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Following is an example of the distribution of Boiler Steam in a flash/heater system. The example does not cover the whole range of refinery designs but it does indicate the relative importance of the factors listed above. % Distribution FT/HX (See Fig. 2) Heat of Reaction Sensible Heat Loss Blow Off Washer Overflow

evaporation (Digestion plus Evaporation) in a THA digest refinery would require a large Evaporator facility at about a 3.6 economy and much more total Boiler Steam than the comparable MHA refinery.

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The thermal energy required for Digestion and Evaporation in a MHA Digestion refinery will be less than for a comparable THA digestion refinery because of the higher average efficiency (economy) for total evaporation.

MHA digest 275°C -110°C = 165 °C evaporation THA digest 140°C -110°C = 30° C evaporation

When energy required for refinery electric power is included, the THA Digestion refinery will have a significantly lower electric power requirement. MHA Digestion will require more electric power in order to raise the bauxite slurry to digest pressures as high as 80 atm. (plus the frictional pressure drop in the large number of heat exchangers) compared to THA Digestion pump pressure of less than 10 atm. This will raise the MHA refinery electric power 10% -15% higher than the THA refinery.

For the 275 °C MHA Digestion, the evaporation economy (Fig. 2) is near 7 tons water/ton of steam. To achieve the same total

However, that is not the most important difference in providing energy for electric power. With a 60 atm Boiler in the THA

THA and MHA Digestion widely differ in the amount of evaporation that occurs. For instance, compare the 275°C digest (MHA) to a 140°C digest (THA) with the same product dissolved solids concentration where the temperature difference is indicative of the amount of evaporation:

172

process, electric power can be generated by first passing Boiler Steam through an extraction turbo-generator. In a MHA process an extraction turbo-generator can be used only to a limited extent, and most, if not all, electric power will be generated by condensing turbo-generators at nearly three times the energy required by extraction turbo-generators. Another factor is the type of fuel supply. The discussion so far has a Powerhouse design of any fuel (coal, oil, natural gas) generating steam to process and generating electric power by extraction or condensing turbo-generators. In the case of natural gas fuel, electric power can be generated by gas turbo-generators with waste heat boilers for steam supply. Natural gas fuel to the Powerhouse can avoid the use of condensing turbo-generators for electric power and provide power at the same low energy usage of extraction turbo-generators.

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Productivity or Yield is important in design of a refinery. A productivity increase from 75 to 90 gpl alumina is an increase of 20% in production for the same pipes, pumps, and vessels in the refinery. This is very significant in both new and expanding refineries. Higher productivity will give an overall lower Boiler Steam and electric power demand per ton of alumina produced. However, for MHA Digestion, there will be proportionately lower evaporation in Digestion and more mud wash water which will require more evaporation in the Evaporator facility at a lower economy than Digestion. The heat of reaction (HR) per ton of alumina will be the same, but the sensible heat loss in Digestion will be proportionately lower. Overall, perhaps a 10% decrease in Boiler Steam usage per ton alumina for a 20% increase in productivity. The subject requires a more in depth evaluation than can be covered in this paper but it is certainly an important factor in refinery energy usage.

» T ' t and LSH Feed T ° C

superheated (BPE) 10 °C and must be cooled 10 °C before it can condense in the Heat Exchanger. Cooling of superheated vapor to condensing temperature contributes little to heat transfer (less than 2%) compared to the heat transferred by condensation. In this example, the term 'approach temperature' is the difference of the steam condensing temperature above the slurry heater discharge temperature. Shown in the table in Fig. 3 is the sum of the three temperature drops (38 °C) which is also the temperature rise across the LSH. If the temperature of the slurry reporting to the last HX is lower than 182 °C, Boiler Steam will increase. In the example of Fig. 3 there will be more Boiler Steam than required by the heat balance described in Fig. 2. This indicates that if there were no other requirement for heat in the system (such as HR, SHL, etc.), there will be Blow Off vapor. Fig. 4 depicts Fig. 3 graphically and at the bottom shows the effect that HX area and number of stages will have on Boiler Steam to the LSH.

EVAPORATION An Evaporator facility is usually designed by a contractor with boundary limits provided by the alumina process engineer. This typically results in an evaporation economy of around 3.6 tons water/ton Boiler Steam. As liquor caustic concentrations increase for high yields (productivity), achieving a 3.6 economy becomes more difficult because of a corresponding increase in BPE. Increasing the Evaporator LSH temperature to over 200 °C and adding more stages (effects) could increase the economy to over 7 tons water/ton steam, a very significant reduction of Boiler Steam. Erosion/corrosion issues would need to be addressed. Silica scale on heater tubes at the higher temperatures can be controlled by chemical additive. FACILITY DESIGN An important factor that directly determines the minimum Boiler Steam to Digestion is the nature of the Flash Cooling/ Heat Exchange system itself. This is illustrated in Fig. 3 showing an example of temperatures across the first stage following the digester. Digester discharge at 250 °C immediately flashes to 230 °C, the temperature of both liquor and vapor. The vapor is

173

The driving force for heat transfer is the temperature difference across Digestion and across Precipitation. As indicated in Fig. 5, in order to reduce Boiler Steam to Digestion, the delta T across Digestion must be reduced. However, this will not be practical or efficient without correspondingly lowering the delta T across Precipitation This will require more recuperative cooling in Precipitation in place of heat loss by cooling water.

Digester Temperature (250 ^ )

Flash Tanks Temperature BPE

1Û°C

Approach ♦

§βς

Temperature to LSH {212 °C) *** Heot Exchangers Temperature

Design Case

Doubla Area

Double Stages

Double Both

Hash A T , * C BPE, ° C Approach, C C

20 10 8

20 10 2

10 10 8

10 10 2

LSH ΔΤ, ^C % of Bate

38 100

32 84

28 74

22 50

Fig 4, Erteci of Heat Transfer Arsa and Number of Flash Stages

The Bayer Process is one large heat exchanger. This is illustrated in Fig. 5. There is practically no temperature drop in liquor across Evaporation and sensible heat loss across Clarification will be small. The temperature drop across Precipitation is composed of induced cooling (cooling water) to lower the process temperature of liquor to the precipitators, sensible and evaporative heat loss during the long holding time in Precipitation, and induced cooling during precipitation to further lower the temperature.

•i.30 90Δ !^T

T |

SETTLER

Design for MHA bauxite can reach very high temperatures in the Digester (>260°C). The LSH can elevate Boiler pressure above 70 atm. and up to 90 atm depending on Digest temperature. The corresponding temperature of steam will be above the limit to use process condensate for feedwater because of serious corrosion in Boiler tubing. For this reason, Digestion design using Boiler Steam at high pressure (>70 atm), must return LSH steam condensate to the Boilers. For instance, Boiler Steam to LSH at 100 atm and 315°C will be nearly doubled because almost half the steam heat will be returned as condensate to the Boiler. The alternative of routing the condensate to the Digestion Heat Exchangers instead would require using demineralized Boiler feed water at higher cost and, most likely, an increase in Digestion heat exchange area or number of FT stages. Theoretically, the net thermal affect of returning LSH condensate is the same; it's just the issue of physically returning high temperature/pressure condensate to the Boiler. The return of condensate can be avoided by using a different method of heating liquor before the Digester, such as a molten salt HX (2) in place of the LSH. FINAL COMMENTS There should be increased interest in investing in heat recovery facilities in the event of a doubling or tripling of energy cost relative to other alumina costs. The possibilities that immediately arise are: 1.

Use of natural gas to reduce the heat required to generate electric power in MHA refineries. Refineries that do not have the availability of natural gas may consider using LNG.

2.

Increase the number of stages in Digestion Flash Heat Exchange and at the same time increase recuperative cooling in Precipitation to lower the refinery delta T (see Fig. 5).

3.

With high temperature MHA Digestion (>260 °C), it should be feasible to eliminate the Evaporation facility.

4.

Route Washer Overflow directly to Precipitation. It appears wasteful to heat Washer Overflow for Settler operation and then essentially cool the same stream in Precipitation.

5.

Increase the evaporator LSH temperature and the number of stages (effects) to increase the economy in Evaporation.

I

Fig 5. Temperature Difference across Digestion

REFERENCES

and Precipitation

1. D.J.Donaldson, Journal of Metals, Sept. 1981, page 37. 2. Vereint am Werk (VAW house organ) July 1968

174

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

OPTIMIZATION OF HEAT RECOVERY FROM THE PRECIPITATION CIRCUIT Michael Simpson1, Sushant Hial1 and Rashmi Sangram Singh1 Vedanta Aluminium Ltd., Lanjigarh, Kalahandi Dist., Orissa - 766027, India Keywords: Interstage Coolers, Heat Transfer, Steam Savings solid Al(OH)3 in the precipitation reaction as per the following equation :-

Abstract In the Bayer process, temperature profile across the precipitation circuit plays a major role in maximization of the precipitation yield while maintaining product quality. For this reason, plate heat exchangers are used both at the inlet to Precipitation and between precipitation stages at Vedanta Aluminum's Lanjigarh alumina refinery. The cooling medium for the Heat Interchange Department (HID) is spent liquor, while for the Interstage Coolers (ISCs) of Precipitation, both spent liquor and cooling water are used.

Digestion (Heat) Al(OH)3 + NaOH 4 ""► Al (OH)4" + Na+ Precipitation (Cool ) The primary factors which affect the precipitation yield are well documented by authors such as Hond, Hiralal and Rijkeboer [1], They are: > Initial Liquor Alumina(A) and Caustic(C) Concentration > Holding Time > Temperature > Seed Surface Area > Liquor Impurities such as carbonates

A simple model was built using existing heat exchanger performance data, along with heat and mass balances across the Heat Interchange Department (HID) and Precipitation. This was then used to determine feasible modifications for improving heat recovery in the Precipitation circuit. The results obtained have indicated a process steam reduction of 2 % can be achieved with minor modifications.

To maximise precipitation yield, temperature drop is an important parameter that can be controlled within the Precipitation Department. For this reason, interstage coolers have been provided between the precipitation tanks to reduce the temperature of the slurry. Approximately 60% of current precipitation train flow passes through the ISCs, in order to maintain velocity in the correct range through the coolers. The cooling medium for the 1st ISC is spent liquor from hydrate seed thickener overflow, while the cooling medium for other ISCs is cooling water from the alkaline cooling tower. The flow of these cooling media are monitored and controlled to achieve the desired temperature profile across the precipitation circuit.

Introduction In the Bayer process, the temperature profile across the precipitation circuit plays a major role in maximizing the liquor productivity. The supersaturation of the liquor increases as the temperature decreases. Thus the enhancement of precipitation rate is assisted with the help of inter stage coolers across the precipitation unit. The cooling media used in the ISCs are spent liquor and alkaline cooling water. Presently, heat gained by the cooling water is not utilized as the water loses its heat in the alkaline cooling tower. This loss can be prevented by replacing cooling water with spent liquor as an alternative cooling media in ISC - 2. Further heat can be recovered from the precipitation circuit by drawing the inlet slurry to ISC - 1 from further upstream tanks such as the last agglomerator. The recent installation of new ISCs as part of a debottlenecking project will enable the spent liquor temperature to be kept below 55°C consistently. This liquor therefore becomes a good heat sink for hot precipitation slurry, compared to original design when it was at 60°C.

Heat Balance Determination By utilising the well known heat balance equations below [2, 3, 4], and a series of iterations, the final spent liquor temperature can be determined for given pregnant liquor to HID and spent liquor to ISC temperatures Q=UxAxdTLM (1) Q=Mh xCphXiThiTho) (2) Q= Mc xCpcx(Tco.Tci) (3)

Process Description

Where for any heat exchanger: Q=Heat transferred between the streams (kW) U= Effective overall heat transfer coefficient (kW/m2/°C) A= Heat transfer area(m2) dTLM= Log-mean temperature difference (°C) Mh= mass flow rate of hot stream (kg/s) Cph= Specific heat capacity of hot stream (kJ/kg/°C) Τω = Inlet temperature of hot stream (°C) Tho = Outlet temperature of hot stream (°C) Mc = mass flow rate of cold stream (kg/s) Cpc= Specific heat capacity of cold stream (kJ/kg/°C)

The objective of the Precipitation unit is to produce Alumina "Hydrate" (Alumina tri-hydrate or gibbsite) suitable for downstream calcination to produce sandy metallurgical grade alumina. This is achieved by prolonged contact between pregnant liquor and suspended hydrate seed, under conditions conducive to crystal growth and precipitation of alumina from solution. The chemistry of precipitation, which is the reverse of digestion, is quite simple. In digestion, Al(OH)3 dissolves in NaOH to form a complex anion and this complex ion comes out of solution as

175

Tci = Inlet temperature of cold stream (°C) Tco = Outlet temperature of cold stream (°C)

their appropriate heat transfer coefficients (U value). The different scenarios considered are as follows:

Existing Heat Balance

i) ISC - 1 slurry input from last agglomerator Tank:High precipitation temperatures are only required to aid the agglomeration process. Instead of taking slurry input to ISC-l from the second growth stage, we can utilize slurry from the last agglomerator without impacting on the precipitation sizing control. In this case there will be 6.7°C rise in spent liquor temperature across ISC - 1 as shown in Fig 3. The inlet slurry temperature will be 77.9°C and the outlet spent liquor temp will be 61.7°C. The net spent liquor temp after HID will be 83.7°C as shown in Fig 4. A further advantage of this move will be a reduction in the solids concentration through the ISC which is expected to improve its life. The liquor will be of a higher scaling potential however and caustic wash frequency might need to be adjusted. New piping will be required from second precipitator tank (normally last agglomerator) along with third precipitator (alternate last agglomerator) to the ISC-l.

Slurry

Slurry 1054

m3/hr

1054

m3/hr

69.5

DegC

64.2

DegC

ISC-l Spent l i q .

Spent Liq. 1250

m3/hr

59.2

DegC

<— - -

1250

m3/hr

55.0

DegC

Fig 1 : Spent liquor temperature increased by 4.2°C after passing through ISC - 1

Slurry

Slurry 1054

m3/hr

1054

m3/hr

77.9

DegC

69.6

DegC

250 rnB/hrL. 59.2 1250

DegC

I I

irâ/hrl/

i

83.0

SL

I

10Ö0

φ

1250

m3/hr

88.9

DegC

55,2

DegC

1086

iii3/hr

m

DegC

ISC-l

Spent Liq.

1250

Spent Liq.

m3/hr|

t—~~

DegCj

1250

m3/hr

55.0

DegC

HID PGL

PGL 1200

rn3/hr

1Ö1E

DegC

™™

ii

*»

>

>

Fig 3: Spent liquor temperature increased by 6.7°C after passing through ISC - 1

114 nn3/hr 1Ö2.8

DegC

1250

Fig 2: Spent liquor temperature increased to 83°C after passing through HID

83.7

In the existing system, slurry feed for ISC-l is from the fourth tank which is normally the second growth tank. The spent liquor temperature is raised by 4.2°C across the ISC as shown in Fig. 1. As shown in Fig 2., the spent liquor temperature also increases by 23.8°C in passing through the HID. In the existing scenario, there is also a temperature raise of 10 °C in the alkaline cooling water which is used as the cooling medium in ISC - 2.

m

m3/hr

61.7

DegC

1000

φ(

1250

ntf/hr

89.2

DegC

61.7

DegC

φ\

SL

SL 1

HID PGL

PGL

Proposed Modifications with Schematic Diagram

1200

rn3/hr

1086

ί#Γ

102.8

DegC

76.Ö

DegC

'

To recover the heat through spent liquor it is proposed to: (1) Use spent liquor as cooling media in ISC-2 (2) Feed slurry from an upstream precipitation tank having higher temperature Detailed analysis was done across HID and Precipitation by using the historical performance data of the heat exchangers and hence

>

114 πή/hr 102.8

DegC

Fig 4: Spent liquor temperature increased to 83.7°C after passing through HID

176

Fig 7: Spent liquor temperature increased to 83.7 °C after passing through HID ii) Spent Liquor as cooling media in both ISC - 1 & 2:For this scenario, there will be 6.7°C rise in spent liquor temperature across ISC - 1 and 2 as shown in Fig 5&6.The net spent liquor temp after HID will be 83.7°C, shown in Fig 7. Here the spent liquor flow will be in series i.e. the outlet from ISC - 2 will be the inlet to ISC - 1. Spent liquor piping already passes by the ISC-2 to feed ISC-1 and hence this modification will be simple. Analysis of the additional pressure drop from passing through two ISCs shows that the net positive suction head required for the HID pumps will still be met easily. Slurry

iii) Spent Liquor as cooling media in both ISC - 1 & 2 and ISC 1 slurry input from last agglomeratonThis is a combination of the actions from cases(i) and (ii). There will be 9.2°C rise in spent liquor temperature across ISC - 1 and 2 as shown in Fig 8&9.The net spent liquor temp after HID will be 84.8°C , shown in Fig lO.The spent liquor flow will again be in series ( i.e. the outlet from ISC - 2 will be the inlet to ISC - 1) and the ISC - 1 slurry input will be from last agglomerator. Slurry

Slurry

1054

m3/hr

1054

m3/hr

66.4

DegC

62.3

DegC

Slurry

1054

m3/hr

1054

m3/hr

64.9

DegC

613

DegC

y

t

ISC-2 ÌSC-2 Spent Liq.

Spent Liq.

1250

m3/hr

58.3

DegC

<--»-"-

<—. —.

125?

m3/hr

55.0

DegC

Slurry 1054

m3/hr

1054

m3/hr

69.9

DegC

65.6

DegC

<----

1250

m3/hr

5S.3

DegC

Degc| SL k-»1

SL 1000

m3/hrf··-

89.2

DegC]

m3/hr|

64.2

Degcf

m3/hr

51.7

DegC

1018

— 1

I I I SI I

1000

nâ/hff

1250

m3/hf

90.0

DegC

64.2

DegC

m3/hrL/

HID

ρα m3/hr 76.0

Degcl

114

mllk\

DegC

DegC|

PGL

ra3/hrr

m3/hr

64.2 84.8 1250

1250 57.9

c— - » — ■ ·

HID 1200

m3/hr DegC

Spent Liq.

1250

1250

P6L

DegC

1054 71.6

250 m3/hrk

m3/hr|

83.7 JegCJ

m3/hr

55.0

Slurry m3/hr| DegC]

250 m3/hrL 1250

1250

Fig 9: Spent liquor temperature increased by 6.3°C after passing through ISC - 1

Fig 6: Spent liquor temperature increased by 3.4°C after passing through ISC - 1

61.7

<--«-

ISC-1 Spent Liq.

<-

DegC

<- . . .

Spent Liq.

Spent Liq. DegC

5X9

79.3

ISC-1

61.7

m3/hr

Slurry 1054

Slurry

m3/hr

1250

Fig 8: Spent liquor temperature increased by 2.9°C after passing through ISC - 2

Fig 5: Spent liquor temperature increased by 3.3°C after passing through ISC - 2

1250

Spent Liq.

Spent Liq.

102.8

177

PGL

1200

m3/hrl·

1086

m3/hr

102.81

Degc|

m

Degcj

114

m3/hr

102.8

DegC)

Fig 10: Spent liquor temperature increased to 84.8 °C after passing through HID

At a price of $11.1/t steam, the cost savings vary from $147,000-$375,000 per annum and should be adequate to give a good return on investment as only minor modifications are anticipated to be required.

Expected Benefits

1.

Table 1: Cost saving analysis with different modification options Existing

A

2.

m 1

R

f5C-l&2wfià SL series flow Cenditiöß

ÌS€-l*Éh

$t

isc-iyet

SC-l&2«äJi

3.

iSCÖöing

from Last a^Q SL series flow media with ISC

4.

i slurry inlet from last aggio |

12»

1250

83.0

83.7

83.7

84.8

129294

13Ö447

130M5

132225

Heat Gain w,r.t existing, KJ/Sec

1153

1150

2331

Net Mass of Steam Saving, T/hr

152

L52

3.86

0

404

403

1027

0

147467

147126

St Flow, Mi/k

1250

SLTemp.afterHID.DegC HeatAKi/Sec

Safing,$/d3y 1 Saving, $/Annum

1250

374867 1

SAVINGS/ANNUM

m

Existing

Graph 1 : Savings / annum for different scenarios.

Conclusions Up to 3.86 t/hr of steam can be saved by modifying ISC-2 to utilise spent liquor at 55°C as the cooling medium and taking slurry input to ISC-1 from the last agglomerator. Further detailed analysis of the mechanical and operational feasibility along with the cost of the modifications for each scenario will be carried out.

178

References Roelof Den Hond, Iwan Hiralal and Ab Rijkeboer " Alumina yield in the Bayer Process - past, present and prospects" , Light metals 2007 , 3 7 - 4 2 Don W. Green, James O. Maloney, "Perry's Chemical Engineers" Handbook - seventh Edition - Sections 5-12,1114 Nicholas P. Cheremisinoff, "Handbook of Chemical Processing Equipment" , 1 - 6 2 Max S. Peters, Klaus D. Timmerhaus and Ronald E. West, "Plant Design and Economics for Chemical Engineers" Fifth Edition, 643 - 748

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ALUNORTE GLOBAL ENERGY EFFICIENCY Reiner Wischnewski, Cleto Maues de Azevedo Jr., Emerson L. S. Moraes, Arthur Barros Monteiro Alunorte - Alumina do Norte do Brasil S.A., Rodovia PA 481, km 12. Distrito de Murucupi, CEP: 68.447-000. Barcarena-PA. Brasil Keywords: Bayer process, Energy efficiency, Heat integration, Water balance Abstract

18

Alunorte is the largest alumina refinery in the world with a production capacity of close to 6.3 Mtpy. The plant has a specific energy utilization of 8 GJ per ton of alumina and is one of the most energy-efficient plants in the world. The high energy efficiency is achieved by good process design, the utilization of state-of-the-art technologies, good operation and the processing of high quality bauxite. The technologies which are applied in Alunorte to contribute to the global energy efficiency of the plant are reviewed and heat integration and water balance are discussed.

16

1u

c •2 12 to

>»

2> 8 5 !

3rd quartite

Squartile

/inquartile

6

2

Alumina is the principal raw material for aluminium. In 2009 the world-wide average utilization of electrical power for primary aluminium production was 15,215 kWh per metric ton of aluminium [1]. The energy for alumina production is not included and has to be added. The average energy utilization for alumina production was 3,311 kWh per ton of alumina in 2009 [2]. In theory 1.9 tons of smelter-grade alumina are required for the production of one ton of primary aluminium. Hence, energyefficient alumina production is an important contribution to reduce the energy utilization of primary aluminium production.

50

75

100

% of alumina world-wide production outside China Figure 1. Specific energy utilization for alumina production. are discussed which in sum result in high global energy efficiency of Alunorte. Main elements are bauxite of high quality, the precipitation concept and the resulting high liquor productivity. Alunorte's design with regard to heat integration, water balance and utilities for steam and power generation is discussed as well as technology and equipment applied. Some factors are mentioned which work against an improvement of energy efficiency in the future. Current projects to address these issues are mentioned and some possibilities for a further increase of energy efficiency are presented. Alunorte has the aim to be among the most energyefficient alumina refineries in the world also in the long-term.

Table I. Energy used for 2009 alumina production [2], Africa and South Asia North America South America East Asia and Oceania Europe Weighted Average Total

Ί i /

'«ft

Introduction

spec, energy (MJ/t) 14,768 11,449 9,319 11,252 16,842 11,922 499,355 TJ

Squartile Alunorte (8.0 GJ/t)

alumina produced (t) 3,225,778 2,804,849 12,226,990 16,511,664 7,117,522

Energy Utilization for Alumina Production Alunorte started operation in 1995 with a production capacity of 1.1 Mtpy. Since then the plant was expanded three times to a total production capacity of close to 6.3 Mtpy. In Figure 2 the development of production capacity is shown and compared to the energy utilization of the plant. In 2009, the first year after Expansion 3, the actual production remained slightly behind the nameplate capacity of the plant. The Expansion 3 project, Alunorte's current performance and an outlook about the expected development of Alunorte are presented in [3].

-

41,886,803

Table I shows the 2009 specific energy utilization for alumina production by region. On average the production in South America utilizes less energy than in any other region of the world. Alunorte contributes significantly to this result. The production capacity of Alunorte is close to 6.3 Mtpa after its third expansion in 2008 [3]. Alunorte produces close to 50 % of all alumina in South America. Its energy utilization was 8.0 GJ per ton of alumina in 2009. This is significantly less than the average of 9.3 GJ per ton in South America and 11.9 GJ per ton world wide. In Figure 1 Alunorte's performance is compared to the energy utilization of other alumina refineries. The graph includes alumina plants for the production of smelter-grade alumina outside China. The statistic illustrates that the world-wide benchmark for alumina production is at about 8 GJ/t. There are few plants only with a similar good performance. In the following, the different factors

After the first two years of operation, the plant uses between 7.5 GJ and slightly above 8 GJ per ton of alumina. In 2006 the best result was achieved. In the following years from 2006 to 2009, i.e. since Expansion 2, the specific energy utilization trends upwards. With Expansion 2 some new technologies were installed. Bauxite is received from Paragominas through a pipeline and dewatered at Alunorte in hyperbaric filters. The filters use compressed air and thus compressor power. Two more turbogenerators were installed to increase the electrical power which is co-generated at site. The losses of cogeneration of electr-

179

Bauxite Quality & Liquor Productivity

J2 E

Ì1

n

The bauxite quality, in particular the content of available alumina, reactive silica and organic components, significantly affect the performance of an alumina refinery. Alunorte processes high quality bauxites from two different sources, namely Trombetas and Paragominas. Table II compares the main characteristics of these bauxites.

■ 1111117.57.0 SI

h- co σ> O) O) G) G) σ> σ>

mimi

8 8 8 8 8 8 8 8 8

Table II. Typical bauxite quality.

S

6.5

m

CNCMCNtNCNCNCNCNCN

ipacity Figure 2. Development of Alunorte's alumina production capacity and specific energy utilization. icity decrease Alunorte's global energy efficiency. Losses of cogeneration are included in Alunorte's overall efficiency while electricity from the grid is accounted as is without adjusting for losses during generation. The plant availability reduced somewhat which is associated with some negative impact on the specific energy consumption as Alunorte became more complex.

Table ΙΠ. Spent Liquor Analysis. A1203

NaOH (as Na2C03) Na2C03 causticity organic carbon NaCl Na2S04 Na2C204

total energy utilization

g

\pii

1

electricity m

m

Il

(%)

(g/i) (g/1) (g/i) (g/i)

112 280 5.6 98 4.4 2.20 0.28 <0.1

The concept of bauxite residue treatment and disposal is an important factor contributing to high liquor purity. Bauxite residue is filtered in drum filters and disposed at high solids concentration by dry stacking in the residue disposal area (RDA). Collected water from the RDA is neutralized and clarified in a water treatment station and released as effluent to the nearby river [4]. There is no water returning from the RDA to the process. This

digestion steam

E

(g/i) (g/i) (g/i)

The content of organic material in the bauxite is very low. This results in a very low concentration of impurities in the plant liquor. In Table III a spent liquor analysis of Alunorte is shown. The concentrations of organic carbon, sulfate and oxalate are low and the causticity of the liquor is very high. The plant is operated with a high caustic concentration. A precipitation productivity of up to 89 g/1 in the newest process lines can be achieved [3].

other . /

Figure 3. Breakdown energy used £ ofcogenerated | at 1Alunorte.

i

Paragominas 48.2 % 4.6% <0.05% 14.9 %

Although the composition is very similar certain differences are observed during processing. The high content of available alumina results in a low specific bauxite consumption. This is beneficial in terms of energy utilization, since the amount of the non-gibbsite fraction of bauxite, which does not contribute to the production of hydrate, is small. A low amount of digestion live steam is required only to heat up this material with a positive effect on the overall energy efficiency. A low temperature digestion process is used at Alunorte as both bauxites contain minimal amounts of boehmite. Their reactive silica content, however, is relatively high. The reactive silica reacts under consumption of caustic soda to desilication product (DSP) and increases the amount of bauxite residue. Generally, some heat is lost from the process to the environment due to the bauxite residue having a temperature of about 60 °C upon disposal. This heat loss at Alunorte, however, is moderate due to the relatively low mud factor, e.g. mass of mud per mass of alumina produced, which caused by the high content of available alumina in the bauxite and the bauxite residue beingfiltratedbefore disposal.

In Figure 3 a breakdown of the energy utilization of Alunorte is shown. It is divided into electrical power received from the national grid, energy required for steam and power generation and energy for calcination. The energy for steam and electricity generation and calcination is based on the lower heating value (LHV) of the fuels. Electrical energy received from the national grid is accounted as received and is not adjusted for losses during generation. Steam at two pressure levels can be generated with the installed high pressure or low pressure boilers. Three turbogenerators work between the high and the low pressure level and can cogenerate up to 65 % of the required electrical power of Alunorte. Process steam is supplied to the process from the low pressure header. In digestion about 1.0 t steam per ton of alumina is consumed and about 0.2 t/t in evaporation. 3 GJ per ton alumina is used for calcination. Other consumers include losses due to inefficiencies of boilers and turbogenerators, internal steam consumption in the boilers and miscellaneous steam consumers in the process such as for filter aid preparation.

Ζ ΖΖ;Γ,Ζ~ evaporation steam

Trombetas 49.1 % 4.1% < 0.05 % 11.5%

available A1203 reactive Si0 2 organic carbon moisture

f

electricity T

180

to dilution with steam. The upper limit for the caustic concentration is given in the spent liquor live steam heaters due to scaling and corrosion. At lower caustic concentrations less bauxite can be charged to the liquor and liquor productivity decreases. These negative effects are minimized at Alunorte due to avoiding direct heating by steam injection. Furthermore, the heat exchange area for spent liquor heating is well chosen, so that the energy for regenerative spent liquor heating is high.

approach avoids the contamination of plant liquor with additional impurities and contributes to maintain clean liquor with high liquor productivity. Liquor impurities are purged from the process with the bauxite residue. No caustizisation unit for carbonate removal is installed. Process Design The concepts for heat integration, water balance and generation of steam and electrical power are the main contributors for direct savings of energy.

Indirect heating of the bauxite slurry is another important contributor to low digestion steam consumption. The digestion temperature is a mixing temperature determined by the temperature and mass flows of the bauxite slurry and the final spent liquor temperature. The higher the temperature of the bauxite slurry the lower the final spent liquor temperature after the last live steam heater can be. This reduces the utilization of digestion live steam and increases the global energy efficiency.

Heat Integration Alunorte has a low temperature dual stream digestion unit. A simplified flowsheet is shown in Figure 4. The pregnant liquor stream is expanded and flashed in five stages from the digestion temperature and pressure to atmospheric conditions. The liquor is released from the blow-off tank (BOT) at a temperature close to the liquor boiling point. The spent liquor is heated in countercurrent flow by regenerative heating. Heat is transferred from the pregnant liquor to the spent liquor stream with vapor from the flash tanks which is condensed in four spent liquor regenerative heaters. Vapor from the blow-off tank is released to the environment. Live steam heaters are used to heat the spent liquor to the required temperature before it is fed to the digestors. flash tanks

life steam

life steam condensate

Figure 5. Indirect Bauxite Slurry Heaters (IBSH).

life steam heaters — Wl·*—j

Two stages of Indirect Bauxite Slurry Heaters (IBSH) are installed at Alunorte. The aim is to achieve the highest possible bauxite slurry temperature. The IBSH units are vertical shell-andtube heat exchangers (Figure 5). The bauxite slurry is heated indirectly with live steam heater condensate. Alunorte operates the units without major problems. As part of Alunorte's continuous efforts to increase heat recovery a project has been launched to add a third stage of IBSH and maximize the bauxite slurry temperature before digestion. This will reduce digestion live steam consumption due to a lower required final spent liquor temperature. Furthermore, a positive effect on the silica scaling in the spent liquor heat exchangers is expected. A good maintenance program is in place to keep the heat transfer coefficients of the heat exchangers in the digestion area at a high average level.

bauxite slurry heaters bauxite

Figure 4. Simplifiedflowsheetof digestion and slurry heating. A conceptual problem of dual stream digestion technology is the difference between heat sink (spent liquor) and heat source (pregnant liquor). The mass flow of pregnant liquor is larger than that of spent liquor by the amount of bauxite slurry which is charged to the digestors. It is not possible to condense all vapor which isflashedfromthe pregnant liquor stream in the regenerate heat exchangers and transfer all latent heat of the vapor to the spent liquor. The excess steam leaves the blow-off tank and is a loss of heat to the environment. In plants with large amounts of excess steam some of it can be transferred to other process areas.

The efficiency of the vacuum flash cooling units is another major contributor to the low energy utilization of Alunorte. Pregnant liquor is cooled in counter-current flow against spent liquor in a flash train. The pregnant liquor inlet temperature is close to the boiling point and the outlet temperature shall be as low as possible. Additional plate heat exchangers are used to achieve the final fill temperatures to agglomeration and cementation tanks. A simplified flowsheet of Alunorte's vacuum flash unit and plate heat exchangers is shown in Figure 6.

At Alunorte the difference of heat sink and heat source in the digestion area is small and with it the loss of heat to the environment. This is based on a good design of Alunorte. There is no direct heating of the spent liquor by steam injection and indirect bauxite slurry heaters are installed. Direct heating increases the difference between heat source and heat sink in digestion, an additional amount of water must be evaporated and requires additional energy and the caustic concentration drops due

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slaking, gland water, rain, etc. Alunorte's digestion concept which avoids dilution of liquor and uses indirect heating only is one of the main reasons for a comparably small liquor dilution.

cooling water to aggi.

pregnant liquor

Trombetas bauxite is transported to Alunorte by ship while Paragominas bauxite is pumped through a 244 km pipeline [5]. At Alunorte the Paragominas bauxite is dewatered in hyperbaric filters [6-8]. The residual moisture of the dewatered bauxite is about 15% and slightly higher than that of Trombetas bauxite with about 11.5 %. The moisture content of the dewatered Paragominas bauxite is determined by the performance of the filters and the characteristics of the bauxite, in particular its particle size distribution. Work is underway to optimizing the operation of the filters in cooperation with the beneficiation plant of Mineraçao Bauxita Paragominas (MBP). The potential for improving the water balance in bauxitefiltration,however, is small. The same is valid for water entering with fresh caustic. It is supplied at a concentration 50 wt.-% NaOH. The caustic consumption is mainly determined by the non-controllable loss due to the content of reactive silica in the bauxite and the controllable losses which are affected by the efficiency of mud washing and the residual moisture in the bauxite residue. Controllable losses are already very small with less than 8 kg of caustic per ton of alumina. However, some potential improvements in the bauxite filtration area identified.

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condensate Figure 6. Vacuum flash cooling and plate heat exchangers. The better the performance of the vacuum flash cooling units the lower is the pregnant liquor inlet temperature to the precipitation plate heat exchangers. This reduces the requirement of cooling water to the plate heat exchangers and recovers more energy for heating the spent liquor. A high spent liquor temperature reduces the required amount of digestion live steam and improves the performance of the evaporation unit. Alunorte has a spare vacuum flash cooling unit so that cleaning and maintenance can be performed without major performance losses in three of the seven lines. A high average heat transfer coefficient of the heat exchangers of the vacuum flash cooling units can be achieved and process disturbances due to cleaning are small in these lines. As measure for higher average heat transfer coefficients Alunorte optimizes the cleaning cycles of the vacuum flash cooling units and will study the option to install spare vacuum flash cooling units for the other four lines as well. For efficiency improvement their capacity could be increased such that lower pregnant liquor and higher spent liquor temperatures can be reached.

Regenerative condensate is used for washing bauxite residue and hydrate. Both together are the biggest source of plant liquor dilution. The net wash of mud is generally good with 0.8 - 1.2 ton of water per ton of mud. The potential for further improvement is small. The situation for hydrate washing is similar good and the potential for improvements small. Dilution of plant liquor from other sources such as rain, water hosing or gland water is moderate. All tanks are closed and water hosing is done with care. There exist some possibility for further reduction of water input to the liquor but the situation is generally good and room for improvement also small.

Water Balance The water balance directly affects the energy utilization of an alumina refinery. It is unavoidable that water from different sources dilutes the plant liquor. This water is mainly removed by evaporation. With higher rates of liquor dilution the demand for evaporation increases and at the same time energy used in form of live steam.

Small water sinks are product and bauxite residue while the largest part of the water is removed by evaporation in three plant areas - digestion, vacuum flash cooling and evaporation. A good maintenance program is crucial to ensure high average heat transfer coefficients in all heat exchangers of these areas. A high heat transfer coefficient is associated with a high rate of water removal from the liquor and low consumption of live steam in the digestion and evaporation areas. Furthermore, spent liquor return temperature increases with higher performance of the vacuum flash cooling units. The demand for digestion life steam reduces and the efficiency of the evaporation area increases which is associated to a lower requirement for evaporation live steam. Work is underway to further improve the performance of the flash trains in the digestion, evaporation and vacuum flash cooling.

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Steam and Power Generation

water leaving

The design of the utilities area for steam and power generation has to match the demand of steam and electrical power. Inadequate design can lead to high losses in this area and potential savings in steam or electrical power in the Bayer process would be compromised by additional losses in the utilities area. Alunorte has chosen a design which allows a sufficient flexibility to achieve overall high efficiency in steam and power generation.

Figure 7. Volume of water entering and leaving the plant liquor. In Figure 7 the water balance of Alunorte is shown. Due to a high liquor productivity and no direct heating with steam in digestion dilution of liquor with water per ton of alumina can be kept at around 2.3 m /t. Main sources are the moisture content of the bauxite, water for mud and hydrate washing, a small amount entering with caustic and various sources such as hose water, lime

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at high caustic concentrations and at a high liquor to precipitation (LTP) A/C ratio. Seed filtration is used in five of the seven lines and in two lines cyclones for seed concentration are installed. With the precipitation concepts chosen in Alunorte high liquor productivities are achieved.

electrical power

In addition, it is important that the individual pieces of equipment of the plant have a high efficiency, operational procedures are good, high availability can be achieved and a good maintenance program is in place. High efficiency motors for agitators, blowers, etc. are installed and pumps with variable speed or variable frequency drives. An electrical power utilization less than 0.2 MWh per ton of alumina is achieved.

digestion evaporation +> other condensate make-up water

Steam consumption is significantly affected by the process design as explained above. But also individual units, in particular the evaporation area, have to perform well. Alunorte's evaporation units have an efficiency of about 4.2 ton evaporated steam per ton of live steam. Some potential for improvement was identified and Alunorte works on a project for further performance increase.

Figure 8. Simplified flowsheet of utilities area for steam and power generation. A simplified flowsheet is shown in Figure 8. Alunorte receives some of the electrical energy from the national grid and cogenerates the remaining electrical energy with turbogenerators (TG) which work between the high pressure (HP) and low pressure (LP) steam headers. The fraction of the high pressure steam (87 bar and 485 °C), which is not required for cogeneration, passes through pressure reducing valves (PRV) and is subsequently cooled to saturated conditions by adding water in desuperheaters (DSH). This amount of high pressure steam can be replaced with steam from the low pressure boilers. Thus, the generation of steam and electrical power are not directly coupled but allow some operational flexibility. The operation of the utility area for steam and power generation can be optimized under consideration of overall efficiency and operational cost. The high pressure boilers have an efficiency of 84-86 % while that of the low pressure boilers is 88-90%. The second advantage of decoupling of steam and power generation is that a design could be chosen which avoids the installation of condensing turbines, resulting in an overall higher efficiency of the utility area.

A major operational issue is scaling. It affects the heat transfer coefficient in heat exchangers and has an impact on the overall availability of the plant in dependence of the frequency for descaling and cleaning. Both are associated with lower global energy efficiency. Alunorte works on the one hand to optimize the cleaning and descaling work and on the other hand to reduce scaling rates. Measures to reduce the solids concentration in spent liquor are one example. Another is the modification of digestion side entry flash tanks to bottom entry which was done before Expansion 1. Today there are no side-entry flash tanks installed at Alunorte. Carry-over of caustic reduced and longer cleaning intervals could be achieved. Benefits can also be achieved from chemical agents to reduce scaling. Calcination is a major contributor to the energy used for alumina production. The theoretically achievable limit is 2.4 GJ/t, when the enthalpy of formation for aluminium trihydrate and alumina are compared. This value does not consider energy used for the evaporation of remaining moisture in the filtered trihydrate and other technical limitations, such as a minimum stack temperature of the flue gas. Alunorte uses circulating fluidized bed calciners which use on average 3 GJ/t of energy. Alunorte received the energy efficiency award 2010 for the two newest calciners by the German Energy Agency which operate at 2.79 GJ/t [10].

A high boiler feed water temperature is beneficial for the overall efficiency for steam generation. Some of the boiler feed water is preheated in the calcination area. Potential for further improvement is identified in the rate of live steam condensate from the process. The fraction which does not return is replaced by make-up water. Due to its lower temperature the power plant efficiency decreases with the fraction of make-up water. Alunorte works on a project which will increase the condensate return rate. Some more potential for efficiency improvement exists when higher steam parameters are chosen. 160 bar and 540 °C are typical ranges for circulating fluidized bed boilers [9]. However, this can only be implemented as part of the construction of a new boiler. Furthermore, it needs to be carefully studied if these boilers are suitable as utility boilers of an alumina refinery.

Conclusions Alunorte's is one of the most energy-efficient alumina refineries in the world. The main contributors to this performance are discussed, such as the utilization of high quality bauxites, a good process design, the use of energy-efficient calciners and efficient steam and power generation. The heat integration concept and water balance of the refinery are presented. A good heat integration concept is chosen and Alunorte uses indirect heating only and avoids dilution of plant liquor with steam. Other liquor dilution is generally small. Net wash for example is small. The layout of the utilities area for steam and electrical power generation allows a good operational flexibility and has a generally high efficiency. Alunorte is aware of some factors with a negative impact on energy efficiency. Slowly decreasing bauxite quality, a higher rate of cogeneration of electricity or the ageing of the plant are examples. Alunorte addresses these issues and works

Technology, Operation & Maintenance The potential for direct savings of energy in precipitation (from the first precipitation tank onwards) is limited. Indirectly, however, energy is saved when the productivity of the liquor is maximized. High liquor productivity is important for low specific energy utilization. The higher the liquor productivity the lower is the amount of liquor which has to be heated up in the digestion area and cooled down in the precipitation area per ton of alumina produced. The specific steam consumption in digestion and thus specific consumption of boiler fuels is low. Alunorte is operated

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on a number of projects to implement process improvements, e.g. in bauxite dewatering, bauxite slurry heating, precipitation and bauxite residue filtration. Furthermore, cleaning cycles for descaling of heat exchangers are optimized. Alunorte will continue to identify potential for improvement of energy efficiency and aims to be one of the most energy efficient alumina refineries in the world - also in the long term.

Metals 2011, Lindsay, S., Ed., TMS (The Minerals, Metals & Materials Society).

References 1. International Aluminium Institute (IAI), "Electrical Power used in Primary Aluminium Production", online content, "https://stats. world-aluminium.org/iai/stats_new/formServer.asp?form=7", date of issue Sept 2, 2010. 2. International Aluminium Institute (IAI), "Energy used in Metallurgical Alumina Production", online content, "https://stats. world-aluminium.org/iai/stats_new/formServer.asp?form=8", date of issue Sept 2, 2010. 3. Daryush A. Khoshneviss, Luiz Gustavo Correa, Joaquim Ribeiro Alves Filho, Hans Marius Berntsen and Ricardo Carvalho, "Alunorte Expansion 3 - The new lines added to reach 6.3 million tons per year", In Light Metals 2011, Lindsay, S., Ed., TMS (The Minerals, Metals & Materials Society). 4. Jorge Aldi, "Achieving Excellence in Liquid Effluent Treatment at Alunorte", In Light Metals 2009, Bearne, G., Ed., TMS (The Minerals, Metals & Materials Society). 5. Ramesh Gandhi, Mike Weston, Mani Talavera, Geraldo Pereira Brittes and Eder Barbosa, "Design and Operation of the World's First Long Distance Bauxite Slurry Pipeline", In Light Metals 2008, De Young, D.H., Ed., TMS (The Minerals, Metals & Materials Society), 95-100. 6. Ayana Oliveira, Juarez Dutra and Jorge Aldi, "Alunorte Bauxite Dewatering Station - A Unique Experience", In Light Metals 2008, De Young, D.H., Ed., TMS (The Minerals, Metals & Materials Society), 85-87. 7. R. Bott, T. Langeloh and J. Hahn, "Filtration of Bauxite after Pipeline Transport: Big Challenges - Proper Solutions", In Proceedings of 8th International Alumina Quality Worhhop, Armstrong, L, Ed., AQW Inc.,2008, 319-323. 8. A. Campos, R. Bott, "Determination of a suitable dewatering technology for filtration of bauxite after pipeline transport". In Light Metals 2008, De Young, D.H., Ed., TMS (The Minerals, Metals & Materials Society). 9. Stephen J. Goidich and Ragnar G. Lundqvist, "The Utility CFB Boiler - Present Status, Short and Long Term Future with Supercritical and Ultra-Supercritical Steam Parameters", Proceedings of Circulating Fluidized Bed Technology VII Proceedings of the 7th Int. Conf. on Circulating Fluidized Beds, Grace, J.R., Zhu, J-X., de Lasa, H.I., Eds. Can. Society of Chemical Engineering, Ottawa, May 2002. 10. Michael Missalla, Hans Werner Schmidt, Joaquim Ribeiro Alves Filho and Reiner Wischnewski, "Significant Improvement of Energy Efficiency at Alunorte's Calcination Facility", In Light

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

OPPORTUNITIES FOR IMPROVED ENVIRONMENTAL CONTROL IN THE ALUMINA INDUSTRY Richard Mirtina, John Kildea, Everett C. Phillips, Wayne Carlson, Bruce Keiser, and John Meier Nalco Company; 1601 W. Diehl Rd., Naperville, IL 60563-1198, USA Keywords: Bayer process, mercury, emissions, alumina, coal Abstract Alumina production from bauxite offers a unique set of environmental concerns that affect air, water, and solids. Governments and industry have recognized that reductions in plant emissions and environmental impacts are necessary. The alumina industry is not the only industry that has been subjected to, and responded to, such regulatory scrutiny over the past decade. A number of industry sectors are actively developing innovative ways to control a broad range of potential environmental hazards. Several of these technologies may have direct application in alumina refineries. Methods to significantly reduce mercury emissions in both air and water have recently been developed for use in a range of industries. This paper reviews some of these new technologies now in commercial use in non-alumina plants and considers how they may be applicable within the alumina industry. Introduction The emissions to air and water from alumina refineries are coming under increasing scrutiny by both the industry itself and the communities in which they operate. Such environmental concerns are common to a variety of industries1"6 and alumina producers can learn much by looking outside the industry to other industrial processes where emissions problems are being successfully addressed. In recent years, Nalco has developed a number of strategies designed to reduce air and water emissions of a variety of hazardous contaminants. These technologies have been initially targeted at processes unrelated to alumina refining. Nonetheless, the issues that many of these technologies address - emissions of mercury, sulfur oxides (SOx), and nitrogen oxides (NOx) - are also commonly found as issues in the alumina industry.

bauxite, 500-700 ppb levels are common in Jamaican bauxites, and mercury levels as high as 1200 to 2000 ppb have been reported in some bauxites in Suriname.8 Regardless of the initial level, due to the large volumes of bauxite that are processed in the production of alumina, significant quantities of mercury can accumulate at various points in the Bayer process, ultimately resulting in mercury emissions to air and water of up to several tons per year. Additionally, depending on the source of the lime and raw caustic that are also added to the Bayer process, these too can contribute significant inputs of mercury to the system. Fate of Mercury in the Bayer Process Figure 1 outlines the major sources of mercury input and output in a typical Bayer plant. It is interesting to note mat the potential output streams are dominated by gas phase emissions. While clearly there are some differences, emission to air is also the predominant issue in most coal-fired power plants. As a result, the potential for the use or adaptation to Bayer plant operations of the existing control technologies used in coal-fired power plants is apparent. The mercury contained in bauxite exists as a mixture of different inorganic and organic compounds and oxidation states. Regardless of its initial form, virtually all of the mercury is converted to elemental mercury (Hg°) by the highly reducing environment of the digestion phase of the Bayer process.10 A portion of this mercury can then be emitted to the air via digestion vent gases, especially within the first few stages of flash cooling, and some mercury may be collected in the downstream condensate systems, Figure 1.

This paper focuses on the issue of mercury in the alumina industry and then considers a number of recent developments in the area of mercury emissions control and reviews the benefits and downsides of their implementation in the coal-fired power industry. Consideration is also given to how some of these technologies, which are based on both chemical and engineering solutions, might be applied or adapted within the alumina industry to resolve specific problems associated with mercury in the Bayer process.

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The major source of mercury in Bayer process operations is the bauxite ore used. While trace levels of mercury are distributed throughout the geosphere in soils, ores, and mineral deposits at concentrations typically in the range from 20 to 150 ppb,7 mercury levels in bauxite can vary quite drastically between different sources and even between different deposits from a single source. Mercury levels as low as 20-100 ppb have been found in Boke

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185

Following digestion, a large portion of the mercury may be carried within the red mud slurry. The extent to which the mercury will partition to the red mud solids or to the liquor will be influenced by many factors such as the digestion conditions used, the liquor chemistry and the extent to which mercury oxidation will occur, and the type and quantity of process additives used. These will ultimately influence the concentration of the mercury present in the pregnant liquor. Mercury losses to red mud residues have been estimated to be as high as 0.5 lb/day or more, and these losses may account for 15-60% of the mercury input to the process.7 Upon precipitation of the alumina trihydrate from the liquor, trace levels of mercury can become incorporated in the trihydrate crystals which will be vaporized and emitted into the air during the downstream calcination process. Similarly, trace levels of mercury may be incorporated into oxalate solids that are either coprecipitated with the trihydrate or precipitated by side-stream oxalate removal. Thus, depending on the oxalate removal operations in a specific refinery, additional mercury can exit the process through residue disposal or potentially via air emissions if the oxalate residues are sent to a downstream furnace for lime recovery. The oxalate cake at one refinery was reported to have as much as 2000 ppb mercury and it contributed to a loss of 15% of the input mercury.7 Completing the cycle, small amounts of the mercury may be stabilized in the liquor, which will then be recirculated to evaporation and then back to digestion where it can be lost via air emissions through the barometric condensers or, as mentioned before, it may become a contaminant in the process condensate streams. Digestion condensates have been reported to have soluble elemental mercury concentrations in the order of 20 μg/L.11 Mercury Control Options in the Bayer Process In the course of processing the very high volumes of vent gases and or condensate that are produced from evaporation and digestion at a typical refinery, elemental mercury often accumulates in the traps on condensers and heat exchangers to the point where it can be drained off and collected. This can yield several hundred grams of metallic mercury (Hg°) per day which can be collected and isolated for proper disposal.

stabilization of mercury in solution decreases as excess sulfide is slowly converted to sulfate in spent liquor and residual mercury species will be reduced back to the elemental form in digestion. Zinc, mercury, and other heavy metals may also be precipitated from digestion condensates or pregnant liquor by treatment with a series of Nalco dithiocarbamate or dithiocarbonate products.13 For example, the additive(s) could be added just prior to mud filtration and the resulting residues may be removed from the process with the filter cake through the red mud disposal operations. This is similar to current practice in power plants for the control of mercury levels in industrial wastewater as outlined below. Furthermore, the use of these additives could substantially reduce or eliminate the need for sulfide addition and the corresponding large quantities of organics (and often liquor poisons) that are introduced to the process as a result of their use. There also exists the potential to treat calciner stack gases with the injection of mercury sorbents in between the point where the alumina is collected and the final particulate control device. In addition, the co-injection of a mercury oxidant product, again such as currently used in other industries, could also be considered as a means to lower the sorbent feed rate. The mercury sorbents would of course become mixed with the alumina fines and this would prohibit the recycling of the fines back into the calcination cycle as is usually done. However, such a strategy could still prove to be the most economically and technically feasible, especially for refineries where the mercury tends to partition to the alumina. Some of these control options may likely require consideration of the overall needs of the Bayer process to enable appropriate implementation. They may also require modification for specific plant conditions. However, the successful implementation of such strategies in other industries indicates the potential for improved control using these techniques. In particular, a number of these strategies have been successfully implemented within coal-fired power plants and the lessons learned from this industry can provide useful insights for alumina producers looking to reduce emissions. Mercury Control in Coal-Fired Power Plants Mercury is naturally found in coal in concentrations ranging from 20 to 1000 ppb and coal-fired power plants account for 30% of the global anthropogenic mercury emissions.14 All forms of mercury present in the coal decompose during combustion into the highly volatile elemental form (Hg°) which can readily evade capture by existing air quality control devices typically found at power plants. However, in contrast to alumina refineries where mercury is emitted almost exclusively in the elemental form, in coal-fired power plants a significant portion of the mercury can subsequently become oxidized to an ionic form (Hg2+) depending on the coal type and process conditions. In this oxidized state mercury is more readily captured by fly ash and sorbents and/or scrubber liquors. As a result, it is advantageous to maximize the conversion of elemental mercury to the oxidized form to enhance capture. However, this is not the only strategy. The extent to which the mercury becomes oxidized and is removed from the flue gas is highly dependent both on the type of coal being fired, and the operational conditions employed at the plant.15 This means that mercury emission rates can be highly plant specific and an intimate understanding of the factors that affect mercury transformations and partitioning is crucial to developing a successful control strategy for any given plant.

In addition, gas volumes from digestion vents tend to be much smaller than those from calciners and oxalate furnaces, such that these can potentially be much more readily treated by standard techniques to remove mercury. For example, the vent gases can be chilled to further condense mercury vapors and then passed through an appropriate adsorbent, such as activated carbon columns to remove the mercury.7"9 Indeed, the removal of elemental mercury from digestion condensates by passage through a column containing mercury sorbents has already been demonstrated on a lab scale.11 With respect to mercury in Bayer process liquors, it has been discovered that the use of sodium hydrogen sulfide to precipitate problematic zinc salts can dramatically affect the partitioning behavior of the mercury in the red mud and liquor.10 If a sufficiently high concentration of sulfide ion is maintained in the liquor, the mercury will remain in solution, presumably in the form of [HgS2]2",12 which will help to prevent it from precipitating with the alumina and the oxalate residues. While this can potentially reduce mercury emissions in calcination or in oxalate removal, the

186

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Figure 2. Mercury control technologies and their respective points of application for the control of mercury emissions from coal-fired power plants. As a result, like alumina refineries, there is no single, universal solution for controlling mercury emissions from coal-fired power plants. Rather, a wide range of complementary control technologies can be employed to cater to the broad range of process conditions and plant operations in facilities where better control of mercury emissions is required. By understanding the process, where the problems are and what form they take, a suite of various control measures can be selected and employed together, as necessary, to effectively target the needs of each specific site (Figure 2). While the details of the various control strategies can be complex, a brief overview of some of the key technologies that have been developed by Nalco and are now in use within the industry are presented below.

state (Hg2+) rather than the elemental form (Hg°). To facilitate the conversion to the oxidized state, boiler additives can be applied to increase the relative proportion of oxidized mercury in thefluegas. For example, Nalco has developed a solution product that can be applied directly to the coal prior to combustion. It can be added to the coal at the pulverizers, or injected directly into the furnace. In numerous commercial trials, the fraction of mercury in the oxidized state has been increased to as high as 90%.17 By using such mercury oxidants the performance of mercury sorbents, including activated carbons, is significantly enhanced. This can be a major benefit to plant operations since, by using this technology, a reduction in sorbent feed rates (and thereby operating costs) can be achieved. In addition to reducing sorbent costs, this can also result in the preservation of a plant's ash sales since high carbon feed rates will render the fly ash unsuitable for use as a pozzolan in concrete manufacture. By reducing sorbent feed rates or eliminating the need for sorbents entirely, the value and quality of the fly ash sold by the plant to the concrete market can be maintained.

Coal pretreatment One option to limit the emission of mercury and other contaminants from a power plant is to prevent them from entering the process in the first place. While beneficiation of coal to enhance thermal output has been long practiced, the removal of minor elements such as mercury is more problematic. Nalco has developed a technology known as MagMill™ which can remove a large fraction of many of the heavy metals (e.g. Hg, Se, As) present in coal. This is completed in a dry process before the coal is burned. A combination of crushing, belt separation, and high powered magnets removes abrasive minerals like pyrites. By targeting the removal of pyrites, a high degree of mercury removal can be achieved since mercury in coal is known to be associated with pyrite occurrence.16 The result of processing the coal with MagMill is a product that is a cleaner, less abrasive coal with lower sulfur and heavy metal content.

While the fly ash is nominally a "waste product", the value can be quite substantial for those plants that are able to sell their fly ash. For example, a typical 500 MW plant generating 150,000 tons of fly ash annually that can be sold for $20/ton, risks $3 million per year in lost revenue by injecting activated carbon as a sorbent to control mercury emissions. In addition, the plant may incur an additional fly ash disposal cost of at least $1.5 million per year as a result. Mercury sorbents One of the most promising technologies to recently emerge for the control of flue gas mercury emissions is the use of mercury sorbents. Powdered sorbents such as activated carbon or engineered inorganic sorbents can be injected into the ductwork upstream of a plant's particulate control device to capture mercury.19 In this way, mercury is removed from the flue gas and

Mercury oxidation / speciation control Capture of mercury in downstream particulate control devices and/or scrubbers can be enhanced if the mercury is in the oxidized

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becomes commingled with the fly ash which can then be safely disposed of or, depending upon the quality of the ash, used in various applications including concrete, pavement, structural backfill, and brick manufacture. Activated carbon injection (ACI) is by far the most established technology in this area with over 60,000 MW of commercial bookings at over 150 coal-fired plants in North America as of June of 2010.20

mercury capture, however the target of 90% capture was only reached using an ACI rate well above 3 lb/MMacf.

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The application of mercury oxidation technology as described above is especially well suited to those plants that have wet flue gas desulfurization (FGD) units installed for SOx control, where the ionic form of mercury readily dissolves in the aqueous liquor phase. With the bulk of the mercury shifted from the gas phase to the aqueous phase, a coagulant product can be applied to sequester and precipitate the mercury from FGD liquors without any concomitant measurable increase in the mercury content of FGD byproducts. This also has the effect of suppressing the phenomenon of FGD mercury re-emission whereby oxidized mercury in the aqueous phase is reduced back to the elemental form and emitted from the stack.21 Mercury removal from wastewater In order to meet challenging industrial wastewater mercury discharge limitations, companies often turn to precipitation aids. For this application Nalco has developed polymeric chelants with an exceptionally high affinity for mercury. Upon binding mercury from solution, it forms large precipitates thatflocculate,settle, and are readily filtered to consistently attain extremely low mercury levels in the parts-per-trillion range. Case Studies The various technologies outlined above can be applied, either individually or in many cases in concert, to reduce and control mercury emissions. These strategies have been applied in a number of commercial power plant operations with excellent results. Two examples of how these technologies have been applied, each outlining how different control strategies can be used to achieve the desired results, are presented as case studies. Case study #1: Site description: 600 MWe pulverized coal-fired boiler firing subbituminous coal with low NOx burners, overfire air, and cold-side electrostatic precipitator (ESP).

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The overall aim of the operation was to increase mercury capture without impacting plant operations. Therefore, simply increasing the activated carbon injection (ACI) rate was not an option, as this would have compromised the quality of the fly ash produced at the plant, which was being sold for use in concrete. Thus, a 90% mercury capture with an ACI rate of less than 2 lb/MMacf was sought.

In this case, two separate issues needed to be addressed. First, improved capture of the mercury from the gas phase was required. Similar to the previous example, MerControl 7895 was applied to the coal feed to help achieve this. Additionally, the discharge of mercury in the plants wastewater was strictly controlled. While the improved capture from gas could be achieved, further containment and control of the mercury from the aqueous environment was required. For this purpose, a polymer chelant product (e.g., Nalmet® 1689) was used to treat the FGD wastewater.

The "normal" operation involved use of a halogenated carbon added to the outlet of the air pre-heater. Under such an operating mode, using an ACI rate of 2 lb/MMacf, the reduction in mercury was less than 80% (Figure 3). Increasing the ACI rate improved

188

Figure 4 shows the stack emissions of total mercury (in both elemental and oxidized forms) plotted across a range of dose regimes of the MerControl 7895 product. The same data are plotted as a percent of mercury capture. Application of MerControl 7895 at a rate of 265 mg/kg of coal resulted in greater than 85% mercury capture and achieved the target of 0.008 lb Hg/GWh. The net result of this treatment was that the bulk of the mercury reported to the FGD liquor. 100%

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Figure 6. A plot of total mercury concentration of the FGD wastewater clarifier solids as a function of mercury capture from air. As capture of mercury from air increases, the concentration of mercury in the solids increases proportionately due to the application of Nalmet 1689. The use of coal as a power source for Bayer plants is common with coal accounting for approximately 20% of the energy used in the production of alumina.22 In many cases, the Bayer refinery operates their own coal-fired power plant as part of the overall refinery operations. Therefore, as noted above, such power plants can be a significant source of mercury emissions and in such cases, the control technologies described above, will clearly have a direct fit to the industry.

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Atmospheric emissions from alumina refineries are increasingly a problem with mercury being a common issue across the industry. However, similar issues are being faced by a number of other industries and a range of potential solutions are being developed and used for air emission and wastewater control generally, and in particular for mercury capture and control.

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189

9. Alcoa,'Emssions & Waste', (2010). http ://www. alcoa. com/sustainability/en/info_page/operations_e nvemissions.asp

Studies of various Bayer refineries have shown that both the amount of mercury emissions and the points in the process at which they are emitted vary quite drastically from refinery to refinery.7 The complexity of the problem stems from the various sources of bauxite and the differences in operating conditions employed at particular refineries. As is the case with coal-fired power plants, the issue of mercury emissions is so highly plant specific that a detailed knowledge of the various process steps and chemistries as well as the ability to accurately measure mercury in various process streams will be essential to delivering a customized solution for each individual refinery. It is expected that the aforementioned mercury control technologies in use at coal-fired power plants can be adapted to meet the needs of the alumina industry. Whether this will involve the removal of mercury from digestion liquors or the oxidation and sorption of gas phase mercury from digestion vents and calciner stacks, or some combination thereof, will ultimately depend on the particular needs of the refinery in question.

10. A. B. Livingston, 'Estimation of a mass balance for mercury at a Bayer alumina refinery, using thermal decomposition, amalgamation, and atomic absorption as an assessment method, M. S. thesis', (Texas A&M University, 2005). 11. M. Mullett, J. Tardio, S. Bhargava and C. Dobbs, 'Removal of mercury from an alumina refinery aqueous stream', Journal of Hazardous Materials, 144 (2007), 274-82. 12. H. G. Wildman, Process of recovering mercury', US Patent 1,896,876, (1933). 13. J. T. Malito, 'Process for removing heavy metals from a caustic fluid stream, US Patent 6352675', (2002). 14. A. P. Jones, Hoffmann, J. W., Smith, D. N., Feeley, T. J., Murphy, J. T., 'DOE/NETL's Phase II Mercury Control Technology Field Testing Program Updated Economic Analysis of Activated Carbon Injection, Preparedfor the U. S. Department of Energy', (2007).

References 1. D. G. Streets, Q. Zhang and Y. Wu, 'Projections of Global Mercury Emissions in 2050', Environmental Science & Technology, 43 (2009), 2983-8.

15. A. Kolker, C. L. Senior and J. C. Quick, 'Mercury in coal and the impact of coal quality on mercury emissions from combustion systems', Applied Geochemistry, 21 (2006), 182136.

2. A. Cain, 'Mercury Releases from Industrial Ore Processing', US EPA, (2005). http://www.epa.gov^s/reports/stakesdec2005/mercury/Cain2. pdf.

16. B. Toole-O'Neil, S. J. Tewalt, R. B. Finkelman and D. J. Akers, 'Mercury concentration in coal-unraveling the puzzle', Fuel, 78 (1999), 47-54.

3. E. D. Thoma, C. Secrest, E. S. Hall, D. Lee Jones, R. C. Shores, M. Modrak, R. Hashmonay and P. Norwood, 'Measurement of total site mercury emissions from a chlor-alkali plant using ultraviolet differential optical absorption spectroscopy and cell room roof-vent monitoring', Atmospheric Environment, 43 (2009), 753-7.

17. B. A. Keiser, G. Finigan, J. Meier, J. Shah and J. Lu, '#145, Mercury Emissions: Demonstration of Air and Water Quality Management', MEGA Symposium (Baltimore, MD, 2010).

4. J. B. Milford and A. Pienciak, 'After the Clean Air Mercury Rule: Prospects for Reducing Mercury Emissions from CoalFired Power Plants', Environmental Science & Technology, 43 (2009), 2669-73. 5. D. Wolf, 'Understanding the Industrial Boiler MACT Rule', HPAC Engineering, (2010).

18. K. H. Pedersen, A. D. Jensen, M. S. Skjoth-Rasmussen and K. Dam-Johansen, Ά review of the interference of carbon containing fly ash with air entrainment in concrete', Progress in Energy and Combustion Science, 34 (2008), 135-54. 19. H. Yang, Z. Xu, M. Fan, A. E. Bland and R. R. Judkins, 'Adsorbents for capturing mercury in coal-fired boiler flue gas', Journal ofHazardous Materials, 146 (2007), 1-11. 20. Institute of Clean Air Companies (ICAC), 'Updated Commercial Hg Control Technology Bookings (June 2010)'. http://www.icac.com/files/members/Commercial_Hg_Booking s_060410.pdf

6. US EPA, 'EPA Sets First National Limits to Reduce Mercury and Other Toxic Emissions from Cement Plants', (2010). http://yosemite.epa.gov/opa/admpress.nsfe77fdd4f5afd88a385 2576b3005a604ffef62balcb3c8079b8525777a005af9a5!Open Document

21. G. Blythe, J. Currie and D. DeBerry, 'Bench-scale Kinetics Study of Mercury Reactions in FGD Liquors', Final Report — DE-FC26-04NT42314, (2008).

7. C. Dobbs, C. Armanios, L. McGuiness, G. Bauer, P. Ticehurst, J. Lochore, R. Irons, K. Ryan and G. Adamek, 'Mercury emissions in the Bayer process - an overview*, Proceedings of the Seventh International Alumina Quality Workshop (Perth, Western Australia, 2005), 199-204.

22. International Aluminium Institute, 'Form ESO 12 - Energy Used in Metallurgical Alumina Production', (2009). https://stats.worldaluminium.org/iai/stats/formServer.asp?form=8

8. Bauxiet Instituut Suriname, (2006). http://www.bauxietinstituut.com/files/Environment%20%20an d%20Technology%20-%20Suralco%20LLC.pdf

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ALUMINA REFINERY WASTEWATER MANAGEMENT: WHEN ZERO DISCHARGE JUST ISN'T FEASIBLE.... Lucy Martin1 and Steven Howard1 ^echtel Australia, 100 Brookes Street, Brisbane, Queensland 4006, Australia Keywords: "Alumina Refinery, Waste water, Environmental Design Criteria" of the refinery will be strongly influenced by the investment required for materials transportation. Project NPV will in turn be impacted by the operating costs associated with that location. As commodity prices increase and accessible deposits become depleted, the need to develop inland ore bodies, and consequently refineries, is increasing.

Abstract Management and treatment of liquid effluents are determinant considerations in the design of alumina refineries. Rainfall, evaporation rate, proximity to the coast, process design and layout, ore mineralogy, the local environment, and potential impact on contiguous communities are all integral to the development of an appropriate refinery water management strategy. The goal is to achieve zero discharge of liquid effluent to the environment. However this is not always the most feasible solution under the extreme rainfall conditions in tropical and subtropical locations. This paper will explore the following issues for both inland and coastal refineries: • • •

Bauxite is formed through weathering of lateritic rock under high rainfall in tropical and subtropical environments. Figure 1 shows regions currently undergoing lateritic weathering. These regions tend to have well defined wet and dry seasons, which forces refinery designers to address the "feast or famine" issue associated with water availability. Zero discharge of effluent is ideal for sustainability, but is often impractical for water impoundment.

Methods to reduce and control refinery discharges Treatment design criteria Socioeconomic aspects relating to surface water use in settlements adjacent to the refinery

Bauxite also contains organic and inorganic impurities which increase the costs and environmental risks associated with its processing into alumina. This is primarily due to the solubility of the various minerals and organic compounds in the Bayer process liquor, and the tendency of the impurities to accumulate therein. Demonstrated treatment methodologies for the large scale effluent neutralization required for discharge into surface or marine waters primarily involve the use of either seawater alone, or seawater supplemented by spent sulfuric acid, a byproduct of the refining process. When considering refineries close-coupled to inland bauxite deposits, the cost of pumping seawater must be set against the cost of bauxite and alumina transportation.

Introduction With an increased global focus on sustainable development from governments, international financing institutions, mining companies, and communities, new and expanding mining operations can face a higher level of environmental and sociological regulation and scrutiny than in the past. Many legislative frameworks allow existing operations to continue to meet a less stringent standard than would be expected on a greenfield facility. Therefore this paper will pay most attention to those regions expected to become future alumina sources and the associated technical barriers to sustainable and viable projects.

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In summary, this indicates that: 1. Water availability for refinery consumption will be highly variable from season to season 2. Water discharge will be required during both "extreme" or "normal" rainfall events. 3. In inland locations, water effluents may discharge into streams potentially used for drinking and primary industries (farming, animal husbandry, and fishing) 4. Typical refinery waste water treatment techniques, such as seawater neutralization, may not be feasible.

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Refinery Water Balance Water supply and demand An alumina refinery that utilizes the Bayer process will consume 2.0-2.3 tonnes of raw water per tonne of alumina produced. The actual rate will depend upon the bauxite quality, the process design, demand for nonprocess applications (e.g, for potable water), and the extent water is recycled within the facility. About 10 percent of the total intake is accounted for by free moisture in the bauxite feed, and in the 50 percent caustic soda solution, the primary process reagent.

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Figure 1. Lateritic Bauxite Sources (modified from Freyssinet et al [l]) The relationship of a refinery location to its bauxite deposit and transport infrastructure directly affects material handling equipment costs and resultant project NPV. The optimum location

Figure 2 depicts the inputs and outputs of water across the outer system boundary between the refinery, including the residue

191

disposal area, and the external environment. This assumes one hundred percent diversion of potential 'run-on' to the site.

various facilities. These measures minimize the probability of releasing highly concentrated, potentially toxic material which can only be discharged under the most extreme circumstances. Environmental control measures for an alumina refinery must also include facilities to handle contaminated runoff, due to rainfall catchment within the facility perimeter. One or more stormwater ponds (SWPs) would also be provided, and must be considerably larger than the process spill ponds. An SWP is intended to impound rainfall runoff collected from nonprocess areas, which will mobilize relatively minor amounts of surface contaminants during the initial onset of rain. After a short period, typically about one hour, the runoff quality is similar to that found in neighbouring areas outside the refinery perimeter. The impounded SWP water is impure, but is suitable for recycling into the process for various duties, reducing the intake of raw water. Finally, the impoundment area constructed for the permanent storage of bauxite residue and other solid wastes, the bauxite residue disposal area (BRDA), also accumulates contaminated water, which may be returned to the process in order to recover the soda content.

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Figure 2 Refinery Water Balance There are only two sources of water that supply the remaining ninety percent of the total demand: raw water from natural surface and subsurface sources, and rainfall. This means that a significant proportion of the water intake is noncontrollable, and under certain circumstances may far exceed the capability of the refinery to control the quantities involved. With the notable exceptions of the People's Republic of China (PRC) and Russia, most refineries have been established in tropical and subtropical regions, in close proximity to the principal lateritic bauxite provinces. These are not only exposed to high seasonal rainfall, but are also at risk of extreme and unpredictable flood events. New capacity planned for Brazil, Guinea, and SE Asia will be faced with the same issues.

Under normal operating conditions, the Bayer plant liquor inventory is controlled within narrow limits by varying the input rates of caustic soda, raw water and in-plant evaporation. Operating procedures should be aimed at keeping the process spill ponds empty, and the water levels in the SWP and BRDA as low as possible, during periods of little or no rain. The onset of heavy and/or sustained rainfall, typical in tropical locations, gives rise to large volumes of site runoff, and potentially run-on if this is not effectively diverted. This can rapidly fill the available impoundments, which have a finite limit despite the risk analysis and major investment involved in their provision. In order to handle situations that exceed the maximum impoundment capacity, the regulatory option often exists to impound only the so-called "first flush" runoff, after which all subsequent catchment would be allowed to bypass the SWPs and be released to the environment. Exceptional rainfall events would defeat even this strategy (assuming that it were permissible) and the refinery would then be forced to release contaminated water to the environment. Such a release may be diluted to mitigate the alkalinity of the effluent, but would still result in contaminating receiving waters above background levels, probably in violation of licence provisions.

Environmental control requirements It has long been recognized that effective environmental management is critical to the viability of any project. Stringent environmental control standards and industry best practice with regard to operations and maintenance must be reflected in the refinery design criteria.

Community demand, including agriculture, for water may become a significant component of total consumption. Any shortage or degradation of the community supply may become a significant issue. These factors need to be considered in the overall conceptual design of the water supply system.

The focus of this paper is the management of liquid effluent, any release, either treated or untreated, of which is potentially harmful to the environment and consequently deleterious to the wellbeing of the community. The Bayer plant itself handles a large volume of process liquor, the bulk of which is an aqueous solution of caustic soda containing dissolved aluminium, silica, and many other organic and inorganic impurities, including trace metals such as Molybdenum and Vanadium, which occur in the bauxite.

The conclusion is that despite careful scenario planning and investment in major infrastructure, sooner or later an extreme event will cause an unacceptable environmental incident. Therefore, the location of the refinery must be such that the receiving waters are able to withstand the impact in a sustainable manner. This explains why most major refineries are located near the ocean, which also facilitates the application of seawater neutralization, the only current demonstrably effective liquid effluent treatment technology.

The refinery is designed to contain the live liquor inventory, with minor spills within operating facilities being promptly returned into the process. Provision is also made to intercept larger, accidental spillages which could result from equipment failure or abnormal events, such as the loss of electrical power. Spill ponds are installed at strategic locations, and have sufficient capacity to handle the contents of one or more of the largest tanks in the

Refinery design principles The unavoidable accumulation of contaminated water due to rainfall, reclamation constraints, and severe limitations on release to the environment have the following consequences for the design of the refinery:

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1.

areas is perceived to improve employee safety by minimizing the hazard of caustic liquor exposure, but this may be achieved by proper safety procedures and training, whilst minimizing water usage.

Refinery environmental policy that mandates compliance with relevant regulatory standards becomes a major driver for selection of the refinery and BRD A locations.

This policy also has a profound impact on refinery design features and the operational procedures that enable the standards to be met.

4.

A policy mandating compliance with liquid effluent standards is consistent with that for particulate emissions or any other form of environmental impact. The difference between implementation for particulates (e.g., the installation of dust collectors) and that for release of impounded water is one of relative complexity for the latter and requires coordination of the total refinery design.

Water consumption may be allocated to two broad categories: noncontrollable, governed by the combined requirements for essential services, and controllable, which allows for some degree of flexibility as to the quantity used. Most of the process requirements fall into the noncontrollable category, for example, boiler feedwater, flocculant, and lime preparation. All are examples of end uses directly related to the refinery production rate and to the demand for raw materials and additives. Domestic water demand is closely related to the number of employees. Effective control depends upon the installation of water saving devices such as tap restrictors, employee education regarding conservation, and the prohibition of usage for nondomestic purposes.

Facility design criteria must address the (water) environmental policy, ranked among safety and NPV considerations, and evaluate this before detailed design can proceed. Some objectives may be compatible, for example minimizing the caustic content of final stage mud washer underflow results in a significant reduction in operating cost. 2.

Departures from conventional refinery and waste impoundment design become necessary, and can add to the mitigation costs of potential effluent management problems.

The final disposition of water usage in this category is important in limiting the potential environmental impact. Applications such as dust mitigation can help to dispose of excess impounded catchment under dry season conditions. Others, such as vehicle washing, may create additional problems caused by runoff turbidity or contamination by hydrocarbons and other chemical agents.

A refinery location that will experience extreme rainfall events should incorporate features to minimize both the volume effects and possible contamination of rainfall catchment. This facilitates the segregation of concentrated process liquor (that cannot be released) from relatively innocuous nonprocess runoff that can be impounded, recycled during dry periods, and released subject to meeting licence conditions.

The supply of water to end users beyond the outer system boundary should be from an independent source of supply. This measure imposes a physical constraint on consumption and avoids uncontrolled influence on the infrastructure provided for the refinery itself.

Waste impoundment design must satisfy all of the following criteria: • Effective, permanent sequestration of red mud waste that will continue to generate alkaline leachate over the longer term • Recovery of supernatant liquor and impoundment drainage back into the refinery process. • Sequestration of solid-phase organic and inorganic waste from process liquor purification facilities • Disposal of other environmentally sensitive wastes (e.g., ash, scale, and waste acid) 3.

Introduction of measures to enforce economic use of controllable process and domestic water can significantly reduce usage.

5.

Several tiers of impoundments must be established to segregate process liquor, contaminated effluent and rainfall catchment.

• Process spill containment A "process spill" is defined as any form of liquor, slurry, or solids released during routine operational tasks within the plot limit of any facility. Spills of this nature are normally of low quantity and are promptly collected and returned into the process. The concentrations of process chemicals are too high to permit release to the environment.

Measures must be adopted to minimize the controllable intake of raw water, particularly during the wet season, thereby reducing the total inventory to be managed.

There is a possibility that large process spills may occur due to equipment failure, such as a pipeline rupture. The refinery design should incorporate process only spill ponds for the purpose of intercepting and returning material that may overtop the limited facility containment capacity. The contents of spill ponds cannot be released, and should be recycled to the refinery via large sump relay tanks (SRTs), which provide additional surge capacity and operational flexibility.

Due to the probable excess of rainfall over evaporation in the regions under discussion, raw water intake must be strictly controlled to minimize the total volume under management. It follows that the consumption of raw water in the Bayer process should be minimized by using as much water as possible that is already within the outer system boundary. Raw water is needed for a number of essential process duties, treated for domestic (potable) water, and for the fire control infrastructure. The number of water entry points into the refinery must be minimised, and all volumes accurately metered. Strong design discipline is necessary to ensure consistent water control policies and practices are enforced. Some measures may conflict with custom and practice employed elsewhere, but capital and operating costs associated with treatment of this water may be excessive. For example, using hoses for housekeeping in process

Process design modelling should incorporate provision for large process spills, to ensure total containment facilities and return rates into the process are adequate, and that defined events of this nature do not compromise the refinery's rated production and other key parameters. •

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Runoff management

"Runoff is defined as rainfall catchment from the refinery plot limit that must be monitored for contaminants and handled accordingly. Runoff falls into two categories: •

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temperature and turbidity) associated with the effluent itself and the background conditions in the receiving waters. • Final Stage Mud Washer Underflow Slurry The residue slurry contains solid phase compounds that will release highly alkaline leachate over a long time frame due to unavoidable contact with supernatant liquor and rainfall. This tendency cannot be controlled at source, due to the presence of sodium aluminium silicates and calcium compounds formed in the Bayer process. The only control, at the margin, is to minimize the concentration and quantity of liquor disengaged from the residue, and to reduce direct contact of the consolidated mud with liquor or rainwater.

Catchment within process facility plot limits. This may be relatively dilute, but is still far too contaminated for release. It must be returned to the process via the SRT system, and will have negative impacts on the process and energy demand. These effects must also be considered for the refinery mass and energy balance. The impact must be minimized by reducing the process catchment footprint, i.e., by roofing large tanks and routing rainwater catchment to surface drains outside the facility plot limit. Catchment from nonprocess areas. Runoff will mobilize soluble contaminants such as oil, dust, and dirt, the concentrations of which may be appreciable depending upon such factors as paving and drainage design, control of fugitive dust from the refinery, and general housekeeping standards. Under tropical rainfall conditions, the initial (or "first-flush") nonprocess runoff will exhibit short term contaminant loadings that will be sampled and the flow directed to a large Storm Water Pond (SWP).

The major controllable variables for residue disposal are the density of the slurry, and the alkaline concentration of the liquor discharged with the slurry into the BRDA disposal area. Paste thickening is the most effective means of preparing very dense, immobile slurry which will release little or no liquid as it further consolidates under its own weight. The remaining design issue is to minimize the final stage mud washer underflow total soda, usually measured as g/L Na2C03. A typical target is 5 g/L, estimated from the number of washing stages to be installed, the ratio of wash water to mud, and other factors. The target is estimated by the steady state refinery mass balance, and by making certain assumptions related to equipment reliability, washer stage efficiency, etc. Under actual conditions, major departures from steady state operation and other assumptions cause excursions in last washer soda, by as much as a factor of 10, to 50 g/L or more. This implies that the number of installed washing stages and the sparing of pumps should go beyond that suggested by steady-state modelling; otherwise the environmental control objectives will be unachievable.

In practice, the volume of runoff from "average" rainfall intensity can be far greater than the SWP capacity, as costs to construct these facilities to collect the rainfall volumes experienced in tropical and sub-tropical locations can be prohibitive. This can mean that all but a small fraction of nonprocess runoff must be released, and the refinery must be designed with this in mind. The water diverted to the SWP will be reclaimed as "process water", diverted to the BRDA buffer zone, or (as a last resort) treated and released. There is a tendency to work with "average" rainfall data that can be highly misleading, and could result in erroneous assumptions underlying the planning for the capacities of spill ponds, SRTs, the SWP, and the BRDA. A credible environmental control strategy should be based upon scenario planning for abnormal events, such as 1:100 year return rainfall events.

Discharge Criteria Where bauxite mining and alumina refining are emerging, local and national legislation and associated compliance monitoring are often not sufficiently developed to address contaminants of interest associated with these activities. Consequently, drinking water, primary industries and ecosystems may not be adequately protected by the existing environmental legislative framework.

• BRDA design The BRDA is one of only two impoundments from which lowlevel contaminated water may be released (the other being the SWP), provided that the regulatory and/or best environmental control practices can be addressed. It may be possible to optimize the total capital investment for the two impoundments by linking them, such that the BRDA provides the final and somewhat larger capacity for low level effluent.

Many international mining and minerals processing companies have implemented sustainable development policies which require them to not only meet the requirements of in-country legislation, but to also consider the use of the best available technologies and international guidelines. For projects seeking external financing, the International Financing Corporation (IFC) has developed specific Environmental Health and Safety (ES&H) guidelines that are "technical reference documents with general and industryspecific examples of Good International Industry Practice (GIIP)". These guidelines are used by the IFC for project appraisals. IFC EH&S guidelines have been developed for both mining and base metal smelting and refining [2,3]. However these may not include all contaminants of interest relevant to alumina refinery effluent discharge. For example, the base metal and smelting guidelines state that toxicity should be considered on "a case specific basis" and do not provide guidelines for discharge of heavy metals associated with alumina refining.

The design of the BRDA must prevent contamination of impounded runoff outside the deposition areas for red mud and the other solid wastes that have special sequestration criteria. Liquor that is either released from the residue slurry or accumulates from rainfall directly onto the mud deposition area must be recycled to the refinery. A buffer zone between the active solid waste deposition zones and the outer BRDA perimeter should be provided to allow for abnormal rainfall events that may temporarily exceed the active zone's capacity. This liquor, and any other leachate from other waste disposal areas, should not be allowed to mix with impounded low level water.

Assuming that local legislation is not sufficiently defined, guidelines from international organizations and other countries with highly developed legislativeframeworkscan also be used to develop robust design criteria for effluent discharges from the project site.

The remaining area of the BRDA may be used to store and reclaim, or to treat and release, excess water. The final treatment and disposition method will be site-specific, and will depend upon a number of factors (alkalinity, toxic ionic species, BOD,

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An alumina refinery will generate large volumes of liquid and solid phase alkaline effluent from a variety of sources, predominantly waste mud and other process byproducts discharged with the mud, including:

Design criteria for effluent disposal must consider the environment to which the effluent will be discharged - a primary consideration is whether the discharge is to fresh water or the marine environment. Discharge quality requirements to marine water are often less stringent than the discharge to fresh water due to various factors, including higher naturally occurring baseline concentrations of compounds in marine water compared to freshwater, increased dilution of effluent when discharged to open water, and less sensitive receptors from both an ecosystem and downstream user perspective in saline water.

• • • •

Surface water is used for drinking, recreation, and primary industries in many countries. In developing countries it is more likely to be used for these purposes without treatment. Strict and enforced industrial zoning legislation is required where industrial effluent is discharged to control population influx and resultant water usage; however this is often lacking in the regions currently experiencing growth in bauxite mining and associated refining. In these situations it must be considered that human and animal ingestion of surface water could take place at any point outside of the industrial fenceline, therefore drinking water quality standards, as a minimum requirements, must be attained at inland refineries discharging to streams and rivers.

Diluted process liquor entrained with the bauxite residue (red mud) Insolublefractionof the bauxite (iron and titanium oxides) Desilication product (DSP) - hydrated sodium aluminium silicates Calcium compounds, e.g., tri-calcium aluminate, calcium oxalate.

The other effluent streams will be generated by contamination of rainfall runoff from the refinery site, and possible spillage and release of process materials to impoundments. Effluent that is released from the waste mud stream will invariably exhibit pH levels above 12. Both the liquid and the solid phase sources of alkalinity (hydroxide ions) must be reacted with a neutralizing agent. The rate at which the neutralization reaction proceeds will vary greatly depending upon the reactants involved, the pH, concentration, and temperature. Principles of Effluent Treatment The primary objective is to reduce the immediate and longer term environmental risk of the (solid or liquid) waste stream by reducing the alkalinity to the minimum practicable level. However, low alkalinity is not the only consideration. Potentially serious soluble pollutants such as aluminium and other elements (e.g., molybdates, vanadates and arsenates) must also be targeted.

An alternative is to provide a separate secure drinking water supply to the local population and educate them regarding withdrawal of water, which may limit the treatment costs for effluents. If this can be achieved and is accepted by permitting authorities, then it may be possible to discharge at a higher concentration and allow for mixing until drinking water, or other international standards, are achieved at a compliance point downstream of the discharge location. Mixing zone calculations need to take into account the lowest surface water flowrate when calculating dilution, as this can often mean that discharge criteria to achieve compliance at a the edge of a mixing zone need to be more stringent during the dry season. This applies to impurity concentrations, temperature, and turbidity (visual impact)

Options for neutralization include reaction with seawater, dilute sulfuric acid (or a combination of both), and carbon dioxide. A third option, carbonation, is in the early stages of development and little process performance data is available to suggest this as a viable option at this phase of development. Application of either seawater or acid neutralization must recognize that a certain fraction of solid phase alkalinity is released over long periods of time, weeks or months, depending upon the prevailing conditions and the composition of the residue. In practical terms, there is no such thing as complete neutralization, due to the limited treatment time available.

Non-Governmental Organizations (NGOs) are an increasingly prominent force in the enforcement of environmental and sustainable development policies. There is an increasing public demand for independently verified information and action, due to the poor performance of some regulators and companies. Therefore, in countries without sufficient environmental legislation to protect the health of local communities, their livelihoods, and the surrounding flora and fauna, large corporations are often under pressure to implement stringent environmental and sustainable development policies to protect their reputation.

Treatment with sulfuric acid invariably involves the utilization of spent dilute acid used for cleaning refinery heat exchangers. During the cleaning process, the acid dissolves scale deposits which may contain additional pollutants. A corrosion inhibitor will also be employed, the nature of which must be assessed if the reacted acid is to be released into the environment.

In summary, development of project specific effluent discharge criteria to meet sustainable development policies will require a vigorous social, health, and environmental impact assessment of existing and potential future uses, and a "one size fits all" solution cannot be applied. However, generally speaking, discharge quality requirements to marine rather than aquatic environment are less stringent and consequently more achievable.

The chemical and physical properties of the particular bauxite to be processed have a determinant impact upon the process design of the Red Side of the refinery, upon the BRDA design, and upon the technology selected for effluent treatment. Significant effort and cost must be invested in the characterization of the bauxite, and the same attention must be paid to the environmental control requirements.

Effluent Neutralization Technologies

Basis of Design The Refinery Basis of Design must specify the following fundamental criteria: Refinery location. . • Will it be practicable to utilize seawater neutralization? Climatic conditions - rainfall and evaporation.

In this context, the term neutralization refers to the reduction of the level of alkalinity in the stream to be treated. Neutralization processes are required for effluent that is to be permanently sequestered, or is to be released to the environment where there is no other alternative.

195

hydrotalcite precipitation will occur at the outfall, creating a visible plume.

•

Will it be necessary to release effluent to the environment, and what quantities are likely? Receiving waters for effluent release. • What are the social, agrarian and environmental factors associated with effluent release? • What effluent discharge standards should apply? • What are the critical parameters for the outfall point? Waste streams to the BRDA • What tonnages of solids and liquid waste are to be discharged? • What will be the composition of the red mud? • What variability in waste loadings and concentrations may be expected? • What is the probable toxicity profile, and does mitigation technology exist?

It should be noted that attempts have been made to augment seawater neutralization with sulfuric acid. This significantly alters the chemistry such that a much lower pH is necessary to remove aluminium from the solution. Removal of oxalate is favoured at high pH values, so that the addition of acid may be counterproductive and costly. Alternative Technologies Other technologies for industrial water treatment, such as membrane treatment and ion exchange, are untested on alumina wastewater chemistry and at the scale discussed in this paper. If these technologies were proven to be effective, the volumes that would require treatment in tropical and sub-tropical locations would result in significant increases in capital and operating costs for these facilities.

Acid Neutralization Residue is mixed with dilute acid which reacts immediately with soluble alkalinity, producing a rapid but temporary drop in pH. This is commonly achieved by adding (waste) acid after the last mud washing stage. Attack by residual acid (if any) on the solid phase alkaline content occurs over a much longer time frame, and may lead to a gradual increase in pH. It is therefore impractical to neutralize the solid phase component prior to residue disposal. Post-neutralization of waters released from the BRDA may be necessary.

Conclusion In tropical and subtropical climates, coastal and inland refineries will discharge effluent under normal and extreme circumstances. To minimize the amount of effluent requiring treatment and discharge, the environmental design features and operational controls for management of raw water intake, and the water balance detailed in this paper should be incorporated into the alumina refinery design.

Acid neutralization produces a dilute sodium sulphate solution, which, if released, may give rise to environmental impact in its own right, such as algal blooms or local concentrations exceeding background levels, or that specified for potable water (<250 mg/L). Careful assessment must be made to establish that sufficient acid will be available for primary/secondary residue or effluent treatment, that probable peak effluent discharge rates can be handled, and potential heavy metal contamination of spent acid is acceptable.

Inland refineries pose additional challenges due to the stringent discharge criteria applied to inland waterways. Discharge criteria can be imposed, such as government legislation or financial institution requirements, or prescribed by company internal sustainable development and environmental policies. In either circumstance, effluent discharge requirements are often defined according to the expected use of the water body by contiguous communities. The most demonstrated treatment method for large scale refineries is seawater neutralization. Transport of seawater to an inland refinery and the reverse for discharge can put a strain on the project capital and operating costs, and may not be feasible.

Seawater Neutralization The important reaction in seawater neutralization is the precipitation of hydroxyl ions by reaction with magnesium (Mg++) ions present in the seawater. Again, the rate of reaction varies greatly - rapid in the liquor phase, much slower with calcium compounds and DSP. The presence of sulfate ions will inhibit the reaction with DSP.

Other water treatment methodologies are not proven on large scale refineries required to meet stringent inland water requirements and may require expensive reagents to treat the effluent. These can increase the risk profile of the project, drive up capital and operating costs and may result in breaching of discharge design criteria, with consequential downstream impacts.

A major advantage of seawater neutralization is the precipitation of aluminium ions during the formation of hydrotalcite, the primary reaction product from the soluble alkalis and magnesium. Hydrotalcite formation has also been shown to remove vanadium, molybdenum and phosphorus provided that the pH is in the range 8.0-10.0.

Acknowledgements Thanks to Bob McCulloch and Bill Imrie for their valuable comments and Rosemary Dotlic, Daniel Hayes & Eric Tlozek from the Bechtel Communications team for their assistance in formatting and editing this paper.

Seawater neutralization of residue allows immediate effluent discharge, provided turbidity criteria can be attained, eliminating the need for separate containment and management of liquor. By providing excess seawater, alkaline runoff generated by the slow dissolution of alkaline compounds can be neutralized and released.

References 1. Ore-forming Processes Related to Lateritic Weathering; P.H Freyssinet, C.R.M Butt, R.C Morris and P. Piantone. Society of Economic Geologists Inc. Economic Geology 100th Anniversary Volume p681-722 2. IFC General Environmental, Health, and Safety Guidelines: Mining, December 10, 2007 3. IFC Environmental, Health, and Safety Guidelines Base Metal Smelting and Refining April 30, 2007

System design must ensure that the seawater supply and discharge capacity always exceeds the magnesium demand of excursions in residue alkalinity, due to process problems. If the pH is allowed to rise, some of the trace metals in the hydrotalcite will revert to the soluble phase and impact on receiving water quality. If excess alkalinity cannot be precipitated within the BRDA, additional

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

HIGH PURITY ALUMINA POWDERS EXTRACTED FROM ALUMINUM DROSS BY THE CALCINING—LEACHING PROCESS LIU Qingsheng u , ZHONG Chunming1, FANG Hui1, XUE Jilai 2 2

falculty of Material and Chemical Engineering, Jiangxi University of Science and Technology, Ganzhou 341000 , CHINA; School of Metallurgical and Ecological Engineering, University of Science and Technology Beijing, Beijing 100083, CHINA Keywords: aluminum dross, alumina, extraction, calcining-leaching process order to decompose the mullite phase present in the Al dross. Due to the low cost and stability of sulfuric acid, this acid is used as a lixiviant during the recovery of A1 2 0 3 from sintered mixed dross. The impurity ions in the leaching liquor were removed by means of EDTA complexation, and washing with distilled water. This is a low temperature process that does not consume too much energy, and high purity A1 2 0 3 is obtained.

Abstract A processing technology was developed for alumina extraction from Al dross with the calcining-leaching process. The aluminum dross was mixed with soda and sintered at 1173K to yield soluble aluminates. Subsequently the sintered dross was leached with sulfuric acid to produce a solution containing aluminum . The unwanted metal ions, including Fe 3+ and Na+, were removed by ethylene diamine tetraacetic acid (EDTA) and water washing. Then NH4HCO3, was added controlling the crystallization of NH4A10(OH)HC0 3 , and the drying and calcining process was carried out, resulting in ultra fine A1 2 0 3 powders with high purity. The characteristics of the A1 2 0 3 powders were examined by means of XRD and SEM. The extraction efficiency of A1 2 0 3 can

Experimental The primary raw materials of the experiment were Al dross, soda and H 2 S0 4 . The compositions of the Al dross are listed in table 1. Al 39.77

surpass 98 % by optimization of the calcination and lixiviation processes. Well-dispersed fibriform A1 2 0 3 powders were obtained by calcining at 1000°C and the purity of the ultra fine A1 2 0 3

Table 1 Chemical Composition of the Al dross(wt%' M K Ca Fe Ti Si O 8 Na 1.33 0.56 1.59 4.98 4.57 0.55 0.40 41.59

The recovery of A1 2 0 3 from Al dross is based on application of hydrometallurgical processes such as acid or base leaching, purification, precipitation and calcination. The procedures of the experiment are illustrated in Fig.l. The crystalline phase and microstructure were characterized by X-ray diffraction (XRD) and scanning electron microscope (SEM).

powders was more than 99.36% . Introduction Billions tonnes of Al dross produced during the casting or remelting of aluminum have accumulated over the years. The majority of this dross is disposed off in landfill sites, causing serious pollution of the environment!!]. Finding methods for producing useful materials from the dross is a very important task for society.

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Aluminium dross contains mostly aluminium, silicon and other metal oxides and salts such as KC1 and NaCl. Thus it has the potential to be a rich alumina source, and is a good raw material suitable for production of aluminum and its chemicals. Different types of dross have been employed as raw materials by several authors. Dross includes considerable amounts of aluminum alloys, and it has been utilized as a deoxidsing agent in steel making, as a refractory material, and as a cement material [2-5]. Attempts have been made to extract alumina from the alloy dross by adopting either pyro- or hydrometallurgical methods [6,7]. Amer carried out the dissolution of waste dross in two steps, in order to produce a highly pure aluminum sulfate[8]. However, these processes have found little practical application because of the highly corrosive nature of the concentrated acid and alkali involved, and because these processes also constitute an environmental hazard.

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The objective of this work is to facilitate the recycling of Al dross and produce high purity A1 2 0 3 . In this paper, the calciningleaching method was used to recover A1 2 0 3 from Al dross. In the proposed process, Al dross is mixed with soda prior to sintering in

Α12Ω2. Fig.l Flow chart of extraction of alumina from Al dross

197

b. Sintered mix powders Fig.2 XRD patterns of raw fly ash and calcined ash

Results and discussion

SEM analysis of the sintered raw Al dross and sintered dross also gave similar results, as shown in Fig.3. SEM image of sintered dross exhibited mainly as clumps. They were formed by reactions of sodium hydroxide, aluminum, and silicon oxide in the sintering process. These clumps were irregular in shape. However, Glassy material was found in the SEM image of raw Al dross, as shown in Fig.3(a). This was the evidence of existence of a mullite phase in the raw material.

Sintering the mixture of Al dross and soda The main objective of sintering the mixture of Aluminum dross and soda was to break the crystalline mullite phase (3Al203-2Si02) and the mesh framework of the glass phase, rendering free aluminate for leaching [9]. The following reactions occur during the calcining of fly ash and soda[10]: NqCQ=NqO+Cqî

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Figure 2 shows the XRD patterns of Al dross and the sintered mixture of Al dross and soda. From this figure we can see there are characteristic peaks of mullite in the raw material, while after sintering, these peaks disappeared and many new characteristic peaks were found. So the stable mullite crystalline phase was successfully broken-up by the sintering. NaA102 is the major component of the sintered mix. 2CaOSi02 and Na2OFe203 are the minor components. The soluble aluminum materials are dissolved. The process also produced insoluble residues called red mud.

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The sintering parameters such as sintering temperature and sintering time were considered to find the optimal activation efficiencies (rendering the maximum free aluminates). As shown in Fig.4, 1173K was the optimum temperature. At this temperature, the sintered material had the highest alumina extraction efficiency. Sintering below or above this temperature would produce abnormal sintered products, which led to lower alumina extraction efficiency.

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The purpose of the leaching process is to leach the soluble Al ions (in the form of sodium aluminates). The possible reactions that take place in the leaching of the raw Al dross are as follows :

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The leaching parameters such as the concentration of sulfuric acid, leaching time and leaching temperature were considered to find the optimal extraction efficiency of AI2O3.

80 °C the extraction efficiency is low. Increasing the leaching temperature can accelerate the progress of the leaching reactions and improve the aluminum extraction efficiency.

The results presented in Fig.5 and Fig.6, indicate that the efficiency of aluminum species extraction increases with an increase in both acid concentration and leaching time. Increasing the acid concentration can accelerate the leaching reaction and achieve high extraction efficiency. However, when acid concentration reached a certain value, the extraction efficiency may not improve. The optimum acid concentration was 0.9mol/l at which the alumina extraction efficiency reached the maximum. To some degree, prolonging the leaching time can increase the aluminum extraction efficiency,as prolonging the leaching time will allow the completion of the leaching reactions. Dissolution temperature/K" 100

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The leaching liquor obtained from the previous processes contained Al , Na+, Fe3+, Ca2+, Mg2+, and other impurities. In order to obtain high purity A1203, the impurity ions had to be removed. The selective precipitation of ions by pH value adjustment was found to be unsuccessful in separating Fe3+ and Al3+[11]. We chose ethylene diamine tetraacetic acid (EDTA), a complexing agent, to remove the Fe3+ and other impurity ions. Selectivity arises because the complexing ability of each metal in the leaching liquor with EDTA is different [12,13].

Φ

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H2S04 concentration/(/T7oi/m/)

Fig.5 Effect of H2SO4 concentration on alumina extraction efficiency The dissolution time was defined as the period that started at the time when the sintered mixture was introduced into the acid and ended when the liquid was separated from the red mud. This was the time period that the dissolution solution actually contacted the sintered mixture product. The longer the dissolution process is run, the greater the loss of the alumina extraction efficiency. This was due to the dissolution of Na2OFe203 as well as the side reactions. Since the particle size of the sintered mixture was very small, a short dissolution period actually gave the best result. The exact relationship between the dissolution time and aluminum extraction efficiency is shown in Fig.6.

The pH value of the leaching liquor was adjusted to 3.0 before the EDTA was added, and then the solution was continuously stirred for 30 minutes. The color of the solution turned from dull yellow to bright yellow. NH4HCO3 was subsequently added into the solution to precipitate the aluminum. At the same time the solution was quickly stirred. After filtration, white precipitates and yellow solution were obtained. These results show that the Fe3+ contained in the leaching liquor was effectively removed. The other soluble impurities absorbed in the precipitates could be removed by three cycles of washing. A white gelatinous precipitate was obtained and then dried in a microwave oven for 5 min and the powder calcined at 1000°C for 2 h. Active γ-Α12θ3 powder was obtained. The XRD pattern and SEM images of the A1203 product are shown in Figs.8 and 9. The purity of the A1203 product was as follows: A1203 99.36%, MgO 0.51%, CaO 0.06%, Fe203 0.04%, and Si0 2 0.03%. The result indicated that further investigation of strategies for Mg removal is necessary to achieve higher purity A1203.

Dissolution time//)

Fig.6 Effect of dissolution time on alumina extraction efficiency Temperature was the most important factor affecting alumina dissolution from the sintered mixture. From the results presented in Fig.7, it can be seen that when the temperature is lower than

199

3. H.N. Yoshimura and A.P. Abreu, "Evaluation of aluminum dross waste as raw material for refractories," Ceramics International, 34(3)(2008), 581-591. 100 [■

4. El-Katatny, E.A., Halany, and S.A., Mohamed, "Surface composition, charge and texture of active alumina powders recovered from aluminum dross tailings chemical waste," Powder Technol, 132, 2003 , 137-144. 5. LIU Qingsheng, "Preliminary research on preparation Al-Si-Ti alloy with aluminum dross as electrolysis materials". TMS,EPD congress, 2010, 639-644. 6. B. Dash and B.R. Das, "Acid dissolution of alumina from waste aluminium dross," Hydrometallurgy, 92(2008), 48-53.

10

20

30

40

50

60

70

80

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7 B. R. Das, et al., "Production of alumina from waste aluminium dross," Minerals Engineering, 20(2007), 252-258.

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8 Amer, A.M., 2002, "Extracting aluminum from dross tailings," /. Metals, 54(2002), 72-75.

Fig.8 XRD pattern of A1 2 0 3 product

9. MIKITO UEDA and SHIRO TSUKAMOTO, "Recovery of aluminum from oxide particles in aluminum dross using AlF-NaFBaCl2 molten salt," Journal of Applied Electrochemistry, 35(2005), 925-930. 10. XIAO Y. Y AN, "Chemical and electrochemical processing of aluminum dross using molten salts," METALLURGICAL AND MATERIALS TRANSACTIONS B, 348, 2008, 348-363. 11. MATJIE R H, et al., "Extraction of alumina from coal fly ash generated from a selected low rank bituminous South African coal," Min Eng, 18(2005), 299-310. 12. AKIRA O, et al., "Leaching of various metals from coal into aqueous solutions containing an acid or a chelating agent," Fuel Process Technol, 85(2004), 1089-1 102.

Fig.9 SEM images of α-Α12θ3 Conclusions

13. IYER R S and SCOTT J A, " Power station fly ash—a review of value-added utilization outside of the construction industry," Resour Conserv Recycl, 31(2001), 217-288.

(1) Al dross can be used can be used as a bauxite substitute in the soda sintering process for alumina production. Optimum conditions for a desilicated ash sintering were 1173K and 60 min. (2) Optimum conditions for sintered for Al dross dissolution were an H 2 S0 4 concentration of 0.9mol/l at 358Κ for 1.5h. (3) The AI2O3 purity was improved by means of EDTA complexation of other impurity metals and washing with distilled water. Finally, High purity (Al 2 0 3 >99%) active α-Α1203 was obtained. References 1. O. Manfredi and W. Wuth, "Characterizing the physical and chemical properties of aluminum dross," JOM, 49(1997), 48-51. 2. F.A. Lopez et al., "The recovery of alumina from salt slags in aluminium remelting," Canadian Metallurgical Quarterly, 33(1994), 29-33.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Effect of Calcium/Aluminium ratio on MgO containing calcium aluminate slags 1

Wang Bo1, Sun Hui-lan1, Guo Dong1, Bi Shi-wen2 Hebei University of Science and Technology; 70 Yuhua East Rd; Shijiazhuang, Hebei, 050018, China; 2 Northeastern University; No.l 1, Lane 3, Wenhua Road; Shenyang, Liaoning, 110004, China Keywords: calcium aluminate slag, CaO ratio, MgO, C20A43M3S3, alumina leaching 12CaO»7Al203 and to inhibit the formation of C2oA13M3S3 [1315].

Abstract MgO is the main impurity in calcium aluminate slags. The presence of MgO will change the nature of the alumina bearing phases and decrease the alumina leaching properties of the slag. In order to remove or reduce the negative effect of MgO, the method of changing the C/A (CaO/Al203, molar ratio, excluding CaO in 2CaOSi02) of the calcium aluminate slag was studied and the effect on leaching mechanism was also analyzed. The results showed that the formation of the quaternary compound 20CaO13Al2O3-3MgO-3SiO2 (C2oA13M3S3) is inhibited with increasing C/A in the calcium aluminate slag, and the MgO crystallized in the form of periclase under these conditions. There is an optimal C/A to improve the alumina leachability of the calcium aluminate slag. The increase of C/A could not remove the negative effect of MgO on leachability completely, and the optimal C/A of the slag increases with increasing MgO content.

Previous research demonstrates that the alteration of the CaO content could improve the alumina leaching properties of the slag. So the alteration of C/A (CaO/Al203, molar ratio, excluding the CaO of 2CaOSi02) of the calcium aluminate slag, to remove the negative effect of MgO, was studied and mechanism of this effect are discussed in this paper. Experimental 2.1 Materials CaC03, Na2C03, NaOH, Si0 2 and MgO used in the experimental studies were chemical pure reagents. Al(OH)3 used in the experimental was an industrial pure reagent. The A/S (Al203/Si02, mass ratio) of the calcium aluminate slag was 1.3.

Introduction

2.2 Smelting of calcium aluminate slag

Calcium aluminate slag is obtained from blast furnaces when smelting iron-bearing bauxite. The ideal components of calcium aluminate slags are 12CaO-7Al203 and y-2CaOSi02 [1, 2]. The slag can react with a sodium carbonate solution and its alumina leaching rate can reach 85%. This method realizes the comprehensive utilization of iron and alumina values in the ore [3-5].

Samples with different C/A were smelted in a MoSi2 resistance furnace, and the container was a graphite crucible. Another corundum crucible was used outside of the graphite crucible in order to prevent the break-out of the slag which could disintegrate and destroy the graphite crucible during cooling, thus contaminating the furnace. The smelting temperature was 1500°C and the holding time was one hour. The sample was taken out at 400 °C from the resistance furnace, and its cooling rate was 5°C/min.

The existence of impurity MgO which comes from iron-bearing bauxite and lime, will affect the alumina leaching rate of calcium aluminate slag during the industrial process [6-8].

2.3 Leaching of calcium aluminate slag

Eremin [9] studied the effect of MgO on calcium aluminate slag and pointed out that the compound 6CaO-4Al203MgOSi02 (CwAxMySz) would be formed in the slag when MgO was present. This compound does not react with the sodium carbonate solution. Wang Bo[10, 11] studied the mechanism of MgO effects on such slags systematically. The results showed that MgO will dissolve into the crystal lattice of 12CaO-7Al203 and forms limited C2oA13M3S3 when the MgO content is less than 1.0%. Above this level, a large amount of C2oA13M3S3 was formed which decreased the alumina leaching rate. The results also showed that this quaternary compound could react with sodium carbonate solution but its alumina leaching rate was very low [12].

The sodium aluminate solution obtained from leaching the slag was treating using the carbonization precipitation process, and the circulating mother liquid is then used to leach new calcium aluminate slag. The feasible conditions for alumina digestion are: leaching temperature 75 °C, leaching time 2h, L/S ratio 4.5, caustic alkali concentration 7g/L and sodium carbonate concentration 120 g/L. The leaching experiments are carried out in magnetically stirred constant temperature water bath. After leaching and dry filtration, the filtrate is used to analyze the composition of the solution, and the filter residue is washed and dried for analysis.

Therefore, it is clear MgO will affect the alumina leachability. The presence of MgO could not be avoided because it is the main impurity of the slag, so how to remove or reduce the negative effect of MgO became the focus of this research.

2.4 Analysis The contents of A1203 and Na20 in samples and filtrate were analyzed by a chemical method. Phase components of the calcium aluminate slag were identified by X-ray diffraction (PANalytical PW3040/60). A Malvern laser particle analyzer was used to analyze the particle size of the slag.

There are two methods to remove the negative effect of MgO from the phase components of the calcium aluminate slag. One is to make MgO form a more stable component which contained no, or little A1203. The other method is to stabilize the compound

201

Results and discussion 3»! Effect of C/A on self-disintegrating property of slag The self-disintegration of calcium aluminate slag can reduce the energy consumption during leaching and alumina production, and this is a very important characteristic of calcium aluminate slags. The percent content of granularity which is lower than 74um in samples is defined as the self-disintegrating rate. Slags with different MgO contents were cooled slowly and then well-mixed in order to analyze the granularity of the slag. The particle size results analyzed by Malvern 2000 are shown in Fig. 1. 100 w

98 S

96

-

^^^

■^^'* [■ f^-"-""

^„---'"^

ö0

%

92

a

^<^C—===·

^χ^

J^-"-""

-m- MgO 0% * Mg01% - A - MgO 3% ψ MgO 5%

\-

90 1

1.4

1

1.5

1

1.6

1

1.7

1

1

1

1.8

1.9

2.0

OA

Fig. 1 Granularity of slag with different C/A and MgO content The results of granularity showed that the self-disintegrating rate of calcium aluminate slag with different MgO content and C/A were good, and they were basically higher than 95%. The selfdisintegrating rate was lower than 95% only when C/A was very low. It can be seen that when the C/A of slag was fixed the selfdisintegrating rate decreased with the increase of MgO content, and also when the MgO content was fixed the self-disintegrating rate increased with the increase of C/A. The self-disintegration of slag was caused by the crystal transformation from ß-2CaOSi02 to y-2CaOSi02. With the increase of MgO content, self-disintegration of slag decreased because of the increase of C20A13M3S3 content. A certain amount of Si0 2 was consumed when the quaternary compound C20A13M3S3 was formed. Therefore, the content of 2CaOSi02 decreased in the slag and the self-disintegrating rate of the slag decreased. The content of C2oA13M3S3 decreased with the increase in C/A when the MgO content wasfixed,so the self-disintegrating properties were improved. But the content of Si0 2 in C2oA13M3S3 was low, so the decrease of 2CaOSi02 was also low and had little effect on self-disintegration rate of the slag. Therefore, the selfdisintegration rate of the slag was still sufficient to meet the demand of the alumina leaching process. 3.2 Effect of C/A on alumina leaching property of slag The alumina leaching experiment on the calcium aluminate slags was carried out in order to study the effect of C/A on the alumina leaching properties of MgO containing slags. The leaching conditions were as in section 2.3, and the results are shown in Fig.2.

Fig.2 Leaching rate of containing MgO slag with different C/A Fig.l shows that the alumina leaching rate of slag increased at first and then decreased with increase of C/A. The highest alumina leaching rate of a slag which did not contain MgO appeared when C/A was 1.7. The highest alumina leaching rate of slag whose MgO content was 1% appeared when C/A was 1.8. The higher alumina leaching rate of slag whose MgO content was 3% appeared when C/A was between 1.8 and 2.0, and then the leaching rate decreased smoothly. The highest alumina leaching rate of slag with MgO content of 5% appeared when C/A was 2.0, and a maximum rate did not appear under these conditions. Therefore, the alumina leaching rate of calcium aluminate slag decreased with increasing MgO content when C/A was 1.7. The increase of C/A could improve the leaching property of slag to some extent, and the optimal C/A of the slag also increased with increasing MgO content. In addition, the variation of C/A could not remove the negative effect completely, because the optimal alumina leaching rate decreased with the increase in MgO content. 3.3 Effect mechanism of C/A on MgO containing calcium aluminate slag The increase of CaO can reduce the negative effect of MgO on leachability of calcium aluminate slags, so the mechanism of effect of C/A on MgO containing slags was analyzed by XRD. The existence of the quaternary compound C20A13M3S3 was also investigated. The XRD spectra are shown in Fig.3~Fig.5. Fig.6 and Fig.7 showed the relative intensity of independent characteristic peaks of 12CaO«7Al203 and C2oA13M3S3 as a function of MgO content and C/A.

was 1.8. Crystalline MgO or its mineral forms, was not found in the slag under these conditions. - - 20CaQ13AI 203-3MgO3Si 02 90

C/A=1.8

-

»

·

■-

80 70

^

60

c "03

C/A=2.0

LOLAJI 10

20

30

50

- ■ - C/A=1.8 - · - C/A=2.0

40

40

50 60 70 80 2 Thet a Fig.3 XRD spectra of slag with MgO=l%

30

90

■

20 L_

,

1

1

_l

MgO content /%

Fig.6 Relative intensity of characteristic peaks of 12Ca07Al 2 0 3

2 CaO· Si 0 2 12CaO-7AI 2 0 3 20CaO13AI 203-3MgO-3Si 0 2

14

MgO

C/A=1.8

-·

12 A 10 •>9

1

8

d>

6 4

Fig.4 XRD spectra of slag with MgO=3%

0 - 2CaOSi 0 2

C/A=1.8

30

40

50

60

70

1

..1

_J

Fig.7 Relative intensity of characteristic peaks of Q0A13M3S3 withC/A=1.8

- 20CaO13AI 2 0 3 -3^ÖO3Si 0 2

20

1

MgO content/%

- 12CaO7AI 2 0 3

10

•

80

Fig.6 and Fig.7 show the relative intensity of characteristic peaks of 12Ca07Al 2 0 3 and C2oA13M3S3. Relative intensity is defined as a peak intensity ratio between the strongest independent characteristic peaks of 12Ca07Al 2 0 3 or C20A13M3S3 of slag and the strongest peak of slag. The main phase in the calcium aluminate slag whose MgO content was 3% and C/A was 1.8, were y-2CaO«Si02,12CaO»7Al203, and C2oA13M3S3. The intensity of the characteristic peaks of Q0A13M3S3 is strong showing that a large amount of C2oA13M3S3 was generated. A great deal of 12CaO»7Al203 transformed into C2oA13M3S3 under these conditions. This is why alumina leaching rate decreased. When the C/A of slag increased to 2.0, compound C2oA13M3S3 disappeared and 3CaO»Al203 was not found either. Meanwhile periclase was found in the slag. So the extra-addition of CaO decomposed the quaternary compound C2oA13M3S3 and that was the reason why the alumina leaching rate increased. The reaction between C20A13M3S3 and CaO is shown below:

90

2 Thet a

Fig.5 XRD spectra of slag with MgO=5% The compound C2oA13M3S3 which is difficult to leach was found in the slag whose C/A was 1.8, but it could not be found in the slag whose C/A was 2.0. The intensity of characteristic peaks of 12CaO7Al203 were clearly enhanced when the C/A of the slag increased. Thus some of the C2oA13M3S3 transformed into 12Ca07Al 203 which is easy to leach, with the increase of C/A of the slag. Some 3CaOAl 2 0 3 was also found in slag when the C/A was 2.0, and that was the reason why the alumina leaching rate of slag whose C/A was 2.0 was lower than that of a slag whose C/A

7(20CaO13Al2O3-3MgO-3SiO2)+58CaO-+ 13(12CaO»7Al203)+21(2CaOSi02)+21MgO

203

The analysis of a slag whose C/A MgO content was 5% was similar to that of a slag whose MgO content was 3%. The only difference between them was the intensity of the characteristic peaks of periclase (MgO).

Northeastern University: Natural Science, 29(11)(2008), 1593-1596. [7]

Therefore, the increase of C/A of calcium aluminate slag improved alumina leaching properties and decreased the negative effect of MgO on teachability of the slag, but it could not remove the negative effect of MgO completely. Thus the mechanism of this effect needs to be studied further.

TONG Zhi-fang, BI Shi-wen, YU Hai-yan and WU Yusheng, "Leaching kinetics of non-constant temperature process of calcium aluminate slag under microwave radiation," The Chinese Journal of Nonferrous Metals, 16(2)(2006), 357-362.

[8]

D.J.Connor, Aluminium extraction from non bauxitic materials, (Sydney: Aluminium-Verlag Gmbh, 1988), 230250.

[9]

N.I.Eremin, "Investgations on the complex processing of bauxites," (Bauxite-Alumina-Aluminum, Symposium of ICSOBA . Budapest: Research Institute for Non-Ferrous Metals, 1971), 329-335.

[10]

WANG Bo, YU Hai-yan, MIAO Yu, SUN Hui-lan, BI Shiwen and TU Gan-feng, "Effect of MgO on leaching and self-disintegrating property of calcium aluminate slag," Light Metals, (4)(2008), 11-13.

[11]

WANG Bo, YU Hai-yan, SUN Hui-lan, BI Shi-wen, TU Gan-feng and GENG Hong-juan. "Effect of MgO on crystal and leaching property of 12CaO-7Al203," Minning and Metallurgical Engineering, 28(05)(2008), 68-71.

[12]

WANG Bo, YU Hai-yan, SUN Hui-lan, BI Shi-wen and TU Gan-feng, "Synthesis and AI2O3 leaching property of 20CaO»13Al2O3eMgO«3SiO2," The Chinese Journal of Nonferrous Metals, 19(2)(2009), 378-382.

[ 13]

I.Kapralik and F.Hanic, "Studies of the system CaO-A1203MgO-Si02 in relation to the Quaternary phase Q," Transation of British ceram society, 1980, 128-133.

[14]

JIANG Feng-hua and XU De-long, "Influence of trace compositions on the formation of Q phase in high-alumina cement system," Journal of the Chinese Ceramic Society, 33(10)(2005), 1276-1279.

[15]

MENG Tao, YANG Li-qun and XU Xian-yu, "Studies of CaO-Al203-MgO-Si02 system in relation to the formation and hydration of phase Q," Bulletin of the Chinese Ceramic Society, (3)(1998), 21-34.

Conclusions 1) MgO content and C/A of calcium aluminate slags had little effect on self-disintegration rate of the slag. The selfdisintegrating rate of most samples was above 95%. 2) The increase of C/A could decrease the negative effect of MgO on leachability of the slag, but it could not remove the effect completely when the MgO content was less than 5%. The optimum C/A of the slag increased with the increase of MgO content. 3) The addition of CaO could decompose the quaternary compound C2oA13M3S3, and MgO crystallized in the form of periclase. Acknowledgements The authors greatly acknowledge the financial support of the National Nature Science Foundation of China (Project No: 50674028), and the foundation of Hebei University of Science and Technology (Project No: XL200921). The authors express their profound gratitude to the editors and reviewers of TMS. References [1]

L.K.Barr, "Alumina Production from Andalusite by the Pedersen Process," (Stockholm: Almqvist & wiksell international, 1977), 64-70.

[2]

J.Grzymek, A.Derdacka and Z.Konik, "Method for obtaining aluminum oxide," U.S Patent : 4149898 , 19782-21.

[3]

ZHANG Jing-dong, LI Yin-tai, BI Shi-wen and YANG Yihong, "Research on integrated utilization of high-ferrum bauxite in Guigang Guangxi," Light Metals, (8)(1992), 1618.

[4]

BI Shi-wen, YANG Yi-hong, LI Yin-tai, ZHANG Jingdong and DUAN Zhen-ying, "Study of alumina leaching from calcium aluminate slag," Light Metals, (6)(1992), 1015.

[5]

J.Grzymek, "Complex Production of aluminium oxide and Iron from laterite raw materials applying the calcium aluminates polymorphism," TMS Annual Meeting and Exhibition, Light Metals, 1985, 87-99.

[6]

WANG Bo, YU Hai-yan, SUN Hui-lan and BI Shi-wen, "Effect of material ratio on leaching and self-disintegrating property of calcium aluminate slag," Journal of

204

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Study on extracting aluminum hydroxide from reduction slag of magnesium smelting by vacuum aluminothermic reduction Wang Yaowu, Feng Naixiang, You Jing, Hu Wenxin, Peng Jianping, Di Yuezhong, Wang Zhihui School of Metallurgy, Northeastern University, Shenyang, Liaoning 110004, P.R.China Key words: magnesium smelting; aluminum hydroxide; reduction slag; Ca02Al203; Abstract Magnesite + dolomite Aluminum powder

The reduction slag from magnesium smelting by vacuum aluminothermic reduction using dolomite and magnesite as raw materials consists mainly of A1203 and CaO. The alumina is more than 67% of the slag and is mainly in the form of CaO · 2A1203. The process of producing aluminum hydroxide from the slag by leaching with a mixture of sodium hydroxide and sodium carbonate was studied. This showed that the leaching rate of alumina is more than 86% when leaching temperature is 95°C, and L/S is 5, and leaching time is 2h. The magnesia-alumina spinel which generated in the reduction process and 3CaOAl203#6H20 which is generated in the leaching process, are the two main causes of loss of alumina. The chemical composition of aluminum hydroxide products obtained after carbonation precipitation can meet the quality standard of the alumina plant.

C02^

] Calcination] I Mixture \* |Briquetting| i A , I Vacuum reduction \—*> Magnesium Reduction slag_ Na2C03 NaOH

T

I Leaching I -

7

Decomposing Al(OH)3 Figure 1 The flowchart of the new magnesium production method

Introduntion A new magnesium production method has recently been introduced, which is based on two steps, the first of which is producing magnesium with vacuum aluminothermic reduction, and the second step is the extraction of aluminum hydroxide from the reduction slag. Figure 1 gives the flowchart of the new magnesium production method. The process of producing magnesium uses dolomite and magnesite as materials and aluminum powder as the reductant. After the reduction process, magnesium and a reduction slag can be obtained. The main phase of the reduction slag is Ca02Al 2 0 3 . The content of A1203 is more than 67% and the content of Si0 2 is below 1%. For a mass ratio of A1203 to Si0 2 is more than 60, the slag is suitable for leaching to produce aluminium hydroxide. This paper describes the study of extracting aluminum hydroxide from the reduction slag of this magnesium smelting process.

Experimental Materials The experimental materials is a reduction slag of magnesium smelting using dolomite and magnesite as materials and aluminum powder as reductant, the addition of materials and reductant follows the reaction equation (1). The magnesium smelting experiment was carried out at 1140°C, and the aluminum power excess is 5%. The chemical composition of the reduction slag is listed in Table 1 and Figure 2 shows the XRD pattern of the reduction slag. CaO+6MgO+4 Al = 6Mg + CaO · 2Al203

Table 1. The chemical composition of reduction slag (%) MgO

CaO

Si0 2

A1203

Al

Fe 2 0 3

6.13

20.86

0.89

67.54

3.50

1.08

205

(1)

£

4000 l··

40

2θ/(°)

Figure 2 The XRD spectra of reduction slag As shown in Table 1 and Figure 2, the main phase of the reduction slag is Ca02Al 2 0 3 , and there are also small amounts of CaOAl203, MgOAl203, MgO and Aluminum powder. The main chemical composition is A1203, CaO, unreacted MgO and Al.

The leaching experiments were carried out in a constant temperature water bath, and used an Erlenmeyer flask which is sealed with a condenser and stirred with a magnetic rotor. Before the experiments began, the leaching solution of mixed NaOH and Na2C03 was heated to a given temperature, then added into the reduction slag. After leaching for a given time, the leaching solution was filtered rapidly and washed three times by hot deionized water, and the alumina concentration of the filtrate was measured by chemical titration, and the alumina leaching rate of reduction slag was calculated.

Experimental principle Alumina in the reduction slag is mainly in the form of Ca02Al 2 0 3 , and it cannot be leached by a sodium hydroxide solution because in the leaching process, the CaO of the reduction slag would react with sodium hydroxide to form Ca(OH)2 which would then react with NaAl(OH)4 to form tricalcium hydroaluminate (3CaOAl203-6H20)[l,2] (equation 2). Because tricalcium hydroaluminate is stable in sodium hydroxide solution[3], a lot of A1203 entering into the leaching process is lost. The CaO of the reduction slag would react with sodium carbonate prior to sodium hydroxide to form CaC03 which has lower solubility and is more stable than Ca(OH)2 . The formation of CaC03 inhibits the formation of tricalcium hydroaluminate and reduces the loss of Al203[4], The leaching principle of CaO· 2A1203 and CaO A1203 are listed as follows[5, 6]:

The filtrate without the wash liquid was a sodium aluminate solution, and it was decomposed by carbonation in the water bath at 80 °C. The finish of decomposition was determined by the decomposed rate. When the decomposion reached 90%, it was stopped, filtered and aluminum hydroxide could be obtained. Results and discussions

In the reduction slag, the aluminium is present in two main species, one is the alumina which exists in CaO-2Al203 and CaOAl203, the other is elemental Al (unreacted aluminum 3Ca(OH)2 + 2NaAl(OH)4 = 3CaO · Al203 · 6H20 + 2NaOH (2powder). ) The elemental Al can enter into the leaching solution CaO · Al203 + Na2C03 + 2H20 = 2NaAl(OH)4 + CaC03 O) completely and alumina can not. To distinguish the two parts, the leaching rate was described as the leaching rate of CaO · 2Al203 + Na2C03 + 2NaOH + H20 = NaAl(OH)4 + CaC03alumina (4) alumina and the aluminum leaching rate was described the leaching rate of element aluminum. The conditions effecting the In the leaching process, elemental Al ( unreacted aluminum alumina leaching rate are mainly leaching temperature, leaching powder ) in the reduction slag can react with sodium hydroxide time, the content of sodium hydroxide and sodium carbonate. and also with sodium carbonate, but the reaction rate of elemental Al with sodium hydroxide is much faster than with sodium carbonate, so most of the elemental Al react with sodium hydroxide. Elemental Al is fully solution in the leaching sloution. 2AI + 2NaOH +6H20= 2NaAl(OH)4 + 3H2

Effect of leaching temperature on alumina leaching rate The results of effect of leaching temperature on alumina leaching rate are shown in Figure 3, and the experiments were carried out under the conditions of leaching L/S 5, leaching time

(5)

Experimental Method

206

2h, content of sodium hydroxide 100g/L, and content of sodium carbonate 120g/L.

temperature 95°C, leaching time 2h, content of sodium hydroxide 100g/L

-

90

|

80

•

■—'"'

M

2

^ ^

i_

·

·

*

I The alumina leaching rate I | — · — The aluminum leaching rate |

h

1 —"—The alumina leaching rate 1 1 — · — The aluminum leaching rate 1 60

65

70

75

80

85

90

100

95

The content of sodium carbonate/g/L

Temperature/°C

Figure 5 Effect of content of sodium carbonate on the alumina leaching rate

Figure 3 Effect of leaching temperature on the alumina leaching rate

It can be see from Figure 5 that the content of sodium carbonate have little effect on the alumina leaching rate when the content of sodium carbonate is between 80 and 130 g/L especially over 100 g/L. In the leaching process of the reduction slag, there are at least five reactions in the leaching solution, they are the formation of sodium aluminate (equations 3, 4 and 5) , the formation of tricalcium hydroaluminate (equation 2) and the decomposition of tricalcium hydroaluminate (equation 6). There are two main role of sodium carbonate, one is reaction with CaO to form CaC03, the other is decomposing tricalcium hydroaluminate| 7,8 j.

The effect of leaching temperature on the alumina leaching rate is very notable, the leaching rate of alumina in the reduction slag increases with increasing leaching temperature and levels off when the leaching rate reaches 90°C. Effect of leaching time on alumina leaching rate The results of effect of leaching time on alumina leaching rate are shown in Figure 4, and the experiments were carried out under the conditions of leaching L/S 5, leaching temperature 95°C, the content of sodium hydroxide 100g/L, and content of sodium carbonate 120g/L.

The content of sodium carbonate depends on two factors, one is that the content of sodium carbonate must exceed the theoretical amount to avoiding the forming of tricalcium hydroaluminate and the theoretical amount is about 80g/l when the L/S is 5, the other aspect is that since the decomposition of tricalcium hydroaluminate is a reversible reaction, thus the higher the content of sodium the better in order to decompose tricalcium hydroaluminate. According to the results shown in Fig.4, the optimal content is 100 g/L.

'■ •

\s 20

40

60

80

The alumina leaching rate The aluminum leaching rate

100

120

140

160

180

3CaO · Al203 · 6H20 + 3Na2C03 :

200

ì 2NaAl(OH)4 + ANaOH + 3CaC03

(6)

Leaching time/min

Effect of the content of sodium hydroxide on alumina leaching rate

Figure 4 Effect of leaching time on the alumina leaching rate As shown in Fig.4, the leaching rate of alumina in reduction slag increases with the extension of leaching time and levels off when the leaching time extends 2h.

The effect of the content of sodium hydroxide on alumina leaching rate is showed in Figure 6. The experiments were carried out under the conditions of leaching L/S 5, leaching temperature 95°C, leaching time 2h, content of sodium carbonate 100g/L.

Effect of the content of sodium carbonate on alumina leaching rate The results of effect of the content of sodium carbonate on alumina leaching rate are shown in Figure 5, and the experiments were carried out under the conditions of leaching L/S 5, leaching

207

As shown in Figure 6, the content of sodium hydroxide has great effect on the alumina leaching rate. The alumina leaching rate increases at first, and then decreases with the increase of sodium hydroxide concentration, and there is a maximum at a content of 75g/L. Phase analysis of Leaching slag

- The alumina leaching rate - The aluminum leaching rate [

60

65

70

75

80

85

90

95

100

The main chemical composition of the leaching slag is listed in table 2, and the XRD pattern of the leaching slag was analysed in Figure 7.

105

The content of sodium hydroxide/g/L

Figure 6 Effect of content of sodium hydroxide on the alumina leaching rate Table 2 The main chemical composition of the leaching slag A1203

MgO

CaC0 3

Si0 2

Fe 2 0 3

others

17.88

6.38

70.53

2.81

0.79

1.61

24000 r

Li

22000 [

A-CaC03 o - MgOAl203

20000 [ 18000 [

V-MgO 0 - 3CaOAl203.6H20 - 2CaOSi02

;

16000 [ 14000 [ 12000 1 [ 10000 [

0

8000 [

2000 [

ΔΔ

V

6000 [ 4000 [

Δ

CO

Δ

°0

of 1

-2000

i

1

2θ/(°)

Figure 7 The XRD of the leaching slag As shown in Figure 7 and Table 2, the main phase of the leaching slag is calcium carbonate, and there are some other phases such as magnesia-alumina spinel (MgOAl203), tricalcium hydroaluminate (3CaOAl203-6H20), and calcium silicate (2CaOSi02). It is also found that the Ca02Al 2 0 3 and CaOAl203 are decomposed completely.

In the sodium aluminate solution which result from leaching of the reduction slag, the alumina concentration is 120-150g/L, and Si0 2 concentration is 0.2-0.3 g/L, and the ratio of A1203 and Si0 2 is more than 500. Since the A/S is high, the sodium aluminate solution can be decomposed by carbonation without desilication process. The composition of the aluminum hydroxide obtained is shown in Table 3.

Table 3 The composition of aluminum hydroxide[9] A1 2 0 3 Fe2Os Si0 2 0.02 AH-1 64.5 0.02 the national AH-2 0.04 0.04 standards of 64.0 china 0.06 AH-3 63.5 0.06 Al(OH)3 of this project 64.3 0.03 0.01

The chemical composition of the aluminum hydroxide products used reduction slag of magnesium smelting as materials can meet the quality standard of the alumina plant. In addition, the

Na 2 0 0.4 0.5 0.6 0.43

sodium aluminate solution is suitable to producing high added value non-metallurgical alumina.

208

Conclusions (1) The main phase of the reduction slag from magnesium smelting by vacuum aluminothermic reduction is Ca02Al 2 0 3 , and the Ca02Al 2 0 3 can be decomposed completely by leaching with alkali solution to form CaC03 and sodium aluminate solution. (2) The leaching rate of alumina from the slag is more than 86% and the leaching rate of element aluminum is more than 90% when the reduction slag leached at 95 °C for 2h. The magnesia-alumina spinel which is generated in the reduction process and 3CaOAl203»6H20 which is generated in the leaching process were the two main pathways which caused loss of alumina. (3) Aluminum hydroxide conforming to the national standards of China could be produced after carbonation precipitation. References 1. M. Gawlicki, W. Nocu-Wczelik, M. Pyzalski. Studies on the hvdration of calcium aluminates, Journal of Thermal Analysis. 1984, 29 :1005-1008. 2. Xiao Wei , Liu Wei, Zhang Nianbing. Leaching kinetics of calcium aluminate sinter. Light Metals.2009, (8):23-26. 3. Yuan Jiongliang, Zhang Yi. Study on desilication process of sodium aluminate solution by adding tricalcium hvdroaluminate. Nonferrous metals. 2003, 55 (l):60-65. 4. C. Ostrowski, J. Zelazny. Solid solutions of calcium aluminates C^A, C12A7 and CA with sodium oxide. Journal of thermal analysis and calorimetry. 2004, 75: 867-885. 5. He Runde. The extracting of alumina from powdered calcium aluminate. Light metals. 1997, (6):20-25 6 Fan Juan , Ruan Fuchang. Review on the extracting methods of aluminum from powdered calcium aluminate. Applied chemical industry. 2002, 31(1): 1-5. 7. Liu Guihua, Li Xiaobin , Peng Zhihong. Reaction behavior between calcium oxide or calcium hydroxide and aluminate solution with heavy caustic soda. The Chinese journal of nonferrous metals. 2000, 10(2):266-269. 8. Qiu Zhenzhuo. Extracting calcium aluminate by sodium hydroxide. Light Metals.1999, (4):18-22 9. The national standards of china. Aluminum hydroxide. GB/T 4294-1997.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

APPLICATION OF THERMOGRAVIMETRIC ANALYSIS FOR ESTIMATION OF TRI-HYDRATE ALUMINA IN CENTRAL INDIAN BAUXITES - AN ALTERNATIVE FOR CLASSICAL TECHNIQUES Y. V. Ramana and Rajesh C. Patnaik Vedanta Aluminium Limited, Lanjigarh, Kalahandi, Orissa-766 027, India Keywords: Bauxite, tri-hydrate alumina, gibbsite, thermogravimetry. Trihydrate alumina content in bauxite is generally estimated by a bomb digestion system in the plants of the alumina industry using the Bayer process. Though this analytical method provides the required levels of accuracy, the large number of steps involved in the process, starting from digestion, filtration and titration to complete one analysis, becomes tedious and time consuming. If there are a large number of samples to be processed within a stipulated time to meet the process requirement, completion of the targeted number of samples in time is a challenge. This necessitates looking into quick alternative methods with equal levels of accuracy and dependability.

Abstract Tri-hydrate alumina or gibbsite (A1203.3H20) content in bauxite is a fundamental quality parameter in the Bayer alumina process using low temperature digestion. Classical techniques available for its estimation are time consuming and prone to standard and non-standard sources of error. Extensive studies on analysis of samples from Central Indian bauxite sources using thermogravimetry over varied temperature ranges and comparison with that of the data obtained from classical techniques, have revealed that loss of molecular water at different temperatures in thermo-gravimetry provides a meaningful tool to correlate with tri-hydrate alumina content by applying relevant correction factors derived from the experimental data at different concentrations of tri-hydrate alumina. The studies also have found that the thermogravimetric analysis can be used as a very fast and dependable technique with higher levels of accuracy over classical methods and is free from other interferences. The accuracy levels of the method developed were checked using reference standards.

Since the above conditions are always prevalent at Vedanta, alternative instrumental techniques which are very fast, with competitive levels of accuracy to wet chemical analysis, particularly thermal decomposition techniques, have been tried. This is a reasonable approach as the thermal techniques to study the dehydration and dehydroxylation behaviours of bauxites have been widely documented [1]. It is well known that thermal decomposition of aluminium minerals in bauxite is a factor of time and temperature. Dehydroxylation of gibbsite occurs in the temperature range of 260-380 °C with peak losses at 320-340 °C.

Introduction Central Indian bauxite sources contain three principal hydrates of aluminium, i.e., gibbsite or trihydrate alumina (A1203.3H20 or Al(OH)3)), boehmite and diaspore or monohydrate alumina (A1203.H20 or AlOOOH) with gibbsite often being the predominant mineral. The other major minerals are aluminogoethite, hematite, kaolinite, anatase, rutile and quartz. Alumina extraction using low temperature digestion in the Bayer process results in dissolution of the gibbsitic form of aluminium hydroxide where other iron and titania minerals remain undigested. Kaolinite (Al4[Si4Oi0](OH)8) shows significant dissolution in sodium aluminate liquors, subsequently forming a desilication product, whereas the effect on quartz is less significant. All the minerals present in the bauxite, and their chemical-mineralogical characteristics, will have some influence either on process control, or on the methods of estimation of primary pararmeters.

The established general decomposition pattern of gibbsite, A1203.3H20 is as below: A1203. 3H20—>A1203· x H 2 0 (as residual hydroxide) + y H2Oî Using thermogravimetry, this property of losing hydroxyls as water molecules at elevated temperatures from the gibbsite mineral present in the bauxite, and converting the weight loss into an equivalent alumina content, is accomplished. The values obtained are then compared and equated with the data previously generated using classical techniques of analysis for the same samples. From the proportions of released and residual hydroxides lost as water molecules, a relationship is established at a particular temperature by comparing the thermal decomposition data with wet chemical data and by applying a correction factor if required.

Determination of the trihydrate alumina content in feed bauxite is of paramount importance, as the total process control is directly or indirectly affected by its proportion in the bauxite and also the accuracy levels in its estimation. Vedanta Aluminium Ltd., at Lanjigarh is presently using varied sources of bauxites for production of alumina and a Central Indian source is one of them. Determination of the chemical and mineralogical composition of these sources with acceptable limits of accuracy, and with the necessary speed, is of pivotal importance to facilitate required blending and homogenising so as to feed a uniform quality of bauxite to the plant.

Materials and Methods 1. Bauxite sample collection and preparation Bauxite samples were collected from the mines of M/s. CR Mittal from Central Indian region. A total of 29 samples have been collected and prepared using Indian Standard [2]. Apart from the mine samples, a total of 10 international reference standards of bauxite have been used for comparison.

211

a combination of released and unreleased (bound hydroxyls) water molecules from the bauxite. The instrumental settings followed for the final analysis of bauxite samples are detailed in Table 1.

2. Chemical analysis of bauxite LOI of the sample is determined by subjecting 1 g of sample to 1000 °C for one hour and from the weight difference between before and after heating [3]. The major elements; alumina, silica, ferric oxide and titania in bauxite samples are also determined using the respective Indian standard test methods of [4-7]. Reactive silica in bauxite is obtained by deducting the percentage of nonreactive silica from the percentage of total silica [8].

Since the percent loss of mass corresponds to the water molecules associated with trihydrate alumina in bauxite, the weight of trihydrate as alumina is calculated as below: ,, .

3. Trihydrate alumina determination using wet chemical techniques Using an M/s. Intronics bomb digester, a known quantity (1.3 + 0.0001g and 0.65 + 0.0001g) of bauxite are heated under pressure, in a solution of sodium hydroxide (165g/l as Na2C03) respectively at 145 °C and 240 °C for gibbsite and boehmite; the insoluble matter is separated by filtration. Alumina present in the filtrate is given by titrating the equivalent of the hydroxide ions (OH) released by fluorine in a chelating solution containing sodium gluconate [9]. Mass percentage of A1203 is calculated as gibbsite for the solution at 145 °C and as gibbsite and boehmite combined at 240 °C. The difference in the values between the two determinations is expressed as boehmite.

Covers

Unit

Moisture

355 °C LOI

-

No

No

Start Temp

°C

25

107

End Temp

°C

107

355

Ramp Rate

°C/Min.

10

20

Ramp Time

Min.

00:08

00:12

Hold Time

Min.

00:20

00:20

Total Time

Min.

00:28

00:32

Max Time

Min.

00:00

00:00

-

None

None

Flow Rate

ml/Min.

Off

Off

Comparator

%

0.1

0.1

Final weight

-

at constancy

at constancy

Atmosphere

(kmàMass-MoisteeMassi* 100 initial Mass

Gioiste Mass-355 °C I d Mass)»100 — Moisture Mass

mxuÂ%*

ö

MoistufeMas£-J55^LOIMassFl^i..^* litt Moisture Mass

Molecular ratio of A1203 : H 2 0 Factor for residual hydroxidefractionof 3H20

4. Trihydrate alumina determination using thermogravimetry A thermogravimetric analyser, model TGA-701 of M/s. Leco Corporation is used for the determination of loss of mass in bauxite. Step Name

„

Released fraction of 3H 2 0 Unreleased fraction of 3H20 %of unreleased 3H20 %of released 3 H20 Molecular mass of released 3 H20 Factor for unreleased | 3H 2 0

= 1.889 = 1.06

2.89

2.87

2.85

2.83

2.81

2.79

2.77

0.11

0.13

0.15

0.17

0.19

0.21

0.23

3.67

4.33

5.0

a

6.33

7

7.67

96.3

95.7

95

943

93.7

93

92.3

52.0

51.7

51.3

50.6

50.2

49.9

1.04

1.05

1.05

1.07

1.08

1.08

■ ■

Table 2. Basis of gibbsite thermal dehydroxylation The initial step in the thermal decomposition of gibbsite is the diffusion of protons and the reaction with hydroxyl ions to form water. From the comparative analysis of thermal decomposition data with wet chemical data of trihydrate, it is observed that there is a certain fraction of unreleased water molecules characteristic of either the source of bauxite or the concentration levels of trihydrate present in the bauxite. The factor for the unreleased fraction of water molecule is calculated with different combinations of released and unreleased proportions of three water molecules available with each alumina molecule for completion of the total quantity of water. Out of several combinations, as displayed in Table 2, a fraction of 0.17

Table 1. Test conditions for bauxite observed during thermogravimetric analysis A number of samples are analyzed using different instrumental conditions in the range from 100 to 420 °C to arrive at the most common and nearest temperature where the wet chemical values are comparable with the scanned data derived by applying the necessary correction factors. These are calculated on the basis of

212

method is free from such interferences. Other variables like heating rate, external pressure, water vapour pressure, sample particle size etc., do not have a significant influence on these determinations [10].

unreleased molecules and its corresponding factor (1.06), is considered to be most appropriate for the calculation of actual trihydrate content, when compared with the values of trihydrate determined by classical techniques, for the Central Indian bauxite sources.

is.*». U,fe

Results and discussion

!

The data obtained by conducting wet chemical analysis of the major elements and trihydrate alumina content in 29 samples collected from a bauxite source of Central India (M/s. CR Mittal) has been furnished in Table 4. Data on trihydrate alumina obtained from thermogravimetric analysis and boehmitic monohydrate data obtained from wet chemical techniques using a 240 °C bomb digestion of the same samples, is also furnished for comparison. The data on the other possible interfering parameters in wet chemical analysis, like reactive silica and boehmitic alumina, are also furnished in the table to study their effect Interpretation of the above data reveals several interesting factors related to the thermogravimetric analysis and its application to the determination of trihydrate content in bauxites. Data on the wet chemical analysis of trihydrate alumina content is furnished in column 10 (ranging from 24.10 to 48.27%) and the respective values obtained by using thermogravimetric analysis are furnished in column 12 (ranging from 24.18 to 47.22%) in Table-3. Out of the total 29 samples analysed, only four samples are highlighted in the Table- 3 as showing a difference of > 1% THA as under column 13, wherever the actual THA content is more than or equal to 45 % in the bauxite. It is interesting to note that except the above four samples, all other samples have a THA of < 40% and variation is < 1%. The data also indicates that as the THA content is increasing from 40% and above, the variation between the two methods is showing an increasing trend indicating that the factorial value needs to be corrected with increasing THA content. It is also clear from the table that the variation is very insignificant if THA value is < 35% as reflected in majority of the samples studied. These observations reveal that the thermogravimetric data obtained for bauxites having < 35% trihydrate alumina for the source studied can be a perfect match with the wet chemical data and it can be used as a direct alternative technique. Similarly, up to 45% THA content, the thermogravimetric method can be accepted with 1% variation. Above 45%, it is necessary to make corrections to the factor derived from the loss of mass.

Samp!« Legation

m

TiOì

SiOì (H)

L.OÎ

1

2

3

4

5

6

7

8

Î

61 SI

Pit-AS!

31.04

4SI

3.60

3S.65

2

6132

Pa-E/S2

9.S4

8.33

0.36

56.35

3

61 S3

Pfc-D'Sl

16.66

7.60

LÔ4

4

6184

Pit-OSS

16.76

5.48

2.90

R-SiOj VtttTBA

m

TGA-THÄ at3S5*C

»iff. Eeeiumt« (M-12)

m

u 1

9

1β

11

12

13

20*1

3.26

33.47

31.72

3363

0.16

1.04

23.83

0.7:9

35.10

32.38

34.35

0.25

19.46

48.04

25.79

0.93

45.07

41.S4

44.35

0.72

100

49.15

24.85

2.56

41.30

38.37

40.67

0.63

4.77

5

6185

Pft-OS2

6.13

6.42

0.Î1

57.72

2S.57

0.00

48.27

44.55

47.22

LOS

90S

6

6186

PÜ-A/S2

25.72

7.48

2.39

4110

22.50

2.10

37.90

35.63

37.74

0.16

0.00

7

6187

PÎI-C/S5

9 SS

6.37

1.15

5512

26.63

102

44.74

41.09

43.56

LIS

8.99

8

618S

PÎÏ-D/S3

22.16

5.3S

1.24

4537

24.73

1.12

4133

39.26

41.61

0.72

OSO

9

6189

Pit-C'Sl

24.82

5.54

0.3S

43 67

24.09

0.33

41.39

3S

Î>5

40.57

0.52

0.60

10

6190

P&-D/S2

13.1$

6.74

1.27

51.30

26.76

1.12

46.78

43.03

45.61

1.17

2.81

1!

6191

Pà-E-S2

14.97

6.Ö9

0.59

51.31

25.39

0.79

41.73

3S.63

40.95

0.78

8.55

12

6192

Pà-E-S3

19.84

5.S7

1.22

4S.23

24.08

1.09

39.93

37.4S

39.70

0.23

6.20

13

6193

Pìt-C:·^

19.04

6.60

no

49.39

23.25

0.91

37.58

35.34

37.46

0.12

9.94

14

6194

Pfc-IKÔÎ

9.5Î

6.42

063

60.53

21.34

0.56

25.73

24.07

25.51

0.22

33 52

15

6195

H t - H M Ö 20.32

5.56

4.18

49.S0

19.33

3.73

24.10

22.8Î

24.18

-0.ÖS

21.24

5.57

O.il

27.il

16

6196

»-1ΚΌ3

16.30

1?

6197

PÌHK/04

15.20

5.86

Î.93

56.33

18

6198

Pit-lK/OS

10.67

6.06

0.91

60,25

19

9134

Pit-04

40.34

3.98

5.57

31.70

20

9267

Pit-C

20.5?

6.95

1.20

21

9268

Pà-A/SS

14.92

7.84

22

9269

Pfi-E-Si

16.04

23

9270

Pìt-C'S2

24

9271

Pit-ES2

25

9272

Ptt-AS3

32.85

5.71

LSI

36.99

26

9273

PÌÌ-AS4

16.97

6.14

0.54

4S.32

27

9274

Prt-AS2

32.02

4.45

2.18

3S.50

23

9275

Pit-ASl

12.42

7.Î4

1.05

50.12

29

9276

Pìt-A'Sé

25.96

4.32

9.22

39.22

2.05

24.60

23.10

24.49

19.84

1.72

25 J 3

24.17

25.62

-0.39

2S.76

21.27

0.S2

27.10

25,73

27.27

-0.17

31.87

17.29

3.59

26.79

25.41

26.93

-0.14

0.00

4649

23.3S

1 07

38.79

36.78

3S.99

-0.20

5.10

1.12

49.13

25.53

1.00

42,44

40.69

43.13

-0.69

2.97

7.40

Î.SS

5145

21.95

1.42

32.71

31.05

32.94

-0.23

16.32

7.30

6.37

0.84

56.02

23.31

0.75

47.62

43.80

46.43

1.19

4.67

15.53

7.09

1.43

4S.95

22,77

1.2S

35.27

33.42

35.43

-0.16

10.54

20.68

1.3S

34.32

32.53

34.48

-0.16

O.OC

26.20

0.49

43.92

41.96

44.48

-0.56

0.00

21.22

1.96

34.92

33.19

35.18

-0.26

O.OO

27.54

0.93

46.69

44.60

47.28

-0.59

0.00

19.82

7.64

29.21

27.68

29.34

-0.13

» ■ * >

2.30

54.SÎ

20.09

Table 3. Data on chemical composition and trihydrate alumina determined from wet chemical and thermogravimetric techniques on Central Indian bauxite samples.

The other advantages of thermogravimetric analysis over the classical techniques, is that the former facilitates multiple sample analysis i.e., up to 20 samples at a time resulting in higher throughput, whereas the later is confined to cumbersome time consuming individual analysis.

Verification with international reference samples To validate the above test data, a set of ten international reference bauxite samples with known trihydrate content have been analyzed under the same set of experimental conditions. The data of these samples with different combinations of released (2.89, 2.87, 2.85, 2.83, 2.81 and 2.79 H20) and respective unreleased (0.11, 0.13, 0.15, 0.17, 0.19, 0.21 and 0.23 H20) water have been tabulated in the Table-4. The data indicates that out of the ten, four standards namely; BXT 05, 06, 07 and 09 show a direct match with the reference value within a 1% limit of variation. By applying a factor of 1.06 derived from the release of 2.83 H 2 0 molecules, as established in the above experimental case of

In the normal course, when wet chemical techniques are used for this analysis, parameters like boehmite and reactive silica effect the determination of trihydrate alumina. To verify this effect, the analysis of boehmite and reactive silica and other parameters were conducted and furnished in the table along with trihydrate data. Reactive silica ranged from 0.33 to 7.64% (column 9) and boehmite ranged from 0 to 33.52 (column 14) in the experimental samples. These parameters are not showing any significant influence on the respective values of trihydrate alumina determined by using thermogravimetry, confirming that the

213

Central Indian bauxite samples, five standards, i.e., BXT 02, 04, 08, 09 and 11 match with less than 1% variation. The other standards namely; BXT 01, and 03 show more than 1% variation between wet chemical and thermogravimetric determinations even after applying the factor. Since the mean value of THA is considered as the standard wet chemical reference and there is a chance of variation up to 1% in actual trihydrate value for an individual standard, these interpretations are considered as indicative but trends are close to the actual data observed for experimental samples. Sample ID

THA THA·,* value

BXT 01 54.3

BXT 02

m

BXT 03 51

BXT 04

415

BXT 05

414

BXT 06

45i

BXTÔ* 39.?

BXT 08

4L?

BXT 09 26.4

BXT 11 44.?

Rei

355 *C (Tri* eut factor)

TM

Va!« 54.3 +A05 46.5 +/-Ô.7 51.0 +/-04 415 +/-0.4 414 +/-0.6 45.5 +/-0.6 39.7 +/-0.6 41.7 +/-0.4 264 +/-1.0 44." +/-0.9

2.89 H : 0

187 H 2 0

2.85 H 2 0

183 H 2 0

181 H 2 0

The data obtained with some Central Indian bauxite sources showed very good agreement between the two methods. An excellent comparison is noticed up to 35% of trihydrate alumina in the bauxite and up to 45% it is comparable with acceptable limits of variation. Above 45%, the data needs appropriate correction based on the proportion of molecules of water released/unreleased. The test data obtained is verified by comparison with that of international reference bauxite samples. Similar behavior and trends were observed, indicating validity of the experimental data. The results also show scope for adapting the technique for other sources.

2.79 H 2 0

Dia. UM Dift 1.045 Dift 1.053 Dift 1.06 Dift 1.068 Dift 1.075 Dift

49.28 -5.02 51.15 -3.15 51.5

18 51.89 141 5124 -2.06 52.63 -1.67 5198 -1.32

4123 -3.27 44.8? •L63 45.18 IM 45.52

45.8« -5.14 47,6

Acknowledgement The authors are grateful to Dr. Mukesh Kumar, CEO and Mr. Bimalananda Senapati, Head of Refinery for their support. The authors also thankfully acknowledge the infrastructural help provided by the management of Vedanta Aluminium Ltd., Lanjigarh.

058 45.82 -Ô.6S 46.17 -0.33 46.47 -0,03

■3.4 4752 •3.08 43.29 -171 48.61 -2.39 48.98 -102 49.3

-L?

References 40.42 -108 4156 -0.54 4214 -016 42.56 0.06 4185

ÔJ5 43.17 0,6? 43.45 Ô.95

1. Lodding E., 1969. The gibbsite dehydroxylation fork. Schwenker RF and Gran, PD (eds.), Thermal Analysis, Volume-2. Inorganic materials and Physical Chemistry. New York: Academic Press pp. 1239-1250. 2. Indian Standard: Methods of sampling bauxite (First Revision), IS: 1999-1987 Page 1-15. 3. Indian Standard: Methods of chemical analysis of bauxite, Parti- Determination of loss on ignition (First Revision) IS: 2000 (Part-1) - 1985, Page 1-5 4. Indian Standard: Methods of chemical analysis of bauxite, Part 2- Determination of silica (First Revision) IS: 2000 (Part2)-1985, Page 1-6. 5. Indian Standard: Methods of chemical analysis of bauxite, Part 3- Determination of alumina (First Revision) IS: 2000 (Part-3) - 1985, Page 1-6. 6. Indian Standard: Methods of chemical analysis of bauxite, Part 4- Determination of ferric oxide (First Revision) IS: 2000 (Part-4)-1985, Page 1-6. 7. Indian Standard: Methods of chemical analysis of bauxite, Part 5- Determination of titania (First Revision) IS: 2000 (Part-5) - 1985, Page 1-5. 8. IBM Manual, 1979. Manual of procedures for chemical and instrumental analysis of ores, minerals and ore dressing products, Page 11-12, Issued by the Controller, Indian Bureau of Mines. 9. Ibrahima Sory Cisse and Jiwen Ge, 2010, Determination of bauxite's phases by the bomb digest method at Kamsar laboratory ISO 9002 (Guinea), Journal of American Science, 6 (6), 139-145, 2010. 10. Paulik F, Paulik J, Naumann R, Kohnke K and Petzold D, 1983. Mechanism and kinetics of the dehydration of hydrargillites, Part I & II. Thermochimica Acta, 64, 1-12

4199 0.59 44.62 121 4452 152 4517 187 45i? 3.1? 45.91 3il 46.21 3.81

44.9

M

40.36

ÙM 41.89 119 42.18 148 42i

46.61 1.11 4652 1.42 471«

39.08 -162 40.57 -1.13 40.84 -0.86 41.15

25.6?

4.n 26.65

4158 -112 44.2

L7I 47Ì9 109 47.95 245 4817 2.7?

18 42.78 3.08 43.1

3.4 43.39 3.69

455 41.42 -018 41.74 0.04 41Ö1 U l

125 26.83 0.43 27.03 0.63 2 ? l i &ß 27.42 1.02 27.6

11

-Oi 44.5 -01 44.84 0.14 45.13 (MS 45.48 0.78 45,77 1.07

Table 4. Thermogravimetric data of international reference Standard bauxite samples These observations show that the trends noticed with Central Indian bauxite samples are mostly applicable to the international reference standards and vice versa, hence confirming that the observations for Central Indian case can also be applied for other sources within the defined ranges of trihydrate concentration. The observations also indicate that wide variations and deviations in case of BXT 01 and 03 standards are most likely due to the very high concentrations of trihydrate present, i.e., > 50 % and they may need to have a different correction factor, depending on the release of water molecules at that concentration of trihydrate. A similar effect is also observed in case of Central Indian sources with more than 45% trihydrate. Conclusion Thermogravimetric analysis based on the loss of mass of water molecules is found to be a very fast and alternative technique for assessing trihydrate alumina in comparison to classical methods of analysis which are tedious and time consuming.

214

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DETERMINATION OF OXALATE ION IN BAYER LIQUOR USING ELECTROCHEMICAL METHOD Seval Turhan1'2, Betül Usta1'2, Yücel Çahin1'2, Oktay Uysal2 Anadolu University, Faculty of Science, Department of Chemistry, 26470 Eski§ehir; Turkey 2 ENTEKNO Industrial, Technological and Nano Materials Ltd., Anadolu University, Yunus Emre Campus, Anadolu Teknopark No: 104,26470 Eskiçehir; Turkey 1

Keywords: Bayer liquor; Oxalate; Pencil graphite electrode, Differential pulse voltammetry Abstract chromatography [7-10], high performance liquid Chromatograph [11-14], and gas chromatography [15-17]. In many factories, the oxalate was determined by titration with a standard potassium permanganate solution. This method often resulted in low precision and accuracy. Furthermore, it takes much time and many agents. Voltammetric techniques have not found use for this purpose. Recently, there has been a considerable effort in the development of voltammetric methods for the determination of some biological and industrial samples. It is generally believed that direct redox reactions of these species at the surface of usual electrodes are irreversible and therefore require high overpotential [18]. Moreover, the direct redox reaction of these species at the unmodified electrodes takes place at very similar potentials and often suffers from a pronounced fouling effect, which results in rather poor selectivity and reproducibility. In this study, we investigated the electrochemical determination of oxalate ion at pencil graphite electrode (PGE) by using differential puls voltammetry in Bayer liquor. A linear relationship between oxalate concentration and current response was obtained with good reproducibility of the current . Also real Bayer process samples were analysed by DPV and the results obtained compared with titration by potassium permanganate.

The Bayer process can be summarized as the digestion of bauxite with caustic liquor and the subsequent precipitation of hydrated alumina [1]. Most bauxite contains organic compounds in various amounts. Depending upon the digestion conditions, 510% of the organic carbon is converted to sodiumoxalate [2]. When sodiumoxalate, if not controlled in Bayer process, builds up to a certain level of supersaturation, it precipitates out in the hydrate precipitator tank. This co-precipitation affects the quality of alumina [3]. In this study, we investigated the electrochemical determination of oxalate ion by using differential puls voltammetry in Bayer liquor. A linear relationship between oxalate concentration and current response was obtained with good reproducibility of the current. Introduction The Bayer process can be summarised as the digestion of bauxite with caustic liquor and the subsequent precipitation of hydrated alumina [1]. Bauxite often contains organic matter in various amounts. Typically digestion of the bauxite during the Bayer process more man half of this organic carbon is extracted into the liquor. With recycling of the Bayer liquor the concentration of humic substances and their degration and oxidation products builds up an equilibrium concentration depending on the amount of organic carbon present in the bauxite. Because of this, in the Bayer process, the organic matter builds up to an equilibrium level, typically to 10-30 g/L of organic carbon content, determined by the outputs and the inputs [4]. Besides the bauxite input, some organic matter comes from other sources, such as process water, red mud flocculants or antifoams. The outputs are complex through the adsorption on the red mud or adsorption on the product. The accumulated organic matter and its breakdown products are known to cause numerous process problems. As a result, the liquors darken, and notable amounts of carbonate and oxalate are formed. Depending upon the digestion conditions, typically 5-10% of the organic carbon is converted to sodium oxalate [2,3]. Sodium oxalate has been shown to be very harmful with regard to alumina productivity and particle size [3, 5]. When sodium oxalate, if not controlled, attains a certain level of supersaturation, it precipitates out as fine needles in the hydrate precipitator tank. This coprecipitation affects the alumina product quality and productivity in many ways [3]. A variety of techniques have been utilized for the identification and determination of sodium oxalate and organic substances in Bayer liquors, these have included titrimetry [6], ion

Experimental Chemical and Reagents Sodium sulfate (99.0%, Aldrich), potassium permanganate (99.0%, Aldrich), and sodium oxalate (99.5%, Aldrich) reagents are commercially available as analytical grade and used without further purification. Stock solutions of oxalate solution were prepared daily from by using ultra-pure deinoized water. Bayer liquors and alumina hydrate were obtainaed from Seydiçehir Eti Aluminyum A.§. (Turkey). Apparatus Electrochemical experiments were carried out by a conventional three-electrode system. Pencil graphite (PG) was used as working electrode. Pt wire and saturated calomel electrode (SCE) were used as an auxiliary and reference electrode, respectively. Differential pulse voltammetry (DPV) measurements were performed by Autolab PGSTAT 100 Potentiostat/ Galvanostat with GPES 4.9 Version conversion software (EcoChemie, The Netherlands).

215

Electrochemical Detection of Oxalate Ion

Real Sample Analysis

The electrochemical detection of oxalate ion was performed in aqueous solution of 0.1 M Na2S04 using potential puls between 0.0 V and +1.5 V with a scan rate of 50 mVs"1 for three scans. All electroanalytical measurement were made at room temperature.

In order to verify the reliability of the electrode, it was applied for the determination of oxalate in Bayer liquors (pregnant liquor and spent liquor) and alumina hydrate samples. The oxalate concentrations in the samples were determined using the standard addition method and direct interpolation in the linear regression. Obtained results for oxalate ion concentrations from different Bayer liquors and alumina hydrate samples are given in Table I. Also oxalate in Bayer real samples determinated by using titration method. Obtained results are given Table 2

Result and Discussion Preparation of Calibration Curve for Oxalate Ion A series of DPVs was recorded at various concentrations of oxalate to determine its calibration curve. The response of the PG electrode to oxalate was found to increase with increasing oxalate concentration. Fig. 1 shows DPV curves recorded on the PG electrode in the presence of various oxalate concentrations in the range of 0.10-lOmM. Differential pulse voltammograms of oxalate had shown only a single peak (oxalate). The figure 2 shows the calibration curve obtained by measuring the DPV peak current intensity (in μΑ) vs. the concentration of oxalate (in mM) with a correlation coefficient of 0.9984.

0.250

0.500

0.750 E/V

1.000

1.250

Table I. Experimental Results of Oxalate Ion Samples [Oxalate] [Oxalate] with DPV with titration Pregnant liquor 33792 mg/L 48000 mg/L Spent liquor 21472 mg/L 26000 mg/L Hydrate alumina 0.122 mg/gsamDle 0.20mg/g samDle 1 Conclusion Electrochemistry has many advantages for making it an appealing choice for industrial oxalate ion analysis. Electrochemical method is a high specific and analitical method which is likely to get preference in sampling small volumes and complex sample matrices. The determination of sodium oxalate in several Bayer samples was undertaken and the results have been summarised in Table 1. As can be seen compared to results, titration method is higher than DPV method. Because of titration method often resulted in low precision and accuracy. In this study, PG electrodes was performed for the first time in different Bayer liquors and alumina hydrate. This method provide a simple and rapid technique to determine oxalate ion in high alkali media. Also, this method takes less time when compared with other methods. Further studies on the variations of electrode, electrolyte, linearity of response and quantitative properties of DPV method, etc. are underway. The developed method is straightforward and suitable for the determination of the oxalate ion in Bayer samples.

1.500

Fig. 1. .Differential pulse voltammograms of oxalate at various concentrations in 0.1 M Na2S04 at the PGE. Scan rate was 50mVs_1.

Acknowledgements The authors are grateful to Seydiçehir Eti Aluminyum A.C. (Turkey) for providing Bayer samples. References

10

1. T.G. Pearson, The Chemical Background of The Aluminium Industry (London, UK, Monograph No:5, Royal Institute of Chemistry, 1955), 137-143. 2. G. Lever, Travaux (1983), 13, 335 3. R. Calalo and T. Tran, "Effects of sodium oxalate on the precipitation of alumina trihydrate from synthetic sodium aluminate liquors " Light Metals, (1993), 125-133. 4. B. Gnyra and G. Lever, "Review of Bayer organics-oxalate control processes,"L/g/tf Metals, ( 1979), 151. 5. N Brown, and T.J Cole, "The behaviour of sodium oxalate in a bayer alumina plant," Light Metals, (1980), (15)105-117. 6 A.I. Vogel,, and G.H. Jeffery, Vogel's Textbook of Quantitative Inorganic Analysis (Harlow: Longman, 1986), 4th ed. 7. Grocott, S.C., Jefferies, L.P., Bowser, T., Carnevale, J., and Jackson, P.E., J. Chromatogr. A, 602 (1-2) (1992), 257.

12

[Oxalate]/mM Fig. 2. The analytical curve is linear in the range 0.10lO.OOmM of oxalate concentration with a correlation coefficient (tf2) of 0.9984.

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8. Jackson, P.E., J. Chromatogr. A, 693, (1) (1995) 155. 9. T Bowser,, and S.C, Grocott, Ion Chromatography in the Alumina and Aluminium Industries, Proc. 2nd Int. Alumina Quality Workshop, Perth, 1990. 10. Whelan, T.J., Shalliker, R.A., Mclntyre, C, and Wilson, M.A., Ind. Eng. Chem. Res.,, 44, (9), (2005) 3229. 11. Whelan, T.J., Kannangara, G.S.K., and Wilson, M.A., Ind. Eng. Chem. Res., 42 (26) (2003), 6673. 12. Whelan, T.J., Wilson, M.A., and Kannangara, K. (Proc. Aust. Organic Geochemistry Conference, 2002), 77. 13. Susie, M. and Armstrong, L.G., J. Chromatogr. A, 502 (1990), 443. 14. Xiao, J.B., Chen, X.Q., Jiang, X.Y., and Wu, S.D., Ann. Chim., 96 (5-6) (2006), 347. 15. Baker, A.R., Greenaway, A.M., and Ingram, C.W., "A Microwave Digestion-Based Determination of Low Molecular Weight Organic Acids in Bayer Process Liquor." Talanta, 42. (10) (1995), 1355. 16. Smeulders, D.E., Wilson, M.A., and Armstrong, L., "Insoluble organic compounds in the Bayer process,". Ind. Eng. Chem. Res., 40 (10) (2001), 2243-2251.. 17. Wilson, M.A., Ellis, A.V., Lee, G.S.H., Rose, H.R., Lu, X.Q., and Young, B.R., "Low-temperature products" Ind. Eng. Chem. Res., 38 (12) (1999), 4663. 18. R.N. Adams, "Probing brain chemistry with electroanalytical techniques,"^««/. Chem. 48 (14) (1976) 1126A-1138A.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINA and BAUXITE

Alternative Alumina Sources Poster Session SESSION CHAIR

James Metson University of Auckland Auckland, New Zealand

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

The effect of ultrasonic treatment on alumina leaching from calcium aluminate slag Sun Hui-lan1, Wang Bo1, Guo Dong1, Zhang Xue-zheng1, Bi Shi-wen2 ^ebei University of Science and Technology; 70 Yuhua East Rd; Shijiazhuang, Hebei, 050018, China; 2 Northeastern University; No. 11, Lane 3, Wenhua Road; Shenyang, Liaoning, 110004, China Keywords: calcium aluminate slag, alumina, leaching, ultrasonic 2.1. Materials and apparatus

Abstract

The material used in the leaching process is a calcium aluminate slag which is obtained by treating iron-bearing bauxite with a blast furnace. Table 1 shows the chemical composition of the slag, while Fig.l shows the XRD results from the slag. The main phases in the slag are 12Ca07Al 2 0 3 and y-2CaOSi02. After grinding, the particle size of the slag is below 74um.

The effect of ultrasonic treatment on the alumina leaching properties of calcium aluminate slags was investigated. The effects of the main leaching parameters such as ultrasonic power, leaching time, leaching temperature, and the concentration of sodium carbonate on alumina leaching rate were determined. The mechanism of ultrasonic effects on leaching is also discussed. It was found that the increase in leaching rate was slight, but leaching conditions were improved significantly when ultrasonic treatment was used in the leaching process. Leaching temperature and sodium carbonate concentration decreased by 30°C and 60g#L_1 respectively with the addition of ultrasonic treatment. Ultrasonics would not promote the decomposition of γ2CaOSi02. The agglomeration of leaching residue was prevented by the cavitation and mechanical effect of ultrasonic treatment.

Element Content (Wt.%)

Table 1 Chemical analysis of slag A12Q3 SiQ2 CaO TiQ2 22.24

16.05

54.89

1.23

MgO 2.12

l-Y -2CaO SÌO2

Introduction

2

With the rapid development of the steel and aluminum industries, the shortage of suitable resources of iron ore and bauxite has attracted considerable attention [1]. Therefore, the utilization of low grade ores becomes imperative. Leaching alumina from resources which are low grade and refractory, such as ironbearing bauxite, red mud and fly ash, has become a research focus in alumina production [2-4]. Iron-bearing bauxite is treated by the process of "Blast furnace Sintering-". After smelting, a calcium aluminate slag is obtained, and then it is used to leach alumina [57].

2r-12CaO 7A120^

1,2

3-CaTi03 4-MgO

1,2

1 1

1,3 2,3

22

r^M wj 20

1 ill

40

34 2 60

2 80

2Theta

Fig.l XRD spectra of calcium aluminate slag

However, there are many problems in alumina leaching of calcium aluminate slag. The required concentration of sodium carbonate in the leaching solution is high, leaching time is long, and the alumina leaching rate is low [8-9]. Solving these problems effectively could improve the comprehensive utilization value of iron-bearing bauxite.

The ultrasonic cleaner (KQ-100) used in the experiments is produced by Kunshan ultrasonic instrument Ltd. The frequency of the ultrasonic wave is 40 kHz, and the power, which is adjustable, is between 40 and 100W. Stirring of the solution is achieved using a high speed mixer. The tri-mouth glass flask is fitted with a thermometer, stirrer and a reflux condenser. Scanning electron microscopy (SHIMADZU SSX-550) is used to observe the surface morphology of the slag.

Ultrasonic wave treatment has advantages in leaching processes of ores which are difficult to treat [10-12]. Both the leaching rate of refractory ores and the leaching conditions are improved with the cavitation effect of the ultrasonic wave [13, 14]. The application of ultrasonics in hydrometallurgy is concentrated in Heavy metal and precious metal metallurgy, but little research has been carried out in alumina production, especially in treating iron-bearing bauxite using ultrasonics.

2.2. Procedure The sodium aluminate solution obtained from the slag is treated with the carbonization precipitation process, and circulating mother liquid is used to leach new calcium aluminate slag. The conditions of the leaching solution for alumina digestion are listed below: caustic alkali concentration Nk=7g»L"1, a^l.6 (molecular ratio between Na20 and A1203), liquid-solid ratio L/S=5. Other conditions are variable in different experiments such as sodium carbonate concentration, leaching time, leaching temperature, and ultrasonic power.

Therefore, in order to find an effective method to improve the leaching rate and to decrease the concentration of sodium carbonate and leaching temperature enhancing leaching of calcium aluminate slag with ultrasonic treatment was studied. Experimental

221

The leaching solution is preheated to the required temperature and then poured into the tri-neck glass flask in which the weight of calcium aluminate slag is 10 g. The leaching process begins when the speed of mixer is up to 300r*min"1.

treatment could improve the leaching conditions, but could not increase the maximum leaching rate. In addition, the increase of the useless bubble and the formation of a sound barrier, were the reason why the leaching rate changed slightly when the ultrasonic power was higher under both conditions.

After leaching and dry filtration, the filtrate is used to analyze the composition concentration of solution, and the filter residue is washed and dried for analysis.

3.2 Effect of leaching time on leaching rate of slag

Results and discussion

Leaching rates with different leaching times were investigated under the following conditions: Nc=120g»L"\ T=75°C and 45°C. Ultrasonic power, fixed according to section 3.1 was 120W in the following experiments. The results are shown in Fig.3.

3.1 Effect of ultrasonic power on leaching rate of slag Variations in ultrasonic power will cause large changes in the sound intensity and the maximum radius of cavitation bubbles. These changes will affect the leaching process of the slag.

-

The effect of ultrasonic power on leaching rate of slag was investigated under the following conditions: 1), sodium carbonate concentration (Nc^^Og'L"1, leaching temperature (T)=75°C; 2), Nc= óOg'L"1, T=45°C. Leaching time (t) was 80min, and other conditions were as shown in section 2.2. The results are shown in Fig.2.

-

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-

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h ul t r asoni e ho ut ul t rasoni c h ul t rasoni c hout ul t rasoni c

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1

60

80

100

Leaching time /min

The leaching rate of slag with ultrasonic and without ultrasonic treatment increased slightly with increasing leaching time when the leaching temperature was 75°C. That was because the reaction speed was fast at this temperature.

/ i

1

40

wi wi wi wi

Fig. 3 Effect of leaching time on leaching rate

-*-75°C, 120g.L_1

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80

120

160

200

With ultrasonic assistance, alumina leaching rate was improved visibly with increasing leaching time when leaching temperature was 45 °C. Extended leaching time had little effect on leaching rate. That was because the temperature and the initial leaching rate were low, so the cavitation effect was significant when time was less than 80min. When leaching time was over 80min most of the alumina was leached from the slag, so leaching rate increased slightly. The optimal leaching time was 80min, under this optimal condition the leaching rate with ultrasonic was close to the leaching rate which was obtained with or without the existence of ultrasonic when the temperature was 75 °C.

Ultrasonic power /w

Fig. 2 Effect of ultrasonic power on alumina leaching rate The alumina leaching rate of the slag was improved slightly with ultrasonic treatment, compared with that of slag without ultrasonics, when ultrasonic power was below 180w under condition 1. But if ultrasonic power was increased to 200W, the leaching rate could be improved by 1.5%. That was because the viscosity of the solution decreased and the movement of grains became violent when the power was high. Therefore, the leaching reaction was easy to carry out. For ultrasonic treatment at high temperature, cavitation occurred readily, but the cavitation intensity decreased with increasing temperature. In other words, because the leaching rate was high without ultrasonic and the cavitation effect decreased at high temperature, the effect of ultrasonic on leaching rate was slight.

Therefore, although ultrasonic could not increase the maximum leaching rate when leaching temperature was 45 °C and leaching time was 80min, it has decreased the required leaching temperature by 30°C.

The leaching rate increased significantly with increasing ultrasonic power under condition 2. The leaching rate changed little when ultrasonic power was up to 80W. The alumina leaching rate was improved by 5% with ultrasonic treatment. The optimal leaching rate obtained under condition 1 was close to that obtained under condition 2. This suggested that the ultrasonic

3.3 Effect of Nc on leaching rate of slag Nc(defined in section 3.1, passage 2) was high in the leaching process without ultrasonics, and this would increase the consumption of sodium carbonate. Experiments with different Nc were carried out under the following conditions: t=80min, L/S=5. The ultrasonic power was 120w. The results are shown in Fig. 4.

222

Leaching rate of slag without ultrasonic improved with increasing leaching temperature. Leaching rate with ultrasonic assistance was better than that without ultrasonics. When leaching temperature was 45°C and 75°C leaching rate of slag with ultrasonic increased by 10% and 4%. Effect of leaching temperature on leaching rate of slag with ultrasonic was not obvious. This meant that the addition of ultrasonic treatment weakened the effect of leaching temperature on the leaching rate of the slag.

94 92 90 SS ""■ 88 Φ

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—■—75°C —·—75°C —A—45°C —▼—45°C

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Generally, the increase of temperature will decrease the viscosity and promote molecular motion and Mass Transfer of liquid systems. So leaching rate without ultrasonic increased obviously with the increase of leaching temperature. When there was ultrasonic treatment, the vibration and cavitation effect also accelerated the speed of molecular motion and mass transfer in the liquid system. And under this condition the increase of temperature could not promote the leaching process unceasingly, so the effect of temperature was weakened.

with u l t r a s o n i c without ultrasonic with u l t r a s o n i c without ultrasonic

_i_

_l_

_L_

_L_

_JL_

80

90

100

110

120

Sodium carbonate concentration /g.L"1

Fig 4 Effect of Nc on leaching rate Fig.4 shows that with the increase of Nc, alumina leaching rate increased. When leaching temperature was 75 °C and Nc increased from 60 g^L"1 to 120 g'L"1 the alumina leaching rate without ultrasonic increased by 3.3%, but leaching rate with ultrasonic increased by 1.3%. The results indicated that Nc had little effect on leaching rate and the existence of ultrasonic weakened the slight effect when leaching temperature was high.

3,5 Effect of ultrasonic on stability of v^CaOSiO? in slag Although y-2CaOSi02 was stable, some of it in slag was still decomposed in sodium carbonate solution. The cavitation effect of ultrasonic treatment would cause instantaneous high local temperatures and pressures. Shock waves and micro jets would be generated which increase the reaction activity of solids [14]. Meanwhile, the shock wave could intensify the relative motion and collision probability of reactants. If the decomposition of γ2CaOSi0 2 was intensified with the existence of ultrasonic, the loss of alumina would increase, and these results were not expected [15]. Therefore, the content of silica in leaching solution was investigated under different ultrasonic power. Leaching time was 80min and L/S of slurry was 5. Other conditions and results are shown in Fig.6.

Nc had a great effect on alumina leaching rate when leaching temperature was 45 °C and ultrasonic was not used. Leaching rate increased obviously with the increase of Nc. But with the addition of ultrasonics, Nc had a little effect on leaching rate. The use of ultrasonic not only increases the leaching rate but also weakens the effect of Nc on leaching rate. The leaching rate of slag with ultrasonic under 45 °C and 60 g^L"1 was close to that of slag without ultrasonic under 75°C and 120 g·!/ 1 . In another word, under the circumstances that the extraction rate is guaranteed not to fall, the existence of ultrasonic could not only decrease leaching temperature by 30°C but also decrease Nc by 60 g#L"1.

1.4

3,4 Effect of leaching temperature on leaching rate of slag

1.2

The effect of leaching temperature on leaching rate of slag was studied when Nc was 60g#L"\ leaching time was 80min, and L/S was 5. The results were shown in Fig.5.

ί

1.0

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-■-75°C,120g.U 1

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Fig.6 Effect of ultrasonic power on silica content

S 84

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<

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-

Silica content decreased with the increase of ultrasonic power when leaching temperature was 45 °C and Nc was 60geL"1. The silica content changed little when temperature was 75 °C and Nc was 120g»L"1. This suggests that ultrasonic treatment does not promote the decomposition of y-2CaOSi02. Silica content under the former condition was much lower than that under the latter condition. So the influencing factors of decomposition of γ2CaOSi0 2 were leaching temperature and sodium carbonate concentration Ultrasonics did not appear to affect the

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223

decomposition of y-2CaOSi02. On the contrary, the improved leaching conditions decreased the decomposition of y-2CaOSi02.

sodium carbonate concentration decreased from 120 g»L to 60 g-L"1. (2) Ultrasonic didn't promote the decomposition of y-2CaOSi02. On the contrary, the improved leaching condition decreased the decomposition of y-2CaO#Si02.

3.5 Effect mechanism of ultrasonic on alumina leaching property of slag

(3) The effect mechanism of ultrasonic was that ultrasonic enhanced the dispersion of solid particle and prevented the agglomeration of leaching residue.

In addition to the cavitation effect, a mechanical effect would be generated by ultrasonics. The shock wave whose speed could reach lOOOmms"1 damaged and sheared the surface of solid particles, this helps to sustain a high reaction activity. In order to study the mechanism of ultrasonic treatment, micro morphology of the leaching residue was observed by SEM. The results were shown in Fig.7 and Fig. 8.

Acknowledgements The authors greatly acknowledge the financial support of the National Nature Science Foundation of China (Project No: 50674028), and the foundation of Hebei University of Science and Technology (Project No: XL200923). The authors express their profound gratitude to the editor and reviewers of TMS. References [1]

GUO Jing-jian, "Who will ensure the supply of bauxite in China/' China Metal Bulletin, 42 (2006), 2.

[2]

J.Grzymek, A.Derdacka and Z.Konik, "Method for obtaining aluminum oxide," U.S Patent : 4149898 , 19782-21.

[3]

J.Grzymek, "Complex Production of aluminium oxide and Iron from laterite raw materials applying the calcium aluminates polymorphism," TMS Annual Meeting and Exhibition, Light Metals, 1985, 87-99.

[4]

ZHANG Jing-dong, LI Yin-tai, BI Shi-wen and YANG Yihong, "Research on integrated utilization of high-ferrum bauxite in Guigang Guangxi," Light Metals, (8)(1992), 1618.

[5]

BI Shi-wen, YANG Yi-hong, LI Yin-tai, ZHANG Jingdong and DUAN Zhen-ying, "Study of alumina leaching from calcium aluminate slag," Light Metals, (6)(1992), 1015.

[6]

SUN Hui-lan, YU Hai-yan, WANG Bo, MIAO Yu, TU Gan-feng and BI Shi-wen, "Leaching dynamics of 12CaO#7Al203 ," The Chinese Journal of Nonferrous Metals, 18(10)(2008), 1920-1925.

[7]

TONG Zhi-fang, BI Shi-wen, YU Hai-yan and WU Yusheng, "Leaching kinetics of non-constant temperature process of calcium aluminate slag under microwave radiation," The Chinese Journal of Nonferrous Metals, 16(2)(2006), 357-362.

[8]

D.J.Connor, Aluminium extraction from non bauxitic materials, (Sydney: Aluminium-Verlag Gmbh, 1988), 230250.

[9]

WANG Bo, YU Hai-yan, SUN Hui-lan and BI Shi-wen, "Effect of material ratio on leaching and self-disintegrating property of calcium aluminate slag," Journal of Northeastern University: Natural Science, 29(11)(2008), 1593-1596.

[10]

L.B.Sukla, K.M.Swamy and K.L.Narayana, "Bioleaching

P ^ ^ ^ ^ ^ ^ ^ ^ ^ ^ ^ ^ ^ ^ ■ t e M M M M M M

Fig.7 SEM image of slag with ultrasonic wave ( X 300 ]

Fig.8 SEM image of slag without ultrasonic wave ( X300 ) The dispersion of leaching residue was better when ultrasonic was used in leaching process. The particle size of most of the residue was small. But the dispersion of leaching residue was worse when ultrasonic was not used in the leaching process. The agglomeration phenomenon was serious under these conditions. In addition, other changes such as cavities caused by shock waves and micro jets were not found in the images. Therefore, the effect mechanism of ultrasonic which improved the leaching process was that ultrasonic enhanced the dispersion of solid particle and prevented the agglomeration of leaching residue. Conclusions (1) Ultrasonic could not increase the alumina leaching rate greatly, but it could improve the leaching condition obviously. The leaching temperature decreased from 75 °C to 45 °C, and

224

of sukinda latente using ultrasonics," Hydrometallurgy, 37(3)(1995), 387-391. [11]

K.M.Swamy, L.B.Sukla and K.L.Narayana, "Use of ultrasound in microbial leaching of nickel from latérites," Ultrasonics Sonochemistry, 2(1)(1995), 5-9.

[12]

E.Sayan and M.Ayramoglu, "Statistical modeling and optimization of ultrasonic-assisted sulfuric acid leaching of Ti0 2 from red mud," Hydrometallurgy, 71(3)(2004), 397401.

[13]

A.S.Slaczka, "Effect of ultrasound on ammonium leaching of zinc from Galmei Ore," Ultrasonics Sonochemistry, 24(1)(1986), 53-55.

[14]

XUE Juan-qin, MAO Wei-bo, LU Xi, LI Jing-xian, WANG Yu-jie and WU Ming, "Dynamics of ultrasoundassisted oxidation leaching of nickel sulfide concentrate," The Chinese Journal of Nonferrous Metals, 20(5)2010, 1013-1020.

[15]

LI Xiao-bin, ZHAO Zhuo, LIU Gui-hua, ZHOU Qiu-sheng and PENG Zhi-hong, "Behavior of calcium silicate hydrate in aluminate solution," Transactions of Nonferrous Metals Society of China, 15(5)(2005), 1145-1149.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

THEORY AND EXPERIMENT ON COOLING STRATEGY DURING SEEDED PRECIPITATION Zhanwei Liu 12, Wenmi Chen \ Wangxing Li 1 2 School of Metallurgical Science and Engineering, Central South University, Changsha, Hunan, 410083, China 2 Zhengzhou Research Institute of Chalco, Zhengzhou, Henan, 450041, China Keywords: Seeded precipitation, Cooling way, Sandy alumina temperature of precipitation very low. So the a k (molecular ratio of Na2Ok to A1203) of the pregnant precipitation liquor is low and the supersaturation is high. Secondary nucleation occurs and the fining of the product particle size is caused, making industrial production fluctuate. This situation is more likely to happen when the caustic concentration is high and the seed coefficient is not very high.

Abstract The η-t curve is derived from the dynamic equation of seeded precipitation. The curve has the characteristic of a hyperbolic function and the experimental results conform to the characteristics of the curve. The η-t curve indicates that the precipitation ratio rises rapidly at the preliminary stage of precipitation, the precipitation ratio ascends slowly at middle stage of precipitation and the precipitation ratio increases more slowly at the final stage of precipitation. On the basis of this character of the η-t curve and the technology features of seeded precipitation in producing sandy alumina in Chinese alumina plants, various parameters are determined in the precipitation process as a result of intermediate lowering of temperature.

Some researchers have performed the experiment such that the caustic concentration of sodium aluminate liquor is 125g/l, the a k is 1.71, the seed coefficient is 2.30, the initial temperature of precipitation is 55°C, the final temperature of precipitation is 40 °C and the liquor was precipitated for 72h. They found the product particle size didn't become fine. Given closer analysis we know that because the concentration of the pregnant precipitation liquor is lower, a k is higher and the seed coefficient is relatively high, thus the fining of the product particle size is avoided.

Introduction The precipitation temperature is one of the most important factors influencing the precipitation ratio in the process of seeded precipitation. Based on the η-Τ curve, the precipitation ratio is highest when the temperature is somewhat lower than 40 °C. In industrial production, the temperature of the pregnant precipitation liquor after leaf filtering, is higher (90~95°C), so the liquor must be cooled down and then precipitated. For cooling, the textbook and many papers have shown that we should cool down quickly at first and then slowly. There are two main cooling strategies in industrial production. One way is that the pregnant precipitation liquor is cooled to around 60 °C directly and then precipitated. In general, this route is always taken by plants producing powder alumina. But this way has also been adopted by Pechiney and Pingguo Aluminium Company producing sandy alumina.

Things would be different if the precipitation conditions are slightly more demanding. When the caustic concentration of sodium aluminate liquor is 150~160g/l, the a k is 1.55. This sodium aluminate liquor was precipitated at constant temperature with a different seed coefficient. The product particle sizes with different amounts of seed are shown in Figure 1. 50 < 45 ,

Έ 40 Z2. 35

^

1

30 25

They must add in a great deal of seed, and so the solid content in the first precipitator reaches 750~850g/l. In order to prevent the product from becoming fine, an advanced automatic monitoring system is fitted. The other approach is an intermediate lowering of temperature, that is to say, firstly the pregnant precipitation liquor was cooled down to a temperature range between 70-^75 °C for agglomeration for some time, and then the intermediate lowering temperature is adopted to cool down the liquor for further precipitation. This way is used by plants producing sandy alumina with two-stage processes that is represented by the Aluminum Company of Switzerland.

ω 20 o

£

15 10 r 5 ! 50

55

60

65

70

75

Tenperature( °C )

Figure 1. The product particle sizes with different amount of seed

Using The Strategy Of Single Stage Cooling Or Intermediate Lowering Of Temperature

It can be seen from the Figure 1 that when the seed addition is high (300g/l and 500g/l, for example), the product particle sizes don't become fine at a precipitation temperature 60 °C, but the product particle sizes begin to become fine at precipitation temperature 55 °C; when the seed addition is 150g/l, the product

In general, if the strategy of single stage cooling is adopted, the initial temperature of precipitation must be cooled to a low temperature, such as 60 °C or under, in order to make the final

227

particle sizes are significantly fine at precipitation temperature 55 °C . When seed addition is low, such as 50g/l, even if the precipitation temperature is 70 °C, the product particle sizes also become fine, with the fining of the product particle sizes more serious at precipitation temperature 65"Cor lower.

A=\ +

1

at

C3t

η = ^ί + ^

(Λ-Α,Υ

Λ c, =· '' A.-4

(i)

C3=^

= KJ + C

The Character Of The η-t Curve In Laboratory Tests

When t=0, A=Ao, we can solve for C as follows:

c=-

We have done the precipitation testing in the laboratory to simulate plant conditions. The caustic concentration of the sodium aluminate liquor is 155g/l, the a k is 1.55. This sodium aluminate liquor was precipitated at constant temperature with different seed coefficients. The seed additions are 300g/l and 500g/l respectively with the seed obtained from the Henan Branch China Aluminium. The η-t curves are shown in Figures 2 and 3 respectively.

1

Λ-Λ

Where AQ is the concentration (g/1) of A1203 in sodium aluminate liquor. So we can solve for A as follows:

- A

K0(A0-Ae)

We will further explain the character of the η-t curve by the laboratory test.

(A-Ae)

1 A-Ae

=

The η-t curve equation (4) is a hyperbolic function, that is to say, when the seed coefficient is fitted (if the seed coefficient is too low, the induction period will be prolonged). The characteristic of the η-t curve is that the precipitation ratio rises rapidly at the preliminary stage of precipitation, the precipitation ratio ascends slowly at the middle stage of precipitation and the precipitation ratio increases more slowly at the final stage of precipitation.

(2)

Solving the differential equation (2) as follows:

-

(4)

Where:

Where A is the instant concentration (g/1) of A1203 in the sodium aluminate liquor, Ae is the equilibrium concentration (g/1) of AI2O3 in sodium aluminate liquor, t is the precipitation time (h), and Ko is precipitation rate constant. Equation (1) is transformed as follows:

-'<*-*>-**

Α,-Λ

c2t

Researchers in China and abroad have carried out many studies about the dynamics of seeded precipitation in sodium aluminate liquors, and similar dynamic equations are put forward. Next, the η-t curve is derived from the dynamic equation of seeded precipitation. Take following precipitation rate equation for example to deduce the η-t curve.

K0(A-Ae)2

A

= l-±-

The Theory Of Intermediate Lowering Temperature

=

(3)

It then substitutes into the precipitation ratio formula, equation (3), and we can find the n as follows:

The major bauxite resource is diaspore ore in China. The digestion conditions of this bauxite ore are more exigent, and the concentration of the pregnant precipitation liquor is high, that is always between 140g/l and 160g/l. If the strategy of one-stage cooling is adopted to make the initial temperature of precipitation drop to 60 °C or even lower, the fining of the product particle size is difficult to avoid.

~

Λ-Λ

i+K0t(A0-Ae)

1

-A- K t

A ι+* 0 ι(Λ-Λ)

-.

228

the precipitation ratio is about as high as that in the temperature range of 55-70 °C when the seed addition is 500g/l, the precipitation ratio is much higher than that in the temperature range of 55-70 °C after precipitating 12 hours when the seed addition is 300g/l. The Determination Of Various Parameters Of Intermediate Lowering Temperature By Using The Change Rules Of The η-t Curve The Determination Of The Initial Temperature Of Precipitation

0

10

20

30 Tira* h)

40

50

Because the precipitation ratios are about the same for the first 12 hours in the temperature range of 55-70 °C , it is proper to determine the initial temperature of precipitation in the temperature range of 65-70 °C which is a bit higher within the temperature range of 55-70°C. By doing this, the precipitation ratios are not affected and the fine particles of aluminium trihydroxide can be agglomerated very well.

60

Figure 2. The η-t curve with 300g/l seed addition

The Determination Of The Time Of Intermediate Lowering Temperature

0

10

20

30 Tìne(h)

40

50

Because the precipitation ratios are about same in the abovementioned temperature ranges, and the gaps become too big for different temperatures with the longer times of precipitation, we should cool down after precipitation for 12-14 hours. In this way the liquor can be precipitated at a low temperature in time, and the precipitation ratio is improved. If the time of intermediate lowering of temperature is too late, the liquor is still precipitated at high temperature, so the precipitation ratio is influenced. The later it is cooled down, the greater the precipitation ratio is influenced. It is not appropriate to cool down too early, for example after precipitating 7-8 hours, if so, the product particle size may become fine and scale is readily produced on the interstage cooling equipment.

60

Figure 3. The η-t curve with 500g/l seed addition

The Modes Of The Precipitator Before The Intermediate Lowering Of Temperature

The η-t curves in Figures 2 and 3 are completely in conformity with the change rules of the above-mentioned η-t curve derived from the dynamic equation. The precipitation ratio rises rapidly at the preliminary stage of precipitation. It can be seen from Figures 2 and 3 that in the temperature range of 50-70 °C, the precipitation ratios reach about 30% for the first 12 hours which is the 65 per cent of the total precipitation ratios for 60 hours when the seed addition is 300g/l; the precipitation ratios reach about 35% for the first 12 hours, which is 70 per cent of the total precipitation ratio for 60 hours when the seed addition is 500g/l. But the precipitation ratio ascends slowly over the middle stage of precipitation and the precipitation ratio increases slower and slower at final stage of precipitation. The precipitation ratios are about same at different times for the first 8 hours in the temperature range of 50-70 °C; the precipitation ratios are also about same for the first 12 hours in the temperature range of 55-70 °C ; the precipitation ratios are different at different temperatures after precipitation for 24 hours, the gap increases with longer time of precipitation, and the precipitation ratios are higher with decreasing temperature. If the precipitation temperature is too high (75 °C), the precipitation ratios are much lower than that in the temperature range of 50-70 °C at the preliminary, middle and final stage of precipitation; if the precipitation temperature is too low (50°C),

Seeded precipitation is performed in batch operation or continuous operation. Currently, continuous operation is adopted in most plants. The disadvantage of continuous operation is that the pregnant precipitation liquor with the lower a k flows continuously into the precipitator to rapidly mix with the precipitated liquor ( a k is higher) in this precipitator, so the a kis increased immediately and the precipitation ratio is decreased. Of course, the precipitation ratio of the liquor flowing from this precipitator is also decreased. From the characteristics of the above-mentioned η-t curve, we know that the precipitation ratios for the first ten hours or more, are 65-70 per cent of the total precipitation ratios for the full precipitation time. So it is at the initial stage of precipitation (more specially, it is at the first precipitator) that the seeded precipitation performed in continuous operation influences the precipitation ratio remarkably. But there are no obvious influences on the middle and final stages of precipitation. To a certain extent, the disadvantage of continuous operation can be overcome by adopting the strategy of intermediate lowering of temperature. That is to say, the pregnant precipitation liquor flows into the precipitator from the top which is stirred interlayer, so the a k of the pregnant precipitation liquor is not significantly increased,

229

then the precipitated slurry flows out from the bottom of the precipitator and was pumped into interstage cooling equipment.

These include the time of intermediate lowering of temperature, the initial temperature of precipitation and the cooling extent of intermediate lowering of temperature and so on. This has certain direct significance to the determination of the temperature decreasing strategy in the process of seeded precipitation.

The Technical Projects On The Intermediate Lowering Of Temperature Cooling The Pregnant Precipitation Liquor By Plate Heat Exchangers

Reference 1. Wangxing Li, "The Impact of Temperature on Precipitation Ratio in the Process of Seeded Precipitation," Light Metals, (5) (1998), 14-18.

The pregnant precipitation liquor was cooled down to the temperature ranges between 65/~~70°C from the temperature 95 °C by using plate heat exchangers. It is calculated that it will be possible to save 200,000 tons of steam every year, generating 10 million yuan every year in economic returns, if the spent seed precipitation liquor was preheated by the way of cooling down the pregnant precipitation liquor.

2. Yuling Wang, Huaizhong Zhu and Liwei Zhou, "Application of Intermediate Lowering Temperature in the Process of Seeded precipitation," Advanced Ceramics, (2) (2005), 27-28. 3. Junzhao Fan, "The Technology of Producing Sandy Alumina Using Sodium Aluminate Solution with High Solid Content," Non-ferrous Smelting, (3) (2002), 50-54.

The Location Of Interstage Cooling Equipments In The Series Of Precipitators

4. Jinqing Chen et al., "Depth Analyzing of Seeded Precipitation and Discussion on Intensify Method" Non-ferrous Metals, (6) (2003), 30-34.

It has been mentioned above that it is appropriate to cool down the liquor after precipitating 12-14 hours. So the location of interstage cooling equipment will depend on the retention time of the liquor in each precipitator. If the retention time of the liquor in each precipitator is range of 10~12h, we should begin to cool down the liquor after it flows out from the first precipitator. Of course, we should begin to cool down the liquor after it flows out from the second precipitator, if the retention time of the liquor in each precipitator is in the range of 5~8h.

5. Chongyu Yang, The Process Technology of Alumina, (China, NY: Metallurgical Industry Press, 1993).

The Tube Heat Exchangers As The Interstage Cooling Equipment Because there are a lot of Al(OH)3 solids in the liquor, the plate heat exchangers are easily blocked, so the plate heat exchangers are not suitable for this application. Tube heat exchangers and vacuum heat exchangers are preferred equipment. The tube heat exchanger is simple and practical, which was developed in Hungary and used successfully in industrial production. The temperature of the slurry is 65~70°C as mentioned above, and it needs to be cooled to around 50 °C, thus the cooling requirement is considerable. So if it is not enough to cool down by using one-stage cooling equipment, we can adopt a multi-stage series to cool down the liquor to the set temperature. To conclude, the temperature of the pregnant precipitation liquor after leaf filtering is 90 ~ 95 °C . It was cooled down to temperature ranges between 65 ~ 70 °C by the plate heat exchangers and flows into the first precipitator from the top, then the precipitated slurry flows out from the bottom of the precipitator and was pumped into multi-stage tube heat exchangers to cool down to temperature ranges between 50~55 °Cand then was feed into follow-up precipitator series to continue precipitation. The temperature of the precipitated slurry in the last precipitator is 40~45°C. Conclusion The characters of the η-t curve are studied by the way of the basic theory of precipitation, combined with experiments. This characteristic curve elucidates the rationale for adopting the strategy of intermediate lowering of temperature in the process of seeded precipitation and the technical parameters are determined.

230

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EXTRACTION OF ALUMINA FROM RED MUD BY DIVALENT ALKALINE EARTH METAL SODA ASH SINTER PROCESS S. N. Mener1, A.K. Rout1, B. K. Padhi2 1

Department of Chemistry, Kiit University, Bhubaneswar, Orissa, India, Pin-751024.

2

R & D Department, Nalco, Damanjodi, Dist.-Koraput, Orissa, India, Pin -763008. Email Id: [email protected] Key Words: Red Mud, Sintering, Leaching and Divalent alkaline earth metal oxide. soda and alumina. Recoveries of 95% Na20 and 70% A1203 were obtained [5]. Alyanak et al. [6] studied on the recovery of aluminium from the red mud of the Seydisehir Aluminium Company Limited. They prepared a mixture composed of 26.6 wt% red mud, 37.2 wt% CaC03 and 36.2 wt% Na2C03. The mixture was sintered at 900°C for 2 hours then leached in water. Increasing the sintering temperature as high as 1150°C decreased the alumina extraction efficiency.

Abstract A novel method for extraction of alumina from red mud using a divalent alkaline earth metal oxide soda ash sinter process followed by thermo chemical leaching in suitable caustic concentration at 105°C for different times is presented in this paper. During sintering Sodalite/Cancrinite, Gibbsite and Boehmite in red mud is converted to sodium aluminate as well as stable phases of divalent alkaline earth metal ferrite, titanate and silicate. In the above novel economic process an extraction of alumina from red mud of 97.64%, 98.7% and 99.5% was achieved using CaO, MgO and BaO respectively. The morphological study of sinter products and leached products was studied using SEM and XRD.

In another article, a gallium recovery method was applied to bauxite ores, during alumina production by the Bayer process. Gallium movement and possible locations of gallium recovery in the Etibank-Sedisehir Aluminium Plant and some results for gallium concentrate production by precipitation were presented [7]. S. Uday [8] had worked on the recovery of iron, titanium and aluminium values from red mud by calcining it, preferably between 800°C to 1350°C and smelting the material with a carbonaceous reducing agent in an electric furnace, thus obtaining molten iron and a slag containing substantially all the titania, alumina and silica.

1. Introduction Red mud is a by-product in the manufacture of alumina that contains mainly iron oxide (54-65%) with significant amounts of silica, alumina, calcium oxide and titania, dispersed in a highly alkaline and caustic liquor [1,2]. The treatment and disposal of bauxite residue is a major operation and may account for 30-50% of operations in an alumina refinery [1]. The red mud has been accumulating at a rate of 70 million tonnes annually throughout the world [3]. For a given aluminum production rate, the quantity of red mud generated during the alumina extraction process varies significantly depending on the original properties of the bauxite and the operating conditions of the Bayer process, and in particular, the process temperature [4]. Depending on the factors, between 1 and 2.5 tonnes of bauxite residue is generated per tonne of alumina produced [1]. The disposal of the red mud is associated with pollution problems. Disposal of any solid waste is associated with space or real estate near the industry, and the cost of disposal and pollution, which are now crucial factors. Obviously, these three factors are also associated with the red mud disposal. As the red mud contains a large amount of valuable chemicals, there is a need for developing a technology for the recovery of at least some of the important chemicals [3].

The red mud has also been investigated for the making of artistic glazes in ceramic. For this purpose, no additives and some fluxing materials such as read lead, ulexide, cryolite, borax, soda ash, etc. were added to the red mud in different proportions. Finally, the optimum conditions forming novel glazes were determined [9]. The red mud was also investigated to find out whether it can be used as a construction material or not. For this purpose, the red mud, a waste material of Aluminum Plant in Seydisehir, was mixed with clay, which was taken from the same region. This mixture was pressed and sintered at between 1000 and 1150°C. It was stated that construction materials produced from the mixture had higher compressive strength and lower water absorption ratio compared with the related standards [10]. In another study, the structure of the red mud deposits of Seydisehir, Turkey was investigated by simple statistical analysis. It was shown that the ratio of the iron oxide % (D i) in the red mud to the mean iron oxide % (D2) was found to fît a linear function represented by plotting D\ID2 versus K λΙΚ 2 in a double logarithmic plot. The relation in the form of a power function was D 1= 1.0092D2/r10021/j:2-0022[ll].

Sulfided red mud is active as a hydrogénation catalyst, due to its iron sulfide content, and has been used for the hydrogénation of organic compounds, and the liquefaction of coal and biomass [2].

The thermal behaviour of the red mud was investigated in the literature. An extensive characterization was performed by thermal and X-ray diffraction analyses. Dried red mud was substantially inert up to 900°C, the loss of H 2 0 from aluminium hydroxides and of C0 2 from silico-alumino-carbonates being the only detectable effects [12]. Several methods for extracting alumina as well as for recovery of other metal values from alumina-bearing materials

In the literature, a lot of different articles that explain the properties and the investigation of the red mud tailings can be found. One such reference investigated the red mud under caustic concentrations of 10-30% to determine the extraction of soda and alumina as a function of caustic ratio, CaO: Si0 2 ratio, water content, temperature and reaction time. Hydrothermal treatment of the red mud with lime at high temperatures of 300°C was found to be an effective method for the recovery of

231

adding divalent alkaline earth metal oxide (i.e. CaO, MgO and BaO) and Na2C03 to the red mud and sintering at temperatures of 900-1100°C [14]. The silica in the red mud reacts with divalent alkaline earth metal oxide to form the relatively inert divalent alkaline earth metal silicate as shown in equation: 1. The sodium, which is in the red mud via soda ash or sodium carbonate, reacts with the alumina and forms the soluble sodium aluminate as given in equation:2 [15]. The Fe203 contents of red mud reacts with sodium carbonate to form sodium ferrite as given in equation: 3.The Ti0 2 which is in the red mud reacts with sodium carbonate forms sodium titanate as given in equation: 4. The sodium titanate reacts with divalent alkaline earth metal oxide forms divalent alkaline earth metal titanate and free Na20 as given in equation: 5. The sodium ferrite reacts with divalent alkaline earth metal oxide forms divalent alkaline earth metal ferrite and free Na20 as given in equation: 6.

have been studied [13]. The United States Bureau of Mines has been evaluating possible alternatives to the use of bauxite for several years. These include acid leaching processes and sintering processes applied to clays and anorthosites. Fly ash from burning regular coals has also been investigated extensively as a potential source for alumina. The processes studied for the extraction of alumina from fly ash include variations of existing sintering and leaching processes. There are now a few millions of tonnes red mud stored in Red Mud ponds in India. These wastes have not been investigated yet in any industrial process or used for material production as an additive. The aim of this study was to investigate these muds for maximum alumina extraction by a divalent alkaline earth metal oxide soda ash sinter process followed by thermo-chemical leaching. 2. Materials and Methods

Si0 2 +MO -> M2Si04 + 2H20 A1203 + Na2C03 -► 2NaA102 +C0 2 Fe 2 0 3 + Na2C03 -♦ 2NaFe02 +C0 2 Ti0 2 + Na2C03 -> Na2Ti03 +C0 2 Na2Ti03+MO -»MTi0 3 +Na 2 0 2NaFe02 + 2MO -> M2Fe205 +Na20

Red mud and crushed Lime (CaO) used in experimental studies were supplied from, Damanjodi, Dist-Koraput. The MgO, BaO and Na2C03 used in experimental studies were supplied from Qualigens fine Chemicals, Bombay. The major elements were analysed by classical methods as well as X-Ray Fluorescence Spectroscopy (Model PW 2400). The minor elements were analysed by AAS. The chemical composition of the red mud is shown in Table 1. X-Ray Diffraction (XRD) analysis of sintered red mud and leached red mud were carried out with a Diffractometer using CuKa (Ka= 1.54186 A) Rigaku Dmax 2200, Japan. The morphological behaviours of sinter red mud and leached red mud were studied by SEM (Model:430, Electron Microscopy, United Kingdom).

The product is then leached in an alkaline solution or water as given in equation: 7 and the sodium aluminate solution is directed to the precipitation stage of the Bayer process. NaA102(s) + M2Si04(s) + 2NaFe02(s) + 3H20 -» NaAl(OH)4(aq) + 2NaOH(aq) + M2Si04(s) + Fe203(s) (7) To produce high aluminium and sodium extractions the decomposition of divalent alkaline earth metal silicates must be as low as possible. In practice the amount of silica extracted in the leaching step is too high, as the M2Si04 is not sufficiently stable under the leaching conditions [16] and silica extractions of around 15-20% [17] occur during leaching. This ties up sodium and aluminium with the formation of Mg3Al2(OH)12, Ca3Al2(OH)12 or Ba3Al2(OH)12, Hydro garnet and DSP. The opportunity exists for a process to treat red mud to separate aluminium and sodium from silica, producing an adequately stable insoluble phase for leaching and therefore yielding greater extractions.

An amount of 100 g of red mud, 10 g of divalent alkaline earth metal oxide (i.e. CaO, MgO and BaO) and 15 g / 20 g/25 g of Na2C03 was mixed thoroughly and pulverised in a pulveriser, then the above mixture was sintered at temperatures of 900, 1000, 1100°C for 1 to 4 hrs in a ceramic crucible in a high temperature muffle furnace. An amount of 100 g of red mud, 15 g of divalent alkaline earth metal oxide (i.e. CaO, MgO and BaO) and 10 g/20 g /25 g of Na2C03 was mixed thoroughly and pulverised in a pulveriser, then the above mixture was sintered at temperatures of 900, 1000, 1100°C for 1 to 4 hours in a ceramic crucible in a high temperature Muffle furnace. An amount of 100 g of red mud, 20 g of divalent alkaline earth metal oxide (i.e. CaO, MgO and BaO) and 10 g/20 g/25 g of Na2C03 was mixed thoroughly and pulverised in a pulveriser, then the above mixture was sintered at temperatures of 900, 1000, 1100°C for 1 to 4 hrs in a ceramic crucible in high temperature muffle furnace. After sintering the sinter products were leached with 80 gpl caustic at 105°C for 1 hr in an autoclave of 2.5-litre capacity. The leached red mud was filtered after the leaching. The amount of A1203 in the leach solutions was determined by using a titrimetric method.

An alternative method of divalent alkaline earth metal oxide Na2C03 sinter is presented in this paper to produce a M : Si ratio of one in the sinter product[16]. The potential benefits are to halve raw material costs and to give a more stable insoluble phase. Also the extra sodium needed could be provided by spent liquor creating the potential for organics removal through liquor burning in the sinter step. The removal of organic impurities in sintering followed by leaching at 105°C is due to formation of TCA or Layered Double Hydroxide [Hydrotalcite: Mg6Al2(OH)16 C03.4H20)].

Table 1. Chemical Analysis of Red Mud €&C9BÉCtf

AbO>

FesQs

THfe

SiOr

N*0

CaO

VaO*

IfgQ

<*0

iS-IS

5Ï-57

3-5

S-Î2

4-6

8.8823®

Ü14Ù.2Ì

0.13-

UaO

loia*

02»

11-13

(l) (2) (3) (4) (5) (6)

However, thermodynamic calculations and laboratory tests show that MSi03 does not form under sinter conditions [18]. From the present experiments it was thought that the following divalent alkaline earth metal oxide sinter reaction was feasible due to the presence of sodalite in the red mud formed by the reaction of kaolin and sodium aluminate liquor.

3. Results and Discussion Methods of aluminium and sodium recovery from red mud have been around for many years due to the requirement of some refineries to process very high silica bauxites. The most common method to process red mud is the "divalent alkaline earth metal oxide Na2C03 sinter process". The process involves

Na2O.Al203.2Si02 ( Sodalite ) +2 MO + 4Na2C03 -* 2Na2MSi04+2NaA102+4C02+2H20 (8)

232

A similar lime sinter process is referred to in a paper discussing the integration of coal combustion with lime sintering [19]. Reaction (8) shows that soluble sodium aluminate and sodium alkaline earth metal silicates are produced. The sodium associated with sodium alkaline earth metal silicate must also be recovered in the leaching step. The objective of this project was to investigate the use of a M : Si ratio of one, in a divalent alkaline earth metal oxide, Na2C03 sinter process for use as a red mud treatment for the recovery of aluminium and sodium and to compare this with the current lime soda ash sinter system of M:Si = 2.

RM+MgO+Na2C03(1 :.1 :.25)-900-4 4000-

LO

g- 3000 -

'S

1 2000 e

1 2 3 4 5 6 7 8 9

Temperature in°C

Time in hours

900 1000 1100 900 1000 1100 900 1000 1100

1 3 4 4 4 3 1 4 1

2.50

I? 3.50

4.50

5.50

d-value

Figure 1: MgO Soda Ash Sinter Process with highest Extraction Efficiency 98.70% at 900°C.

Table 2. Extraction Efficiency of divalent alkaline earth metal oxide soda ash sinter process at different condition. Condition RM+MO (CaO, MgO, BaO)+ Na2C03 1:0.10:0.20 1:0.15:0.25 1:0.20:0.25 1:0.10:0.25 1:0.15:0.20 1:0.20:0.25 1:0.10:0.25 1:0.15:0.25 1:0.20:0.25

i 1 Irl

1.50

The extraction efficiency of divalent alkaline earth metal oxide soda ash sinter process is shown in Table 2.

SI. no.

co

Extraction 1 Efficiency (%)

RM+CaO+Na2C03(1 :.2:.25)-1100-1

Θ

61.79 66.66 97.64 98.70 97.22 96.50 99.14 99.50 95.57

Figure 2: Lime Soda Ash Sinter Process with highest Efficiency 97.64% at 1100°C.

*Sl.No.l to 3: RM + CaO + Na2C03 sinter process, Sl.No.4 to 6: RM + MgO + Na2C03 sinter process, and Sl.No.7 to 9: RM + BaO + Na2C03 sinter process. In the synthetic MgO Na2C03 sinter stage, calcination at 900°C for 4 hrs with red mud, MgO and Na2C03 in a ratio of 1:0.1:0.25 was found to produce the most stable Sodium Magnesium Silicates, Di-Magnesium Silicates, Magnesium Ferrite (SEM Fig.6) and Magnesium Titanate (SEM Fig.7). This was confirmed by XRD (as shown in Fig .1 and Table 3) and maximized the extraction of alumina at 98.7%. The lowest extraction efficiency of 47.84% obtained in the MgO Na2C03 sinter process with red mud MgO Na2C03 ratio (1:0.15:0.10) sintered at 1000°C for 1 hr is due to the formation of insoluble magnesium aluminium silicate hydrate and sodium aluminium magnesium silico titanate [NaAlMgSiTiOx].

In the synthetic BaO Na2C03 sinter stage, calcination at 1000°C for 4 hrs with red mud, BaO and Na2C03 in a ratio of 1:0.15:0.25 was found to produce the most stable Sodium Barium Silicates, Di-Barium Silicates, Barium Ferrite (SEM Fig. 8) and Barium Titanate (SEM Fig.9) confirmed by XRD (as shown in Fig. 3 and Table 3) which maximized at 99.5% the extraction of aluminium and sodium. The lowest extraction efficiency of 32.56% obtained in the BaO Na2C03 sinter process with red mud BaO Na2C03 ratio (1:0.2:0.1) sintered at 1100°C for 3 hrs is due to the formation of insoluble Barium Aluminium Silicate Hydrate and Sodium Aluminium Silicate. RM+BaO+Na2C03(1 :.15:.25)-1000-4

In the synthetic CaO Na2C03 sinter stage, calcination at 1100°C for 4 hrs with red mud, CaO and Na2C03 in a ratio of 1:0.2:0.25 was found to produce the most stable Sodium meta aluminate, Sodium Calcium Silicates (Fig.4.a) Di-Calcium Silicates (Fig.4.a), Calcium Titanate (Fig.4.b) and Calcium Ferrite (Fig.5.a and 5.b) confirmed by XRD (as shown in XRD Fig. 2 and Table 3). This procedure maximized the extraction of alumina at 97.64%. The lowest extraction efficiency of 38.35% obtained in the CaO Na2C03 sinter process with red mud CaO Na2C03 in a ratio (1:0.1:0.15) sintered at 900°C for 2 hrs is due to the formation of insoluble calcium aluminium silicate.

B

1800

5.50

Figure 3: BaO Soda Ash Sinter Process with highest Extraction Efficiency 99.50% at 1000°C.

233

The existing CaO Na2C03 sinter product ties some calcium up with iron due to the formation of Ca2Fe205 (d value= 2.70,2.67,2.10,1.84, XRD Fig. 2 and Table 3), which have Whisker like structures as shown in the SEM images in Fig.5.a and 5.b. This meant that extra calcium was added to ensure the maximum amount of silica was reacted with the calcium and not with sodium or aluminium. At 1100°C, instead of sodium aluminate and sodium ferrite forming, Na2AlFe04 formed in products with iron present. For the synthetic new lime soda ash sinter process without iron it was found that a two stage caustic or water leach (105°C/60 min., 105°C/240 min.) was required to gain maximum Na and Al extractions due to the decrease in Al in solution over time as shown in Figure 10. It was found that in an MgO Na2C03 sinter process sintering at 900°C for 4 hrs with red mud MgO Na2C03 in a ratio of 1: 0.1:0.25 followed by leaching, formation of Sodium aluminate, MgTi03 and Mg2Si04 maximise alumina extraction up to 98.7%. In a BaO Na2C03 sinter process sintering at 1000°C for 4 hrs with a red mud BaO Na2C03 ratio of 1:0.15:0.25 followed by leaching and formation of Sodium Aluminate, Ba2Fe205 (SEM Fig.8) and BaTi03 (SEM Fig.9) give maximum alumina extraction efficiency of 99.5%.

When Iron is introduced to the system there are effectively two soluble phases (NaA102 + NaFe02) in which the ratio of soluble sodium to soluble aluminium increases. The results in Figures 12 and 13 show the benefit of the new alkaline earth metal Na2C03 sinter process in terms of aluminium extraction. Much of the sodium and all of the aluminium is extracted from the existing alkaline earth metal sinter process in the first stage. The second stage merely completes the leaching of sodium. In the new red mud alkaline earth metal sinter extraction the resultant solids have a sodium content of 0.36 % compared to the 6.00 % sodium content of the red mud. The aluminium levels have been reduced from 17.8% to 0.42%. A more stable sodium alkaline earth metal silicate would increase these extractions and could be achieved with the optimisation of sinter temperature and leach conditions.

This allowed the removal of the majority of the aluminium before the decomposition of the sodium divalent alkaline earth metal silicate and subsequent reaction to form DSP, Hydro garnet or Mg3Al2(OH)12, Ca3Al2(OH)12 or Ba3Al2(OH)12. To ensure leaching conditions were the same for both systems the two-stage process was also employed for the existing lime sinter products. The results of extractions for the sinter products without iron are displayed in Figure 11.

Fig.12.New vs E x i s t i n g S y n t h e t i c CaO N a 2 C 0 3 Sinter

Fig.13 New Vs E x i s t i n g Red Mud CaO N a 2 C 0 3 Sinter 10

20 Tim e(hrs)

Fig.10 Single Stage Extraction over time

100

<

Mag-

2.00 KX

E H T * 15.00 kV

Fig.11.New vs Existing S y n t h e t i c CaO N a 2 C 0 3 S i n t e r w i t h o u t Iron

f°»m'

Detector = SE 1 Date :11 Jul 2007

Figure 4a: Formation of Sodium Calcium Silicate and DiCalcium Silicate (2.00 KX)

234

Figure 4b: Formation Calcium Titanate (4KX) J

*c

i "XL' % MM

Figure 5a: Formation of Calcium Ferrite (4KX)

Figure 7: Formation of Magnesium Titanate (4.05KX)

Figure 5b: Formation of Calcium Ferrite 10K

Figure 8: Formation of Barium Ferrite (4.69KX)

235

2. In the MgO-Na2C03 sinter process alumina extraction achieved was 98.7%. 3. In the BaO-Na2C03 sinter process, the extraction efficiency achieved was 99.5%. But from an economic point of view, the lime-soda sinter process is best and suitable method for extracting the remaining alumina from Bayer's process waste residues (Red Mud). References

Figure 9: Formation of Barium Titanate (3.88KX) Table 3. Formation of major phases confirmed by XRD during sintering of red mud followed by leaching Sinter Process Red Mud + CaO +Na2C03

Red Mud + MgO+Na2C03

RedMud + BaO + Na2C03

Formation of Major phases Dicalcium Silicate Calcium Titanate Calcium Ferrite NaA102 CaSi03, CaV206 CaMn03 Calcium Aluminum Silicate Dimagnesium Silicate Magnesium Titanate Magnesium Ferrite NaA102 Dibarium Silicate

Characteristic d spacing (A) 2.21

Barium Titanate Barium Ferrite NaA102

2.13,2.24 1.63,2.43,2.63,2.78 2.13,2.23

1

2.70,2.52,2.87 2.70,2.67,2.10,1.84 2.13,2.59,2.64,2.95 2.70,2.67,2.21 2.67 2.64 2.53 2.70,2.53,2.09 2.96 2.96 2.78,3.15

4. Conclusion 1. By adoption of the sintering method followed by alkaline leaching, the alumina extraction achieved was 97.64% in LimeSoda Sinter process and impurities like Si02, V 2 0 5 and Mn0 2 are removed from the liquor. These impurities are removed from the crystal lattice of Hematite as well as Sodalite and Cancrinite to form insoluble materials like CaSi03 (d value= 2.70, 2.21), CaV206 (d value= 2.70, 2.67) and CaMn03 (d value= 2.67) confirmed by XRD.

1. Q. D. Nguyen and D. V. Boger, Int J Miner Procès, (54) (1998), 217-233. 2. J. Alvarez, S. Ordonez, R. Rosai, H. Sastre and F.V. Diez, Appi CatalA: Gen, (180) (1999), 399-409. 3. B. K. Mohapatra, M. B. S. Rao, R. Bhima Rao and A. K. Paul, "Characteristics of Red Mud generated at Nalco Refinery, Damanjodi, India", (Light Metals 2000 as held at 129th TMS Annual Meeting; Nashville, TN; USA, 2000), 161-165. 4. L.Y. Li, Waste Management, (21) (2001), 525-534. 5. P. J. Creswell and D. J. Milne, (Proceedings of the Annual Meeting on the Light Metals AIME, TX, USA,1982), 227-238. 6. C. Alyanak, C. Oktaybas, H. Sesigur and E. Acma, (Proceedings of the International Metallurgy and Materials Congress, Istanbul, Turkey, 1995), 91-97. 7. H. Yuzer, E. Avci, O. Emrullahoglu, E. Gencer and T. Haser, (Proceedings of the International Metallurgy and Materials Congress, Ankara, Turkey ,1993), 215-223. 8. S. Uday, Metallurgical and Chemical Process Limited (Hamilton, Canada), U.K.Patent GB 843607(1960). 9. Z. Mete and A. Cam, (Proceedings of the Chemistry and Chemical Engineering Congress, Istanbul, Turkey, 1992), 1924. 10. Kara M., Ekerim A. and Emrullahoglu O. F., (Proceedings of the International Metallurgy and Materials Congress, Istanbul, Turkey, 1995), 1435-1440. 11. Sahin S, Hydrometallurgy, (47) (1998), 371-376. 12. V.M. Sglavo, R. Campostrini, S. Maurina, G. Carturan, M Monagheddu, G. Budroni and G. J. Cocco, Euro Ceramic Soc, (20) (2000), 235-244. 13. A.K. Mehrotra , P.K. Bashnor and W.Y. Svrcek, Cart Jour ChemEng.,(57) (1979), 225-32 14. W. R. King, "High Caustic Mud-Sinter process for high silica weips Bauxite", Non-Metallic Materials Research, Kaiser Aluminium and Chemical Corporation, (RR 80-71,Project 7991-60310,1980). 15. P. Leci and A. Guidi, "Pyrogenic Attack of Bauxite" extractive Metallurgy of Al, (1) (1962), 231-249. 16. A. Hartsborn, "Lime Sinter with Calcium: Silica ratio: :l",Comalco, (RTS Technical Note 9392, 2000). 17. Q. Likuan, "The Recovery of Aluminium by Soda-Lime Sinter Process", (Zheng Zhon Conference Proceedings 1993). 18. Osborne J., "Thermodynamic Analysis of Lime Sintering of De-Silication Products ", (Rio Tinto Technical Note, 27731/1997). 19. V. L. Rayzman and I. K. Filipovich, "Integrating coal combustion and red mud sintering at an alumina refinery," 70M,51(8)(1999),16-18.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DISSOLUTION KINETICS OF SILICON FROM RED MUD IN PURE WATER Xiaohui Li, Kai Huang, Hongmin Zhu* School of Metallurgical and Ecological Engineering, University of Science and Technology Beijing, 100083, China (* Corresponding author: Tel: +86-10-62332267; Email: [email protected] ) Keywords: red mud; dissolution kinetics; liquid layer diffusion; Noyes-Whitney equation atmosphere and so decreasing the greenhouse effect.

Abstract

Haraguchi et al proposed that the utility of steelmaking

In this study, the dissolution kinetics of silicon from sintering

slag as the nutritive elements to promote the growth of algae

red mud in pure water was investigated. Experimental factors

[8,9]. Now in this study, we have tried to check the feasibility of

such as particle size and stirring speeds were studied to test their

using red mud as a nutrient source. Our experiments indicated

effect on the extraction fraction of silicon from red mud. It was

that it is also quite viable to use red mud for the significant

found that the silicon extraction increases with stirring speed

promotion of the growth of algae and the results will be

and decreasing in the particle size of red mud. Based on the

published later. In the process of the production of alumina, a

experimental results, the Noyes-Whitney equation was adopted

huge amount of red mud is produced causing serious

to describe and explain the dissolution behavior of the red mud

alkalization of the soil and polluting the underground water after

in pure water, and the experimental data fit the model quite well

long term storage, and the treatment and comprehensive

in the whole process of dissolution.

utilization of red mud is a troublesome problem. Though a large number of studies have been conducted on the comprehensive

Introduction

utilization of red mud, the research progress is not significant.

Thousands of phytoplankton live under the surface water

However, taking accountant of the fact that red muds are also

of the ocean and constitute the primary productivity of the ocean.

rich in Fe and Si, we tried to investigate its dissolution behavior

The phytoplankton breeds rapidly under appropriate conditions,

preceding its real application in promoting the growth of algae

and the growth of phytoplankton would consume huge amounts

of the ocean.

of carbon dioxide through photosynthesis [1]. Of the global emissions of carbon dioxide, about 50% of those from human

Experimental

activities are captured the oceans [2]. Phytoplankton growth

The red mud sample used in this study was donated by

needs not only C, N, P nutrient sources, but also Fe, Si and other

Shandong Aluminum Company. Samples were dried, crushed

trace elements. Martin et al found that deficiency of iron can

into smaller particles by ball milling and sieved to the different

inhibit the growth of algae [3,4]. However, there are many

particle size fractions. Particles with the size range of

oceans that lack trace elements such as Fe, Si. It is reported that

39-185μπι were sent for analysis of chemical composition by

about 20% of the ocean area has excessive N, P nutrients, but

X-ray fluorescence spectroscopy.

the phytoplankton biomass is very low, and this area is known

The dissolution experiments were carried out in the plastic

as high nutrient-low chlorophyll (HNLC) [5]. The researchers

containers. The suspension was agitated with a Teflon stirrer,

found that the iron limits phytoplankton productivity in the

driven by a magnetic stirring apparatus. A water-bath heater was

HNLC, thus the deficiency of iron affected the carbon dioxide

employed to provide the stable heating conditions at the

transportation from the upper sea water to the deeper [6,7]. If

predetermined temperature. When the temperature reaches the

iron was added in the HNLC, the phytoplankton growth could

required value, a certain amount of red mud particles was added

be promoted, consuming excessive N, P nutrients and

and the dissolution reaction was started. After the different

accelerating the carbon transporting from the ocean surface to

reaction time intervals, the solution was sampled, filtered and

depth and ultimately reducing the carbon dioxide level of

the clear supernatant solution was sent for analysis of the silicon content by UV visible spectrophotometer (UV2000). Taking

237

into account the different factors to test the extraction efficiency

interface between the particles with the solution. It was shown

of silicon from the red mud, dissolution time intervals from 0 to

that the decreasing of particle size can provide greater diffusion

120 min were examined.

area, which made the extraction of silicon increase.

In order to investigate the dissolution behavior of red mud,

4

we have carried out long time dissolution trials up to 100 hours.

3.5

An SEM instrument was also used to observe the morphology variation of the solid particles.

t

Results and discussion Chemical composition

3 2.5

O 1.5

From the X-ray fluorescence spectrum, the composition of

1

red mud can be determined as shown in Table 1. It was shown

0.5

that the main metal elements include iron, silicon, aluminum

0

and calcium etc, in which these elements mainly existed in the form of their oxide. According to our knowledge, all these

— · — 80rpm —♦— 160rpm

0

20

as silicon from dissolution in the leach liquor would be quite

4

Table 1.Chemical composition of the red mud sample (wt.%) Si0 2

MgO

P2O5

MnO

Γ1Ο2

A1203

Cont

36.01

29.61

21.44

6.49

1.71

1.7

1.31

1.73

120

0.5 g/L, T=20°C)

low.

Fe 2 0 3

100

concentration in the leach liquor (particle size range: 46~74μπι,

predicted that the concentration of the interesting elements such

CaO

60 80 time /min

Fig.l. Effect of stirring speed and time on the silicon

oxides belong to the sparingly soluble phases, so it can be

Cmp

40

3.5 3

Effect of stirring speed Fig.l shows the variation of extracted silicon concentration

E a a

2.5

U

1.5

2 —♦— 39~46μ m

1

in the leach liquor from the red mud with different dissolution times, for the samples under different stirring speeds of 80rpm,

0.5

and 160rpm. It can be seen that the silicon in the red mud could

0

be extracted rapidly into water in the first 60 minutes, then the dissolution rate would gradually decrease, after 120 minutes it

0

—♦— 74~185μ m

20

40

60 80 time / min

100 120

nearly reached an equilibrium state. It was found that increasing

Fig.2. Effect of particle size and time on the silicon

of the stirring speed could promote the extraction of silicon

concentration in the leach liquor (T=20°C, 0.5 g/L, 300 r/min)

from the red mud.

Kinetics analysis

Effect of particle size

In order to study the changes of surface morphology before

Fig.2 shows the variation of extracted silicon concentration

and after long term dissolution, samples were observed by SEM

in the leach liquor from the red mud with dissolution time for

after dissolution for lOOh. As shown in the Fig.3, there seems

the samples with different particle sizes of 74~185μπι and

almost no difference in the microstructure for the samples after

39~46μπι. It was found that the decrease of the particle size

long term dissolution in pure water. Considering that the

could enhance the extraction of silicon from the red mud, i.e.,

extraction of silicon in pure water is quite low as shown in Fig.2,

the smaller the particle size, the higher the extraction rate of

the maximum amount of silicon extraction from the red mud

silicon. With decrease of particle size, the corresponding

reached only 3.57ppm. It can be concluded that the dissolution

particle surface area will increase, leading to a larger contact

238

of silicon from red mud in the pure water was a very slow and is

particles, the variation of the silicon concentration extracted into

a nearly homogeneous dissolution process.

the water may follow the liquid layer diffusion control model as shown in equation (1) which was derived from the classic Noyes-Whitney equation [10], and at t=to, C=0 , Csurface=Cs, the

«

corresponding initial conditions of equation (1) can be derived.

,n<

(1)

FTF ) = T' + c ·

In which, D—diffusion efficiency of silicon through in the bulk liquid layer; C—silicon concentration in the leach liquor at the moment of t; Cs— water solubility of silicon compound under the experimental conditions; A—contact interface area between the red mud particles and water; δ—liquid layer thickness; Cerror—the error of experimental data.

3.5 3 ~

2

·5

"

2

g

1.5 1

Fig.3. Scanning Electron micrograph photos of red mud sample

0.5

before and after dissolution lOOhours, (a) before, (b) after.

0

0

20

40

When red mud particles were put into the pure water, the

60 80 time / min

100

120

particle surface reached saturation solubility rapidly and a liquid

Fig.4. Fitting curves based on the liquid layer diffusion equation

layer will be established around the particle in the dissolution

for different particle sizes of red mud

process. Between the two sides of the liquid layer, a

(T=20°C,300r/min,0.5g/L)

concentration gradient will appear between the solid particle surface and bulk solution, which makes the elements diffuse

Ψ

from the particle surface through the liquid layer into the bulk

3

solution. The experimental results suggest that the dissolution is

2.5

quite fast at the beginning, then the extraction of silicon will gradually decrease and after

— · — 80rpm ♦ 160rpm

•

? 2

120min it nearly reaches

equilibrium. It was found that the extraction of silicon from the red mud is increased with faster stirring. Because in this case,

^y^m

^

the thickness of liquid layer will gradually become thinner and the extraction of silicon will consequently increase. So the

^/+

1 0.5

observed dissolution seems follow a process of dissolution of

n

silicon through a saturated solution layer to the bulk solution, and it can be defined as limited by liquid layer diffusion. Under the different conditions of the dissolution of red mud

239

20

40

60 80 time /min

100

120

1

Fig.5. Fitting curves based on the liquid layer diffusion equation

References

at different stirring speeds (particle size ranged ΐη:46~74μπι,

[1] Liu Q, Environmental Chemistry (Beijing, Chemical Engineering Press, 2004), 127-133.

T=20°C, 0.5 g/L)

[2] Zhang ZB, Liu LS, Advance of Ocean Chemistry (Beijing , Chemical Engineering Press, 2005), 65-67.

Based on the dissolution experimental data of silicon from

[3] Martin JH, Broenkow WW, Fitzwater SE, et al., "Yes, it does:

red mud in pure water at different stirring speed and particles

A reply to the comment by Banse," Limnol. Oceanogr, 35 (3)

sizes, the fitting results were calculated and depicted in Fig.4

(1990), 775-777.

and Fig.5, respectively. All the curves fitted quite well according

[4] Martin JH, "Glacial-interglacial C0 2 change: the iron

to the equation (2). It was found that the whole process of

hypothesis," Paleoceanography, (1990), no.5:l-13.

dissolution was in agreement with the results of our proposed

[5] Leblanc K, Hare CE, Boyd PW et al., "Fe and Zn effects on

assumption, suggesting that the whole dissolution process of

the Si cycle and diatom community structure in two contrasting

silicon was controlled by liquid layer diffusion step. Moreover, this simple diffusion model can be used to describe this kind of

high and low silicate HNLC areas," Deep Sea Res .1, (2005), 52.

dissolution process where the extraction rate of substances are

[6] Martin JH, Coale KH, Johnson KS, et al., "Testing the iron

quite low, there is negligible change of the particle size, and the

hypo thesis in eco- systems of the equatorial Ocean," Nature,

particle surface reaches saturation solubility in a very short

(1994), no. 371: 123-129.

period of time.

[7] Coale KH, Johnson KS, Fitzwater SE, et al., "A massive

Conclusions

phytoplankton bloom induced by an ecosystem scale iron

The dissolution process and kinetics of silicon dissolution

fertilization experiment in the equatorial Pacific Ocean," Nature,

from red mud in pure water was studied. It was found that the

(1996), no.383: 495-501.

silicon extraction was a quite slow and is a nearly homogeneous

[8] Haraguchi K, Taniguchi A, "Effect of simultaneous

dissolution process with the change of particle size negligible

enrichment of dephosphorization steelmaking slag and treated

after the long dissolution times. In this kind of dissolution

municipal sewage on growth of coastal phytoplankton

system, the concentration of silicon would reach a saturation

assemblage," ISU Intern., 8(4) (2003), 430-437.

equilibrium state around the particle surface in a very short time,

[9] Haraguchi K, Suzuki K, Taniguchi A, "Effects of

that is, a saturation layer was quickly established close to the

steelmaking slag addition on growth of marine phytoplankton,"

phase interface, and its thickness was affected by stirring speed

ISU Intern., 43(9)(2003), 1461-1468.

and so on. Based on the dissolution experimental data for silicon

[10] Noyes, A.A., Whitney, W.R., "The rate of solution of

from red mud in pure water, the models are fitted very well

solid substances in their own solutions," J. Am. Chem. Soc,

according to the equation, suggesting that the whole dissolution

19(1897), 930-934.

process of silicon follows the liquid layer diffusion model.

240

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

The effect of cooling rate on theteachabilityof calcium aluminate slags Wang Bo 1 , Sun Hui-lan1, Zhang Xue-zheng1, Bi Shi-wen2 ^ e b e i University of Science and Technology; 70 Yuhua East Rd; Shijiazhuang, Hebei, 050018, China; 2 Northeastern University; No.ll, Lane 3, Wenhua Road; Shenyang, Liaoning, 110004, China Keywords: calcium aluminate slag, cooling rate, self-disintegrating, alumina leaching The cooling rate will not only affect the components and structure of the slag, but also affect the self-disintegration rate of the slag[3]. If the cooling rate is fast, the glass phase will be formed in the between the crystal grains. The glass phase first will prevent the crystal transformation of 2 C a O S i 0 2 (from ß to γ), and second will cover the surface of 2 C a 0 7 A l 2 0 3 . Therefore, the alumina leaching rate of the slag decreases. Besides, a slag which does not self-disintegrate is hard to crush and grind, and this process will also consume a lot of energy.

Abstract A calcium aluminate slag was synthesized by melting a CaC0 3 , Si0 2 and A1 2 0 3 mixture at 1500°C for lh. The effect of cooling rate on alumina leaching and self-disintegration of the calcium aluminate slag were studied. The mechanism was examined using the methods XRD, SEM, and laser particle size analysis. The results indicated that the phase components, alumina leaching property and self-disintegrating property of the calcium aluminate slag were better when cooling rate was lower than 20°Cmin"1. The formation of 2CaOAl 2 0 3 «Si0 2 and 3CaOAl 2 0 3 decreases alumina leaching rate and self-disintegrating rate when the cooling rate was higher than 20°Cmin"1. The conversion rate from ß-2CaOSi0 2 to y-2CaOSi0 2 was still high when cooling rate was between 20°Cmin"1 and SO'Cmin"1. Therefore, the main reason of the decrease in self-disintegrating rate was the formation of2CaOAl 2 0 3 *Si0 2 .

With regard to the optimal cooling rate of calcium aluminate slags Eremin [4] considered that the proper value was S'Cnnin 1 ~ 6T>min_1. But Zhang Jing-dong suggested that if the cooling rate of slag was slower than lO'Omin" 1 when the initial temperature of the slag was between 1400°C and 1250°C, the alumina leaching rate would be higher. Therefore, studies of the effect of cooling rate of calcium aluminate slags have been carried out by previous researchers, based on actual slags, and the results have been useful [5-7]. However, the optimal cooling system was not determined, and the effective mechanism was unclear according to these studies. Therefore, in order to confirm the phase transformation law and the effect of cooling rate on calcium aluminate slags, a study of cooling rate through XRD, SEM and particle size analysis has been carried out and is reported in this paper. In order to exclude the effect of impurities and other factors, the materials which were used to synthesis the slags were analytically pure.

Introduction The cooling rate of calcium aluminate slags is one of the important parameters in the "sintering - blast furnace smelting alumina leaching" process, which is used to utilize iron-bearing bauxite. During the process of blast furnace smelting, iron and a calcium aluminate slag are separated, and then the slag is subjected to a cooling and self-disintegrating procedure[l, 2]. Calcium aluminate slag is in a liquid state when the temperature is 1500°C. The slag completes the transformation from liquid phase to solid phase during the cooling process. There are many pathways as the melt of calcium aluminate slag cools down. The first is that the melt, which has no time to crystallize when the slag is quenched, exists in the form of a glass phase. The second is that the liquid has no time to react with the solid phase which is generated during cooling process, and then crystallizes independently. The third is equilibrium crystallization according to the phase diagram. Therefore, the cooling rate is the key to optimizing the alumina leaching properties and self-disintegration of the slag.

Experimental 2.1. Materials and apparatus Calcium aluminate slag was synthesized by mixing reagents A1 2 0 3 , Si0 2 and CaC0 3 . The A/S (the weight ration of A1 2 0 3 to Si0 2 ) of slag was 1.3, and the C/A (CaO/Al 2 0 3 , molar ratio, except CaO of 2CaOSi0 2 ) of slag was 1.7. The components of materials were shown in Table 1. Table 1 Components of materials Components CaO A12Q3 SiQ 2 W% 57.25 24.16 18.59

Calcium aluminate slag system is polysystem, so the components of solid, which is crystallized from the liquid phase, are different from that of liquid. If the cooling rate is very fast, it will affect the diffusion of liquid phase and then change the solid composition and crystallization path. And besides, it is important that there is a peritectic reaction between gehlenite and liquid phase, and this reaction is very slow. If the cooling rate is fast, the liquid has no time to react with gehlenite, it will form a new solid/liquid mixture and the effect will be different than if gehlenite had formed. That is, the new liquid system will deviate from the best operating region, and does not crystallize the phase which is optimal for leaching.

2.2 Smelting of calcium aluminate slag The samples with different cooling rates were smelted in a vacuum furnace, and the container was a graphite crucible. A Vacuum furnace was used in the experiments largely due to its thin insulating layer, and thus low thermal mass. It could thus realize rapid heating and cooling. The smelting temperature was 1500°C and the holding time was one hour. The sample was taken out of the vacuum furnace at

241

400 °C. Cooling rates of the calcium aluminate slag between 2°C min"1 and 40 °C min"1 were studied in this paper.

7000

* - Y -2GO S02

2.3 Leaching of calcium aluminate slag

• - 12QO 7AL203 v - GO

£

O Ü

4000

3000

1000

ujy^ 20

40

60

80

2Theta

2.4 Analysis

Fig. 2 XRD spectra of slag with cooling rate=10°C -min"1

The contents of A1203 and Si0 2 of samples and filtrate were analyzed by a chemical method. The Phase components of calcium aluminate slag were identified by X-ray diffraction analysis (PANalytical PW3040/60). A Malvern laser particle analyzer was used to analyze the particle size of the slag.

7000 » - Y - 2CaO Si 0 2 6000

* - CaO Al O

4000

£

3.1 Phase transformation law of slag with different cooling rate

c 8

3000

In order to study the phase transformation law of calcium aluminate slags with different cooling rates, XRD analysis was used in the experiments. The results are shown in Fig. 1 to Fig. 4.

2000

^UJIM^

7000 - Y -2QO Si02 6000

• - 12CaO 7AI 2C>

5000

Results and discussion

2Theta

- 12GO 7ALj03 - QO ALP3

Fig. 3 XRD spectra of slag with cooling rate=20°C min"1

5000

w

Αψ3

5000

The sodium aluminate solution obtained from slag leaching should then be treated by a carbonization precipitation process, and the circulating mother liquor used to leach new calcium aluminate slag. Therefore, the feasible conditions for alumina digestion were: leaching temperature was 75 °C, leaching time was 2h, L/S ratio was 4.5(by weight), caustic alkali concentration was 7g/L and sodium carbonate concentration was 120 g/L. The leaching experiments were carried out in a magnetically stirred, constant temperature water bath. After leaching and dry filtration, the filtrate was used to analyze the composition of the solution, and the filter residue was washed and dried to preserve for analysis.

4000 - γ -2CaO

Si0 2

- 12CaO 7AI 2 0 3 - CaO Al 2 0 3 - 2CaO ΑΙ 2 0 3 · S 0 2

20

IwJJi^^ 40

60

5

3000h

80

2Tteta Fig.l XRD spectra of slag with cooling rate=2°C min"1

^HwllW 2Theta

Fig. 4 XRD spectra of slag with cooling rate>40°C min"1 Fig.l shows that when the cooling rate was 2°C-min"1, the calcium aluminate slag had a proper phase component. The main phases were 12Ca07Al 2 0 3 (alumina is easy to be leached out) and γ-

242

2CaOSi0 2 (slag could self-disintegrate because of its phase transformation). Little CaOAl 2 0 3 (the alumina teachability is lower than 12Ca07Al203) was formed in the slag. When the cooling rate increased to lO'Cmin"1 and 20°Cmin"1 the phase components of slag were similar to the slag whose cooling rate was 2°C-min"1(Fig.2 and Fig.3). 2CaOAl203»Si02 which was difficult to leach was not found under the three conditions. But 2CaOAl203»Si02 could be found in slag when cooling rate was above 40°C-min ^ . 4 ) . ß-2CaOSi02(2.78 , 2.61 , 2.19Â) was not found in slag either when cooling rate was not above 20°Cmin"1. It could be found when cooling rate was above 40°C-min"1, but its amount was limited. Therefore, its characteristic peaks were very low and were not marked in Fig.4.

Therefore, a certain amount of 3CaOAl 2 0 3 and 2CaOAl203»Si02 would appear when the cooling rate was very rapid. This would affect the alumina leaching property and selfdisintegrating property of calcium aluminate slag.

According to the former results when the cooling rate was under 20°C-min"1 the calcium aluminate slag would have an ideal phase composition. If the cooling rate was accelerated continuously, 2CaOAl203»Si02 would appear and effect the phase of slag. Trace quantities of ß-2CaOSi02 was formed when cooling rate was over 40°Cmin"1. And this indicated that 2CaOSi0 2 had a rapid crystal transformation speed, so cooling rate had a little effect on crystal transformation of 2CaOSi0 2 under the conditions studied in this paper. These results accorded with Zhang Xiong [8], who deemed that 2CaOSi0 2 had a higher crystal transformation ratio when cooling rate was between 2°Cmin"1 and 500 °C min"1. He further noted that the transformation rates of y-2CaOSi02 were 100%, 95% and 90% when the conditions were slow cooling (about 2°Cmin"1), natural cooling (about 60 °C min"1) and wind cooling (about 500 °C min"1).

The slag with different cooling rate was well-mixed in order to analyze the granularity of slag. The particle size results analyzed by Malvern 2000 are shown in Fig.6.

3.2 Effect of cooling rate on self-disintegrating property of slag The self-disintegration of calcium aluminate slag can reduce the energy consumption during production, and it is a very important characteristic of calcium aluminate slags. The percent content of granularity which is lower than 74um in samples is defined as the self-disintegrating rate.

In addition, when the cooling rate of calcium aluminate slag was over lO'C-min"1 trace amount of particles (particle size bigger than 1mm) which didn't self-disintegrate completely appeared. XRD analysis was carried out on these grinded particles, and the result is shown in Fig.5.

15

• - 12CaQ

7Ai,0 3

* -CaO

6000

4000

2000

n

'.

30

Fig.6 indicates that the self-disintegrating rate of slag decreased with acceleration of cooling rate. However, the self-disintegrating rate of slag was still higher than 90% when cooling rate was not higher than 20°Cmin"1, this suggested that under this condition the effect of cooling rate on pulverization was little. The results could be verified by XRD (Fig.l to Fig.3). ß-2CaOSi02 and 2CaOAl203»Si02 were not found in slag, so the crystal transformation of y-2CaOSi02 was complete and the particle size was relative stability.

* - 3CàO AL,03 8000

25

Fig.6 Self-disintegrating rate of slag with different cooling rate

12000

10000

20

Cooling rate / eC · min"1

•

ILjJL L L ^ .

If cooling rate increased continually, the self-disintegrating property of slag decreased obviously. When cooling rate was 40°Cmin"1 the self-disintegrating rate was only about 80%. There were two points for the decrease according to XRD analysis. The first was the formation of 2CaOAl203#Si02 (Fig.4) which decreased the content of 2CaOSi0 2 of slag, thus the selfdisintegrating rate decreased. The second was the existence of the trace of ß-2CaOSi02 which decreased the crystal transformation rateof2CaOSi0 2 .

^^LJ

2Theta

Fig.5 XRD spectra of particles without self-disintegrating Fig.5 shows that the major phase of particles which didn't selfdisintegrate was 3CaOAl203, and the minor phases were f-CaO and 12Ca07Al203.

Further, according to report [8] the crystal transformation rate was still higher than 95% even under natural cooling condition (about 60 "C-min"1). So the effect of the existence of ß-2CaOSi02 on pulverization of slag was little.

243

and 3CaOAl203 which were difficult to leach would be formed if the cooling rate increased continually. And because of that the self-disintegrating property and alumina leaching property of slag would decrease obviously.

Therefore, the major reason of the decrease of self-disintegrating rate was the formation of 2CaOAl203#Si02; the minor reason was the existence of ß-2CaOSi02.

(2) The crystal transformation rate was still high when cooling rate was between 20°C-min"1 and SiTC-min"1. Therefore, the main reason of the decrease of self-disintegrating rate was the formation of 2CaOAl203»Si02.

4,3.4 Effect of cooling rate on alumina leaching property of slag The alumina leaching experiment of slag was carried out in order to study the effect of cooling rate on the alumina leaching properties. The leaching conditions and process were as in section 2.3, and the results are shown in Fig.7. 88

84

2

The authors greatly acknowledge the financial support of the National Nature Science Foundation of China (Project No: 50674028), and the foundation of Hebei University of Science and Technology (Project No: XL200921). The authors express their profound gratitude to the editors and reviewers of TMS.

■

86

* 82 "53

Acknowledgements

-

References

■V

80

«50

Ì78

ä

-

[1]

J.Grzymek, A.Derdacka and Z.Konik, "Method for obtaining aluminum oxide," U.S Patent : 4149898 , 19782-21.

[2]

ZHANG Jing-dong, LI Yin-tai, BI Shi-wen and YANG Yihong, "Research on integrated utilization of high-ferrum bauxite in Guigang Guangxi," Light Metals, (8)(1992), 1618.

[3]

BI Shi-wen, YANG Yi-hong, LI Yin-tai, ZHANG Jingdong and DUAN Zhen-ying, "Study of alumina leaching from calcium aluminate slag," Light Metals, (6)(1992), 1015. N.I.Eremin, "Investgations on the complex processing of bauxites," (Bauxite-Alumina-Aluminum, Symposium of ICSOBA . Budapest: Research Institute for Non-Ferrous Metals, 1971), 329-335.

3 76 ■\.

74 72

-

7n

l_

0

5

i

i

1

1

1

1

1

10

15

20

25

30

35

40

45

Cooling rate/ °C · rri n 1

Fig.7 Leaching property of slag with different cooling rate The alumina leaching rate of slag showed a decreasing tendency with the increase of cooling rate (Fig.7). When the cooling rate was in slow cooling stage (<5"C-min-1) the slag had a better phase components and particle size, so the alumina leaching rate was higher (>85%).

[4]

When the cooling rate was in medium cooling stage (ö'C-min"1 ~ 20°C-min"1) the alumina leaching rate decreased but was still above 80%. That was because the main phases were still 12Ca07Al203 and y-2CaOSi02. The decrease of leaching rate was because of the particles which didn't self-disintegrate. The formation of 3CaOAl203 and the existence of f-CaO which didn't react with A1203 or Si0 2 in these particles decreased the content of 12Ca07Al203.

[5]

J.Grzymek, "Complex Production of aluminium oxide and Iron from laterite raw materials applying the calcium aluminates polymorphism," TMS Annual Meeting and Exhibition, Light Metals, 1985, 87-99.

[6]

WANG Bo, YU Hai-yan, SUN Hui-lan and BI Shi-wen, "Effect of material ratio on leaching and self-disintegrating property of calcium aluminate slag," Journal of Northeastern University: Natural Science, 29(11)(2008), 1593-1596.

[7]

D.J.Connor, Aluminium extraction from non bauxitic materials, (Sydney: Aluminium-Verlag Gmbh, 1988), 230250.

[8]

Zhang Xiong, "Quantative control of C2S transformation", Journal of The Chinese Ceramic Society, 23(6)(1995), 680-684.

When cooling rate was in fast cooling stage (20°Cmin1 ~ öCTC-min"1) the alumina leaching rate decreased obviously. The main reason was the formation of 2CaOAl203#Si02 and 3CaOAl203; the minor reason was the decrease of the selfdisintegrating rate. The two reasons interacted with each other, and decreased the alumina leaching property obviously under this condition. Conclusions (1) The phase component, self-disintegrating property and alumina leaching property of calcium aluminate slag was good when cooling rate was slower than 20°Cmin"1. 2CaOAl203»Si02

244

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PREPARING POLYMERIZED ALUMINUM-IRON CHLORIDE WITH RED MUD LUGuilin1'2, YUHaiyan1, BIShiwen1 School of Materials and Metallurgy; Northeastern University ,Shenyang, Liaoning, 110004, China 2 School of sciences; Shenyang University of Technology, Shenyang, Liaoning, 110870, China Keywords: Red mud, Flocculant, Polymerized aluminum-ferric chloride, Basicity, Hydrolyzing

Abstract A new method is proposed for the preparation of polyaluminium ferric chloride (PAFC) with red mud. The dependence of pH of the PAFC on basicity, is discussed and the difference in the products of the process of red mud hydrolysis from products in the process of sodium hydroxide hydrolysis is studied. The results show that basicity increases with increasing pH, the preferred pH value is 3.5, and basicity is up to 85.55%. Comparing the appearance of PAFC with that of crystalline aluminium chloride, polyaluminium chloride by SEM, can indicate that the distribution of materials is not uniform on the surface of PAFC. Energy Dispersive analysis indicates that the content of aluminum, iron, and oxygen accounts for 60.73% in the white areas, 43.95% in the black areas in the process of red mud hydrolysis; whereas the content of aluminum, iron, and elements in the white areas is 49.74%, 23.45% in black areas in the product from sodium hydroxide hydrolysis.

Experimental Materials and Equipment Waste red mud, hydrochloric acid and sodium hydroxide were used as starting materials. The waste red mud was obtained from Shandong aluminum factory in China. Compositions and content of red mud are shown in Table 1. The equipment is shown in Table 2. Table 1 Compositions and content of red mud Composition

Alumina

Iron oxide

Silica

Titania

Others

Content(%)

20.54

26.73

25.21

1.85

25.67

Table 2 names and types of equipment

Introduction Red mud is a waste byproduct generated during the production of aluminafrombauxite. Amounts of red mud vary with the different quality and processing of ores containing aluminum. If bauxite is used as starting raw materials, the product of per ton alumina will result in approximately 1 ton of dry red mud, and will produce about 7 tons dry red mud if nephelite is used as the starting raw material. Today, worldwide production of red mud is estimated to be about 150 million tons each year, and production of red mud is more than 4 million tons in Chinail]. So red mud occupies largeland areas, and pollutes the environment12"51. In order to solve the above problems, various means of utilization of red mud have been suggested by worldwide metallurgy workers, for example, red mud was used as catalyst16"91, plastics padding110"111, cement[12], production of castings1131, glass-ceramics1 415] and so on. Red mud is predominantly a mixture of oxides of iron, aluminium and titanium with relatively smaller amounts of silica and magnesia. Utilizing the components of aluminium and iron in red mud to prepare flocculant-polymerized aluminum-ferric chloride (PAFC) has been reported[161. The processes is that red mud reacts with Hydrochloric acid, the mixture is filtered, the filtrate is neutralized and hydrolyzed with sodium hydroxide. This process will not only consume a lot of sodium hydroxide, but also the utilization rate of hydrochloric acid is very low and content of byproducts (mostly NaCl) is quite high. Because red mud contains iron oxide, alumina, sodium aluminate and other alkaline substances which can be used to react with aluminum chloride, iron chloride and free hydrochloric acid, replacing sodium hydroxide with red mud to neutralize and hydrolyze the filtrate will save a lot of sodium hydroxide, increase the utilization rate of hydrochloric acid and purity of the products, and cut the cost of production.

Names

types

scanning electron microscope

Hitachi S-3400N

Synthetic Principles of PAFC Alumina(Al203) and Iron oxide(Fe203) in waste red mud react with hydrochloric acid to form [A1(H20)6]C13, and [Fe(H20)6]Cl3. Fe203+6HC1+9H20 ►2[Fe(H20)6]Cl3 (1) A1203+6HC1+9H20 ► 2[A1(H20)6]C13 (2) The addition of red mud (or alkali) to the filtrate will increase the pH value of system, when the pH value reachs some limit, [A1(H20)6]3+ and [Fe(H20)6]3+ will hydrolyze to form H+and a series of coordinate ions. The equations are as follow: [A1(H20)6]3+ - 2*L~ _

[Al(OH)(H20)5]2+ + H+

2+

[Al(OH)(H 2 0) 5 ] -^~- [Al(OH)2(H20)4]++H+

(3) (4)

[Al(OH)2(H20)4]+ - 2ST- [Al(OH)3(H20)3] + H+

(5)

[Fe(H 2 0) 6 ] 3 + -^~- [Fe(OH)(H20)5]2++H+ [Fe(OH)(H20)5]2+-™\- [Fe(OH)2(H20)4]+ + H+ [Fe(OH)2(H20)4]+ -22:- [Fe(OH)3(H20)3] + H+

(6) (7) (8)

With the increase of concentration of OH", coordinated water hydrolyzes and substances formed will polymerize. Bridge will form between OH" and OH", [A1(H20)6]C13 and [Fe(H20)6]Cl3 gradually are converted into dipolymer, tripolymer, finally high molecular weight species PAFC [Al2(OH)nCl6-n]m, [Fe2(OH)nCl6-n]m. Synthesis Processes of PAFC Hydrochloric acid was added to waste red mud. It reacted with waste red mud for a period of time, then the mixture of liquid and

245

solid was filtered and washed, filter liquor and filter cake were obtained. The filter cake then reacted with hydrochloric acid, and was filtered and washed once again. Both filter liquors were mixed, and neutralized and hydrolyzed with either waste red mud or sodium hydroxide. If waste red mud is used to neutralize and hydrolyze the filter liquors, the mixture of liquid and solid will be filtered and filter liquor obtained will be concentrated and dried, and PAFC will be obtained; If sodium hydroxide is used to neutralize and hydrolyze the filter liquors, the solution obtained will be concentrated and dried, and PAFC will be obtained. The production process of red mud neutralization and hydrolysis are shown in figure 1.

«—H

red mud

CD 3—content of effective composition (%) ( calculated as it was AI2O3) 0.01699—Mass (g) of A1 2 0 3 equivalent with 1.0 ml NaOH standard solution [C (NaOH)=1.000mol/L] Results and Discussion Effect of P H Value of Solution on Basicity of PAFC Different volumes of sodium hydroxide solution were added to the filter liquors obtained when red mud reacted with hydrochloric acid. The pH value of the filter liquors was adjusted to 1.5, 2.0, 2.5, 3.0, 3.5 respectively. The alumina and iron oxide content was determined for each sample, and then basicity determined. The results of the experiments are shown in Table 3. Table 3 Effect of pH on basicity of production

hydrochloric acid

i—^ 1 filter and washing iv filter liquor

1r

+—\

filter cake

hydrochloric acid

filter arid was]ling

Ψ

^r

_2 r filter cake

—► red mud

filter liquor

1 filter

x— 1

1f filter cake

i

r

filter liquor

drying

1 products

Fig. 1 Production process of neutralized red mud Determining the Basicity of PAFC One third of the mol ratio of OH" and Al3+ in a sample is called the basicity. Generally, the effects of flocculation increase with the increase of basicity. Firstly, potassium dichromate titration was used to determine content of iron oxide ω i, and EDTA titration was used to determine content of alumina ω 2 . Alumina was used to calculate effective content ω 3, ω 3 = ω 2 +0.6384ω i, 0.6384 is a factor in which iron oxide was converted into alumina. Then quantitative hydrochloric acid solution was added to the sample of PAFC, sodium fluoride was used to shelter Al3+, sodium hydroxide standard solution was used to titrate above solution. Use equivalent distilled water(no carbon oxide ) and hydrochloric acid as blank sample and make the same experiment. The formula of basicity W4 is as follow:

(Vo- V)xCx0.01699 mco^

X 100%

basicity (%)

1.5

5.82

2.0

72.27

2.5

76.35

3.0

80.89

3.5

85.55

When pH value of solution is between 1.5 and 3.5, the basicity of products increases with the increase in pH of the solution, but for pH value >3.5, the solution will change into a gel. If the pH value of the solution is low, most of the base will be used to neutralize free hydrochloric acid, and the hydrolyzing degree of A1C13 and FeClß should be low, as the contents of hydroxyls are low. Thus basicity of the product is low when the pH value of solution is less than 1.5. When the pH value of solution is more than 2.0, then some of base reacts with A1C13 and FeCl3, and some of chloride in AICI3 and FeCl3 is changed into hydroxyls. So basicity of the products increases with the increase of alkalinity. When the pH of the solution is more than 3.5, A1C13 and FeCl3 will be changed to deposits of Al(OH)3 and Fe(OH)3.

w

W4=

pH

Comparison of PAFC with Aluminum Chloride and Polymeric Aluminum Chloride In order to determine whether the synthetic products are the target PAFC or not and, simultaneously, determine the composition of the products, crystalline aluminum chloride and polymeric aluminum chloride were chosen to compare with the synthetic products. Their appearance and composition were investigated using scanning electron microscope and energy dispersive analysis. The experimental results are shown in figure 2, figure 3 and figure 4. Conditions for the preparation of the synthetic products were that the liquid-solid ratio of hydrochloric acid to red mud is 4:1, the concentration of hydrochloric acid is 6 mol/L, reaction time is 60 minutes, reaction temperature is about 109°C, the extraction is the second extracting (see fig. 1, pH value of the filter liquors was adjusted to 3.5 with sodium hydroxide solution. The synthetic product- PAFC was obtained after the solution was dried.

(9)

Vo—volume (ml)of sodium hydroxide standard solution consumed by blank sample V— volume (ml)of sodium hydroxide standard solution consumed by sample of PAFC C—concentration of sodium hydroxide standard solution(mol/L) m— mass of sample PAFC(g)

246

Ì >

*

*i «^ .

1%

4

Λ

4i

|»

►

#* *, * ^ * >

/^^^^B m 1

* m

I

:

,, ■·■

S3400n 15 OkV 9 7mm x2 00k SE 6 /18/20C 8 12:3B'

(..

;

'

m

20.0um

(a) SEM of PAFC (a) SEM of crystal aluminium chloride «.s -

39.62 32.42 27.96

OJT AIE CX

-

S.1

OK NaK AUL OK CaK FeK

Al

55.44 26.90 17.66

19.56 10.31 11.74 38.87 01.0S 18.44

34.34 12.60 12.22 30 80 00.76 09.28

0

a

KG*

3.4

-

1.1

-

0.«

-

JL

i.

j

V

~

Î

""""[""■"■""■r

v-ynmnnn

IM

7.M

«.M

S.00

IM

(b) Energy spectrum of crystal aluminium chloride Fig.2 Crystal aluminium chloride

(b) Energy spectrum of white areas of PAFC OK NaK ΛΙΚ OK CaK FeK

11.99 25.06 03.96 51.04 0046 07.50

20981 30.52 04.11 40.31 0032 03.76

(c) Energy spectrum of black areas of PAFC Fig.4. PAFC of pH=3.5 with sodium hydroxide neutralisation. (a) SEM of polyaluminium chloride

According to figure 2(a), figure 3(a) and figure 4(a), the appearance of the synthetic product is much different from that of crystal aluminum chloride and similar to that of polymeric aluminum chloride. The surface of the crystalline aluminum chloride is smooth, and the distribution of materials is uniform. Distribution of materials is not uniform on the surface of PAC and PAFC. The white and black areas occur on their surface. According to the energy spectrum of white areas in figure 2(b), figure 3(b )and figure 4(b), crystalline aluminum chloride contains only Al, Cl, and O; polymeric aluminum chloride contains traces of Fe, and Na in addition to Al, Cl, and O; PAFC contains high levels of Fe, Na in addition to Al, Cl, O. The appearances and compositions of white areas show that the white areas should be PAFC. Comparing compositions of white areas in figure 4(b) with that of black areas in figure 4(c), the content of Fe, and Al in the white areas are more than that of Fe and Al in the black particles. Fe and Al in

(b) Energy spectrum of polyaluminium chloride Fig.3 Polyaluminium chloride

247

According to figure 5(a), the distribution of materials of PAFC obtained in process of red mud neutralizing is more uniform than that in process of sodium hydroxide solution in the figure 4(a). The white areas of materials are quite large in figure 5(a). Comparing figure 5(b) with figure 4(b), content of Al, Fe and O in PAFC obtained in process of red mud neutralizing is higher than that in process of sodium hydroxide neutralizing. The content of Al, Fe and O is 60.73% in the former, and 49.74% in the latter. The content of Na is 0.95% in the former, and 10.31% in the latter. The results indicate that the content of PAFC obtained in process of red mud neutralization is higher than that in the process of sodium hydroxide neutralization in the chosen white areas. The content of by-product NaCl is lower in the former than in the latter in the white areas. Comparing figure 5(c) with figure 4(c), the content of Al, Fe and O in PAFC obtained in the process of red mud neutralization is higher than that in process of sodium hydroxide neutralization in the black areas.The content of Al, Fe and O is 43.95% in the former, and 23.45% in the latter. The content of Na is 3.78% in the area analysed in figure 5(c), and 25.06% in figure 4(c), which also indicates that content of PAFC obtained in process of red mud neutralization is higher than that in process of sodium hydroxide neutralization in the chosen black areas, and the content of by-product NaCl is lower in the former than in the latter in the darker areas.

the white particles account for 30.18%, but is only 11.46% in black particles. The effect of Neutralizing approach and Hydrolysis on the Appearance and Composition of PAFC Red mud was leached with Hydrochloric acid, the filter liquor was neutralized and hydrolyzed to pH=3.5 with red mud. The results of the experiment are shown infigure5.

III,. "Ill &:, 'sa%.

::

* ;ii :

:

:"■'

;

'

Π5 ] (a) SEM of PAFC

The composition of the material in figure 5(c) can indicate whether compositions of the black areas in figure 5(a) are PAFC or not. The proportion of Fe and Al is 29.6%. Subtracting the proportion of the atoms of Na from that of atoms of Cl gives 38.28%, which indicates each Al (or Fe ) can bond with 1.3 Cl atoms. The residual atoms (Cl ) were hydrolyzed into hydroxyl (-OH). The analysis above shows that compositions of the black areas infigure5(c) are PAFC, in addition to a little by-product NaCl.

6.2 -

I2S

3.f

OK NaK AIK CIK FeK

-

C»

28.68 00.95 Î5.67 38.31 16.38

47.32 01.09 15.33 28.52 07.74

KCi« 2.5 -

1.2 -

Al

II IH 1.00

Conclusions

I

(1) Red mud can be used as a starting material to prepare PAFC. (2) The basicity of PAFC increases with increasing pH, and the preferred pH value is 3.5. (3) Comparing the appearance of PAFC with that of crystalline aluminium chloride and polyaluminium chloride by SEM can indicate that the distribution of materials is not uniform on the surface of PAFC. The Energy Dispersive spectrum can be used to analyze the composition of PAFC. (4) The Energy spectrum indicates that the content of PAFC obtained in the process of red mud neutralization is higher than that in the process of sodium hydroxide neutralization; the content of by-product NaCl is lower in the former than in the latter. Process of red mud neutralizing is better than process of sodium hydroxide neutralizing.

Fe 2.00

3.00

4.00

S.00

6.00

7.00

8.00

9.00

10.00

11.00

12.0«

(b) Energy spectrum of white areas of PAFC

4.2 H

l*g?g?»i] OK 3233 03.78 NaK 23.35 AIK 52.27 CIK FeK 082?

22.52 04.SÖ 25.2S 43.08 04.32

KCf*

References 1. JIANG Yijiao, NING Ping, "An overview of comprehensive utilization of red mud from aluminum production", Environmental Science and Technology, (2003), 26(1), 40—42. 2. R. A. A. Blackman, K. W. Wilson "Effects of red mud on marine animals", Marine Pollution Bulletin, (1973) 4(11) 169-171. 3. S. Baseden, D. Grey, "Environmental study of the disposal of red mud waste", Marine Pollution Bulletin, (1976), 7(1), 4-7.

(c) Energy spectrum of black areas of PAFC Fig.5 PAFC with pH value 3.5 in process of red mud neutralizing

248

4. Shaobin Wang, H.M. Ang, M.O. Tadé, "Novel applications of red mud as coagulant, adsorbent and catalyst for environmentally benign processes",_Chemosphere, (2008), 72(11), 16211635. 5. Ekrem Kalkan, "Utilization of red mud as a stabilization material for the preparation of clay liners", Engineering Geology, (2006), 87( 3-4), 220-229. 6. Cakici A I, Yanik J, Uçar S, Karayildirim T, Anil H, "Utilization of red mud as catalyst in conversion of waste oil and waste plastics to fuel", Journal of Material Cycles and Waste Management, (2004), 6(1), 20—26. 7. Juan J. Llano, Roberto Rosai, Herminio S astre, Fernando V. Diez, "Catalytic hydrogénation of anthracene oil with red mud", Fuel, (1994) 73(5), 688-694. 8. Jale Yanik, Md. Azhar Uddin, Kazuo Ikeuchi, Yusaku Sakata, "The catalytic effect of Red Mud on the degradation of poly (vinyl chloride) containing polymer mixture into fuel oil", Polymer Degradation and Stability, (2001), 73( 2), 335-346. 9. Jorge Alvarez, Salvador Ordonez, Roberto Rosai, Herminio Sastre, Fernando V. Diez, "A new method for enhancing the performance of red mud as a hydrogénation catalyst", Applied Catalysis A: General, (1999), 180( 1-2), 399-409. 10. Singh B, Gupta M. "Surface treatment of red mud and its influence on the properties of particulate-filled polyester composites", Bulletin of Materials Science, 1995, 18 (5) : 603-621. 11. Soo-Jin Park, Dong-Il Seo, ChangwoonNah, "Effect of Acidic Surface Treatment of Red Mud on Mechanical Interfacial Properties of Epoxy/Red Mud Nanocomposites",_Journal of Colloid and Interface Science, (2002), 251(1), 225-229. 12. Maneesh Singh, S. N. Upadhayay, P. M. Prasad, "Preparation of special cementsfromred mud", Waste Management, (1996), 16(8), 665-670. 13. Nilza Justiz-Smith, Vernon E. Buchanan, Gossett Olive. "The potential application of red mud in the production of castings", Materials Science and Engineering A, (2006) 420(1-2), 250253. 14. Jiakuan Yang, Dudu Zhang, Jian Hou, Baoping He, Bo Xiao, "Preparation of glass-ceramics from red mud in the aluminium industries", Ceramics International, (2008), 34(1), 125-130. 15. Vincenzo M. Sglavo, Stefano Maurina, Alexia Conci, Antonio Salviati, Giovanni Carturan, Giorgio Cocco, "Bauxite 'red mud' in the ceramic industry. Part 2: production of clay-based ceramics", Journal of the European Ceramic Society, (2000), 20(3), 245-252. 16. ZHANG Lingfen, HUANG Jianfang, MO Yazh, "A study on the preparation of polymerized aluminum ferrum chloride from red mud"._Water Purification Technology 1998,(03):2~4.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Adsorption of Polyethylene Glycol at the Interface of Dicalcium Silicate - Sodium aluminate solution Haiyan YU1, Xiaolin PAN1, Zhongke LU2, Tingting DING1 1

School of Materials Metallurgy, Northeastern University; Shenyang, Liaoning, 110004, China 2

Shandong Branch, Chalco; Zibo, Shandong, 255065, China

Keywords: Polyethylene glycol; ß-CaOSi02; Sodium aluminate solution; Adsorption thermodynamics [3-5]. However, the basic theories, such as the

Abstract

models of adsorption, the laws at the interface of sodium Surfactants, such as polyethylene glycol (PEG), are widely used

aluminate solution and ß-CaOSi02, and the effect of adsorption

to inhibit the secondary reaction of dicalcium silicate (ß-

on the solid-liquid interface were seldom studied. Consequently,

CaOSi02) in sodium aluminate solution. The adsorption isotherm

according to the development of Gibbs adsorption isothermal

of PEG on ß-CaOSi02 in sodium aluminate solution at 80 °C

formula on the solid-liquid interface [6-10], the adsorption laws of

was researched in this paper. It was found that the type of surface

PEG in clinker-sodium aluminate solution systems and the

adsorption of PEG is saturated adsorption, and the adsorption

mechanisms of PEG inhibiting the secondary reaction were first

behavior belongs to "S"-type, according with the multi-molecular

studied in this article.

layer adsorption model of the Freundlich Equation. The surface coverage of PEG close to saturated adsorption derived from the

Experimental procedures

2

specific surface area of ß-CaOSi02 is 0.723 mol/μπι . The relation curves between the surface pressure and the molecular

Dicalcium silicate was prepared by pure calcium oxide and silica.

area of adsorption film were obtained according to the adsorption

Calcium oxide and silica were first mixed with the molar ratio of

results at the solid-liquid interface by Gibbs formula. The

2 to 1, grinded 2 hours in a mortar, and then calcined at 1350 °C

variation of interfacial energy caused by adsorption, as well as the

followed by quenching in the air.

relationship between the relation curves and the type of adsorption

Sodium aluminate solution with caustic ratio (i.e. molar ratio of

was also discussed.

Na20 to A1203) of 1.30 was prepared by the red mud lotion from the Shanxi Branch of the China Aluminum Company, industrial sodium hydroxide and aluminum hydroxide, and analytically pure

Introduction

sodium carbonate. The concentrations of Na20 and Na2C03 of sodium aluminate solution were 36 g/L and 15 g/L, respectively.

In the sintering process of alumina production, there is about 30%

PEG was added to sodium aluminate solution with the

dicalcium silicate (ß-CaOSi02) in the clinker. During the

concentration of ß-CaOSi02 75 g/L. The concentrations of PEG

leaching of clinker, as ß-CaOSi02 reacts with sodium aluminate

in sodium aluminate solution were 0, 80 mg/L, 120 mg/L, 160

solution, both A1203 and Na20 re-enter into the red mud, resulting

mg/L, 200 mg/L, 240mg/L, respectively. The mixed solutions

in the loss of A1203 and Na20 [1]. It was found [2] that additions

were stirred magnetically (600 r/min) for 1 hours in 80 °C water

of surfactants, such as polyethylene glycol (PEG), can inhibit the

bath followed by centrifugal separation. 0.5 ml supernatant fluid

secondary reaction of ß-CaOSi02 with sodium aluminate

was taken out to determine the content of PEG by the method of

solution, which increases the leaching rate of A1203.

spectrophotometry using a 722S Spectrophotometer.

Much of work was done on the additives for inhibiting the

The adsorbed amount per unit weight of adsorbent at equilibrium,

secondary reaction, which was mainly concerned on the selection

Qe, can be expressed as follows:

and application of additives as well as the adsorption kinetics and

251

(1)

Qe=(C0-Ce)/W0

The value

of jirdin C with

different concentrations can be

obtained by graphic integration, and then the value of π with

where C 0 and Ce is the initial concentration and the residual

different adsorbed amounts can be calculated. At very low

concentration of additives (mg/L) in the solution, respectively,

concentration, the relationship between Γ and In c is considered as

and W 0 is the concentration of adsorbent (g/L) in the solution.

a line. Accordingly, the area of each adsorbed molecule (σ) at different

Gibbs Adsorption Isotheral Formula

amounts of Γ is given by the Avogadro constant (NA) as: According to the interface Gibbs formula in the ideal solution

σ = 1/(ΠνΑ)

[11], the basic form of surface tension, γ, under constant

(6)

temperature and pressure is given by: Results and Discussion (2) i=1 A .

Adsorption isotherm

i=1

The adsorption isotherm of PEG on ß-CaOSi02 in sodium

where ^ is the surface molar quantity of i component, As is the

aluminate solution is shown in Fig. 1. The type of adsorption

surface area, μϊ is the chemical potential of i component, T[ is the

isotherm belongs to "S"-type, which indicates that the adsorption

excess amount of i component per unit surface (i.e. the surface

of PEG on 2CaOSi0 2 accords with the Freundlich multi-layer

adsorbed amount or surface concentration (mol/m2)).

model, and follows the Freundlich Equation as expressed in Eq.

As the interfacial tension numerically equals to the interfacial

(7):

pressure (π), the relationship is d π = -d γ. lnß,=lnk +

Thus, άπ = Υ^Γ,RT dinCi

InC.

(7)

(3) where k and n are constants.

For binary systems, if Group 1 is identified as the major component, Γι (1) =0. The adsorbed amount of Group 2 is Γ2(1). Then, Eq. (3) can be written as:

-ày=RTYf

dlnc2

(4)

For the adsorption in the dilute solution, it is reasonable to consider the surface excess amount Γ2(1) as the actually measured amount (apparent adsorbed amount), so Γ 2 (1) is replaced by Γ and c 2 is replaced by c in Eq. (3). For the solid adsorption from liquid, lô

Eq. (3) is integrated on both sides:

π=

γ80-γ$ι=1ΐτ\ΐτά\ηα

2Ô 30 40 50 Equilibmim concentration (nig/L)

Fig. 1 Adsorption isotherm of PEG on ß-CaOSi0 2 at 80 °C

(5)

where γ8ο and YSL is the solid-liquid interface free energy before

The relationship between the residual concentration (Ce) and the

and after adsorbing the solute, respectively.

value of Q/Qe is shown in Fig. 2. The curve can be fitted by the Freundlich Equation as Eq. (8):

252

l n ß , =1.242 In Ce -4.184

interface [14, 15]. Therefore, the adsorption of polyethylene

(8)

glycol is mainly physical adsorption, which is also confirmed by IR detection. In addition, the curve in Fig. 3 changes at the point of 1.09 mg/g, which is approximately considered as the adsorption 1.0

capacity of the covered monolayer.

♦ JS

0.5

3 c

1

R 2 =0.963

y

0.0

r

H

♦

-

yS

-0.5 -1.0 .1 R

2.0

S

1

1

2.5

3.0

i

3.5

·

4.0

4.5

InCe

Fig. 2 Freundlich isotherm of PEG on ß-CaOSi02 0

As shown in Fig. 2, the experimental data can be well simulated

5 σ öuii2/raole) 10

15

Fig. 3 Relationship between the interfacial pressure and

by the adsorption isotherm. The adsorption of PEG at the interface

molecular area

of dicalcium silicate-sodium aluminate solution belongs to "S"type, having the characteristic of multi-layer adsorption.

When the adsorption capacity is 2.45mg/g, the surface coverage is

Adsorption film on B-CaO-SiO?

calculated to be 2.24, which is larger than 1, demonstrating that

The π-σ curve of the interfacial film on ß-CaOSi02 can be

the adsorption of PEG is a multi-molecular layer. Meanwhile, the

obtained from the adsorption results by using Eq. (5) and (6). As

single molecular area is 1.383 μηι2/ιηο1, and the corresponding

shown in Fig. 3, the shape of π-σ curve is similar to the adsorption

coverage is about 0.723 mol/ μπι2.

of insolubles in the liquid surface [12]. Phase transitions take place at the interface of dicalcium silicate-sodium aluminate

Conclusions

solution, which is in accordance with the formation of surface micelles. It is known that polyol compounds are typically non-

It was found in the research that the adsorption of PEG at the

ionic surfactants, with a structural formula of H-(0-CH2-CH2)n-

interface of dicalcium silicate-sodium aluminate solution is a

OH. When the concentration of polymer is lower, the molecules

saturated multi-layer adsorption and the type of adsorption is "S"-

of PEG are adsorbed at the interface of the ß-CaOSi02 by -OH

type in accordance with the Freundlich Equation. The adsorption

groups [13], and no significant changes of interfacial tension

of PEG on ß-CaOSi02 in sodium aluminate solution belongs to

happen.

physical adsorption. The π-σ curve can be drawn according to the

As the concentration of polymer increases, the surface micelles

Gibbs formula, which is available to obtain useful information

(or called semi-micelles) are formed between PEG molecules and

from the shape of the π-σ curve.

adsorbed molecules by the hydrophobic effects of C-H bonds, resulting in a sharp increase of adsorption. At the same time, the

References

interface pressure changes rapidly, while the correspondingly molecular area only changes in a cell range, which is correlative

[1] BI S W, YU H Y. Production technology of alumina [M].

to the formation of surface micelles. The reason of the sudden rise

Beijing: Chemical Industry Press, 2006: 250-251.

in the adsorption capacity is the formation of hemi-micelles at the

253

[2] ZHANG C Z. Research on additive for inhibiting secondary

Foundation item: Project (50974036) supported by the National

reaction in clinker leaching process [D]. Northeastern

Natural Science Foundation of China: Project supported by

University. 2008: 49-50.

Aluminium Corporation of China

[3] REN G K. Technical research of secondary reaction inhibitor and sweetening process [J]. Light Metals, 2008(5): 16-18. [4] LI T C. Technical Research of Secondary Reaction Inhibitor and it's Sweetening Process [J]. Nonferrous Metals, 2002(1): 26-28. [5] SHEU E Y, STORM D A, SHIELDS M B. Adsorption kinetics of asphaltenes at toluene/acid solution interface [J]. Fuel, 1995,74:1475-1479. [6] OLOF S, TOBIAS H, TORGNY S, THOMAS A. Adsorption of delmopinol at the solid/liquid interface - The role of the acid-base equilibrium [J]. Journal of Colloid and Interface Science, 2010, 350: 275-281. [7] PAUL J, JOHN R. The adsorption of a polysaccharide at the talc-aqueous solution interface [J]. Colloids and Surfaces A: Physicochemical and Engineering Aspects, 1998, 139: 27-40. [8] WU Z S, LI C. Kinetics and thermodynamics of b-carotene and chlorophyll adsorption onto acid-activated bentonite from Xinjiang in xylene solution [J]. Journal of Hazardous Materials, 2009, 171: 582-587. [9] SOMASUNDARAN P, KRISHNAKUMAR S. Adsorption of surfactants and polymers at the solid-liquid interface [J]. Colloids and Surfaces A: Physicochemical and Engineering Aspects, 1997, 123-124: 491-513. [10] WANG Y J, SONG Z F. Spectroscopy and Chromatography [M]. Beijing: Peking University Press, 1995: 225-239. [11] HU Y. Modern Chemical and Engineering Thermodynamics [M]. Shanghai: Shanghai Science and Technology Press, 1994: 239-240. [12] HARKINS W D. The Physical Chemistry of Surface Films[M]. New York Reinhold, 1952: Chap2. [13] CHEN H W, ZHOU Z K. The Analyses on the Impact of the clinker

stripping

conditions

on

the

Secondary

Reaction[J].Light Metal, 2001(8):21-24. [14] SCAMEHORN J F, SCHECHTER R S, WADE W H. Journal of Colloid Interface Science, 1982, 85:463-467. [15] ZHU B Y, ZHAO X L, GU T R. Journal of Chemistry Society Faraday Transactions, 1988, 84 (ll):3951-3958.

254

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PRODUCTION OF HEMATITE ORE FROM RED MUD P.K.N.Raghavan1, N K. Kshatriya1 and Katarzyna Wawrynink2 bharat Aluminium Company Limited, Korba (CG) India, 495-684 2

Warsaw University of Technology, Poland

Keywords: red mud, zerosite, hematite, thenardite universally as solid residue. The Jarosite generated in a Zinc refinery is normally associated with free sulfuric acid and ferric iron. The Jarosite mainly contains iron,sulfur,zinc, lead and calcium. Jarosite released from industrial process is complex and its quality and quantity make the of safe disposal complex. The most widely used method for disposal of Jarosite is the Jarofix Process.In this Jarofix process the Jarosite is treated with hydrated lime and then stabilized with ordinary Portland cement before dumping in a pond. At present the long term stability of Jarofix material remains unsatisfactory and when the material does break down it is likely that the trace metals will be released into the surrounding environment with potentially adverse consequences. It would be advantageous if red mud and Jarosite could be transformed into other forms less harmful to the environment. In the present work an attempt is made to evaluate the utilization potential in developing marketable commodities using red mud and Jarosite for safe and effective waste management.

Abstract Red Mud generated as a waste material in the Bayer process is a potential source of metal values. Red Mud is highly caustic residue (pH > 13).Approximately 1-2 tonns of red mud are being generated per MT of alumina produced. Disposal and storage problems will prevail due to the leachable soda content and the volume of material to be handled. Jarosite is a hazardous waste of Zinc refinery and is both difficult and costly to manage and to dispose safely. Jarosite is normally associated with sulfuric acid and Ferric iron in the waste produces additional acid on hydrolysis. Red Mud is source of hydroxyl and sodium ions whereas Jarosite is source of sulphate and hydrogen ions. Reaction between red mud and Jarosite in water will therefore produce a liquid phase containing sodium sulfate and a solid phase dominated by iron hydroxide. The liquid phase was evaporatively concentrated to precipitate salt and the solid phase was calcined to produce a ferruginous solid that could be used as an iron ore. The reaction of red mud and Jarosite produces two marketable commodities (Gypsum) & (Thenardite). In this paper an attempt is made to evaluate procedures for the preparation of iron ore and marketable byproduct salts by reacting the Balco's Red Mud and Jarosite from Hindustan Zinc of Vedanta Group.

Materials and Method Sample collection: The solid residue, Jarosite was obtained from our group company Hindustan Zinc Limited ( HZL) and Red Mud from Bharat Aluminium Co.Ltd.,Korba (India).The samples were oven dried (105-1100C) separately, sieved through 150 micron sieve. For characterization sampling was done from the oven dried sample adopting the conning and quartering method.

Introduction For every tonne of alumina produced in the Bayer processing of bauxite, 1 to 2 tonnes of red mud generated depending upon the quality and grade of the bauxite and, to some extent, the processing conditions. Red mud consists of almost all the elements present in the parent bauxite from which it is derived, many of them in enriched form, in addition to sodium compounds formed by reacting with caustic soda used for digestion, and calcium compounds if lime is added during the processing stages. It is infact a complex mixture of various oxides like Fe203, Ti02, A1203, Si02, Na20 occurring in major amounts. Disposal and utilization of red mud are common problems for all alumina refineries in the world. Industries located far the coastal region (as is the case with alumina refineries in India) face significant disposal problems. To dispose off a huge volume requires enormous civil, mechanical and process efforts causing huge expenditure. All alumina producers are, therefore keen to make effective utilization of red mud. During metallic Zinc extraction from Zinc concentrates huge quantity of Jarosite are released

Characterization of chemical & Mineralogical properties: The major constitutes of red mud were analyzed by XRF PW 2440 Phlips, Netherland. Heavy metals were detected by Atomic Absorption Spectrophotometer (AAS), SOLAAR S 4AA, and Thermo Elemental, USA with SOLAAR software package. The mineralogical analysis were carried out using PANlytical X'Pert Cubix Pro series diffractometer equipped with copper target tube,X'celerator detector and operated at 40kV and 30 mA. Diffraction data were analyzed using PANlytical X'Pert High Score plus Version 2.1.

255

The process for recovery of Hematite and Thenardite by reacting with Red mud and Jarosite is summarized in Fig. 1. Results

W&FER

SOLUTION

SOLID

EVAPORATE

CALCINE

The various major constituents present in Jarosite are given in Table 2 while that of the heavy metals such as Cr, Ni, Cu and Co in Jarosite are shown in Table 3. The % composition of major constituents of Red Mud and Calcined material are shown in Table 4. X- Ray diffractograms of Jarosite, Red Mud, Calcined solid (Hematite Iron) and Thenardite are shown in Fig. 2, Fig.3, Fig. 4 and Fig. 5 respectively.

800 \

The major mineral phase of Jarosite is potassium iron sulfate hydroxide (KFe3(S04)2(OH)6) and iron sulfate hydrate (2Fe203S03.5H20) (Fig.2).The Main crystalline phase of Red Mud are : Hematite (Fe203), Goethite (FeOOH), Gibbsite (Al(OH)3), Boehmite (AlOOH), Cancrinite (3NaAlSi04.NaOH), Quartz (Si02), Anatase (Ti02) etc (Fig.3).

HEMATITE SOLUTION

SOLID

FeA

ZYPSUM CHILLED

THENARDITE Fig.l. Schematic representation of recovery of i»matite from red mud

Table2.Chemical Properties of Jarosite (in %)

Experimental The process for recovery of Hematite from red mud by reacting with Jarosite is briefly outlined below: (i) A mixture of Jarosite along with red mud and water were mixed at different ratio at ambient temperature (280C) with magnetic stirring. The ratio of Jarosite, Red Mud and Water are given in the following Table 1. Table. 1: Various Ratios of Jarosite to Red Mud Used in the Experiment SN

Jarosite: Red

Mud

Water

pH

(weight ratio)

(ml)

1

1:1

300

6.5

2

2:1

300

6.0

3

4:1

300

6.0

To investigate the possible use of these neutralization reactions two parts of Jarosite and one part of red mud were reacted for 24 hours at ambient temperature. (ii)The neutralized slurry was filtered and washed with water. (iii) The filtered solution was concentrated by evaporation and two minerals (Gypsum and Thenardite) could be easily separated Gypsum is much less soluble than Thenardite and can be precipitated first and removed. (iv) The solid phase was calcined at 8000C for 2 hours to produce a ferruginous feed stock for blast furnaces producing pig iron.

256

Si

2.45

Fe

24.54

Al

2.12

Zn

4.47

Ca

4.13

S

11.68

Pb

0.52

K

0.30

Na

0.35

Mg

0.28

Mg

0.16

um

s asm

Cr

165

Ni

81

Cu

980

Co

42

Table3· Heavy Metals in Jarosite (in ppm) The Chemical Composition of BALCO's Red Mud is given in Table-4. Constituents %

Red Mud

Calcined Material

m]

I . .11

u;y wUäJi.... ...

T T TT.riTTTT r

10

20

Ί âû

40

5

" ' T 0Û

0611KM" [*21iieB3

taras

Fig.2. X-Ray diffractogram of Jarosite

Fe203

38.38

53.13

Si02

7.21

7.10

Ti02

17.37

6.87

A1203

20.61

Trace

Na20

6.39

0.26

P205

0.35

0.15

LOI

8.03

2.19

Table4. Chemical Properties of Red Mud and

L-.n ;

Calcined Material (in %) Couts

30

40 Posi Don [ïïheta] Couts .

QMZ

Fig.3. X-Ray diffractogram of Red Mud

Fig.4. X-Ray diffractogram of Solid Calcined Material (Hematite)

257

Chief Executive Officer and Whole Time Director, Mr. Gujan Gupta during the progress of this work. The authors would like to thank the Management of Vedanta Resources for allowing us to publish this paper.

Counts ■i prepgrd

References 1. Gaines et al, Dana's New Mineralogy Eighth Edition Wiley, 1997. 2. K L Bhat, K A Natarajan,T Ramchandran, Electroleaching of Zinc leach residue, Hydrometallurgy, 1987,18,pp.287-303

jy^uhi'mmekßJ

Π-tüs^i-— 23

30

40

■""!

5C

" T " ' §0

"T1 70

PcsHion f2~hets]

Fig.5. X-Ray diffractogram of Thenardite Conclusions Every Bayer process alumina refineries produce very high volumes of Red Mud, disposal of which in open fields has the potential to cause great environmental damage. It not only causes harm to vegetative sources, but also extends great health risks to both humans and animals. Jarosite wastes generated from the hydrometallurgical processes contain significant quantities of compounds of sulfur, lead, zinc, iron etc. The presence of toxic substances in Red Mud and Jarosite make these waste hazardous and causes serious problems in for their disposal. Due to weathering there is a release of toxic elements in soluble form which can ultimately contaminate the soil, ground water and aquatic life due to improper management of such hazardous wastes. The present work shows that it is possible to efficiently and cost effectively neutralize red mud from BALCO Alumina Refinery by mixing with Jarosite waste from Hindustan Zinc refinery (HZL). Neutralization of red mud using Jarosite will reduce both the costs and the risks associated with the long term storage unneutralized red mud. This neutralized material can be stored without liability. Neutralization process will provide a variety of cost and operational benefits to BALCO and HZL including (1) Helping conserve fresh water (2) Creating a soil medium for revegetation (3) Reducing future and stock piled caustic red mud and Jarosite (4) Potentially resulting in the production of marketable sodium salts (5) Creating a raw material hematite iron ore. Acknowledgement The authors acknowledge the constant encouragement from our Chief Operating Officer, Mr. Bibhu Prasad Mishra and

258

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

ORGANIZERS

Mohd Mahmood Aluminum Bahrain- Alba Manama, Bahrain Abdulla Ahmed Aluminum Bahrain- Alba Manama, Bahrain Charles Mark Read Bechtel Corporation Montréal, Canada

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

Enviroment- Emissions / Anode Effect I SESSION CHAIR

Robert Baxter Bechtel Corporation Montréal, Canada

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

HF MEASUREMENTS INSIDE AN ALUMINIUM ELECTROLYSIS CELL Karen Sende Ösen1, Thor Anders Aarhaug1, Asbj0rn Solheim1, Egil Skybakmoen1, and Camilla Sommerseth2 2

^INTEF Materials and Chemistry, NO-7465 Trondheim, Norway Department of Materials Science and Engineering, NTNU, NO-7491 Trondheim, Norway Keywords: HF levels, aluminium electrolysis cell

the relative contributions to HF emission, and to locate the main HF generation sites. The industrial measurements have been carried out in the gas channel between individual cells and the main duct system. From such measurements, essential information regarding the importance of certain parameters such as crust integrity has been obtained [4]. Still, the fluoride concentration in the duct represents the average of the conditions at the top of the cell.

Abstract HF emissions to the working atmosphere may still be a problem for the aluminium industry. The objective in the present work was to study how the HF evolution is distributed between feeder holes, other openings in the crust, gases diffusing through the crust, fumes from the secondary alumina residing on top of the crust etc. A movable "gas sniffer" connected to a Tunable Diode Laser was used to measure the HF concentrations at the above mentioned locations. The stationary HF level in an open flaming feeder hole was approximately 9000 ppm, when measured a few cm above the bath surface. In comparison, when the probe was positioned 5-10 cm above a crust area with good integrity, the HF concentration was in the range 5-10 ppm. The results support the notion that most of the HF evolves from open feeder holes and the tapping hole.

Useful and more direct information can be obtained by doing measurements locally on top of the cell. The main purpose of the present work was to gain knowledge about how the hydrogen fluoride evolution is distributed locally between the feeder holes, secondary alumina, openings in the crust, and gases diffusing through the crust as illustrated in Figure 1. To make measurements locally at the top of the cell, a portable sniffer based on online Tunable Diode Laser monitoring was constructed.

Introduction HF emissions to the working atmosphere continues to be a problem, and new challenges concerning emissions to both internal and external atmosphere emerge as a consequence of the ongoing cell technology development. In this situation, it is important to know precisely which factors affect the formation of hydrogen fluoride, and where it evolves. Hydrogen fluoride is generated when fluorides present in the bath or in the vapour phase react with moisture, f AlF,(diss) + H 2 0(g) = 2HF(g) + lAl 2 0 3 (diss) NaAlF4(g) + H 2 0(g) = 2HF(g) + 1 Al 2 0 3 (s) + -|Na3AlF6(s)

Figure 1. Schematic representation of water sources and possible HF-evolution sites (represented by the red arrows) in an aluminium electrolysis cell.

(l)

Design Criteria for "HF-sniffer"

(2)

Based on the data from the literature [5], as well as calculations using the fluoride evolution model by Haupin and Kvande [2], the total fluoride evolution from a rather large modern cell, such as the cells in Hydro's plant at Sunndals0ra (SU4), will be in the order of 30-40 kg F/t Al. Between half and two thirds of this is due to HF. Depending on the gas suction rate, the concentration of HF in the duct will be around 300 ppm (volume), or about 250 mgF/Nm3. This value represents the average situation on top of the cell. The maximum possible gas concentration can be estimated by assuming that all HF is formed in the bath and follows the cell gases out from below the crust. 20 kg HF/t AI then corresponds to about 3.4 vol% HF (24 000 mgF/Nm3) in the

It has long been established that two of the main sources of water are the structural hydroxyl in the primary alumina and the moisture content in the air [1, 2, 3]. A certain amount of HF is also generated from electrochemical oxidation of hydrogen in the anodes. It is also a well known fact that the HF formation rate increases with increasing temperature, decreasing bath ratio, and decreasing alumina concentration [1,2]. Quite a lot of industrial measurements and a few laboratory studies have been carried out during the recent years to quantify

263

in this case, since it turned out that a certain minimum suction rate was necessary to get a good reading.

cell gas coming up from, e.g., an open feeder hole. This can be regarded as the maximum concentration that can be expected. Formation of 20 kg HF/tAl is equivalent with 0.48 wt% water in the alumina, if all water reacts to HF. The total amount of water in the alumina is normally much higher, which indicates that most of the water is not available for reaction to HF. Furthermore, the equilibrium in Eq. (1) is not completely displaced towards the right hand side. Figure 2 shows the concentration of HF and H 2 0 in the cell gases at 960 °C ("normal" bath composition) as a function of the content of water in the alumina. In these calculations, it was assumed that all water is available for reaction with the bath. 1

1

1

1

Figure 3 shows a principal sketch of the set-up. The instruments were placed on a table close to the wall by the emergency exit. It was then out of the way for all motorized traffic, and the magnetic field was moderate. The tunable diode laser (TDL) was a HF and H 2 0 Lasergas II Single Gas monitor from Neo. The PFA tubings (diameter 1/4") between the laser and the probe were approximately 20 m long and flexible. The probe could then be transferred manually from one measuring site to the next. Gas was sucked continuously through the probe and into the laser measuring cell with the help of a pump placed at the end of the line at a rate between 5 and 15 1/min. N2 (g) was purged through the path of the laser beam between the emitter and the receiver outside the actual measuring cell to eliminate the effect of humidity in the ambient air that otherwise would be present. The laser was connected to a computer, and the values recorded every 10th second. During some of these measurements, a thermocouple (Pt-PtRh 10%, type S) was placed at the probe tip. The purpose of this was to document the vertical position of the probe. Figure 4 a shows a picture of the probe tip with the thermocouple, and Figure 4 b shows how the probe was inserted through a hole in the cover.

1

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sample gas

Inert gas into outer circuit

2.5

Tunable diode laser (measuring cell 10 cm)

Inert gas out from outer circuit

"Gas watch" counter for the gas volume that

Water in alumina / wt%

Figure 2. Equilibrium concentration (Eq. (1)) of H 2 0 and HF in the cell gases as a function of the H20content in the alumina, assuming that all water in the alumina is available for reaction with the bath. 11.5 wt% A1F3, 4.5 wt% CaF2, 3.5 wt% A1203, 960 °C, 93 % current efficiency.

Vacuum pump

S

Figure 3. Gas sniffer, as it was used in the 2nd campaign. Improvements from 1st Campaign: New laser (higher upper limit), shorter measuring cell (shorter response time and higher upper limit), inside coated with PTFE, sapphire glass instead of fused silica, pressure probe installed, and thermocouple at the probe tip.

Experimental Set Up and Procedure The measurements were performed at an end cell at SU4 Hydro Sunndal in Norway during two campaigns, one in October 2008 and one in May 2010. The laser used in the 1st campaign had a more sensitive measuring range and was therefore suitable for measurements only at sites where the HF concentration was below 750 ppm, such as on top of the crust far away from any openings. The second campaign was performed with modified equipment. The measuring cell was coverered with PTFE (Teflon ®) to provide better protection towards HF than the steel used earlier. The cell was also made shorter. Furthermore, a sapphire glass window was used instead of fused silica. The shorter measuring cell resulted in a quicker dynamic response and contributed to a higher upper concentration limit since the physical principle utilized in TDL technology involves a correlation between concentration and optical path length. A pressure probe was placed in the measuring cell. This improved the instrument's capability to handle underpressure, something which was essential

Figure 4. a) Probe tip with thermocouple. b) Probe inserted through a cell cover.

264

Above Crust and in Open Tap Hole and Feeder Hole. When the probe was placed as shown in Figure 5, the HF stationary HF concentration varied between 5 and 10 ppm; see Figure 7. When the probe was positioned above the feeder hole as shown in Figure 6, the concentration was in the range 100-200 ppm.

Figure 5. Probe positioned on top of the crust.

100

150 time/minutes

Figure 7. HF and H 2 0 levels above crust (probe positioned as shown in Figure 5). Figure 6. Probe positioned above feeder hole.

Figure 8 shows how the concentration varied as the probe was moved from above a blocked feeder hole to an open tap hole close to the bath surface. The probe was then flushed with N2, then placed in an open feeder hole close to the bath surface. Keeping the probe in the open tap hole or the open feeder hole made the concentration rise abruptly to above the measuring range of the instrument. It should be noted that the probe was positioned only a few centimeters above the bath surface, and not as shown in Figure 6. The flat signal when placed in the open tapping hole is simply a signal overload. The reason for the slightly stepwise increase when held above the open feeder hole is that the probe had to be yanked away every time the crust breaker came down before feeding. This resulted in a small dilution effect.

To verify the system, measurements were performed in the duct (the measuring site was located at a place close to the cell where only the off gas from one particular end cell is passing) where the expected concentration range could be calculated to be around 300 ppm as mentioned earlier. During one of these periods, a NaOH containing filter was placed at the end of the line. This was done to confirm that the online measurements were in consistency with the well established Sintalyzer fluoride analysis method. Then the probe was placed on top of the crust as shown in Figure 5, and further hand held above an open tapping hole and above and inside an open flaming feeder hole. Figure 6 shows the probe placed above an open flaming feeder hole. Results 1st Campaign Duct Measurements. A more convenient short measuring probe was used during the first duct measurements. This probe was kept in place over night for approximately 16 hours and 40 minutes while recording the HF and H 2 0 concentrations continuously. To trap the HF passing through the laser, several filters soaked with NaOH were placed at the end of the line and kept there for about 16 hours and 13 minutes, that is, during almost the entire time period. The subsequent Syntalizer analysis performed on the filters showed an average concentration of 213 mg/Nm3 HF in the gas passing the laser during the measurement The average HF concentration calculated from the laser measurements done at the same time interval was 253 ppm, which coreesponds to 210 mg/Nm3.

0.0

20.0

40.0

60.0

80.0

100.0

time/minutes

Figure 8. HF and H 2 0 levels at various locations inside the cell. 2nd Campaign

The values recorded by the laser varied from slightly below 200 ppm to slightly above 400 ppm, clearly correlating with the feeding cycles as would be expected.

The main purpose of this campaign was to obtain a quantitative value of the HF levels in the open feeder holes and tap holes.

265

Duct Measurements. As in the first campaign, to confirm that the system was without large leaks, and that the values obtained were reliable, various measurements were done in the duct. Figure 9 shows HF and H 2 0 levels when a short probe and the long probe were placed in the duct with and without dust filter. As can be obeserved, when used with a dust filter, the long probe exhibited a larger response time than when used without the dust filter (after 4 minutes the measured value is still far below the expected stationary value). This was probably caused by the fact that a larger pressure drop is built up across the dust filter. When the dust filter was removed, it took about 4 minutes to obtain a "reasonable" value (approx. 300 ppm). The rest of the measurements were therefore conducted without using the dust filter.

As shown in Figure 10, the highest temperature (approximately 876°C) was measured inside the open feeder hole, a few centimetres above the bath surface.The bath temperature at 14:00 was measured to be 954 °C. When held in the open tap hole, the bath surface could not actually be observed visually. The temperature being "only" between 600 and 700 °C, might indicate that the probe was further away from the bath surface than when held in the open feeder hole, or it might indicate that the temperature above the bath is higher in a flaming hole due to the oxidation of CO. 1.8 1.5 1.2

3

O

N2 in outer circuit

in the duct with "long probe" - without dust filter

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3? 0.9

2

0.6

%,5

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14:15

14:25

14:35

14:45

14:55

15:05

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Figure 11. HF and H 2 0 levels in an open tap hole and in an open feeder hole plotted together with the course of events, a) 14:20-14:30: in open tap hole, b) 14:3014:34: break, c) 14:34-14:43: in open flaming feeder hole d) 14:43-14:44: break e) 14:44-14:52: back in open feeder hole, f) 14:52: break

Figure 9. Measurements of HF and H 2 0 in the duct. Measurements in Open Holes After the duct measurements, a thermocouple was attached to the tip of the probe. The probe was transferred to an open tap hole, and then to a flaming open feeder hole. The measurements are shown in Figures 10 and 11. The figures represent the same time period, i.e, the same events, but presented somewhat differently. Experience from the last campaign had shown that there was a big concentration gradient above the open holes. Therefore, to get as high a reading as possible, the probe was held only a few centimetres above the bath surface. However, since the crust breaker regularly came down (every 2-3 minutes), the probe had to be pulled up and away from the path of the crust breaker. This to a certain extent contributed to dilution of the gas sucked into the probe.

The highest HF value measured was more than 9000 ppm (0.9%), and this was obtained when the probe was held in an open feeder hole (periods c and e in Figure 11). During the same periods, the water vapour level reached a low value when the HF values were high. This suggests that a large part of the moisture is converted to HF according to Eq. (2) at the prevailing conditions; high content of fluoride vapour and high temperature. When the probe was removed from the open feeder hole, the concentration of water vapour rose to a higher level. This is especially pronounced in period f in Figure 11. Why the concentration exceeds its original background level is uncertain.

time

Dynamic Response Time. The dynamic response time is here defined as the time between an imposed change on the system and the first response recorded. The thermocouple has a very fast dynamic response compared to the HF and H 2 0 signal (and is also logged every two seconds), so its dynamic response could be used as a time basis for calculating the response in the other recorded items. In this way, the dynamic response time for the HF and H 2 0 signals are estimated from Figure 12 to be 21-33 seconds and 10-11 seconds respectively; the probe is taken out from an open feeder hole at 13:36:18 and the measured HF concentration starts to decline between 13:36:40 and 13:36:51. H 2 0 starts to increase between 13:36:18 and 13:36:51. The probe is put back in at 13:38:46 and the HF concentration starts to increase between 13:39:05 and 13:39:17. The onset of water level decrease is between 13:38:44 and 13:38:54.

Figure 10. HF and H 2 0 levels in an open tap hole (14:20-14:30) and in an open feeder hole (14:34-14:52) plotted together with the temperature measured.

Stationary Response Time. The stationary response time is here defined as the time it takes for the system to obtain a new stationary value (where this is expected) after a change has been

14:35

14:45

266

be around 4.5 %, more than four times as much as the actually recorded amount. This may indicate that even though there, in principle, would be overpressure and a gas flow directed upwards in the open flaming feeder holes due to the gas evolved underneath the anodes, air may be sucked into the holes and possibly also underneath the crust. Possibly, this can be related to waves and splashing at the bath surface. The observation made in this work, that there is a large difference in HF concentration when measured "above" the hole as opposed to "inside" the hole also supports this. A follow up on the present work was later done at Alcoa Mosj0en, where levels up to 2.6 % HF were detected in crust openings; this is reported elsewhere [7]. The feeder holes at Alcoa Mosj0e had a smaller diameter and were more cylindrically shaped than the ones observed at Hydro Sunndal.

imposed. For HF, this was estimated from the duct measurements to be approximately 4 minutes for increasing concentrations and approximately 2 minutes for decreasing concentrations. For H 2 0, the current data do not provide enough information to estimate the stationary response time. The response time of the laser itself is believed to be around one second, so the discussion above refers to the whole system.

Conclusions The stationary HF level in an open flaming feeder hole was approximately 9000 ppm, when measured a few cm above the bath surface. In comparison, when the probe was positioned 5-10 cm above a crust area with good integrity, the HF concentration was in the range 5-10 ppm. The results support the notion that most of the HF evolution emits at the open flaming feeder holes and tapping hole. The gas sniffer equipment worked according to the purpose, and this method has proven a useful tool for these types of measurements. Future campaigns will be carried out, both at other smelters but also with respect to other kinds

Figure 12. Close up of HF and H 2 0 levels as they change when the probe is taken out from an open feeder hole and put back in. Summary and Discussion

Acknowledgement

Table I shows a summary of the HF levels measured at the various sites. It seems rather clear that the HF gas generated exits almost entirely through the crust openings. Keeping the probe above a blocked feeder hole gave no indication of HF desorption from secondary alumina. The fact that surface adsorbed fluoride does not re-evolve during ore feeds to the pot, has also been reported by others [4,6].

The present work was funded by The Norwegian Research Council and the companies involved in the project Resource Optimation and Recovery in the Materials Industry (ROMA): Hydro Primary Metal Technology, Alcoa Norway, S0r-Al, Elkem, Eramet, Vale Manganese Norway, Fesil, Finnjord, and Alstom. Permission to publish the results is gratefully acknowledged. We would also like to express our gratitude towards Jan Olav Polden at Hydro Sunndal for providing vital equipment and for his assistance and help whenever needed. Thanks are also due to Chris Torjussen and his colleges at NEO for valuable advice in connection with the laser measurements.

Table I. Summary of HF levels. Location

Range

Comments

Above crust ("far" |awayfromfeeder hole) Above open feeder hole Duct

5-10 ppm

October 2008

5-200 ppm

October 2008

200-400 ppm, fluctuate according to feeding cycle 50 ppm Above blocked feeder hole Just above bath, open 3000-5000 ppm tap hole 1 In open feeder hole 0.90-1.0% 1 just above bath

References 1.

W. E. Wahnsiedler, R. S. Danchik, W. E. Haupin, D. L. Brackenstose and J. W. Colpitts, "Factors Affecting Fluoride Evolution from Hall-Heroult Smelting Cells", Light Metals 1978. pp 407-424

2.

W.E. Haupin og H. Kvande, "Mathematical Model of Fluoride Evolution from Hall-Heroult Cells", Light Metals 1993. pp 407-411

3.

Margaret Hyland, Edwin Patterson and Barry Welch, "Alumina Structural Hydroxyl as a Continuous Source of HF", Light Metals 2004. pp 361-365

4.

Michael L. Slaugenhaupt, Jay N. Bruggeman, Gary P. Tarcy and Neal R. Dando: "Effect of open holes in the crust on gaseous fluoride evolution from pots" Light Metals 2003. ppl99-206

October 2008 October 2008 May 2010 May 2010

If the suction rate during the duct measurements is approximately 7000 Nm3/h (it was not measured accurately for the actual cell during the campaign), the HF generation rate can be calculated to be around 27 kg HF/ton AI. If the raw gas that exits the feeder hole were to be completely undiluted, the HF concentration would

267

5.

Edwin Patterson: " Hydrogen fluoride emissions from aluminium electrolysis cells" PhD-thesis, The University of Auckland, 2002

6.

Neil R. Dando and Robert Tang: "Fluoride evolution/emission from aluminium smelting pots impact of ore feeding and cover practices" Light Metals 2005. pp 363-366

7.

Camilla Sommerseth, Karen Sende Ösen, Thor Anders Aarhaug, Asbj0rn Solheim, Egil Skybakmoen, Christian Rosenkilde and Arne Petter Ratvik, Light Metals 2011 (current issue)

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

LASIR™-R - THE NEW GENERATION RoHS-COMPLIANT GAS ANALYZERS BASED ON TUNABLE DIODE LASERS Alak Chanda1, Gervase I. Mackay1, Keith L. Mackay1, John T. Pisano2, Jean-Pierre Gagne3, Pierre Bouchard' unisearch Associates Inc., 96 Bradwick Drive, Concord, Ontario, Canada L4K 1K8 university of California at Riverside, 1084 Columbia Avenue, Riverside, CA, USA 92507 3 STAS, 1846 Rue Outarde, Chicoutimi, Qc, Canada G7K 1H1 Key Words: Tunable Diode Laser Spectroscopy, Hydrogen Fluoride, HF, emission monitoring, process control, aluminum smelter, gas analyzer the energy is absorbed at that wavelength. Since the amount that a molecule vibrates or rotates at a specific wavelength will be dependent on its capability to absorb energy, it will behave differently at different temperatures and pressures. This alludes to the fact that both temperature and pressure compensations for line intensities are required for any spectroscopic measurement. Such compensations can be made either by measuring the property changes of the molecules directly (for simpler molecules) and applying a mathematical correction or by preparing a compensation database empirically (for complex molecules) with laboratory experiments performed under controlled operating environments.

Abstract The laser-based optical gas sensor using Tunable Diode Laser Absorption Spectroscopy (TDLAS) is rapidly gaining favor wherever high sensitivity, real-time measurement and freedom from interferences are required [1,2], specifically for the measurement of HF in primary aluminum smelters and other process gases at various metal smelters. It eliminates the problems associated with extractive gas sampling techniques. A first generation of equipment designed by Unisearch Associates and based on the near-IR TDLAS appeared on the market in the mid-90s. It has since been improved significantly by employing fast scan measurement techniques. Now, the system is more compact, robust, easier to operate and calibrate, and it can be simply audited. The utilization of fast scan measurement techniques with the state-of-the-art high speed electronics and sophisticated software optimization algorithms not only provides enhanced stability and sensitivity but expands the dynamic range of the measurements to five orders of magnitude. This new generation RoHS-compliant instrument (Restriction of Hazardous Substances) is also very inexpensive compared to other gas analysis instruments. The technology has found applications in many industries that require fast, accurate measurement of trace emissions for both process control and environmental regulatory requirements. These include, but not limited to, HF emissions in aluminum smelters, NH3 slip measurements for DeNOx scrubbers in the Power Industry, HC1 and HF emissions monitoring in the Cement Industry and in Incinerators and for process control in Steel Smelting. This paper describes the new generation TDLAS-based gas analyzer and, as an example, its use in primary aluminum smelters.

The advantage of TDLAS over other techniques for concentration measurement is its ability to achieve fast, very low detection limits of the order of parts per billion with virtually no interference from the presence of other gas molecules. The LasIR™ is also widely being used by various aluminum smelters worldwide to measure gaseous hydrogen fluoride [3, 4]. Hydrogen fluoride (HF) is emitted during the aluminum smelting process. To avoid HF leakages in the work area, the electrolytic cells (pots) are hooded and the emitted gases are captured and vented through scrubbers. Hydrogen fluoride is a very toxic gas and represents a major health and safety hazard both for workers in the pot rooms and for neighboring environments. In addition, there is also an economic incentive to recovering this fluoride. The Tunable Diode Laser technology is presently used in over a hundred smelters in the world as an emissions monitoring device used to adjust emission control practices and equipment. Description of Technology

Introduction

The gas analyzing system consists of a tunable diode laser based analyzer and a small optical head that launches the laser light through a medium (that may contain the species of interest to be measured) and receives the transmitted laser light at the other end of the measurement medium. The analyzer itself can be located in a control room, as far as 1 km away from the measurement location. Afiber-opticcable and a coaxial cable link the analyzer to the optical head. Data are logged on to a compact flash disk built inside the TDLAS-based gas analyzer and/or an external PC via the Ethernet or RS232 outputs. Data is also transmitted to the plant database.

A basic TDLAS setup consists of a tunable diode laser light source, light transmitting optics, an optically accessible absorbing medium through which the light beam can be transmitted, receiving optics and detectors. The emission wavelength of the tunable diode laser is tuned over the characteristic absorption lines of a species in the gas in the path of the laser beam. This causes a reduction of the measured signal intensity, which can be measured by a detector and then used to determine the gas concentration. Most molecules are unsymmetrical and polyatomic. When the molecule becomes excited by absorbing near infrared energy, it increases its vibrational and/or rotational frequency. It is this tendency of the molecule to absorb energy that leads to spectroscopic identification of the molecule by the wavelength of absorption and the concentration of the molecule by how much of

The quantification is based on Beer-Lambert Law that relates the absorption of light to the properties of the material through which the light is traveling. It states that when a radiation of

269

frequency I(v) passes through an absorbing medium, the intensity variation along the path of the beam is given by: I(v)=I0(v)e-a(v)cl

and aligned probe model is also available that can be mounted at any stack/duct for short-term or long-term measurements.

[1]

where,

i-600m

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is the transmitted intensity of radiation after it has traversed a distance / through the medium, (in milliwatts) I0 (v) is the initial intensity of the radiation, (in milliwatts) o(v) is the absorption cross-section of the absorbing species, (cm2/# of molecules) C is the number density (concentration) of the absorbing species, and (in # of molecules/cm3) / is the path length or the traverse distance through the medium (in cm). I(v)

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Remote measurement of gases using TOLAS provides the concentration that actually represents a value averaged across the entire path. This gives in situ measurement an advantage over point sampling in which a probe is inserted to draw the gas out of the duct, especially when concentration gradients or heterogeneities exist. The path averaged value obtained from in situ sampling is better representative of the overall concentration within the process.

Probe Measurements

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Figure 1. Schematic of TDLAS-based gas analyzer and the optical head configurations that can be connected to the analyzer byfiber-opticand coaxial cables. Portable stand-alone models are also shown. The TDLAS-based gas analyzer consists of appropriate hardware and software that require no field calibration. This eliminates the requirement of having calibration gas cylinders available on site. For regulatory purposes, compact hand-held audit modules are provided that allow the user to periodically audit the system as a whole (analyzer and optical heads) when desired or use the periodic automated audit feature. Automated audit results can be stored and displayed on screen. Alternatively, where auditing is to be performed by flowing a known level of gas on the path of the optical beam, an optional add-on module is available that can be configured as an inline or a stand-alone offline audit assembly. Calibration, although not required, can also be performed with this add-on module where regulatory requirements have to be fulfilled. With the use of a TDLAS-based analyzer, fast, real-time, in-situ measurements of gas concentrations linear to five orders of magnitude (sub-ppm to percent levels) can be made under various industrial environments.

The use of a laser source coupled with a fiber-optic cable allows the light to be optically multiplexed to make simultaneous measurements of the species of interest at different locations. With the TDLAS-based gas analyzer, measurements can be made from 1 to up to 16 different locations with a single analyzer. In addition, multi-species can also be measured with the same analyzer where the absorption signals of the species lie relatively close to each other and within the tunable range of the laser. Measurement of gases in explosion-proof environment can easily be done by using optical heads that require no electrical components (with configurations that transmit and receive the optical signal viafiber-opticcables). Unisearch has recently introduced a new generation of TDLAS-based gas analyzers that are RoHS-compliant and available in either rack-mount or table top models as standards. With the advent of high speed electronic data processing and measurement techniques, the costs of these types of TDLAS based analyzers are significantly reduced. The gas analyzers come in a compact form that makes them suitable for aluminum smelter environments. They are available in configurations from simple portable single-channel analyzers to systems capable of measuring up to 16 channels that can monitor both multiple locations or different gas species. All the measurement channels are independently controlled, which permits a single multi-channel analyzer to simultaneously measure gas levels in duct/stack, long-path ambient air, extractive sampling and any such combination. Each channel can handle very different gas levels with no interference with other channels. The possible measurement configurations are shown in Figure 1. The system can also be configured as a stand-alone portable gas analyzer with the optical head integrated with the electronics. This light-weight (< 5 kg), energy-efficient (consumes less than 20 W) model can be mounted on a tripod stand and, with the use of a retro-reflector array, allows measurement of various gases in an open-path mode to up to several hundreds of meters. A stand-alone, pre-configured

HF Gas Measurements as Process Control in Primary Aluminum Smelters One of the primary requirements of a gas analyzer, especially when used as a Continuous Emissions Monitoring system, is the accuracy and linearity of the measurements. This data is concurrently important when a system is used for process control applications such as maintaining scrubber efficiency and improving work practices, as well as reducing emissions. Various test and certification agencies exist that are able to verify the functionality of such analyzers. One such certification agency is the well-recognized German TÜV. Approval requires the use of two essentially identical analyzers measuring the same gas. Laboratory and field (industrial plant) tests are conducted in order to verify the manufacturer-stated specifications. These include, among many other parameters, the linearity, accuracy, precision and the operation of the analyzer under various environmental conditions. Figure 2 shows the linearity and accuracy of two TDLAS-based gas analyzers for low levels (0 to 2.5 ppm-m) of HF gas

270

measurement. As can be seen from the plot, the measured HF gas levels were almost identical between the two analyzers. AH Alumina Feeds Returned To 100%

HF measurements with two LasIR analyzers - TUV testing

— a —LaslR#2 y=1.002x +0.011 R* = 0.999

10:25

12:25

14:25

16:25

18:25

2025

Time (hh:mm)

Figure 3. Variation in HF emissions as measured by a TDLAS-based gas analyzer as a function of alumina feed to the dry scrubber.

Actual HF Concentration {mg/m3)

Figure 2. HF measurements with two TDLAS-based LasIR™ during TÜV certification.

A number of installations use TDLAS-based systems to monitor scrubber efficiency. In a typical installation, a single controller is used to monitor the HF concentrations at both the inlet and the outlet of the scrubbers. Figure 4 shows a plot of such measurements at a newly installed scrubber. Measurements were made in an in-situ mode. This plot shows that the new scrubber was able to remove more than 99.8% of the H. This information also allows the scrubber operator to maintain the desired reduction of fluoride emissions from aluminum smelting by adjusting the alumina feed rate (and the secondary alumina recycle rate) to the dry scrubber to optimize the capture efficiency of the gaseous fluoride byproducts. The measurement also indicates to the operator the surface adsorption inefficiencies that arise from losses of scrubber material such as bags either breaking or leaking, or saturation of the reactive material.

In aluminum smelting process, aluminum metal is extracted from its oxide alumina, generally by an electrolytic process called the Hall-H6roult process. As such, aluminum smelters use enormous amounts of electricity and are often located very close to large power generating stations. Primary aluminum is produced by the electrolytic reduction of alumina (A1203), dissolved in a mixture of molten cryolite (Na3AlF6) and A1F3. Since HF is emitted during the process, the electrolytic cells (pots) are hooded and the emitted gases are captured and vented through scrubbers. Hydrogen fluoride is a very toxic gas and represents a major health and safety hazard both for workers in the pot rooms and for neighboring environments. In addition, there is also an economic incentive to recovering this fluoride. Dry scrubbing technology is used at most of the world's aluminum smelters. The hydrogen fluoride is chemisorbed directly on to the alumina from the hot exhaust gas. Fresh (primary) alumina is injected directly into the raw exhaust gas stream, which mixes and reacts with the alumina. Then the reacted alumina, as well as the particulate fluorides and other particulate materials, are removed from the exhaust gas stream by bag filtration. The alumina collected (secondary alumina), which contains almost all thefluoridesand particulates emitted from the reduction process, is then fed to the pots. Thus, the entire process operates as a closed loop for the captured cell emissions. It has been found that manipulations in the feed of alumina to the dry scrubber altered the levels of HF emitted from the scrubber outlet that is usually vented to open air. Therefore, a continuous in-situ measurement is a useful tool to optimize the feed of alumina to the dry scrubbers, and/or the secondary alumina recycle rate.

Also of importance is the accuracy and reliability of the measurements when made under an industrial environment. The gas temperature, pressure, humidity and operating environmental parameters can constantly change as a function of time. In order to validate the measurements, a continuous 5-day comparison study of the HF gas data was made between those reported by a TDLAS-based gas analyzer and the standard wet chemical method. Results are shown in Figure 5. Apart from the response time where the gas analyzer shows sharper peaks, the results were virtually the same.

Average = 218 mgini3

Scrubber Inlet

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An example of the control of alumina feed is shown in Figure 3, obtained from tests conducted at an installation site. The higher HF levels were produced by reducing the primary alumina feed to the scrubber. In just a few minutes, the levels rapidly rose from 0.47 mg/Nm3 to 0.8 mg/Nm3. A few minutes later, when the HF levels appeared to reach stable values, the recycle rate for injected secondary alumina was reduced. The concentration rose to almost 1.8 mg/Nm3. To evaluate the scrubbing response, all parameters were set to normal. Once all the alumina feeds were turned on, an exponential decay in HF emissions was observed.

E

o

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1 Scrubber Outlet

03 =

Average = 0.42 mg An3 1

i

1 12:00

TIME

Figure 4. HF measurements at the inlet and the outlet of a dry scrubber; the scrubber efficiency is determined to be better than 99.8%.

271

increased beyond 50.5%. The results are shown in Figure 8 and indicate that the measurements are linear over the entire measured range. Such systems become increasingly useful in industries such as nuclear material processing plants.

Wet Chemical Method

HF measurement at the top of a roof line in an aluminum smelter

3.0

I 2.5 E

2 3 No. of days

^t

Figure 5. HF measurements at the inlet and the outlet of a dry scrubber. The scrubber efficiency is determined to be better than 99.8%.

2.0

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Figure 6 shows the measurement of HF gas in the exhaust manifold of a single pre-baked anode electrolytic cell. The measurements were made in a path length of 0.5 meter and approximately 1 meter away from the end of the pot. Periodic reproducibility as a result of automated alumina feed and other related processes can easily be seen on the expanded plot.

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20:00

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1:00 Time (hh:mm)

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Figure 7. Measurement of HF at the top of a pot room of an aluminum smelter.

Single Pot {Pre-baked Anode} Duct HF Concentration Measurement

Online Linearity Measurement of HF Test Gas at a Metal Fluoride Processing Plant 58 5% 45.5 %

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18:15

18:30

16:45

17:00

Time ihr.min)

Figure 6. HF off-gas measurement from a single pot 1 meter away from the pot end.

Figure 8. Linearity measurement of HFfrom0% to 50%. Conclusion

Modern cells fitted with hooding to capture the emissions are usually operated at a slight pressure. Process interventions such as the manual alumina feed, changing of anodes or metal tapping and suppression of anode effects, all require opening of the hoods. Such interventions and any leakage in the pot hooding itself result in the emission of HF into the pot room, which is ultimately vented through the roof of the pot room. By measuring the HF emissions in the pot room and correlating their occurrence with specific pot tending activities, improved work practices can be developed and better control of current efficiency can be achieved. Figure 7 shows the levels in a pot room measured with a gas analyzing system. The sampling time was 5 seconds and the path length was 400 meters.

The Tunable Diode Laser based gas analyzer is being used as a tool for making in-situ measurements of emissions and flue gases in various industries. An example of its use specific to, and not limited to, aluminum smelters has been provided. Tunable Diode Laser Absorption Spectroscopy affords the most interference-free analytic method for measuring gas concentrations where fast and accurate response is desired. Such systems are particularly well suited to process control applications. These are small in size, can be designed as calibration-free, rugged and virtually maintenance-free, making it easy to install and operate. The use of fiber-optic cables permits the instrument to be located in an environmentally suitable area, remote from the actual measurement location. In addition, optical multiplexing permits measurements at a number of locations with a single instrument, making the systems very cost effective. Measurement of gases in explosion-proof areas can easily be done. The data is used for the control and optimization of various process related parameters. This leads to energy efficiency, reduction in production time and improvement of the quality of the products.

There are several processes where the gas level changes over a wide range of magnitude. This can be as a result of a specific step in the process or the process as a whole. In order to verify the dynamic linearity of the analyzer, tests were made in an industrial environment where provisions of controlled variations in the injected HF levels existed. The HF gas level was varied from 0% to up to 50%. In the set up used, the levels of HF could not be

272

The technology has found applications in many industries that require fast, accurate measurement of trace emissions for both process control and environmental regulatory requirements. These include NH3 slip measurements for DeNOx scrubbers in the power industry, HCl and HF emissions monitoring in the cement industry and in incinerators, off gas remediation of H2S in wastewater treatment and nickel smelters, deuterated water leakages in nuclear power plants and for process control in steel smelting, to mention just a few. References 1. John T. Pisano, Claudia Sauer, Tom Durbin and Gervase Mackay, "Measurement of Low Concentration NH3 in Diesel Exhaust using Tunable Diode Laser Adsorption Spectroscopy (TDLAS)", Society of Automotive Engineers Conference, Paper 2009-01-1519, Detroit, MI, February (2009). 2. H.A. Gamble, G. I. Mackay, J. T. Pisano and R. Himes. "Post Combustion NOx Control Process Monitoring using near-IR Tunable Diode Laser (TDL) Spectroscopy to Measure Slip Ammonia", 20th Annual International Forum Process Analytical Technology, Arlington, VA (Washington D.C.), February (2006). 3. Harold I. Schiff, Alak Chanda, John Pisano and Gervase Mackay, "The LasIR™-A Tunable Diode Laser System for Environmental and Industrial applications", PIE International Symposium 3535 on Industrial and Environmental Monitors and Biosensors, Boston, MA, November (1998). 4. Harold Schiff, John Pisano, Alak Chanda, Dave Karecki and Gervase Mackay, "Measurements of Fluoride Emissions in Aluminum Smelters by Tunable Diode Laser Spectroscopy", AWMA Annual Meeting, St. Louis, MO, June 1999, paper 115.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

USE OF SPENT POTLINING (SPL) IN FERRO SILICO MANGANESE SMELTING TMS1, Paulo von Krüger2 *TMS (The Minerals, Metals & Materials Society); 184 Thorn Hill Rd.; Warrendale, PA 15086, USA 2 Universidade Federal de Ouro Preto Keywords: Spent Potlining, Ferro Silico Manganese Abstract period that goes from 2000 to 3000 days, the pot cathode lining fails and must be rebuilt. At this time, the old lining is dismantled. This material is called SPL (Spent Pot Lining) and is constituted of the starting lining material and the components absorbed during the operating pot life. Figure 1 shows the steps of the SPL formation [1]

In this work an evaluation of the possibilities to employ Spent Potlining (SPL) as a component of the burden of a Submerged Arc Furnace (SAF) producing Ferro Silico Manganese Alloy is investigated. On this subject, a characterization of the SPL's most probable components as well as their interaction with the existing species in the ferroalloy furnace were carried out. Additionally relevant features of the ferroalloy smelting were identified and characterized. Those figures were introduced in a thermo chemical program where the SAF operational conditions were simulated in order to check the technical feasibility of that use.

SPL can be easily separated in two well defined fractions: The carbon fraction is constituted of the old cathode blocks and non carbon fraction, constituted of old refractory and insulating bricks, ramming paste, electrolytic bath and a series of components formed by the interaction between those components and metal and air.

The simulation results showed that, on the technical point of view, the SPL carbonaceous fraction is a suitable component of the SAF burden, producing Ferro Silico Manganese alloys.

From a general point of view, the reductant is the chief cost item in carbothermic reduction process. Among these processes are the manganese alloys smelting. Ferro Silico Manganese smelting is an electrometallurgical carbothermic process. In such a process, besides its reducing characteristics, carbon is the main element responsible for the electrical characteristics of the burden.

Although some of the figures generated by simulation were confirmed in several industrial exploratory tests, a more detailed test program is advisable. Introduction The electrolytic cells employed in aluminum winning are lined with carbon that is in contact with the bath and metal and insulating material placed between the former and metallic shell. Through its operating life, the cathode lining changes chemically and physically, that leads to its expansion and deterioration. As a consequence a splitting process grows up and after a

Last, but not least, the ash content and composition have an influence on slag volume and characteristics and, depending on the nature of other components, on alloy quality and furnace operation.

New lining

,JOL

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Lining deterioration

Dismantled k

1 st cut (example) 2nd cut (example)

Figure 1 - SPL formation scheme [1]

275

In this paper the possibilities for employing the SPL's carbon fraction as a burden component on ferroalloys smelting is evaluated. Specifically, the selected ferroalloy is from Ferro Silico Manganese type.

Spent Pot Lining (SPL) Essentials [2] The aluminum pot lining is constituted of a series of components, each with its own purpose.

In a first approach, aiming to check the technical feasibility of that alternative, the main physical and chemical characteristics of the SPL are established and the resultant data will be "introduced" in the Ferro Silico Manganese furnace burden and the furnace behavior with this new component was analyzed. Although this first evaluation is a theoretical approach, it indicates if there are constraints that would impair this application. The next step, industrial scale trials would be essential.

Figure 2 shows the scheme of a typical lining. [5] As can be seen, the bottom lining is constituted of three kinds of non carbon materials and a fourth that are the carbon blocks. Among the non carbon materials, the berth is an alumina layer, the refractory are silico-aluminous bricks and the insulating layer is made of low density and low thermal conductivity bricks. Not characterized in the figure, the side lining is constituted of carbon bearing paste, silicon carbide and ceramic materials

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Figure 2 - Pot lining scheme (typical) Carbon blocks, constituting the top layer of the bottom lining, are by far the most important component of this system. The carbon block assemblies form the cathode of the electrolytic system.

Chemical Characterization SPL material comprises carbon and refractory materials, penetrated metal and bath components and minor amounts of aluminum nitride, aluminum carbide and cyanide salts. There is not an exact composition of SPL; it varies over a wide range, depending on several factors, such as lining design, pot age, etc.

Carbon blocks must meet several requirements: among them are high electrical conductivity, low reactivity and high mechanical strength.

The whole of SPL can be divided in two "cuts"; the first cut corresponds to the carbonaceous part (carbon blocks) and the part of the lining beneath the blocks assembly correspond to the second cut.

Depending on the type of the block, those properties can vary among a relatively ample range, in a new lining. However, during their operating life, the blocks absorb bath components and interact with it and the surrounding environment. Within this process, both the physical and chemical characteristics change. For the purpose of this paper, the properties to be considered are those existing at the end of the lining life.

Component Carbon Total Fluorides Free Alumina Total Aluminum Total Sodium Calcium Quartz P Sulfur

Table I shows typical composition of SPL, as well as those from the first and second cuts. According to what was mentioned before this is just a reference composition

Table I. SPL's typical com ponents and concentrations First Cut (carbonaceous) Total (56%) (100%) 54-64 6-20 0-15 5-15 5-12 0.5-4.0 0-6 0-650 g/t 0.1

33.1 15.7 22.3 15.1 14.2 1.8 2.7 0.3 0.1

Second cut (non carbonaceous) (44%) 18.2 4-10 10-50 12.6 12.0 1-6 10-50 0-300 g/t 0.1

and insulating material, whose properties are extremely variable, both geometrically, physically and chemically. Both cuts are impregnated with bath components and metal.

As suggested in the third drawing of figure 1, the first cut is constituted of big pieces of high resistance carbonaceous materials, from the former cathodes blocks. The second cut is a mix of paste, refractory

276

Although the whole of the spent potlining are together from the pot cathode, fortunately the two cuts are quite different and can be easily separated each other.

As will be seen later, the material that meets the most suitable properties, considering the burden characteristics of a Ferro Silico Manganese furnace, is the first cut, both from the chemical, physical and metallurgical point of view.

The photo of Figure 3 shows pot being dismantled. As can be seen, there is a sharp difference between the carbon blocks (1 st cut) and the other components (2nd cut)

On the other hand, a greater part of the second cut is constituted of loose or powdery material that would seriously impair furnace operation. This suggests that only the carbonaceous fraction (first cut) can be considered for the proposed application.

Figure 3. Section of the lining of an aluminum pot, during its dismantling (light line enhances carbon blocks) The number of reaction is extensive; for the sake of simplicity this paper only considers the most relevant.

In order to evaluate the metallurgical behavior the figures shown in Table I must be detailed, identifying the actual compounds present in the material, specifically in the first cut.

Figure 4 shows the evolution of the main compounds during the pot lifetime. For the purpose of the present work, the reference figures are those at the time of pot shutdown.

For that, the possible reactions between carbon lining an surrounding components, as well as their relative preferences must be known and analyzed.

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Besides these components there are the carbon and quartz. Both the former and the later are diluted during lining life. Based on these aspects, the SPL Carbonaceous fractions were calculated with the resultant values shown in the following list. Naturally, it must be pointed out that these figures can vary from plant to plant. Another remark is that the criolite and sodium aluminate were decomposed in its elementary compounds with the purpose of the metallurgical operation evaluation. SPL reference composition: C 55.98% A1203

7.18% Na20 22.70% Si0 2 1.55%

1.22% 5.00%

Electrical Properties Electrical resistivity is an important parameter both in aluminum and in ferro silico manganese smelting. In the first instance, the resistivity must be as low as possible. For the ferroalloy smelting the operation is carried out in a submerged arc furnace; the high temperature zone depends on the electrode tip

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277

position that depends on the voltage and burden resistance.

graphitization and impregnation. The first one improves the conductivity and the later increases the resistivity. Figure 5 shows this variation both in ambient and operating temperatures.

During its operating life, the aluminium smelting cathode block is submitted to two kinds of processes:

Cathodic lining - resistivity x time

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Figure 5. Cathode block - change of the resistivity during operating life [3] The negative slope in the first section of the curves suggests that during initial cathode operation that graphitization prevails over impregnation of bath species. The positive slope in the final section indicates that the block is already fully graphitized

and only impregnation occurs. Based on that it can be said that at the end of its useful life, the resistivity of a cathode block are nearly the same, doesn't matter if it was originally of amorphous carbon or fully graphitized types [3]

Reduction temperatures - selected oxides

300

600

900

1200

1500

1800

Temperature K

Figure 6. FeSiMn smelting scheme (easily reduced iron is not shown) and reduction temperatures

Production of FeSiMn is carried out by the simultaneous reduction of manganese and silicon, by carbon, in submerged arc furnaces (SAF), as shown schematically in figure 6.

MnO + C = Mn + CO, depends on its activity in slag; the higher the concentration of free MnO, the lower the reduction temperature. On the contrary, the higher the concentration of MnSi04, the higher its reduction temperature (fig 6).

Burden is constituted by manganese ore, manganese bearing slag, quartz (or quartzite) carbon source and fluxing agents. The manganese reduction temperature from its lower oxide, MnO:

Reduction of silicon from silica (Si02): Si0 2 + 2C= Si + 2CO, requires higher temperatures, as shown in figure 6. Therefore, in order to get a more intense reduction of Silicon, the Manganese reduction temperature must

Ferro Silico Manganese (FeSiMn) Essentials[6]

278

be higher, that means higher concentrations of silica in the slag. On the other hand, the more silica in slag, the more viscous is it. As far as high viscosities impairing the mobility of reagents, inhibiting the reduction reactions, one can conclude that the higher the concentration of Silicon in FeSiMn alloy, the more difficult the process (indeed, there is a limit above it that the process is not feasible).

the SPL characteristics presented above, two simulations [5] of the SAF operation with and without SPL in the burden were carried out. The purpose is to evaluate by comparison, the technical feasibility of the former. All other components were the same in both cases. The simulation was carried out by means of a thermochemical program previously developed for the evaluation of ferroalloys smelting procedures. Its validation was made by the comparison with actual operations [5,7].

There are several ways to determine slags viscosities. This can be done experimentally or applying empirical expressions or diagrams. In this paper the hereafter expression [4] was adopted because it considers species encountered in SPL and so its influence can be evaluated. This is a theoretical approach that should be confirmed experimentally, but for the purpose of this work, it is considered acceptable.

Side Reactions Γ3/71 Before going on with the simulation, a thermodynamic analysis was made, aiming to identify the most probable reactions and the resultant compounds. The behavior of the oxides is well known and is the same for both alternatives. The Na20 behavior can be compared to that of K 2 0 that is a very common component in manganese ores.

Ιημ = -12,0 + 0,G525(%Al2O3) + 0,088 l(%Si02)* 0,0409(%CaO)* - 0,042l(%Na20)* - 0,0194(%F)* + 14941/T where (%CaO)* = (%CaO) + (%MgO) + (%MnO) + (%FeO) + (%B203) (%Si02)* = (%Si02) + (%P205) + (%Ti02) + (%Zr02) + (%Fe203) (%Na20)* = (%Na20) + (%K20) + (%Li20) (%F)* = Z(% Fluorides)

On the fluoride side, NaF and CaF2 are quite stable and are incorporated in the slag. Aluminum fluoride (A1F3) may react with silica, silicon or lime, according to the following reactions 4A1F3 + 3Si02 = 3SiF4 + 2A1203 4AlF3 + 3Si = 3SiF4 + 4Al 4A1F3 + 3CaO = 3CaF2 + A1203

Though not being necessarily precise, the qualitative influence of the components on slag viscosity is well posed.

Considering the furnace temperature and the concentrations of the species, it was assumed that all A1F3 would react with Si0 2 , forming volatile SiF4 (first reaction) The formation of CF4 is not likely to occur.

Process Evaluation Based on the features of the FeSiMn smelting and on

Table II. FeSiMn smelting - Results of simulation - conventional procedure Chemical composition Inputs kg/1 alloy Manganese ore 1668 Alloy Slag 22.31% Mn rich slag 910 Si Si0 2 41.61% 334 Quartz 13.45% Fe A1203 23.20% Mn Limestone 230 63.55% CaO 18.19% Al Coke 369 0.00% MnO 16.62% CaF2 Petcoke 155 C 0.70% 0.00% SPL P 0.02% NaF 0 0.00% Outputs kg/1 alloy Electrical Energy Viscosity Alloy 1000 3762 kWh/t alloy 0.58 Slag 929 1

1

In the simulation with SPL it was considered that it was introduced substituting petroleum coke. As far as carbon content of petcoke (>90%) is much greater than that of the first cut of SPL (-54%), the substitution on a weight by weight basis implies in the charging of additional metallurgical coke, to achieve the required carbon. The mass balance of carbon was completed by means of metallurgical coke in order to keep the burden resistivity constant. Table III shows the results of this simulation.

Process Simulation Γ5Ί In the reference process, without SPL, it was considered a conventional operation, using manganese ore and manganese rich slag (from HCFeMn production), quartz and limestone. The reducing agent was a mix of metallurgical coke (70%) and petroleum coke (30%), which is a blend in manganese alloys smelting operations. Table II shows the results of this simulation

279

Table III. FeSiMn smelting - Results of simulation - procedure with SPL Chemical composition Inputs kg/1 alloy Slag Alloy Manganese ore 1678 Mn rich slag 22.31% 39.30% Si 910 Si0 2 Quartz 334 13.45% 23.61% Fe A1203 Mn 63.57% CaO 17.18% Limestone 230 Al 0.00% MnO 15.70% Coke 428 CaF2 0.66% 0.25% Petcoke C 0 P 0.02% NaF SPL 3.60% 155 Viscosity Electrical Energy Outputs kg/1 alloy 3793 kWh/t alloy 0.47 Alloy 1000 984 1 Slag 1 i

1

From the electrical point of view the burden must have suitable resistivity by assuring optimum positioning of electrode tips. For that it would be desirable that the SPL resistivity be similar to that of the substituted reducing agent. Figure 7 shows the change of the resistivity with temperature for charcoal, coke and SPL[2,3 and 6]

Comparing the results from the two tables it can be seen that all but the reducing agent, components proportions are practically the same in both cases. The slag weight is greater when SPL is employed due to its ash content being greater than that of petcoke; this also implies greater energy consumption. On the other hand the sodium and fluoride content of SPL lead to a lower slag viscosity, what enhances the reaction efficiency.

Resistivity x Temperature 350

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Figure 7. Resistivity dependence with temperature - Selected reducing agents According to the above figure, SPL resistivity is the lowest when compared to coke and charcoal. However, is must be pointed out that the figure from SPL corresponds to one resistivity whereas those from coke and charcoal are bulk resistivity's. This suggests that the SPL bulk resistivity, considering the same particle size of that as coke, would be very close to that of this later.

References 1. P. von Krüger, P. Possibilidades de valorizagδo e inertizagδo do Revestimento Gasto de Cubas de aluminio (RGC) . l°Congresso Internacional da Industria do Aluminio - ABAL S. Paulo 2000 2. M. S0rlie, and H. A. 0ye, Cathodes in Aluminium Electrolysis 2nd Edition. Düsseldorf: AluminiumVerlag Gmbh, 1994. 3. M.S0rlie, and H. Grau, Properties Changes of Cathode Lining Materials During cell Operation. Las Vegas Light Metals 1995 TMS - pgs 497 - 506. 4. LA Silva.et al, .Some Empirical Correlations for Evaluating Viscosity of Molten Mold Flux for Continuous Casting of Steel. 1st Process Metallurgy and Structural Engineering of the Int'l Materials Research Congress, Cancun, Mexico - 2000 5. P. von Krüger, Metalurgia - Fundamentos Teoricos e suas Aplicagöes na Redugδo do Aluminio. Apostila de Curso. (internal publicashion- ALUMAR, ALBRΔS e VALESUL, 1995,1997,1998. 6. A. Lucio et al Metalurgia dos Ferro-ligas. Internal Publication - Depto de Engenharia Metalurgica EE/UFMG-1978 7. P. von Krüger, Internal Reports - not published

Concluding remarks The results of the simulations and other features carried out in this work suggest that, in principle, the employ of SPL's First Cut as an auxiliary reducing agent (with additional fluxing characteristics), is a technically feasible possibility. This is a preliminary conclusion based on simulations that must be confirmed by industrial tests and that impact on other parameters such as environment and economics must be investigated.. Initial exploratory industrial tests were made and the results, although essentially qualitative, are promising.

280

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

REDUCTION OF PFC

EMISSIONS AT POT LINE 70 KA OF COMPANHIA BRASILEIRA DE ALUMINIO

Henrique Santos1, Danilo Melo1, Jocimar Calixto1, Jefferson Santos1, Joäo Miranda1 Companhia Brasileira de Aluminio; 347 Moraes do Rego; Aluminio,Sao Paulo, 18125-000, Brazil Keywords: PFC, Anode Effect, Energy Consumption Abstract Following world tendency the Primary Aluminum Industry has been committed to the reduction of PFC emissions in order to reduce the greenhouse effect. Anode Effects is a significant source for such emissions. The Companhia Brasileira de Aluminio (CBA) pot line is equipped with 70 kA VSS Montecatini technology from the end of 60's with side feed pots, has been making efforts and developing actions to improve work practices and the operating process with the aim of reducing both the duration and frequency of anode effects. The anode effect, which was used for decades to assess the process behavior, ensuring their performance and preventing future problems, has been replaced by indicators supported by technological developments and the process computer improvement. Projects were developed and implemented as follow: Liquids control, Resistance control, Feed control, Anode/Cathode control and mainly by team work practices through operating training.

Figure 1. The Companhia Brasileira de Aluminio This paper was based on investigations and testing conducted at the 70 kA Pot Room that has 174 electrolytic cells with side feeding and went through the process of increasing current from 65 kA to 70 kA, from 2002 to 2005.

This paper shows achievements for projects cited, that caused significant impacts by reducing the duration and frequency of anode effects inline with world best practices for VSS technology. A reduction of 66 % of PFC emissions was achieved, from 1.2 t C02e per t Al in 1998 to 0.41 t COs e per t Al in 2009 calculated by Slope Method.

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Introduction The Companhia Brasileira de Aluminio, a company of Votorantim Metais, has an installed capacity to produce 475,000 tons of primary aluminum per year, with nine reduction lines totaling 1,508 electrolytic cells of VSS technology operating at 70 kA, 90 kA and 127 kA.

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Figure 2. Pot Room 70 kA form CBA The anode effect, previously considered essential for evaluating the operational performance of the pot line, has been losing its importance over the years as a best practice for pot operation. The systematic suppression of the anode effect was modified in order to reduce the duration from 1.86 min per cell day to 0.63 min per cell day (Figure 3), and thefrequencyof anode effect from 1.2 AE per cell day to 0.3 AE per cell day (Figure 4). The logic of feeding, which is held at intervals of four hours, has been changed

281

to reach 39 consecutive feeds. But for these reductions were possible, actions were necessary to reduce the number of anode problems, improve the liquid level control and operating procedures.

Figure 5. The Resistance Control System

The next steps for achieving further reductions targeted r improvement of routine operational and process control, because it was necessary to reduce the number of operational problems associated with anode quality.

Figure 3. Average Anode Effect Minutes per Cell Day

Reducing the number of anode problems

Development

The high number of operations related to problems with the anode was a hindrance to continue reducing the number of AEs to the levels proposed. In this, actions were taken in order to obtain an anode quality and stable operation. (See Table I and Figure 6)

Deployment of the reduction control system Up to 1998 the voltage control system was manual hampering an accentuated reduction of anode effects. The implementation of the resistance control system was the first step that led to a significant reduction of anode effect going from 1.1 AE/cell day to 0.7 AE/cell day in 2 years as shown in figure 4.

Table I. Table of anode improvement actions Problem Action Result Excessive anode Reduction of the Reduction of exposure below the size of the gas exposed anode gas collector duct. collector; from 25 cm to 18 cm Reduction of the depth of the pot shell. High number of Improvement of Reduction of fissure in the anode system of studs fissures in the replacement; anode;

Anode Effect Frequency

Increased frequency of stud recovery; Year

Anode quality

Figure 4. Evolution of anode effect frequency at Pot Room 70 kA This result is mainly due to the development of control parameters that became effective in the prediction of anode effects, providing an effective method that would allow reducing the number of anode effects to the levels achieved.

282

Changing composition of anode paste.

Reduced leakage of paste; Improved of the efficiency of stud replacement. Improvement of paste and anode quality.

to kill the AE.

Number of anode problems versus anode height 144

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The results achieved in reducing the emission of PFCs were significant, reducing emissions of 1.20 tonnesC02-e/tonneAl in 1998 to 0.41 tonnesC02-e/tonneAl in 2009 close to the estimated reduction 93% based on 1990 which provides for 0.34 tonnesC02-e/tonne Al in 2020[3] (Figure 7).

Figure 6. Reduction of anode problems Liquids control To achieve an extended period of time without the occurrence of an anode effect and to avoid occurrence of operational problems such as sludge, delay, and others, it was necessary to ensure a stable level of bath as well as an efficient bath chemistry control.

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Table II. Table of improvements of liquids control Problem Action Result Bath level control Implementation of Stability of bath daily measurements of level bath's level after metal tapping Bath chemistry Weekly sampling of Improvement of 1 control bath to measure superheat chemistry. control

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Figure 7. Trends in PFC Emissions per Ton Primary Aluminium Production. Source: 2008 Global Anode Effect Survey Results [3]

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To ensure the proposed results were achieved, actions were implemented with aim of improve operational practices such as to meet the operational standards, which were important to reduce the kill time of the AE. However, the most effective action was the operational training, which allowed the dissemination of knowledge of the automation tool, such as interpreting graphs of resistance and noise behavior, which allowed more rapid detection and efficient operator response, as well as standardization of the actions and awareness of the importance of each task in building the final result. (See Table III)

2 0 0 8 PFC E m i s s i o n s B e n c h m a r k i n g ( V S S )

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PFC Emission· (tonnes C02-*q/tonn« AI)

Figure 8. PFC Emissions Benchmarking (VSS) - Source: International Aluminium Institute - Anode Effect Benchmarking 2008 m The result is due to the reduction in the number of anode effects per year in the order of 75,000 AE/ year in 1990 to 19,000 AE / year in 2009, a reduction of 65000 AE/year which equates to a reduction of about 3.0 GWh/year in energy consumption taking

Installation of visual and audible alarms that provide the operators the information of time

283

into account only the anode effect frequency and a reduction of 66 % of PFC emissions from 1998 to 2009.

Conclusion The biggest challenge for succeeding in reducing PFC emissions was the preparation and awareness of staff to use new tools for operational analysis, to predict and deal with problems without the occurrence of anode effect. The result achieved has motivated the team to strive to improve even more, with the assurance that the problems occur, but can be predicted and studied.

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References

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1. Guimaraes, C. T., et al, Redugδo no Indice de Efeito Anodico Nδo Programado na Redugδo 4 da Albras, (Paper presented at I Seminärio de Reducäo do Aluminio, Säo Paulo, 2003).

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2. Anode Effect Benchmarking 2008, International Aluminium Institute

Anode Effect Frequency (ano<$e effects per
Figure 8. Anode Effect Frequency - Source: International Aluminium Institute - Anode Effect Benchmarking 2008I2]

3. Marks J., Bayliss C, 2008 Global Anode Effect Survey Results (Light Metals 2010), p259. 4. Results of the 2008 Anode Effect Survey, International Aluminium Institute

2008 AEM CD Benchmarkinj{VSS)

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Figure 10. Evolution in PFC emissions, AEM and AEF

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Towards Redefining the Alumina Specifications Sheet - The Case of HF Emissions 1

Linus M. Perander, 2Marco A. Stam, Margaret M. Hyland, Barnes B. Metson ^ight Metals Research Centre, The University of Auckland, New Zealand 2 Aluminium Delfzijl B.V., Delfzijl, The Netherlands

Alumina quality, HF emissions, Dry-scrubbing, Gibbsite, Residual hydroxide, Pore size distribution available for adsorption and the adsorption capacity, however it could be expected that the pore size distribution also plays an important role in providing access to internal reaction sites. The porosity and surface area arises from the structure of the transition aluminas which dominate smelter grade aluminas and reflect the incomplete conversion of gibbsite to alpha alumina in the calcination stage [5].

Abstract For smelting applications, alumina quality is typically defined in terms of chemical and physical properties, with emphasis on impurity elements, surface area, moisture content, particle size distribution and attrition index. However, these properties fail in prediction of the true HF generation potential, as well as the real capacity for HF removal in the dry scrubbers. Using plant measurements and additional laboratory characterization of a number of alumina samples a broadening of how alumina quality is specified is argued for. Measurements of the residual gibbsite/boehmite content and the pore size distribution, coupled with characterization of the alumina microstructure, can be used to predict and understand the generation of HF during feeding and dissolution as well as the ability to capture HF in the dry scrubbers.

Residual hydroxyls are an integral part of the transition alumina structures (and often reported as LOI). The less calcined the alumina, the higher surface area, and correspondingly, the higher the level of residual hydroxides in the structure. Together with other OH sources (such as gibbsite), these represent the main source of HF formation in the electrolyte [6]. In the cell HF is formed in the electrolytic process when the released water (or OH) reacts with the electrolyte (NaAlF3). It is then easy to see that there is a conflicting relationship between generation of HF from residual hydroxide and capturing the HF by increasing the surface area which inevitably results in more structural hydroxyls and increased HF formation in the first place.

Introduction For both regulatory and process stability reasons, reductions in fluoride emissions from aluminium smelters are an essential part of process improvement. The modern, injection type, dry scrubbing system comprises of a reactor (where alumina is injected and comes in contact with the collected cell off gases) and a separate filtration system (where the reacted alumina and other particulates are separated from the gas stream). These reactors are designed to enable good mixing between gas and solids and can be operated at very high efficiencies with over 99.5 % of the gaseous and particulate fluorides typically being captured [1]. With good pot hood collection efficiency, and the dry scrubber recovering and recycling most of the particulate and gaseous fluorides, the operation may almost be regarded as a closed loop between the cells and the scrubber [2].

The surface adsorbed water (represented by the MOI values) on the other hand are rapidly flashed off when the alumina is fed into the electrolytic cell [7]. This may actually be beneficial for the dissolution process as it helps with the dispersion of the material [8]. In this paper four alumina samples with very similar specifications, but which demonstrate very different HF emissions levels, are considered. The measurements currently reported on the alumina specification sheet fail to explain this HF generation potential or accurately predict dry-scrubber performance. By using low-cost, and often already existing, laboratory equipment measurements that better account for this may readily be obtained and used to better predict and control the operation, particularly with regards to emissions. This argues for some redefinition of the alumina specifications sheet, particularly when such specifications are used for example in prediction of emissions.

The chemistry and mechanism of the dry scrubbing process (i.e. the adsorption and reaction of HF with alumina) was explored in depth by Gillespie [3, 4]. Gillespie argued that since the fluoride adsorption capacity is related to both the specific surface area of the alumina and the relative humidity during the adsorption, the reactions must occur in a surface process. The author presented a mechanism which involved several steps, starting with the adsorption of water on the surface of the alumina followed by HF adsorption and acidification of this surface layer which would dissolve the alumina surface and form A102" and AlO2" species and finally the precipitation of oxy- and hydroxyfluorides. Most importantly the reaction was shown to be irreversible under dry scrubbing conditions (temperature and atmospheric), an important finding which enabled new developments of the dry scrubbing process and control strategies to be made.

Experimental The samples examined in this study were obtained during the unloading of the respective alumina shipments and directly from the conveyor belt going to the alumina storage silo. There is only one alumina silo on site which rules out alumina blending as a source of variation. The GTC (Gas Treatment Centre) inlet and outlet HF concentrations are monitored continuously using NEO laser-based gas monitors. This is an infrared single-line absorption spectroscopy measurement which provides a continuous reading of the HF content in the gas stream across the laser path. One laser is mounted in the stack directly after the GTC, and one in each inlet channel to the GTC. The inlet values reported here are averages between these two inlet channels.

Although the mechanisms for the reaction between HF and alumina (in the dry scrubbing process) are relatively well understood, the role of alumina microstructure and porosity is less clear. There is a direct relationship between the surface area

285

TGA experiments were conducted to determine MOI and LOI values, using a Shimadzu TGA-50 apparatus. Approximately 35 mg of each sample was heated, in an open platinum crucible in air, with a constant heating rate of 20 °C min"1 from ambient temperature to 1100°C. A baseline curve was also produced, and subtracted from each sample curve, by recording the weight change when an empty platinum crucible was subjected to the same heating conditions as above.

Table 1. Alumina properties as reported on the specifications sheet for the 4 alumina samples used in this study.

Values for MOI and LOI were obtained based on thermo gravimetric data. As per the ISO standard, the ranges for MOI and LOI were taken as 50-300°C and 300-1000°C, respectively [9]. It should however be pointed out that according to the standard measurement for MOI and LOI values the experiments are performed with a retention time of 2 hours at 300 and 1000°C. The results presented here may therefore not correlate directly to measurements performed according to standard methods but can be used to compare the samples and for calculating the water and hydroxyl content.

Alumina A

Alumina B

Alumina C

Alumina D

LOI (300-1000
0.59

0.51

0.79

0.91

68.0

66.3

75.6

73.0

N/A

N/A

4.0

5.0

1.6

1.2

1.9

1.6

12.0

13.0

4.8

10.87

Bulk Density kgdm-1

1.04

N/A

1.00

0.98

Note that the parameters omitted from the table were not reported on the specifications sheets Table 1 indicates that aluminas A and B have relatively similar parameters (in terms of MOI, LOI, BET-SA and fines content) which is also to be expected as these samples represent different shipments from the same alumina refinery. Aluminas C and D are different compared to A and B, as they both have higher specific surface areas and LOI values. Alumina C also has a significantly lower attrition index than the rest of the samples.

Nitrogen porosimetry experiments were conducted to evaluate the specific surface area and pore volume, as well as pore size and pore size distribution. The measurements were performed on a TriStar 3000 apparatus. Approximately 1 gram of sample was used for the analysis. For the measurements the sample is initially degassed under a nitrogen purge at 120°C for 12 hours. Note that in standard analysis different degassing temperatures may be used. The temperature of 120°C was chosen to avoid the reaction of any gibbsite or boehmite present in the sample. The results obtained here thus represent the whole sample as received (without changing its phase composition). The measurements were conducted under liquid nitrogen temperatures and using nitrogen as the adsorbed gas. The BET method [10] was used to evaluate surface area and the BJH method [11] was used for the pore volume and pore size and distribution evaluations.

Based on the parameters reported in table 1 one could expect aluminas A and B to result in similar HF emissions and C and D to perform equally well but perhaps with a higher background HF emission level due to the higher LOI values. Some of the additional HF generated due to the larger amount of structural hydroxyls present in aluminas C and D could be expected to be offset by the significantly higher BET surface areas. However, as Table 2 indicates, these aluminas did not perform as expected. When alumina A was used the fluoride emissions increased by approximately 60 %, compared to alumina B. Similarly, alumina D resulted in twice the emissions compared to alumina C; far more than what could be expected based on the relative LOI values.

X-ray diffractometry was performed on a Rigaku apparatus using Cu K-alpha X-rays. The analysis was performed using a 2Θ angle of 10-80° and a step width of 0.02°. Rietveld refinement [12, 13] of the diffractograms was then performed using the FuUProf software [14]. The fine sub 40 micron size fraction was separated from the bulk by sieving and analysed separately.

Table 2. Associated hydrogen fluoride emissions (total gaseous per hour andfluorideper cubic meter) and scrubber efficiencies. Fluoride Emission mgm"3 Total Gaseous HF Emission kg h'1 Scrubber Efficiency (%)

X-ray diffractometry with Rietveld refinement allows the alumina phases present in smelter grade aluminas to be quantified. However this technique does not allow the amorphous phases such as rho- or chi-alumina that may affect the performance of the alumina to be measured directly [15, 16]. The highly crystalline phases such as alpha alumina and gibbsite produce sharp, welldefined peaks whereas the less crystalline transition alumina phases produce broader and more diffuse peaks. Several studies report on how the alumina phase composition influences the smelter performance (see for example [17] and references therein).

Alumina A

Alumina B

Alumina C

0.86

0.61

0.56

Alumina D 1.13

1.29

0.92

0.60

1.21

N/A

N/A

99.70

99.48

Since HF is generated in an electrochemical reaction between the bath and any hydroxyl species (H 2 0, OH", NaOH...), a reasonable measure of the HF generation potential is a quantification of the water and hydroxyl content of the aluminas. This is typically done in so called Loss on Ignition and Moisture on Ignition measurements [9], but may also be obtained using thermogravimetric analysis techniques. As will be shown the LOI measurement (300-1000°C) does not account for all the OH associated with the alumina.

Results and Discussion Table 1 displays a selection of properties from the alumina specification sheets for the 4 alumina samples discussed here. It is recognized that other properties are of importance for understanding alumina performance; however, these properties have been chosen as they are most often associated with dryscrubbing and HF generation, which is the topic of this paper.

Figure 1 shows the obtained thermograms for aluminas A and B as well as the thermogram for a standard industrial Bayer gibbsite. It can be seen that the relative weight loss is larger for alumina A, which indicates higher MOI and/or LOI values, and as can be seen from Table 1 the reported LOI value was slightly higher for

286

sample A. The higher LOI value indicates more structural hydroxyls (OH groups retained after the calcination in the Bayer process) and thus a greater HF formation potential. What may also be seen is a sharp weight loss at around 270°C for both aluminas; this indicates the presence of residual gibbsite (or un-calcined material) in the alumina. When heated, gibbsite starts to transform to boehmite and/or transition alumina at around 250°C (depending on the reaction conditions). In the thermogram below (figure 1, blue line) this can be seen as a rapid loss of mass (starting at around 250°C) as the OH-groups in the gibbsite reacts to form H 2 0. The residual gibbsite in the alumina has the potential to react with the electrolyte to form HF, but is obviously not accounted for by the LOI (300-1000°C) value. The MOI (50-300°C) value is typically associated with, and dominated by, the surface water/moisture, although, when gibbsite is present it will be included in the MOI value.

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400

600 800 T e m p e r a t u r e (°C)

1000

|-0.01£g

I

1200

Figure 2. Thermogram for alumina A with the MOI and LOI limits indicated. The blue line is the derivative weight loss (wt-% °C"1) which is a way to better identify specific phase transitions. To estimate the HF generation potential of alumina, the LOI value provides a reasonable estimate of the amount of reactive OH, particularly when the extended range (250-1100°C) is considered. However, when gibbsite and/or boehmite are present in the alumina these components should also be quantified and reported. For this a number of methods have been suggested, such as: DSC [18] Near Infrared Spectroscopy [19] and quantitative XRD, employed in this study and elsewhere [15, 20]. Although readily available, these methods have not been widely adopted in the industry for routine characterisation. Table 3 provides this comparative data for the samples considered in this study.

400

600 Temperature (eC)

Table 3. Additional characterisation data for the four aluminas: MOI (to 300°C) and LOI (300-1000°C and 250-1100°C) values (by TGA), surface area, average pore size and total pore volume (by nitrogen porosimetry), and gibbsite and alpha alumina content (by XRD with Rietveld refinement).

Figure 1. Thermograms for the two of the alumina samples: alumina A (red dashed line, left y-axis) and alumina B (green line, left y-axis) as well as a reference gibbsite sample (blue solid line, right y-axis). Thus, in order to account for the gibbsite as a potential source for HF formation, the temperature limits for the MOI and LOI values need to be re-evaluated. When gibbsite is present in the alumina a more appropriate measurement of the surface water/moisture (less likely to be transported into the electrolyte and react to form HF) is perhaps the weight loss between 50 and 250°C. An LOI value from 250 to 1100°C would then account for both the residual OH and the gibbsite, the primary HF sources. The 'new' limits can be seen in Figure 2. Note, however that although this is perhaps a more useful way to specify the MOI/LOI limits when gibbsite is present, the 'exact' limits for when surface water and/or residual OH is removed cannot be specified as these overlap to some extent. Also, as will be discussed later, not all H 2 0 released from the gibbsite is likely to react with the bath to form HF (the gibbsite dehydroxlation reaction at high temperatures is very rapid and some water will flash off without coming in contact with the bath).

MOI (to 300 °C) wt-% LOI (300-1000
Alumina D I

Alumina A

Alumina B

Alumina C

2.061

1.954

2.478

2.645

0.722

0.682

1.014

1.052

0.990

0.917

1.523

1.448

68.1

69.0

70.7

76.1

5.78

5.70

8.78

11.20

0.210

0.215

0.193

0.207

0.93

0.56

1.98

2.65

8.1

7.6

3.9

6.5

The inevitable effect of residual gibbsite is that a large amount of water (as OH) enters the cell with the alumina. As can be seen from the thermogram in figure 1, approximately 35 wt-% of gibbsite will react to form H 2 0 when heated. This will occur very rapidly when it comes in contact with the electrolyte, inducing a phase transition into the transition aluminas. The phase transition, will be competitive with dissolution but the very rapid formation of large amounts of water vapour will create a volcano effect, carrying fine particles out of the cell and potentially increasing dusting.

287

Assuming that all OH in the gibbsite reacts to form H 2 0 and A1203 (according to reaction 1), 2AI(OH)3(s) + Ei -> Al203(s) +3H20(g)

IM

where ¸χ is the reaction energy, a theoretical increase in the energy demand can be calculated. Based on the amount of water that is evolved the additional (apart from the structural hydroxyl, measured by the LOI tests) HF generated due to the gibbsite in the SGA can also be calculated. To calculate the increased HF generation capacity due to gibbsite in the alumina it is assumed that all H 2 0 will react according to reaction 2: 3H20(g) + 2AIF3(g/s) -> 6HF(g) + Al203(diss)

For comparison two measured HF emission values for aluminas with known MOI and LOI values are included. Not surprisingly the measured values are slightly lower than the calculated ones, but should still provide a reasonable estimate of the HF generation capacity. The measured values were sourced from the literature [7]. In the study dry and 'hydrated' alumina samples were fed to a plant cell and the resulting HF emissions measured. It can be seen that the MOI, or moisture content, has little or no contribution to the HF emissions. Figure 4 displays HF measurements (GTC inlet and outlet levels) when aluminas C and D were used. The higher inlet GTC HF concentration when alumina D was used may be attributed to the slightly higher gibbsite content. However, the gibbsite content alone could not explain the significantly poorer performance of alumina D in terms of HF removal. To understand the performance of these aluminas the porosity and microstructure needs to be evaluated as well. A direct relationship between specific surface area and HF adsorption capacity has been reported [3]. High specific surface area is often considered advantageous in terms of optimal dry scrubber performance, and can often be the only way to deal with increased fluoride emissions. There is also a relationship between surface area and residual hydroxyl content. The residual hydroxyls are retained in the crystal lattice of the transition alumina phases due to incomplete calcination of gibbsite. In other words, higher surface area stems from more of the low order transition aluminas (gamma, chi, rho alumina) which also have more residual hydroxyls retained in the lattice and hence also have a higher HF generation potential.

121

In practice it is not likely that all H 2 0 will react to form HF (as some will obviously exit the cell with the off gases). It should be noted that the HF lost to the environment (due to the incomplete capture) constitutes a direct loss of bath, which will have to be supplemented. Depending on the gibbsite content this may have a significant economic impact if particulate and gaseous emissions are not fully captured. XRD analyses revealed that 0.93 and 0.56 wt-% of alumina A and B, respectively, is gibbsite. Assuming that 1.92 kg of alumina is needed to produce 1 kg of aluminium metal it is then calculated that 18 and 11kg of gibbsite is introduced, for alumina A and B, respectively, to the cell per metric ton of Al produced. Based on the reactions above we can then also calculate that the 0.93 and 0.56 wt-% of gibbsite in the aluminas results in an additional 8.2 and 5.0 kg of HF being formed per ton of Al metal, for alumina A and B, respectively, on top of the 'background' HF from the structural hydroxyls (reflected in the LOI value). The additional HF due to the presence of (0.93 wt-%) gibbsite is about 20% of the total (theoretical) HF. Although these are theoretical maximum HF levels (based on complete reactions) it can be seen that even small amounts of gibbsite in the alumina have the potential to significantly increase the total HF burden.

300

GTC Inlet HF Concentrations and HF Emissions using Alummas C and D (5 days data)

Calculated and Measured HF emissions from structural hydroxyl (LOI) in kg(HF) per T(AI) 50 45 40

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Calculated HF Alumina A

0.8

,

0.9

,

2000

Figure 4. Dry scrubber (GTC) inlet HF concentrations and corresponding HF emissions when using aluminas C and D (5 days data).

,

The pore size distribution and mean pore size, as well as surface area are a measure of the extent of the calcination reactions [5]. As the calcination reactions progress, the pore size grows, but pores coalesce, causing the surface area to decrease as a result. A narrow pore size distribution in a range that allows HF molecules to be easily transported to internal surfaces and reactive sites is desired and results from a well controlled calcination process. Pore size distribution is also an effective way to identify if the

LOI (wt-%)

Figure 3. Theoretical HF generation capacity (calculated based on equations 1 and 2) and measured HF generated by structural hydroxyls in the alumina (numerical values sourced from ref. [7]). In figure 3 calculated (solid line), based on reaction 2, HF generation capacities are plotted as a function of their LOI values.

288

alumina is blended. A mixture of two differently calcined aluminas, or an alumina containing under- or over calcined material will result in a broad pore size distribution or a distinctly bimodal distribution. For a well-controlled calcination process, a narrow pore size distribution centred around 6-8 nm is typically observed, and desired (see figure 5). Higher calcination temperatures or longer calcination times usually result in an increased pore size and highly over-calcined material (more thetaalpha alumina) shows up as a peak at >10 nm pore sizes. Material with a lower degree of calcination results in a lower average pore size and generally shows as a peak in the <2 nm pore size range on the distribution plots. Poor control of the calcination process results in a broader peak in the 4-10 nm pore size range. +

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Figure 6. More than one way to achieve a surface area target. Alumina D displayed a distinctly bi-modal pore size distribution, indicating the presence of under-calcined components as well as over calcined material. From an operational viewpoint the blend of over- and under calcined material is detrimental for a number of reasons. As mentioned previously the low order transition aluminas contain more residual hydroxyls (confirmed by the LOI measurements) resulting in more HF generation upon dissolution. At the same time, the fine pore size is likely detrimental for HF removal. The narrow pores restrict access to internal porosity and readily become blocked (when HF reacts to form oxy-fluorides). This further restricts access to internal sites, thus reducing the capacity and rate of HF absorption. Moreover, the over-calcined components in the sample are likely to have a poor dissolution rate. Thus, the currently used surface area specification, derived under equilibrium conditions is not ideal for predicting scrubbing efficiency and it would be helpful to extended this to include information about the pore size distribution. This information is readily available using most of the commercial surface area instruments by examining the nitrogen adsorption isotherms using the BJH method.

10.0 Pore Diameter (nm)

Figure 5. BJH pore size evaluations for alumina C reveals a 'normal' pore size distribution. As reported in table 3, the surface area measurements revealed that alumina D had a slightly higher surface area than alumina C (76.1 and 70.7 m2/g, respectively) which could be assumed to improve the dry-scrubbing efficiency and offset some of the additional HF that is likely to be generated by the higher LOI content. However, when the pore size distribution plots are examined (figures 5 and 6) some interesting differences between the two samples can be seen. Alumina C shows one peak at a pore size of approximately 8 nm whereas alumina D clearly has a bimodal pore size distribution with one peak centred below the 1.0 nm pore size limit of this technique, and the other centred at a much larger pore size of approximately 12 nm. This would indicate that alumina D contains a large quantity of undercalcined material and that the rest of the sample is over-calcined (rich in theta and alpha alumina). As the small pores contribute more to the total (or average) surface area the alumina D therefore has a larger nitrogen specific surface area. Note that by controlling the calcination conditions, surface areas in excess of 350 mV 1 may be achieved.

Conclusions Plant measurements and additional laboratory characterization of a number of alumina samples was used to highlight some of the shortcomings of the alumina specifications sheet when it comes to understanding alumina performance, particularly with regards to HF generation and emissions as well as feeding and dissolution characteristics. It was demonstrated that residual gibbsite in the alumina results in increased HF levels (which coupled with poor hooding/collection efficiency results in an additional loss of electrolyte) and feeding perturbations. Therefore, knowing the amount of gibbsite in the alumina will help to predict and control the process response to the alumina quality. Several low cost options for determining the gibbsite content in alumina exist (XRD, NIR, DSC or TGA for example) and should be adopted more widely.

289

Nitrogen porosimetry and investigations of alumina microstructure revealed that the low order transition aluminas (gamma, chi, rho alumina) have a limited pore size and plant measurements indicate that this is detrimental for the dryscrubbing process. It seems that the narrow pores, although resulting in large average surface areas, may become blocked thus restricting access to internal surfaces which effects the HF adsorption capacity adversely (at the same conditions). These observations indicate that a pore size distribution specification would be a better requirement for alumina quality (than a single surface area target) as this ensures good scrubbing and may even be energetically favorable in the alumina refinery. The pore size distribution may readily be measured through nitrogen adsorption techniques using existing equipment both in the smelters and alumina refineries.

12.

Acknowledgements

16.

The authors would like to thank Anita Folkers and Karin Bouwmeester for their assistance with the alumina samples and HF measurements, the interpretation of the results and their contributions to this manuscript.

13. 14.

15.

17.

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3. 4. 5. 6. 7.

8. 9. 10. 11.

Well wood, G.A., The practice of dry scrubbing. Light Metals (Warrendale, PA, United States), 2001: p. 371377. Zhang, W., X. Liu, P. McMaster, and M. Taylor, Modeling of impurity balance for an aluminum smelter. Light Metals (Warrendale, PA, United States), 1996: p. 405-11. Gillespie, A.R., Mechanistic Studies of HF Adsorption on Alumina. PhD Thesis, The University of Auckland, New Zealand, 1997. Gillespie, A.R., M.M. Hyland, and J.B. Metson, Irreversible HF adsorption in the dry-scrubbing process. Journal of Metals, 1999. 51(5): p. 30-32. Perander, L., Evolution of Nano- and Microstructure During the Calcination of Bayer Gibbsite. PhD Thesis, The University of Auckland, New Zealand, 2010. Patterson, E.C., Hydrogen Fluoride Emissions From Aluminium Electrolysis Cells. PhD Thesis, The University of Auckland, New Zealand, 2002. Hyland, M., E. Patterson, and B. Welch, Alumina structural hydroxyl as a continuous source ofHF. Light Metals (Warrendale, PA, United States), 2004: p. 361366. Haverkamp, R.G., Surface studies and dissolution studies of fluorinated alumina. PhD Thesis, The University of Auckland, New Zealand, 1992. Aluminium oxide primarily used for the production of aluminium - Determination of loss of mass at 300° C and 1000°C. ISO 806:2004. Brunauer, S., P.H. Emmett, and E. Teller, Adsorption of gases in multimolecular layers. Journal of the American Chemical Society, 1938. 60: p. 309-19. Barrett, E.P., L.G. Joyner, and P.P. Halenda, The determination of pore volume and area distributions in porous substances. I. Computations from nitrogen isotherms. Journal of the American Chemical Society, 1951. 73: p. 373-80.

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Rietveld, H.M., Line profiles of neutron powderdiffraction peaks for structure refinement. Acta Crystallographies 1967. 22(1): p. 151-2. Rietveld, H.M., Profile refinement method for nuclear and magnetic structures. Journal of Applied Crystallography, 1969. 2(Pt. 2): p. 65-71. Rodriguez-Carvajal, J. FULLPROF: A Program for Rietveld Refinement and Pattern Matching Analysis, in Abstracts of the Satellite Meeting on Powder Diffraction of the XV Congress of the IUCr. 1990. Toulouse, France. Perander, L.M., Z.D. Zujovic, T. Groutso, M.M. Hyland, M.E. Smith, L.A. O'Dell, and J.B. Metson, Characterization of metallurgical-grade aluminas and their precursors by 27AI NMR and XRD. Canadian Journal of Chemistry, 2007. 85(10): p. 889-897. Perander, L.M., Z.D. Zujovic, M.M. Hyland, M.E. Smith, L.A. O'Dell, and J.B. Metson, Short- and longrange order in smelter grade alumina - Development of nano- and micro structures during the calcination of Bayer Gibbsite. Light Metals (Warrendale, PA, United States), 2008: p. 29-35. Perander, L.M., Z.D. Zujovic, T.F. Kemp, M.E. Smith, and J.B. Metson, The Nature and Impacts of Fines in SGA. Journal of Metals, 2009. 61(11): p. 33-39. Wehrli, J.T. and A.R. Kane, Application of differential scanning calorimetry for characterization of Bayer process solids. Light Metals (Warrendale, PA, United States), 1993: p. 175-84. Dando, K.R. and N.R. Dando, Rapid, non-destructive analysis of % gibbsite in smelting grade alumina. Light Metals (Warrendale, PA, United States), 2010: p. 531533. Whittington, B. and D. Ilievski, Determination of the gibbsite dehydration reaction pathway at conditions relevant to Bayer refineries. Chemical Engineering Journal (Amsterdam, Netherlands), 2004. 98(1-2): p. 89-97.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DESIGN OF EXPERIMENT TO MINIMIZE FLUORIDE AND PARTICULATE EMISSIONS AT ALUMAR Eliezer Batista1, Edson Montoro2, Paulo Miotto2, Luciano J. P. Souza2 Department of Potrooms and Technology 1 Alcoa Technical Center, 100 Technical Drive, Alcoa Center, Pa 15069 2 Consorico de Aluminio do Maranhäo - Alumar, BR 135, Km 18, Säo Luis, Brazil, Zip code 65065-604 Keywords: Fluoride Emissions, Particulate Emissions, Design of Experiment

was decided to make a full factorial experiment 2k at Alumar to find out the main factors and the size of their impacts. The objectives of this work are to develop statistical models to predict the impacts of operational activities on fluoride and particulate emissions and recommend actions to minimize them.

Abstract Most of Aluminum plants have been struggling to minimize the fluoride and particulate emissions at the main source, which is the pot rooms, in order to reduce the environmental impacts. Nowadays, this challenge has been more difficult to be accomplished because of some reasons, such as: amperage increase, alumina quality deterioration and pot room expansions. Alumar, one of Alcoa's units, following a corporate vision, is continuously searching for alternatives to eliminate or reduce the environmental impact caused by operations. Many actions and studies are in place currently. This full factorial experiment 2k was done with the aim of identifying the main factors and their impacts on fluoride and particulate emissions. The statistical model is showing that the fluoride emission has been affected mainly by Pot Draft, Pot Dressing, and the Usage of Compressed air for Housekeeping with R2 at 82%, and for particulate at 58%. Based on the models, certain actions were recommended to minimize both of these emissions. In addition, this paper describes, step by step, how this kind of experiment can be applied to the Aluminum industry.

Experimental This study was started up at Alumar with a brainstorming of the possible contributors of fluoride and particulate emissions in the potroom. It was done by a very experienced team with people from several areas such as EHS, bath handling area, laboratory and potroom (operators, technicians and engineers). By using statistics and subject matter knowledge, the team hypothesized that the potential causes would be the following: o o o o o o o o

Introduction Alcoa has been recognized by governmental and nongovernmental agencies around the world for their accomplishments related to environmental protection and sustainable growth. Because of this, the understanding and minimization of fluoride and particulate emissions is always a subject of intense interest. Alumar, one of Alcoa's units located in the northeastern region of Brazil, following this corporate vision, is continuously researching the root causes of fluoride and particulate in order to minimize them. Previous works were published about it, such as a study developed to map and determine the impact of each operation (anode setting, tapping, among others) in the fugitive fluoride emissions in the pot rooms [1]. It is already known that the fluoride and particulate emissions are affected by several factors inside the potroom such as bath temperature and ratio (bath chemistry), alumina quality, amperage level, operational activities, pot draft, etc. For example, it was shown the importance of the anode cover integrity to prevent gaseous and particulate fluoride evolution in another Alcoa research [2], which says that the crust integrity was capable of changing the amount of fluoride evolution by more than 1000% while the bath chemistry changes could only change the amount of evolution by 20%. Since fluoride and particulate emissions are impacted by several variables and their effects depend on the interactions of them, it

o o o o o o o o o o

Amperage - load creep Usage of compressed air for housekeeping Increase the number of days between pot dressing Analysis error - Lack of reliability at the Lab Sampling and preparation error - Lack of reliability at the Lab Stuck material and open hole in the ducts of gas Lack of Pot Draft Delays to remove the butts from potroom after anode settings Inefficient cleaning machines for housekeeping High alumina LOI High bath temperature Low bath ratio Excess of Burn-off (Anode failure requiring an extra set) Excess of anode effects Deficient balance of pot draft Pot superstructure cover is not maintained at the right position High amount of fines in the anode cover High percentage of alumina in the anode cover

Variable Definitions The most likely causes were selected from this list, based on experience and previous work of the team. The decision was to study the effect of Pot Dressing, Usage of Compressed Air for Housekeeping and Pot Draft. The pots at Alumar are currently dressed every 6 days by the Pot Tender Operator. This activity is really important to keep the anode cover integrity and consequently to minimize the emissions.

291

Housekeeping is made every shift in a group of pots at Alumar. Compressed air has been used to clean some components of the pots to guarantee the minimal TPM (Total Preventive Maintenance) requirements related to cleanliness. Alumar uses A-398 reactors for gas treatment. Pot draft normally ranges from around 3500 to 4000 ACFM. This reactor has a very good efficiency, but the fugitive emission increases in the potroom if the pot draft is below a minimal level.

Results The experiment was set up and run one time for each combination. Total fluoride and particulate emissions were collected for a period of 8 hours using an isokinetic test, based on US EPA Methods 13 (Determination of total fluoride emissions from stationary sources) and 14 (Determination of fluoride emissions from potroom roof monitors for primary aluminum plants). Every 8 hours the reacted solution of NaOH 0.1N was analyzed, using the ion selective electrode method and the experiment was moved to new settings. Table 3 is showing the results for fluoride emissions and Table 4 for particulate emissions. The experiment was run one time for each combination.

Statistical Experiment The interaction among the variables and their effects on the fluoride and particulate emissions needed to be known for better understanding of the causes. It was decided to make a full factorial experiment 2 3 , three factors with two levels for each one, as can been seen in Figure 1. By definition Factor A is the variable Usage of Compressed Air for Housekeeping, Factor B is Pot Dressing Frequency and Factor C is Pot Draft.

Table 3- Fluoride emission results for each combination 23

(+)— Factor (C)

cC

+

B-

B+

b

a

ab

be

ac

abc

(1) c

0.51

0.622

0.554

0.675

c+

0.188

0.272

0.271

0.652

c

Table 2 presents the level for each variable. The experiment was run with these levels for both response variables (fluoride and particulate emissions).

Low level

High Level

A Compressed Air for Housekeeping

Don't use compressed air

Use compressed air as usual

B Pot Dressing Frequency

Pol dressing every 2 days

Pot dressing every 5 days

C Pot Draft

3000 ACFM

43O0ACPM

B

B+

B-

B+

1.043

1.592

2.282

1.620

0.404

0.434

0.916

2.454

Y=0.468+0.07A+0.087B-0.122C

Table 2- Level definition for each variable Factor

A+

A-

The statistical analysis of experimental data was done using the software DOE-KISS (®-Air Academy Associates, Colorado Springs, CO, USA). The following table (Table 5) is presenting the results for fluoride emissions. The interaction among the studied factors doesn't impact the fluoride emission. F value is above the Sig F (for Alpha = 5%), which means it is a significant model. The factors Pot Draft (C) and Pot Dressing Frequency (B) are really relevant for Fluoride Emissions, while the Usage of Compressed air for Housekeeping (A) is in a "gray zone". Nevertheless, the latter was confirmed as a significant factor using others techniques such as Yates. It is recommended to maintain it as part of the final model. The model has a high strength with R2 at 0.8 and the prediction equation for it can be seen in equation 2 (coded effects).

A+ B

C

+

Table 1- design matrix for experiment 2 3 with all runs B-

B+

C

Xijki = μ + Ai + Bj+ ABy + Ck + AC* +BCjk + ABCijk + Z,(ijk) (1)

+

B

( ♦ )

Each combination was tested once in a period of 8 hours. Table 1 summarizes and gives a better understanding of how the data will be collected. The equation 1 describes the expected response model for this kind of experiment. Table 1 shows a design matrix for experiment 2 3 with all runs.

A-

B+

23

I I Factor (B) (-) (+) Figure 1 - Graphical depiction of full factorial DOE 2

23

B

Table 4- Particulate emission results for each combination

— (+) Factor (A) —

A+

A-

292

(2)

Table 6- Statistical analysis for the variable particulate emission

T a b l e 5- Statistical analysis f o r t h e v a r i a b l e f l u o r i d e e m i s s i o n Multiple Regression Analysts

Multiple R e g r e s s i o n Analysis

'Tltr^T 1 1

1

Factsr Const A B C

Name

Coeflf i 0,46800 Jateamento 0,07000 Rastelamenlo 0,08725 Pot Draft -0,12225

Rsq

0,8283

Adj Rsq

0,6995

StdError

0.1067

F

6.4306

SigF

0,0520

1

Pf? Tail) i Tol ! < 0.0002 0,1371 1 : X 0.0818 1 ; X 0.831'/ 1 : X

Õ-ËÌ Model 1

I Factor I Name

1 I I j

A B

C

Low High Exper

i Jateamento ! Rastelamento

i Pot Draft

-1 -1

- 1 : 1

1 1

0 °

0

I \

1

| j

Source

Nam«

(:*«« Pt2TMt;T«4 1.3431? OJÖOli

-imwma Qi7m

A C

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-03113

0.4205

. · , _ _ ·_

9»m

< 1 [Factor

1 K

0,2231

A B C

1

Prediction Ar$Rs<j I

| Y-hat !Std Error

SS


MS

0,2

3

0,1

Error

0.0

4

0,0

Total

0.3

7

Lower Bound Upper Bound

0,1478809 0,7881191 I

I

S.gF

I

&<►««#

♦•if-

MS

$

0.4

Low High Exper 0 0 0 .· Ã

Prediction Y-hat StdError I

1,343125 0,5921187

99% Prediction Interval

11

I R*(jr*wjfe« '■ 1J9 ' · :

Pareto chart in Figure 2 indicates that the effects Pot Draft (C) and Pot Dressing Frequency (B) are the most important, while the usage of air for housekeeping (A) has a lower impact in the fluoride emissions.

A2

.:....,...vi...,:J

7 ' ♦ · ! " · ♦

I

I . .

Lower Bound Upper Bound

·. . L.

«0,433231 3.119481

. j .:....

j ,j

Pareto chart in Figure 3 indicates that the Usage of Air for Housekeeping (A) is really the most important effect. Pot Draft (C) has some impact but it is much lower than Factor A.

Fluoride Emission Abs. Value of the Factors

n_n Particulate Emission Abs. Value of the Effects

0.7 0.6

|o.5-

JKKHKM

| O

0.4-

J«

|

0-20.1

Comp. Air fur HouseUeping

0

^

l_

Comp. Air for Housekteping

Figure 2 - Pareto chart for absolute values for each effect for fluoride emissions

Draft

Lr

Effect Name

Figure 3 - Pareto chart for absolute values for each effect for particulate emissions.

The results for particulate emissions are as shown in Table 6. The interaction among the studied factors doesn't impact the particulate emissions. As in the fluoride model, the F value is above the Sig F (for Alpha = 5%), which means it is a significant model. The factor Usage of Compressed Air for Housekeeping (A) is the most relevant for this response variable. Pot Draft (C) is in a "gray zone", however, it was confirmed as a significant factor using others techniques such as Yates. Pot Dressing Frequency (B) is not a significant factor for particulate emissions. The model has a reasonable strength with R2 at 0.5, regarding it is an industrial experiment, and the prediction equation for it can be seen in equation 3 (coded effects). Y=l. 34+0.47 A-0.29C


Name

: Miwfmto-ú S 1 Rast*r3«r«nto A 1 Draft Ë-.Á ·| Ë i

I

B ^ ^ - ^ w i l ^ i

T*Wl

«Dr··»«

_'_ ♦:

0468 0,1067064

99% Prediction Interval 1 Recession

factor COM

Conclusions and Recommendations According to the statistical model, fluoride emission can be significantly reduced increasing the Pot Draft (main factor) and improve Pot Dressing as much as possible. The worst condition would be to set the Pot Draft to -1 and the Usage of Compressed Air and Pot Dressing to +1. Considering this condition, the fluoride emission would be 0.74 kg/tAl; and, if the settings were the opposite values mentioned above, the result would be 0.18 kg/tAl. Particulate emission can be significantly reduced if the Usage of Compressed air for Housekeeping is minimized or avoided. Setting the effects to maximize the particulate emission (Usage of Compressed Air to +1 and Pot Draft to -1), will make the particulate emission 2.1 kg/tAl and 0.57 kg/tAl to the opposite settings. The team looked for alternatives to move the factors toward lower emissions.

(3)

293

Fluoride Emission Applying Recommendations

Pot Draft By removing the stuck material and closing the holes in the external ducts of gas, the pot draft can be improved by 500 ACMF. Figure 4 shows the current condition of the duct of gas after more than 20 years of operation and the state of it after the cleaning process.

m A

fl

rfl

^ψ

Λ :;j9B

:

Before Cle*ining

Comet

Condrfion (I>pical)

R*commend»non_2nd ssnphng

Figure 6 - Fluoride emission results under recommendation conditions for a period of 48 hours. Although the model shows 0.23kg/tAl after applying the recommendations for a sampling period of 48 hours, the measured results were 0.34 and 0.37 kg/tAl. It is clear that the implementation of those actions will result in lower fluoride emission.

Afte r Cleaning

Figure 4 - External gas ducts before and after the cleaning process Pot Dressing

Particulate Emission Applying Recommendations

It is possible to improve the pot dressing by reducing the frequency from every 5.3 to 3 days, without increasing the head count. Currently, there are four Pot Tenders by room and they are responsible for the dressing and other operational activities. Their production hours can be optimized if the dressing is done by two operators, and the additional activities are done by the others, as can been in the following figure (Figure 5). Current Condttan - 8 pots dressed by shift by room - Every 5,3 days

JED

Current Condition (Tipical)

Picommendauon. 1 it samphng

Recommendaaon_2nd sampling

Figure 7 - Particulate emission results under recommendation conditions for a period of 48 hours.

Target Condition - 1 4 pots dressed by shift by room - Every 3 days

There is a good agreement between the measured particulate emission and model predictions under the recommendation conditions for a sampling period of 48 hours. As mentioned for fluoride emissions, it is clear that the implementation of those actions will also result in lower particulate emission.

in tuip«$&·* ««nodiijäes

Figure 5 - Pot Dressing Frequency to reduce the emissions.

Acknowledgements Colleagues Joel Cämara, Elisio Bessa and Ciro Kato are acknowledged for their contribution, support and valuable discussion about this subject.

Usage of Compressed Air for Housekeeping It can be reduced by at least 50% through the usage of cleaning machines and changes on the TPM requirements.

References

It was decided to set the conditions according to those recommendations and measure the emissions to check the benefits of those recommendations and to compare the results with the model's predictions. Figure 6 shows the results for fluoride emission and Figure 7 displays the results for particulate emission. The sampling time was 48 hours for both samples.

294

1.

Nagem, N., Batista, E., Silva, A., Gomes, V., Venäncio, L., Souza, L., Understand the Fugitive Emissions at Alumar, TMS, Light Metals 2005, pp. 289-292

2.

Tarcy, G.P., The Effect of Pot Operation and Work Practices on Gaseous and Particulate Fluoride Evolution, TMS, Light Metals 2003, p 193-198.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INNOVATIVE DISTMBUTED MULTI-POLLUTANT POT GAS TREATMENT SYSTEM Geir Wedde, Anders S0rhuus, Odd Bjarn0 Alstom Norway AS, Drammensveien 165,0277 Oslo, Norway Pot Gas Treatment, Emissions, Energy Recovery It is obvious that the GTC size and footprint must fit within a smelter requirements and practical arrangement and it seems that this limit has been reached. Alstom has developed a solution that divides the traditional GTC into smaller sub-centers that are distributed along the pot rooms. These small plants are called Decentralized Distributed Scrubbers (DDS - patented).

Abstract Gas Treatment Centers (GTC) are traditionally arranged in the courtyard between pot rooms and handle vast quantities of pot gas (3- 4 million m3/h) in a large number of filter compartments (1530) with demanding space requirements and challenging control of operation.

Decentralized distributed scrubber (DDS)

Arranging the gas treatment in decentralized installations as Decentralized Distributed Scrubbers (DDS) would obviously save on duct work, alumina handling, storage and civil work. In addition Alstom's DDS integrates silos, heat exchangers, scrubbers, fans and stack into an extremely compact and efficient multi-pollutant control and recovery technology with incomparable footprint. The close integration with pots improves pot gas collection and simplifies the alumina distribution to pots. Module based design allows for short delivery time and early start-up. This paper discusses and reviews the benefits and mitigation of the technical challenges offered by the DDS solution.

Gas treatment plants for pot lines must satisfy the stringent requirements that are forced on them through the continuous production process that governs the aluminium electrolysis. This means a 24/7 operation year round including high equipment redundancies. DDS lav-out and arrangement The difference between the GTC and DDS lay-out in the courtyard between two pot rooms is illustrated in figure 2. The DDS has been laid out close to the pot room walls and connected to a section of pots. A number (12) of identical DDS consisting of equal pot ducting systems, filters, fans and alumina (fresh and enriched) handling systems are shown.

Introduction New greenfield aluminium smelters and brownfield smelter amperage increase or expansions have increasingly demanded larger gas flows to be handled by gas treatment systems leading to an economical challenge as well as space (footprint) issue.

WfM. Ìί^ÌÌίÌψÌ

Gas flows from sections of pots are traditionally handled by Gas Treatment Centers (GTC) arranged in the courtyard between pot rooms. The size of GTCs has dimensions of a large office buildings being up to 150 m long and 30 m wide and rise to levels of 30 m. For operational flexibility and gas and alumina flow considerations the GTC is made up of a number of identical compartments that operate in parallel. The number of compartments could be as many as 30 or above which may be challenging to tune, operate and to detect failures or faults among the individual units.

Figure 2: GTC lay-out (1) and corresponding DDS lay-out (2) The illustrated GTC handles pot gas from a total of 180 pots arranged in two pot rooms with 45 pots in each half of the pot room on both sides of the cross-over passage way. The GTC is designed to operate with a total pot gas flow of 2.5 million m3 per hour. A typical arrangement of DDS would be to handle gas from 15 pots (-210,000 mTh) in each DDS which means a total of 12 small gas treatments plants. This includes allowance for running 2 pots simultaneously on forced draft (~24,000m3/h) for each section. Arrangement and design of a DDS The smallest gas treatment center should include a minimum of 2 reactor/filters and 2 exhaust fans with sufficient redundant equipment for independent operation. Any other combination of

Figure 1: Gas Treatment Center (GTC) for modern pot lines

295

number of pots and compartments is readily adaptable with DDS small compartment sizes and adjustable bag lengths. An arrangement of a DDS is shown in figure 3. Each of the two units is a fully equipped Abart [1] dry scrubber including fresh and enriched alumina storage, fresh alumina metering and transport, reactor/filter, alumina recycling feeder, fan and stack.

the pots through an Alstom Alfeed pot feed system [2] or other feed systems on the market. This saves intermediate silos and alumina transfer points that tend to increase segregation possibilities of the alumina and disturbances of pot operations. The hopper capacity for stored enriched alumina is about 3 tons per compartment and a direct by-pass from the fresh alumina storage to the pot feed system is available ensuring alumina feed to pots at all times. Recycling of enriched alumina within one compartment is achieved with individual feeders with adjustment control. Performance enhancements Each compartment has individual control of gas flow by a dedicated exhaust fan; fresh alumina feed and enriched alumina recycling through frequency controlled drives. Each compartment should run with the same flows and feeds and with the above features DDS has the possibility to individually fine-tune each compartment. This allows the DDS system to easily and accurately achieve equal gas flow with equal alumina feed through each compartment. Each DDS compartment utilizes independent bag cleaning control which allows simultaneous tuning of pressure drop, flow and particulate emission (particulate emission is directly related to pressure drop of a filter).

Figure 3: Arrangement and design of a DDS plant

The direct fresh alumina feed from silo into the gas stream ensures constant feed without any surges and reduces stops for cleaning of airslide systems.

The adaption of fans on top of the fresh alumina storage requires additional stiffening and vibration pads. The mini-stacks are equipped with sound absorbers both at fan inlets and in ministacks.

Overall the DDS has the potential of reduced emissions through optimized adsorbent utilization and enhanced plant stability level.

Benefits of DDS

The DDS is designed to utilize multiple bag lengths, but 8 m bag length is preferred for improved cost efficiency.

The benefits of gas treatment plants arranged as DDS are many but also challenges and some disadvantages are obvious. These are analyzed below with the assumption that fresh alumina is stored in one central silo and the fresh alumina requirement for the DDS is achieved through dense phase transport up to the individual DDS.

Power consumption DDS represent a significant saving in power consumption of the gas treatment installations for a smelter. Pot room ducting, crossover ducts, filter to stack ducts and stack are additional ducting with additional resistance of a traditional GTC. The reduced pot gas pressure drop for a DDS equals about 15 % of the total GTC gas pressure requirement which is a 15 % reduction of the power consumptions of the main fans.

Foot print There is a substantial reduction in foot print area by implementing 12 DDS plants rather than one GTC (figure 2). One DDS unit of 4 compartments takes up approx 90 m2 of area in the courtyard and 12 DDS will then occupy nearly 1100 m2 while a GTC would cover about 2300 m2 which is more than double of the DDS. This saving in space or foot print is of significant value for the smelter.

Steel weight It is easily seen that the duct system has been substantially reduced as the DDS are located close to the pots and use smaller and identical pot duct sections. Ducts from GTC filters to the exhaust fans and the large stack are eliminated. As DDS integrate the required silo capacities, both the fresh and enriched alumina silos can be eliminated. One may expect an increased weight of the DDS filters due to the increased total number of compartments; however, DDS with 8 m long bags has the same weight per unit of gas flow as the Alstom standard GTC equipped with 6 m bag length. With the arrangement as seen in figure 2 the reduction in steel weight of DDS is in the order of 30 % of a comparable GTC.

In addition the flexibility in the location of each DDS in the courtyard allows for an increased optimization of space and pot room and utility building lay-outs Alumina storage and handling Integrated with each Abart filter compartment in a DDS is fresh alumina storage of 25 tons. This capacity reflects close to 2 days consumption of alumina of the pots. From the storage of fluoride enriched alumina in the filter compartment hoppers the enriched alumina may be fed directly to

296

Equipment scope There is a significant saving on alumina transport and silo fluidization, but an increase of alumina feeders. Although GTC has only 6 large fans compared to 48 fans for DDS, the GTC fan cost is about 4 times that of DDS. Filter bags and cages are reduced by the DDS by approx. 20 % because of extended bag length. Electrical items and instruments As each DDS will have to be electrified from a central source, number of power and instrument cables will increase substantially for a DDS arrangement. Motor starters, instruments, plant switches and signals, are proportional to the number of filter compartments which may amounts up to 20 % extra for the DDS system. Although the power tension required is LV - low voltage (< 460 V) and that the cabinet etc for each DDS is located in a common cubicle that can be mounted and furnished in a work-shop and shipped to site, the electrical scope may be more than twice that of the GTC.

B2JI I

I cmntm.cotmx».wiomr%cM}Amßunr\

L.

Figure 5: Installation of work-shop erected DDS modules Erection of several DDS units can be achieved simultaneously and independently which is not the case for a GTC where many items are dependent on each other as for erection of filter compartments, alumina handling to and from these compartments, stairs and platforms, and electrical and instrumentation. Further benefits

TO na«

w i n

Civil work for a GTC installation is a significant part of total cost. The foundations for the heavy silos, filter compartment support, main fans and stack are extensive compared with the simpler arrangements for the DDS.

Figure 4: Basic DDS E&I topology

For smelters with pot line expansions DDS represent a simple and unique opportunity for stage wise implementation as the system connects to a smaller number of pots. The bolted design eases the installation in an operating smelter as magnetic field strength complicate welded construction.

However, DDS simplifies total electrical lay-out and is exampled in figure 4. A GTC will have substations for both MV and LV, as the main fans normally require MV. MV can be removed, and a number of identical cubicles for each DDS installed.

There are projects where the GTC falls on the critical path of the smelter erection sequence. The DDS will allow quick and independent erection and potentially early start-up of metal production.

It is reason to believe that the extra electrical scope of the DDS can be significantly reduced by working close with the smelters utilizing and incorporating the existing electrical system lay-out. Manufacture, transport and erection The size of a DDS compartment is fulfilling the requirement on regulation of transport on roads. DDS is suitable for offsite and offshore pre-assembled, with manufacture in work-shops and transport to site. Cost efficient production of filter panels, hoppers and tops utilizing automatic, robotic manufacturing systems has been demonstrated with success for other products as Alfeed. This allows fast erection in lay-down areas or directly on site as seen figure 5.

Challenges of DDS There are challenges to the DDS solution that sometimes would become opportunities but also a re-think of the smelter operation may facilitate all the benefits that have been addressed above. Pot gas temperatures As the DDS is closer to pots it should experience higher gas temperatures (15-20 °C) than the GTC. It is well established that pot gas temperatures affect the performance of the dry scrubber and in order to meet current requirements the gas temperature at scrubber inlet should be 115-120 °C [3]. This requirement means that the DDS may need additional cooling upstream the filter compartments. If dilution air was to be used a significant quantity of air has to be added and the cost efficiency of the DDS would suffer.

Fluidization of filter hopper and mini-airlifts and pot feed system is generated by one blower/fan only using air from the pot room plant air lines as stand-by. It is also preferable to use plant air instead of installing dedicated air compressor and dryer equipment for bag cleaning air needs.

Integration of heat exchanger Alstom has developed heat exchangers operating on the pot gas upstream the filters [4]. Improvements on this technology are ongoing and a heat exchanger integrated (IHEX) into the inlet duct/reactor of the Abart filter compartment has now been realized. This allows significant savings on the heat exchanger system as support steel, stairs and platforms, and the damper arrangement is combined with the DDS design as seen in figure 6.

Feeders, recycling units and main fans are easily accessible in the DDS and spare units can be stored at the smelter for fast replacement of broken units. During situations of "sick" pots where pot gas temperatures increases and fume evolutions are enhanced the HEX will achieve improved performance and cooling as the temperature differences of pot gas and cooling water increases. Short term high HF evolutions can be controlled with increased fresh alumina feed. Cover material for pots Many smelters and pot technologies use enriched alumina and mixed with bath for covering and dressing pots. For DDS this can be facilitated by inter-connecting several pot feed sections from several DDS units would allow these DDS to run with slightly higher feed equal to the cover material needs. The remaining DDS will have to be slightly starved of fresh alumina feed but the overall emission level from DDS will not be affected.

Integrated c heat exchanger

The additional flow of enriched alumina could be extracted from the pot feed system at an appropriate location into an airlift feeding a cover material silo. Optionally one can feed a tanker with cover material and feed the silo in batches which would then overcome the high capacity discharge needs from silo to pot tending cranes.

Figure 6: The integrated heat exchanger (IHEX) for DDS This innovative solution reduces the pot gas temperature to any desired practical levels (90-120 °C) depending on the purpose of the heat exchanger whether it should only cool the pot gas ahead of the gas treatment plant and /or to recover waste heat for use inside or outside the smelter. The additional 15 (- 20) °C temperature level of the pot gas entering an DDS means a significant contribution to higher value heat (exergy) of about 50 kW/pot. For a 180 pot smelter this amounts to 9 MW in addition to a traditional (GTC) heat recovery potential with IHEX of 15-20 MW.

Emission monitoring The DDS is prepared for the installation of HF and dust measuring equipment, easy accessed from the filter top. To reduce the number of monitors required a system for extraction of gas from each mini-stack into a laser based analyzer has been developed. Emission dispersion

Operation redundancy

Discharge of treated pot gas from GTC stacks would require stack height level above pot room roof louvers to meet stringent emission requirement through emission dispersion modeling. For DDS it would have been detrimental to its design simplicity to have enforced same stack heights as on the GTC.

The DDS is designed to utilize multiple bag lengths, but 8 m bag length is preferred for improved cost efficiency. For special purposes and increased flexibility as when the N-l redundancy principle applies the DDS can be equipped with shorter bag length (5-7 m) which will increase the number of compartments. Connecting pair-wise DDS will add further flexibility and with frequency operated motors of main fans one compartment can be taken off-line with only a slight reduction in total gas flow from pots.

As DDS introduce multiple stacks at locations across the pot line the dispersion of the emissions is significantly improved and DDS emissions can be released at lower levels. The emission through the pot room roof is of similar concentration levels as from the DDS mini-stacks and it would be natural to use similar discharge height level for the mini-stacks.

Even more gas flow capacity is achieved when the HEX is designed with some over-capacity that allows further cooling of the gas.

Magnetic fields The DDS is located closer than a GTC to the pot room and is therefore exposed generally to higher magnetic field strength. This has to be considered especially regarding electronic equipment.

Therefore the combination of size and number of compartments and with a heat exchanger the DDS solution has sufficient gas flow capacity to satisfy traditional redundancy needs of a smelter.

298

The performance of "standard" polyester needle felt filter bags is demonstrating consistent low particulate emission at about 0.9 1.1 mg/Nm3, and the particular test campaign below (figure 8) shows emissions during pulse air cleaning of the filter bags reaching about 12 mg/Nm3 as a peak value and reducing to a value close to zero between each cleaning action. The total emission over several cleaning cycles is 1.04 mg/Nm3

The main fans, however, are located on top of the DDS which increase the distance to the source (electrical bus bars) of the magnetic fields. It has been calculated that the field strength at the DDS main fans is similar to the one at the GTC main fans. The DDS is normally made as a bolted solution, which means considerably less welding and ease of erection if the construction has to take place in the magnetic field of an existing operating smelter.

Filter3 L D 1 2 8 2 S t a n d a r d

Multi-pollutant gas treatment More often S0 2 scrubbing is required for aluminium smelters, in particular the large mega-smelters where S0 2 emission may have a detrimental impact on the local ecology. Wet S0 2 scrubbers are used which absorb S02 into an alkali solution and smelters operating at coastal areas would benefit from using seawater as the absorbent [5]. To accommodate the DDS principle of simplicity and modularization a S0 2 scrubber design has been developed and ongoing tests will confirm the performance which are targeting levels that exceed current S0 2 scrubbers for GTC

Figure 8: Particulate emission peaks during pulse cleaning of "standard"filterbags.

Operational experiences

Filter bags with enhanced micro-denier fibers embedded in the outer filtration surface achieve an exceptionally low peak emission consistently lower than 3.0 mg/Nm3 as seen in figure 9. The total particulate emission is reduced to only 25 % of the "standard" bag emission to as low as 0.25 mg/Nm3.

Experiences with DDS have been obtained since 2007 where 2 DDS plants (figure 7) were installed at an aluminium smelter. Each DDS consists of 2 compartments and handles 115.000 m3/h. The two DDS are interconnected to achieve highest flexibility. Fresh alumina is fed to the four integrated silos in the DDS by an airslide system running from an elevated existing silo. The enriched alumina is fed to 2 pot sections.

Filter! L D 1 2 8 4 M i c r o D e n i e r _ Pulse is avg of 3 pulse cycles I.e. 3 x (

Figure 9: Particulate emission of "micro-denier"filterbags Two test campaigns using manual test equipment, of emission of gaseous fluorides (HF) has been recorded at an average level of 0.26 mgF/Nm3 with F-content of the enriched alumina at 1.80 %. For low HF concentrated pot gas with 1.15 % F-content in enriched alumina the HF emission level was as low as 0.04 mgF/Nm3. Figure 7: Operating DDS installations

Conclusion

Emission performance

The principles of centralization do not always reduce costs and ease the operation of a system. The DDS solution as distributed gas treatment plants located close to the pot room and handling gas from a limited number of pots (10- 30) has many advantages

Particulate emissions has been followed up and monitored regularly as testing of filter bag materials is part of an agreed cooperation with the smelter.

299

and benefits that should be explored further by the aluminium industry. The reduced cost and power consumption in addition to the reduced foot-print of 50 % and the increased potential of waste heat recovery would obviously make this industry more competitive. DDS has flexibility and operational performance that will meet future emission requirements. Particulate emission levels at only 1/10 of traditional level is achieved by applying filter fabric made with micro-denier fibers. Incorporating a SO2 scrubber in the DDS solution makes it a cost efficient multi-pollutant gas treatment technology. References 1. 2. 3. 4. 5.

G. Wedde, P Henriksen "Experiences and developments through 10 years of operation with the Abart dry scrubber" Light Metals 2007 S. Ose, A. Sorhuus, O. Bjarno, G. Wedde " Alfeed, a new alumina feeding system to aluminium pots" Light Metals 2010 A. Heiberg, G.Wedde, O Bockman, S. Strommen " Pot gas fume as a source of HF emission from aluminium smelterslaboratory and field investigations" Light Metals 1999 A. Sorhuus, G. Wedde, K. Rye, G. Nyland „ Increased energy efficiency and reduced HF emission with new heat exchanger", Light Metals 2010 S Broek, B Rogers "The origin and abatement of S02 emissions from primary aluminium smelters" Light Metals 2009

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

FLUORIDE EMISSIONS MANAGEMENT GUIDE (FEMG) FOR ALUMINIUM SMELTERS Nursiani Tjahyono, Yashuang Gao, David Wong, Wei Zhang, Mark P. Taylor Light Metals Research Centre, The University of Auckland, Private Bag 92019, New Zealand Keywords: Fluoride emission, Control, Management, Guidelines, Work practices Abstract

These legislative limits, the health and safety of the employees and the wellbeing of the environment are major drivers for the smelters to reduce their fluoride emissions.

All smelters worldwide operate under strict fluoride emission limits and the reduction of fluoride emissions is further driven by health and environmental considerations. A Fluoride Emissions Management Guide (FEMG) has been written by the Light Metals Research Centre (LMRC) on the invitation of Australian Aluminium Council (AAC), under the Asia-Pacific Partnership (APP) on Clean Development and Climate. The aim of the FEMG is to provide a better understanding of the factors affecting and ways of reducing fluoride evolution and emissions in smelters, and further, to provide smelters with an operational guide for reducing and managingfluorideemissions.

Mechanisms of Fluoride Generation and Escape Pathways The majority of gaseous and particulate fluorides emitted from an aluminium smelter are originally generated in a pot. The amount will depend on cell technology, raw materials, process control, and ambient conditions (temperature and humidity) where the smelter is situated [10]. A typical pre-bake, point-fed pot generates about 15-30 kg total F/t Al, with >99% of these fluorides typically captured in the scrubbing system [10,11].

Reduction in fluoride emission can be achieved by improvement of operational standards in the smelter. This includes operational, control and maintenance practices in the potroom, GTC, and other areas in the smelter that generate fluoride. Along with incorporated audit guidelines and training package, implementation of this guide can, in a short time, lead to significant reduction influorideemissions in aluminium smelters.

Gaseous fluoride (HF) is produced when A1F3 in the bath or as bath vapour (NaAlF4) reacts with a hydrogen source in one or two reactions: 2A1F3 (g or diss) + 3H20 (g) => 6HF (g) + A1203 (s or di SS ) 3NaAlF4 (g) + 3H20 (g) => 6HF (g) + A1203 (s) + Na3AlF6 (s)

(1) (2)

Introduction Particulate fluorides, on the other hand, are generated when bath vapour condenses to form droplets or fine bath particles are picked up and entrained in the pot gases.

Generation of fluorides from the aluminium smelting process is unavoidable with the current technology. Hence, all smelters worldwide operate under strict limits of fluoride emissions and these limits are expected to become tighter with time. Table I provides a comparison of production-based regulatory limits for total airborne fluoride emission in primary pre-bake aluminium smelters for different regions of the world, including the current best practice.

Potrooms and Gas Treatment Centres (GTCs) are often seen as the main sources of fluoride emissions. While it is important that the different factors related to fluoride emissions from these two key areas are addressed and controlled, it is equally important to realise that a fluoride emissions problem may also be related to other plant areas. Examples are the fluorides released from transport/delivery of spent anode butts and the bath processing plant. The mission to reduce and control smelter fluoride emissions spans across all plants within the smelter and is a responsibility of every individual in the smelter.

The majority of larger capacity modern smelters today limit their fluoride emissions to 0.5-0.6 kg F/t Al [1]. These limits are practically achievable, as demonstrated by the performance of Aldel and Tomago smelters (0.46 kg F/t Al in 2005 [2] and 0.56 kg F/t Al in 2007 [3], respectively). Some of the better performing smelters in the world have reported emissions as low as 0.29 kg F/t Al (Portland in 2008) [4] and 0.25 kg F/t Al (Karmoy in 2006) [5]. Table 1. Comparison ofproduction based regulatory limits and the current best practice for total airbornefluorideemissions in the pre-bake aluminium smelting industry. Regulated Airborne Emissions Limits for Total Fluoride (kg F/t Al) World Benchmark 0.2 [6, 7] Europe - Iceland 0.35 [1] Europe / OSPAR 0.6 [8] USA-EPA 0.6-1.5* [9] * The limit of 1.5 kg F/t Al is only applicable to older potlines as specified in the National Emission Standards for Hazardous Air Pollutants for Primary Aluminium Reduction Plants. Region

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Through careful management and control of both operation and maintenance practices around the plant, further reductions in fluoride emissions can be achieved. However, the factors which affect and the ways of reducing fluoride emissions are sometimes not clearly understood, hence there is a need for a management guide to address fluoride emissions in the smelter as a whole. This paper describes the Fluoride Emissions Management Guide (FEMG) which has been prepared by the Light Metals Research Centre (LMRC), The University of Auckland, on the invitation of the Australian Aluminium Council (AAC), under the Asia-Pacific Partnership (APP) on Clean Development and Climate. The paper will focus on the content of the guide and the importance of managingfluorideemissions in the smelters.

anode setting and bath pouring operations are examples of work practices that have a direct effect on fluoride emissions. These operations are generally simpler to target compared to those with an indirect impact, such as the influence of poor anode quality on carbon dust generation in pots, which in turn increases fluoride emissions from more frequent dust skimming and open pot doors.

WhatistheFEMG? The FEMG is a comprehensive guide to managing fluoride emissions in primary aluminium smelters, with a focus on improvements to operational, control and maintenance practices to allow smelters to maximise their environmental performance with their current technologies. It aims to increase understanding of factors controlling fluoride evolution and emission, as well as providing an operational guide to smelters for reducing and managing fluoride emissions. The FEMG covers fluoride emissions from the potroom, GTC and other smelter systems that contribute to emissions, along with incorporated audit guidelines for each area. The guide is specifically prepared for pre-bake, point-fed pot technologies with dry scrubbing technologies.

HDBBBBH Operation

The FEMG is broken down into 6 main chapters, which are: • Introduction & Theory - covers the background of fluoride generation and drivers to reduce fluoride emissions. • Overall Fluoride Emission Management System covers the overall concept and approach for managing smelter fluoride emissions. • Potroom Systems for Reducing Fluoride - specifies the Key Performance Indicators (KPIs) and control points for work practices in the potroom. • Gas Treatment Centre (GTC) Systems for Reducing Fluoride - addresses KPIs and control points for work practices in the GTC. • Smelter Systems Outside the Potroom and GTC details KPIs and control points for other areas in the smelter that affect fluoride emissions. • Fluoride Emission Measurement - lists standard and recommended smelter fluoride measurement methods.

Process Control

Maintenance M

l

Smelter

HF Measurements

· Remedial Actions

Boundary

Figure 1. Overall concept behind thefluorideemission management system. Monitoring Systems Successful management of any process requires good measurements. Measurement systems are important to determine whether the operations are performing within the required specifications, otherwise actions can be taken to rectify the problem. The most effective monitoring system designs include a combination of numbers, charts, colour and alarms, which provide meaningful and relevant information to the users. For monitoring of fluoride emissions, specific equipment and methods, such as US EPA Methods 13A and 14, have been developed to measure the level offluorideemissions as required by law.

Pictures are extensively used to illustrate issues and potential improvements to lower fluoride emission through daily operations. The recommendations contained in the guide have been proven by world-class smelters and are based upon current best practices, whereby if adopted, significant improvement in fluoride emission can be achieved.

Response Systems Once information from the monitoring systems has been received and analysed, response systems can be developed. They include: • Remedial actions • Work standardisation • Staff education and training • Plant and process changes and improvement

Overall Fluoride Emission Management System As previously mentioned, reducing fluoride emissions in the smelter should be the responsibility of all smelter personnel and not be focused only at the potroom and GTC areas. The overall concept of the fluoride emission management system applied in the FEMG is a systematic-interconnected network consisting of four main themes, as illustrated in Figure 1. They are: • Control of work practices within specifications • Monitoring systems • Response systems • Audit systems. Each of these themes has a very important, specific role in reducing fluoride emissions and requires the involvement of all smelter employees. Control of Work Practices When the smelter work practices are not within process and work specification, they can result in a direct or indirect negative impact on the smelter's fluoride emissions. Exposed bath during

302

The response systems can range from an immediate to long term response. They are important for keeping practices within specifications, as well as for operational and financial optimisation. Furthermore, the response systems are utilised to prevent and reduce undesirable process conditions from occurring. Smelter Audit Systems An audit is an on-site review conducted to determine the existing status of a process and whether a certain quality condition has been achieved, ensuring that issues, problematic areas and improvement opportunities are identified and addressed. The level of audits in the smelter can vary from simple (i.e. hooding tightness) to comprehensive assessment (GTC operation). It should be conducted on a regular basis by the smelter operations and management team, and cover the entire systems in the smelter to maintain the integrity and effectiveness of every part of the smelter systems.

collapse, resulting in more open holes in the crust, which again increases duct temperatures in the GTC.

Why is Fluoride Emission Management Important? All smelters worldwide are driven by the same factors to control and reduce their fluoride emissions. Some of the problem areas that smelters face are explored below. High Total Fluoride Emissions Some smelters face difficulties in keeping their fluoride emissions well below the legal limits, which may result in the possibility of heavy fines or even plant closure, should the situation continue to occur. High levels of fluoride emission will also affect the health of employees, particularly those working in areas where fluorides are generated. Skin/respiratory irritation, dental/skeletal fluorosis, and liver, kidney, or heart damage are some of the risks employees face with prolonged high exposure tofluorides[12]. A change in the operational standards is one way of making a difference in the level of fluoride emission in the smelter. By going through each step in every work practice, potential pathways of fluoride evolution can be easily identified. In metal tapping, for example, the fluoride is generated during the opening of tap end doors, as shown in Figure 2. Pot suction is lowered when there are gaps in the cell (including opening of doors and hoods) due to the loss of static pressure in the pot. The reduction of pot suction will lower the pot capture efficiency and hence, results in higher fluoride emission. Furthermore, fumes released from the crucible vent during the metal siphoning process are a major source offluorideemission during metal tapping.

# Hoods Off / Doors Open Figure 2. Schematic diagram showing howfluorideevolution and emission occur when the tap end door is opened during metal tapping procedure [Top], and how pot suction affects the pot capture efficiency and HF total emissions (not to scale) [Bottom].

When the pathways of fluoride generation have been identified, potential improvements can be recognised and change in work practices can be made. As shown in Figure 3, by comparing the smelter's current work practices against the world's best practices, poor work practices can be easily targeted and operational standards can be systematically improved. Audit checklists, like that shown in Table II, are a good tool for smelter managers or superintendents to use for conducting regular audits, ensuring that work practices are maintained within specified standards.

Inlet duct gas temperature is an important operational parameter. Above certain temperatures, the scrubbing efficiency is lowered as particulate fluorides captured on the surface of the bagfilters begin to react, forming HF which is released as a gaseous stack emission. The bagfilters in the GTC also have critical temperature limits, above which they are irreversibly damaged - in some smelters, the gas would then be bypassed from the GTC, sending all pot gases to be released into the environment without scrubbing. Both of these aspects will lead to higher total fluoride emissions and reduction in the operating safety margin of a smelter.

Increase in Amperage The aluminium industry continues to shift towards higher amperage cells to obtain higher levels of productivity and greater economic returns. This shift requires better control and management of fluoride emissions before it can proceed. Increases in line amperage will inevitably lead to higher heat generation from the pots. Poor management of pot heat balance can lead to direct increases in fluoride emissions, for example from more frequent unscheduled anode changes and carbon dusting. Furthermore, if pot heat balance is not managed properly, gas duct temperatures in the GTC can rise excessively, decreasing the efficiency of the gas scrubbing system. Greater heat generation in the pot can also cause increased rates of crust

Table II. Checklist of key areas to reducefluorideemissions during metal tapping. I Check items 1

Only open the tap end door before metal tapping

YES

NO

2

Use a hose to direct the extracted fumes back to pot

YES

NO

3

Cover the tap hole by sprinkling some crushed bath or alumina into the tap hole

YES

NO

Close the tap end door when metal tapping is finished.

YES

NO

\T"

303

BAD Practice

Opening the tap end doors of a group of pots which are to be tapped.

Opening the tap end door of a single pot right before tapping.

Releasing the environment.

Connecting a hose onto the crucible vent and directing the extracted fumes back to the pot.

extracted

fumes

to

the

potroom

Figure 3. Examples of recommended best practice to minimise fluoride emissions during initial steps of metal tapping, presented as comparison of existing work practices and best practice (bad vs. good practice). By installing an online measurement and monitoring of inlet duct gas temperatures, a pre-warning system for high temperatures, as illustrated in Figure 4, can be adopted. Such a system will allow for investigation or action before a bypass from the GTC is activated. Furthermore, regular measurement of inlet duct gas fluoride level allows for detection of long term shifts/fluctuations in fluoride generation from the potroom. This will enable adjustments in scrubbing efficiency, for example, by increasing secondary alumina recycle rate to respond to higher levels of fluoride in the incoming gas to the GTC. Iniet Duct Temperature (deg C)

k/\/\^v\/\/vm Time (days)

Table III. Potential improvements that can be done to reduce fluoride emission due to shift in inlet duct gas temperature. Potential improvement to reduce fluoride emission due to increase in inlet duct gas temperature and fluoride level.

Critical Filter Temperature

120 100

With continual increases in amperage, however, existing GTCs can only cope to certain extent before some capital investment is required to accommodate the increase in the level of fluoride generation. Some potential improvements are shown in Table III.

Low

Cost

Pre

i-warning Temperature

Investigate High Cost

Figure 4. Schematic of continuous temperature measurement of GTC duct gas inlet temperature to detect and treat issue before critical temperature is reached.

304

Use higher temperature bagfilter materials to sustain higher temperature operation and minimise GTC bypass on hot days. Measure and monitor inlet duct gas fluoride level using stack-type testing methods, over regular intervals (e.g. once per fortnight). If at risk of exceeding temperature limits of bagfilters, implementing a heat exchanger system to cool inlet gas ducts before the gas reaches the bagfilters at the GTC. Continuous monitoring (laser) of inlet duct gas fluoride level. Install higher scrubbing capacity (more gas flow) for higher amperage or higher line capacity.

Customisation of FEMG & Identification of Improvements

Customisation and Implementation of FEMG

Using the customisation material and photos, a tailored FEMG for specific smelter's technology and operational, control and maintenance practices can be developed. From the audit review, KPIs and improvement opportunities to reduce fluoride emissions across all areas of the smelter can be identified. This will ensure the content of a customised FEMG and any improvement to work practices is specific and relevant to all the personnel in the smelter.

As potroom and GTC technologies vary from smelter to smelter, it is important for the management guide to be adaptable to any variation in technologies and work practices. The FEMG is customisable to meet the need of any specific smelter. Ultimately, the guide achieves its purpose when the content is being implemented in the smelter. The development of an implementation plan is important to allow the smelter to distribute the relevant information to all its personnel that play a role in fluoride emissions. It will also ensure that maximum gains can be achieved by the smelter in reducing fluoride emissions.

Strategic Planning and Implementation of FEMG It is important to develop a strategic plan to address areas of improvement. The implementation process may include: • Smelter personnel training - training of smelter staff in the identified areas of improvement based on the customised FEMG and training workshops. • Implementation of changes in practices - smelter personnel to apply the necessary changes as per the guidelines in FEMG. • Regular audits - smelter managers to monitor the changes in practices by conducting regular audits. • Fluoride monitoring equipment - fluoride emission monitoring equipment could be used to monitor the changes in the smelter's fluoride emission as improvements to practices are implemented. This can be done by comparing baseline measurements (1-3 months prior to changes) against measurements 1-3 months after implementation of FEMG practices.

An example of the customisation and implementation process for the FEMG is summarised in Figure 5 and consists of several steps. Fluoride Emission Management Guide (FEIV1G) Customisation & Implementation Plan External Audit & Customisation Visit Audit:

Customisation Material:

On Operating, Control & Maintenance Practices as per FEMG

3.g. Photos of good / bad practices, diagrams, etc from Smelter

Improvements to Practices Identified

FEMG Customisation for Smelter

If adopted, smelters stand to achieve significant improvement in environmental performance and operational standards for fluoride emissions reduction. Case Study

Strategic Planning & Implementation of FEMG Training of Staff

Customised FEMG ■ Training Workshops

A smelter case study with the FEMG was carried out by the LMRC. An audit was conducted in two sections of a potroom and in a GTC area in a smelter, using the checklists provided in the FEMG. During the auditing process, some good practices and improvement opportunities were recognised. One example was the use of anode rod seals to minimise emissions from the pathway shown in Figure 6.

Adjust Practices

As per Recommended Improvements

Fluoride Measurements

Monitor Changes with Implementation

Regular Audits

Monitor Changes to Practices

Figure 5. Customisation and implementation process of FEMG. External Audit and Customisation Visit The purpose of an initial audit and customisation visit is to obtain relevant information and data on smelter technology, as well as smelter's operation, control and maintenance practices to customise the FEMG. The information obtained will ideally include photos of bad/good work practices, operational schedules, smelter layout, and other relevant smelter information. Concurrently, operational, control and maintenance practices in the smelter will be reviewed to assess the current practices. The audit will be based on the audit guidelines/checklists and recommended best practices specified in the FEMG.

Figure 6. Schematic diagram offluorideemission from the gap left between the anode rod and the skirt due to damage or missing seal

305

The openings in the cell superstructure around the anode rods are designed to accommodate anode change process. Prior to the audit and training sessions, most of the operators did not understand the role of anode rod seals. Hence, no routine check on the condition of the anode rod seals was done. On average per pot, five seals were missing and seven seals were damaged.

References 1. A. Moras et al, "Modern Potline Gas Treatment Technology for High Amperage Pots - The Alcoa Fjardaal Experience," Light Metals, 2010, pp 239-242. 2. M. Stam et al, "Operational and Control Improvements in Reduction Lines at Aluminium Delfzijl," Light Metals, 2007, pp 243-247. 3. M. Meyer, G. Girault, and J. M. Bertolo, "Development of a Jet Induced Boosted Suction System to Reduce Fluoride Emissions," Light Metals, 2009, pp 287-292. 4. "Alcoa in Australia: Smelter Fluoride Emissions - Point Henry and Portland Aluminium," taken from http://www.alcoa.com/australia/en/info page/environ air.asp 5. H. P. Lange et al, "Innovative Solutions to Sustainability in Hydro," Light Metals, 2008, pp 211-216. 6. S. Lindsay, TMS Alumina Short Course (2004) 7. S. J. Lindsay, "Effective Techniques to Control Fluoride Emissions," Light Metals, 2007, pp 199-203. 8. Emission targets for 2010. OSPAR Recommendation 98/2 on Emission and Discharge Limit Values for Existing Aluminium Electrolysis Plants, www.ospar.org 9. US EPA (1997) "National Emission Standards for Hazardous Air Pollutants for Source Categories; National Emission Standards for Hazardous Air Pollutants for Primary Aluminium Reduction Plants; Final Rule," 40 CFR Parts 9, 60, and 63. 10. E. Patterson et al., "Reducing Fluoride Emissions from Aluminium Electrolysis Cells," (Proc. 7 th Australasian Aluminium Smelting Technology Conference and Workshop, 2001) 11. M. M. Hyland, B. J. Welch, and J. B. Metson, "Changing Knowledge and Practices towards Minimising Fluoride and Sulphur Emissions from Aluminium Reduction Cells," Light Metals, 2000, pp 333-338. 12. U.S. Agency for Toxic Substances and Disease Registry (2003) Toxicological Profile for Fluorides, Hydrogen Fluoride and Fluorine.

Through the training sessions, the importance of anode rod seals on maintaining pot suction and therefore, keeping the fluoride in the pot becomes clear to the operators. These anode rod seals can reduce fluoride emissions by as much as 20% [7]. After the training sessions, regular checks were scheduled on anode rod seals whereby, damaged seals were repaired and missing seals were replaced immediately. Feedback obtained from the operators was generally positive. Operators found the guide to be very practical, particularly the pictures or schematic drawings used to illustrate the problems and the recommendations. The theory and background information contained in the FEMG for each practice provides a better understanding of the importance of each operation - through this understanding, it is easier for operators to implement the FEMG's best practice recommendations. Implementation of the FEMG was also a good opportunity to test the robustness and practicality of the guide. Feedback from the operators and smelter personnel was integrated into the guide, improving the quality and usefulness of the guide. Conclusions The mission to manage and reduce fluoride emissions extends across all areas of the smelter. Every individual in the smelter is responsible and plays a vital role in achieving this goal. Although the potroom and the GTC are likely to be the main areas contributing to fluoride emissions, it is important to realise that other plant areas, such as the bath plant, may also contribute to fluoride emissions. The FEMG provides a structured approach to a better understanding of the factors affecting and the ways of reducing fluoride generation and emissions. The ultimate aim of this customisable guide is to provide practical and technical information to help smelters achieve significant improvements in their environmental performance with their current technologies. Through careful management of operation, process control and maintenance practices across the plant, significant reduction in fluoride emissions and improvement of operational standards can be achieved in the smelters. Acknowledgement The authors would like to thank the Australian Aluminium Council (AAC) - acting on behalf of the Asia Pacific Partnership (APP) on Clean Development and Climate Aluminium Task Force - for their generous sponsorship of the FEMG and Dr. Eng Fui Siew for his extensive contributions to the development of the FEMG.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

Enviroment- Emissions / Anode Effect II SESSION CHAIR

Marco Stam Aluminium Delfzijl Delfzijl, Netherlands

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

On Continuous PFC Emission Unrelated to Anode Effects Wangxing Li, Qingyun Zhao, Jianhong Yang, Shilin Qiu, Xiping Chen Zhengzhou Research Institute of CHALCO, Zhengzhou, China, 450041 China National Engineering & Technology Research Center for Aluminum, Zhengzhou, China, 450041 Jerry Marks J Marks & Associates, LLC, 312 NE Brockton Dr, Lees Summit, MO 64064 Chris Bayliss International Aluminium Institute, London SW1Y4TE, United Kingdom Keywords: Continuous PFC, reduction cells, anode effects, smelters Abstract An investigation was made into operating parameters that could result in continuous PFC emission unrelated to anode effects in two smelters. Anode current distribution was measured in QY smelter. No obvious correlation was observed between the nonhomogenous anode current and continuous PFC emission. Bath temperature and alumina concentration were synchronously measured during continuous PFC monitoring in LX smelter. No continuous PFC emissions were observed when reduction cells are operated at stable bath temperature and alumina concentration. No clear relationship emerged among these factors and further surveys need to be made in other smelters. The effect of feeding mechanism was also investigated. Metal tapping and anode exchange can disturb cells' balance, which result in cell voltage rising, and then may cause continuous PFC emission. It is found that the continuous PFC emission only occur in some particular cells and is different from cell to cell even with the same line current. The continuous PFC emission is only a small portion in total PFC emission.

Investigation Objectives in QY smelter Single-cell studies were performed on two 320 kA cells. Objectives of the study were as follows: (a) Determine the proportion of continuous emissions unrelated to anode effects to the overall PFC emissions, (b) Determine if continuous emissions occur in all cells or in only a few cells. (c)Determine if continuous emissions could take place as a result of non-homogenous current distributions among cell anodes. Investigation Objectives in LX smelter Single-cell measurement was performed on two 176 kA cells which belong to different potlines in LX smelter. The purposes are to determine if bath temperature, feeding mechanism and alumina concentration are the factors affecting continuous PFC emissions. Results and Discussions Investigation in QY Smelter

Introduction PFC survey was performed in CHINALCO smelters (1) using accepted IAI/USEPA measurement methodology. The survey was done in different running statuses such as normally-operated, power-limit, newly-started and new control-mode. PFC emission of normally-operated smelters is close to that of western ones (2). During this survey non-AE PFC emission, however, also called as continuous PFC emissions were first observed. These continuous PFC emissions occurred when no anode effects were taking place. What caused these continuous PFC emissions? Is it universal for all cells or particular for certain cells? Could it be eliminated? A research plan to study continuous PFC emission from Chinese smelters was made in China National Engineering & Technology Research Center for Aluminum/Zhengzhou Research Institute of Chalco (ZRI). Results from the first phase of the study are discussed in this paper.

Gas sample came from two different single cells. Cell numbers were QY-320-6 and QY-320-7. Monitoring duration for QY-320-6 was eleven hours. Two AE peaks were collected (see Figure la). The maximum CF4 emission rate is 0.463 g/s. And CF4 emission is almost zero when there are no anode effects (see Figure lb). Results are shown in Figure 2a from QY-320-7 during six-hour measurement, during which one AE peak was monitored. The maximum CF4 emission rate is 0.848 g/s. During most of measurement period, CF4 emission is close to zero when no anode effects occur (see Figure 2b). Anode Current Distribution Measurement Anode current of cells was usually measured with Millivolt Fork. Voltage drop is measured on every anode rod between two probes at fixed distance apart. The current in each rod is:

Research Plan and Objectives

_ AV; _ AVjA

The existence of continuous PFC emission is like a puzzle. A research plan was developed in ZRI to solve the puzzle. Two smelters were selected to perform studies on continuous PFC emission in the first phase of the plan. One is a pilot plant (code QY), and the other is a common industrial smelter (code LX). QY has two potlines, 320kA cells and 162kA cells respectively. LX has three potlines, all are 176kA cells.

' ~ R, ~ P{T, )L

Where: AVi = mV drop, Ri = resistance of the rod between the fork tines, A = cross-section of the rod, pi = resistivity of the rod (function of rod temperature Tj between the tines), L = distance between fork tines. In cell control practice it is assumed that all the rods have the same

309

temperature. Then, the resistance is the same in each rod and will fall out in the calculation. The sum of measured electric currents in individual rods is then easily normalized to the line current. Measured current distribution in anode rods is normalized to line current in the following way: _ _ < ¢

r

Ä

r

Ó( ^ Line current, kA; (AV)i = Individual rod voltage, mV; N = Number of anode rods. Anode current of two QY-320 cells was measured, calculated and plotted (see Figure 3).

-i

·—i

0

50

«—i

100

'

1

·

150

1

200

«—i

250

1—·

300

1

350

' — i

400

Monitoring duration, min

Figure 2a Original CF4 Emission Trace of QY-320-7

0.5-

0.4-

0.3-

"Φl

2 o 0.2-

0.1 -

0.00

100

200

300

400

500

LJ 600

700

Monitoring duration, min Monitoring duration, min

Figure la Original CF4 Emission Trace of QY-320-6

Figure 2b Continuous CF4 emission of QY-320-7

0.05QY-320-6 X

bC

0.04""«12. 0

i éï.ï

TJH"

— -C:

0.03-

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Graph lb came from graph la by cutting the AE peak which appeared after 660 minutes' monitoring, they took over the same measurement period of QY-320-6. Graph 2b came from graph 2a by cutting the AE peak which came forth after 350 minutes' monitoring, they also described the same measurement period of QY-320-7.

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310

Anode current distributions of QY-320-6 and QY-320-7 were measured once two hours. First current measurement was performed when PFC monitoring had gone on 120 minutes. The mean value of five measurements from QY-320-6 was plotted in Figure 3a. Three measurements were performed on QY-320-7 and the mean value was plotted in Figure 3b. Anode current distribution in QY smelter is similar to anode current distribution variability measured in Western cells (3). There is no obvious relationship between maximum anode current and continuous PFC emission.

It can be learned from investigation in QY smelter that continuous CF4 emissions are only small portion of the total PFC emissions. The amount of continuous PFC varies from cell to cell from near zero to a detectable level. The measured variability of current flow among the anodes is similar to that of Western reduction cells. There was no clear correlation of continuous PFC emission with maximum anode current. Investigation in LX smelter

Percentage of Continuous PFC Total CF4 and non-AE CF4 from two cells were calculated (see Table I). Continuous PFC emission is only small portion of the measured PFC emissions from 320 kA cells. There are different CF4 emissions from different cells even with the same line current.

Single-cell measurement was performed on two 176 kA cells which belong to different potlines in LX smelter. The purposes of this investigation were to determine if bath temperature, feeding mechanism and alumina concentration affect continuous PFC emission. Results from LX176-334 cell Exhaust gas from LX176-334 cell was sampled and monitored continuously for 47.54 hours. No AE peaks were observed. Continuous CF4 emission was almost zero during the whole measurement (see Figure 4).

Table I Proportion of Continuous CF4 Emission

Cell number

QY-320-6 QY-320-7 QY-320-7

Total CF4

188.55 443.74 443.74

Non-AE CF4

AECF 4

183.05 417.29

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311

was 87%. No clear relationship was found between alumina concentration and continuous PFC emission.

During measurement, the change trend of feeding interval is described in Figure 5 together with cell voltage. It is obvious that feeding is under good-control. No clear correlation was found between feeding mechanism and continuous CF4 emission. As compared Figure 5 with Figure 4, it can be seen that detectable CF4 emitted soon after setting new anode. Bath temperature and alumina concentration were measured during PFC monitoring (see Figure 6 and Figure 7). Bath temperature was measured with K-thermo-couple which has autoreading function. Alumina concentration was measured with LECO-RO500C.

300

600

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Running time, min

Figure 7 Trend of alumina in bath during measurement

300

600

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Results from LX176-435 Cell Measurement duration is 45.61 hours for LX176-435 cell. No AE peaks were observed during monitoring as well. Continuous CF4 emission is near to zero during the whole measurement (see Figure 8). Change trends of cell voltage and feeding interval are plotted (see Figure 9). As compared Figure 9 with Figure 8, no clear correlation was found between feeding mechanism and continuous CF4 emission. Metal tapping and anode exchange caused rising of cell voltage. It can be seen that metal tapping could cause a little continuous CF4 emission. Bath temperature of LX176-435 was also measured. Bath temperature changed in a narrow range of 924-933 °C during the whole measurement (see Figure 10). There is no clear evidence to show continuous CF4 emission cause by unstable bath temperature.

1200 1500 1800 2100 2400 2700 3000

Running time, min

Figure 6 Trace of bath temperature during measurement The maximum bath temperature difference in LX176-334 cell was 13 °C . As compared Figure 6 with Figure 4, it seems that continuous CF4 easily emits at lower bath temperature. The influence of bath temperature on continuous PFC needs further investigation. Benefit from good feeding mechanism, alumina concentration in the cell showed stable and in good-control (see Figure 7). The percentage of alumina concentration in the range of 2.0%~3.0% D.00T

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Figure 9 Change curves of cell voltage and feeding interval during measurement exchange and metal tapping disturb the stability of cells, and could cause continuous emissions. Continuous PFC emission from LX176 cells was almost zero. Influences of bath temperature and alumina concentration on continuous PFC emission are not clear in the survey of LX176 cells. Perhaps low line current could play a role in continuous PFC emission. More surveys need to be done on higher amperage cells. Future Plan

0

300

600

900

Anode exchange and metal tapping cause instability of cells in the short-term, which may result in continuous PFC emission. More studies should be done to determine further whether anode exchange or metal tapping cause continuous PFC emission. Continuous PFC emission may take place when bath temperature goes down. More investigation should be carried out to determine the further influence of bath temperature. More detailed survey need to be done to find out if local alumina concentration influences PFC emission. More investigation is needed to determine if higher amperage (>300 kA) is easier to cause continuous CF4 emissions than lower amperage cells like <200 kA. Reduction of PFC emissions requires efforts from smelters all over the world. ZRI would like to strengthen cooperation with IAI in the future. More surveys will be done with the help of a more powerful and flexible equipment donated by Canadian Aluminum Association and Environment Canada as part of the Asia Pacific Partnership on Clean Energy and Climate (4).

1200 1500 1800 2100 2400 2700 3000

Running time, min

Figure 10 Trace of bath temperature during measurement There was a good feeding mechanism, stable bath temperature, quite homogenous alumina concentration in LX176 potlines. Therefore, very low AEF took place in LX176 potlines and the continuous PFC emissions in the measured cells were almost zero. Because anode exchange and metal tapping disturbed the stability of reduction cells, they would have some effects on continuous PFC emission. Perhaps line current in LX smelter was a bit low so that it had very low continuous PFC. No clear correlation was found among bath temperature, alumina concentration and continuous PFC emission.

Acknowledgement This work is supported by The National Natural Science Foundation of China (No. 50974127) and The National Key Technology R&D Program (No. 2009BAB45B03).

Conclusions Different cells have different CF4 emission level even with the same line current. Continuous PFC emission from QY smelter is only a small portion of total PFC emissions. The measured variability of current flow among the anodes is similar to that of Western cells. No obvious relationship was found between maximum anode current and continuous PFC emission. Anode

References 1. Protocol for Measurement of Tetrafluoromethane (CF4) and Hexafluoroethane (C2F6) Emissions from Primary Aluminum

313

Production. U.S. EPA (Washington, D.C.) and IAI (London, U.K.). April 2008 2. Results of the First Survey in CHINALCO Smelters, Zhengzhou Research Institute of Chalco, May 2009 3. Perfluorocarbon Emissions and Primary Aluminum Production, Jerry Marks, February 2010 4. Asia Pacific Partnership on Clean Energy and Climate, Available at http://www.asiapacificpartnership.org/, September 29, 2010.

314

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Monitoring airfluorideconcentration around ALUAR smelter in Puerto Madryn (Chubut Province, Argentina) 1

Jorge Zavatti1; Claudio Lopez Moreno1; Juliana Lifschitz1; Gabriela Quiroga1 ALUAR Aluminio Argentino SAIC, Ruta Nacional N° A010 s/n; Puerto Madryn, Chubut, U9120OIA, Argentina Keywords: Fluoride inmission, air monitoring network, method validation The emission of Total Fluoride (gaseous + particulate) (F) is a significant environmental aspect of the aluminum industry. Therefore the government have established regulations aimed at controlling the environmental performance of smelters through monitoring of F" emissions and inmission.

Abstract The emissions of Total Fluoride (gaseous + particulate) (F) are one of the significant environmental aspects of the aluminum industry. Therefore the government has established regulations aimed at controlling the environmental performance of smelters through monitoring of F" emissions.

The Authority of Chubut Province (Argentina), sets emission limits at 1 kg FVton Al and following recommendations from the World Health Organization, a 16.0 μg F7m3 guideline for inmission has been established [3].

Chubut Province (Argentina), sets emission limits, 1 kg FVton Al and following recommendations from the World Health Organization, a 16.0 μg F/m3 guideline for inmission has been established.

Although measurements were started years before, since 1998 ALUAR has been operating an Air Quality Monitoring Net (RMCA stands for its initials in Spanish), with the purpose of systematically measuring the air Total F concentration in the city of Puerto Madryn and surroundings.

In this paper we present: i)the monitoring network that ALU AR deployed in Puerto Madryn, 12 air sampling stations; ii)sampling procedures and analysis adopted; iii)the results of air fluoride concentration from July 2009 to June 2010; iv)the validation of these results by three scientific institutes of Argentina.

Sampling and analysis are based on ASTM D3267 [4] and ASTM D3270 standards. However some modifications have been made on sampling in order to simplify the procedures. Two samples are taken every week from each and every one of the RCMA stations, sampling periods being of 72 and 96 hours. The results are then recalculated for and average of 24 hours.

The results show that in Puerto Madryn (100000 inhabitants; 2 km to the South of ALU AR's Plant), the median concentration in air was 1.61 μg F/m3 (N = 471 - p95 = 5.32). These values met the inmission guideline level determined by the supervisory authority.

In this context some doubts were publicly raised concerning the accuracy of the information get by ALUAR. These doubts were mainly based on the fact that the sampling was not made following in all the details an international standard. Some objections were also made to the design and location of the monitoring stations

Introduction The ALUAR smelter was originally started in 1974 with 400 SWPB and open pots, working at 150 kA. Starting at 1987 the original pots were transformed to PFPB and gas treatment centers were installed. This and other technical improvements allowed the original potlines to reach now 200 kA, with a high current efficiency and a good general performance [1]. In 1999, ALU AR started its expansion using API8 pots, later upgraded to AP22. This expansion process is still underway.

Therefore, and in order to ensure the quality of the information sent to the authorities and communicated to the residents of the city, Aluar decided to ask to three well known research centers in Argentina (CNEA [5], INQUIMAE [6], PLAPIQUI [7]), to make independent monitoring of the F in air around the smelter.

ALUAR site (42° 44' 21.35" S - 65° 02' 46.73" W) (4 potlines, 736 PFPB cells and 36,000 ton Al/month), is located at 2 km to the North of the city of Puerto Madryn, with 100000 habitants, in the Province of Chubut.

It was required to each Institute to select points close to some of the existing monitoring stations, to design and install sampling stations at these points, to take samples representing at least 1500 hours and to analyze them. The sampling and analysis procedures should be selected by each Institute. The ASTM D3267, ASTM D3268, ASTM D3269 and ASTM D3270 standards were used.

Puerto Madryn is located in an environmentally sensitive area of the Atlantic coast of Argentina, characterized by the presence of flagship species of marine mammals like the Southern Right Whale.

The monitoring by the three Institutes, which made their work in a completely independent way and without exchange of information between them, was made between July 2009 and June 2010. During that time, ALUAR made his own monitoring as usual allowing a direct comparison of the results.

Due to the presence of these species, some of which are endangered, in 1999 the UNESCO inscribed the Peninsula Valdes, very close to Puerto Madryn, on the World Heritage List [2]. The region is also a very important touristic area. The climate of Puerto Madryn is characterized by low rainfall (200 mm/year), prevailing winds from the southwest and an annual average temperature of 13.6 ° C.

315

• Due to practical reasons the ALU AR sampling is made for 72 or 96 hours, while the regulations consider the average concentration for 24 hours. Therefore, the values are affected by a factor >1 to take into account the difference in the sampling time [8] Purposely, both corrections tend to increase the final result. In this way, ALU AR always takes a "safe" position with reference to the adherence to the environmental regulation. The reported concentration could be overestimated but never underestimated. Table 1: CNEA validation of results from the ALUAR-RMCA. Monitoring period from July the 28th to November the 13th of 2009. Research Center: RMCA-ALU AR CNEA Station H (3.8 km to the West of ALU AR) | Median: 1.35 F μg/m3 1 Median: 0.34 F μg/mi 3 3 Percentile 5: 0.21 F μg/m Percentile 5: 0.21 F μg/m Percentile 95: 4.10 F μ/m3 Percentile 95: 2.34 F μg/m3 N:67 N:32 Station E (0.7 km to the East of ALU AR) | Median: 1.03F^g/m3 Median: 1.79F>g/mj Percentile 5: 0.32 F μg/m3 Percentile 5: 0.84 F μg/m3 Percentile 95: 6.79 F μg/m3 Percentile 95: 3.82 F μg/m3 N:31 N:67 Station O (0.8 km to the North of ALU AR) | Median: 0.48 F μg/m3 Median: 1.12 F>g/m^ Percentile 5: 0.31 F μg/m3 Percentile 5: 0.60 F μg/m3 Percentile 95: 3.45 F μg/m3 Percentile 95: 2.60 F μg/m3 N:65 N:31 The number of samples for station differs between CNEA and the RMCA since CNEA took samples of 24 hours of duration, while RMCA Net covers periods of 72 and 96 hours.

Figure 1. ALUAR air sampling network at Puerto Madryn city. Results The results produced by the RMCA show that in Stations A, B, C, N and D located at urban areas in Puerto Madryn city (100000 inhabitants; 2 km to the South from ALU AR's Plant), the median of concentration of F" in air is 1.61 μg /m3 (N = 471 -p95 = 5.32). These values met the inmission guideline level determined by the supervisory authority (16 μg F/m3).

All the air F" concentration values shown in tables 1, 2 and 3 proves that the emission generated by ALU AR operations, do not exceed the corresponding guideline for inmission (16 μg F/m3), set by the environmental authorities of the province of Chubut (Argentina).

The average concentrations of these urban stations do not differ significantly between them and it was not possible to link the temporal evolution of these F concentrations to the wind performance.

At the same time it is worth mentioning that the three research centers that participated in the validation of data systematically report concentration values that are lower than those reported by RMCA in each instance.

Tables 1, 2 and 3 show the results of air F concentration obtained by the Research Centers and those obtained by the ALUAR RMCA, taken simultaneously and at points very close between them, for different periods of monitoring.

Conclusion • The monitoring by three independent and prestigious Institutes has confirmed the correctness of the F" inmission values reported by ALU AR

In most cases the ALU AR values are higher than those obtained by the Institutes. These differences can partially be explained by two reasons:

•The values reported by ALU AR tend to be higher than those obtained by the Institutes. This is explained by the use of corrections that ensure that the concentrations can be overestimated but cannot be underestimated.

• The ALU AR values are corrected in order to take into account the loss of absorbent liquid during the sampling. The ASTM D3267 does not consider any correction for this loss.

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• The results of this study validate the inmission results reported by ALU AR to the authorities since 1998.

References [1] P. Navarro, C. Daviou, L. Daurade. "ALUAR'S al20 PROJECT: A SUCCESSFUL WAY UP TO 200kA". IV International Aluminium Congress (May 18th, 19th and 20th, 2010-Säo Paulo): 594-609. [2] http://whc.unesco.org/en/decisions/2534 [3] World Health Organization. "Air Quality Guidelines for Europe 2nd Edition", WHO Regional Publications, European Series, No. 91 (2000): 143-145. [4] American Society for Testing and Materials (ASTM) (www.astm.org). [5] CNEA http://www.cnea.gov.ar/cac/utt nuevo/unidades 1 .htm [6] INQUIMAE - http://www.inquimae.fcen.uba.ar [7] PLAPIQUI - http://www.plapiqui.edu.ar [8] D. Bruce Turner. "Workbook of Atmospheric Dispersion Estimates". (1994): 4-8 to 4-9.

• The F" concentrations around the ALU AR smelter satisfy the regulations sets by the authorities for the environmental protection • The results confirm that the potroom operating procedures and emission control equipment, originally started in 1988, are still operative for the environmental protection, in spite of the large increase in the production of the plant. Table 2: INQUIMAE validation of results from the ALUARRMCA. Monitoring period from December the 4th of 2009 to March the 19th of 2010. Research Center: RMCA - ALU AR INQUIMAE Station O (0.8 km to the North to ALUAR) Median: 1.66 F μ^ιη Percentile 5: 0.47 F" μg/πrJ Percentile 5: 0.54 F μg/m3 3 Percentile 95: 2.31 F μg/m Percentile 95: 2.84 F μg/m3 N:24 N:26 Station C (3.5 km to the South to ALU AR) Median: 0.84 F μ^πΓ Median: 1.44 F μg/m Percentile 5: 0.46 F μg/m3 Percentile 5: 0.20 F μg/m3 3 Percentile 95: 1.58 F μg/m Percentile 95: 5.15 F μg/m3 N:27 N:30 Station D (3.5 km to the South-East to ALU AR) Median: 1.98F>g/m3 Median: 1.11 F μg/m Percentile 5: 0.73 F μg/m3 Percentile 5: 0.38 F μg/m3 3 Percentile 95: 9.82 F μg/m3 Percentile 95: 1.91 F μg/m N:26 N:27 Both INQUIMAE and RMCA samples were taken covering periods of 72 and 96 hours. Table 3: PLAPIQUI validation of results from the ALUARRMCA. Monitoring period from June the 2nd to August the 9lt of 2010. Research Center: RMCA - ALU AR PLAPIQUI Station E (0.7 km to the East to ALU AR) Median: 1.95 F μg/m3 Median: 2.69 F μg/mJ Percentile 5: 1.17 F μ^πι3 Percentile 5: 0.66 F μg/m3 3 Percentile 95: 6.90 F μg/m3 Percentile 95: 8.13 F μg/m N:20 N:29 Station C (3.5 km to the South to ALU AR) Median: l!53F]Ig/m1 Median: 0.21 F μ^πι' 3 Percentile 5: 0.52 F μ^πι3 Percentile 5: 0.04 F μ^πι 3 Percentile 95: 4.28 F με/m3 Percentile 95: 0.99 F μ^πι N:20 N:27 Station G (2.8 km to the North-East to ALU AR) Median: 0.19 F^g/m 3 Median: 2.46 F μg/πr Percentile 5: 0.09 F μg/πrs Percentile 5: 1.13 F μ^πι3 3 Percentile 95: 0.40 F μg/m Percentile 95: 8.51 F με/πι3 N:22 N:20 Both PLAPIQUI and RMCA samples were taken covering periods of 72 and 96 hours.

317

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

REDUCTION OF ANODE EFFECT DURATION IN 400KA PREBAKE CELLS Wei Zhang1, David Wong1, Michel Gilbert1, Yashuang Gao1, Mark Dorreen1, Mark Taylor1 Alton Tabereaux2, Melinda Soffer2, Xiaopu Sun2 Changping Hu3 Xueming Liang4, Haitang Oin4 Jihong Mao5, Xuehui Lin5 1 Light Metals Research Centre, the University of Auckland, New Zealand institute for Governance and Sustainable Development, US 3 China Nonferrous Metals Industry Association 4 Henan Zhongfu Industrial Co. Ltd, China 5 Northeastern University Engineering& Research Institute, China Keywords: aluminium electrolysis, PFC, anode effect stress to the cell lining from expansion and contraction, melting of the side ledge causing the electrolyte composition to change and increased anode consumption. Other negative impacts include increased hydrogen fluoride emissions, increased operator intervention, exposure of operators to safety risks during manual AE termination, bath spillages and floor damage. The primary aluminium industry is striving to decrease both the number and duration of anode effects [5]. Moreover, with the prospect of a future price on carbon, the emission of these greenhouse gases will impact production costs. It has been calculated that with a C0 2 tax of 15 US $ /ton C0 2 , each anode effect minute per day per cell will increase production cost by about 1.2% [7].

Abstract In order to improve energy efficiency and reduce green house gas emissions, the aluminum smelting industry has been continuously working on reducing both anode effect frequency (AEF) and duration (AED). However, there is still a long way to go to achieve zero anode effect (AE) on very high amperage, low specific power consumption cells due to the added complexity of the process. A new program to quickly terminate AEs has been developed by Light Metals Research Centre, the University of Auckland, in conjunction with the efforts of the Asia Pacific Partnership on Clean Development and Climate (APP) to facilitate investment in clean technologies and to accelerate the sharing of energy efficient best practices. A pilot project was initiated to test an automatic Anode Effect Termination (AET) program on 400kA cells in Zhongfu, China. This paper demonstrates the success of the new anode effect termination (AET) program in killing AEs on this cell technology without conflicting with normal cell operations. The resulting decrease in average anode effect duration (AED) is demonstrated.

The frequency and the duration of anode effects depend upon the technology of the cell, the operational procedures and also upon process control at each smelter. Hence, the amount of CF4 and C2F6 may vary from one plant to the other, according to these parameters. Some types of pot allow more sophisticated interventions to minimize the consequences of an anode effect, such as computer process to control the feed and voltage [8]. Several strategies for AE quenching have been tested and implemented in aluminium smelters, such as pumping or quenching of the anode system, or lowering anodes until they are short-circuited by the metal [9]. Recently, a new AE quenching method has been proposed by Pablo Navarro et al. While a pot is in an anode effect a high proportion of electrical current is conducted from the sides rather than the bottom surfaces of the anodes. This is because the insulating gas layer considerably hinders the flow of current through their bottom surface. This produces an increase in the horizontal current densities and therefore the magnetohydrodynamic (MHD) instability in a pot during the AE. With Navarro's AET strategy, the anode beam is lowered to a particular ACD, at which the MHD instability of the pot is further increased, generating a self-sustained metal wave that short-circuits the anodes, resulting in removal of the insulating gas layer and ultimately termination of the AE. An intense bath circulation is also generated, redistributing alumina within the bath. The particular ACD required to do this is characteristic of each pot technology and depends upon the MHD design. Compared to the traditional AET strategy of physically immersing anodes in the metal pad to short circuit the anodes, this new approach is more energy efficient and faster in killing AEs [9-11].

Introduction The primary aluminium production process has been identified as the largest anthropogenic source of two kinds of perfluorocarbon (PFC): tetrafluoromethane (CF4) and hexafluoroethane (C2F6). PFCs comprise one category of major greenhouse gases with long atmospheric lifetimes and significant global warming potential (GWP). PFCs have an estimated atmospheric lifetime of 10,00050,000 years, and an estimated Global Warming Potential (GWP) of 6,500 and 9,200 times that of carbon dioxide (C02) [1]. Human-made PFCs are very effective absorbers of infrared radiation, so that even small amounts of these gases contribute significantly to the radiative forcing of the climate system [2]. Therefore the aluminium smelting industry has been trying to reduce PFCs emissions [3, 4]. PFCs are generated during anode effects (AEs) in aluminium electrolysis process. Anode effects occur when the concentration of dissolved alumina in the molten bath reaches low values or anode current density is higher than the critical current density [5]. During an anode effect, an insulating layer of gas bubbles appears under the anodes, increasing pot voltage from 4-5V to more than 8V, up to 40-50V, which depend on the type and operating conditions of pots) [6]. In addition to causing environmental damage from the production of PFCs, AEs have other negative impacts. These are increased energy consumption, current efficiency loss, overheating the cell,

The Institute for Governance and Sustainable Development (Washington, DC, USA) initiated a project to develop generic

319

computer software for AE termination in China, which complemented the on-going work of the Asia Pacific Partnership Aluminum Task Force PFC Management Project. Support from this project provided the means for the development of a fast automatic anode effect termination (AET) program based upon the new AE fast kill strategy as discussed in [8]. The program is aimed at terminating AEs as fast as possible to reduce PFC emissions from aluminium smelters. The program has been implemented on 36 cells on the 400kA potline in Zhongfu Smelter, Henan province, China.

Some other factors were also taken into consideration during the beam movement tests. Firstly, there shouldn't be bath spillage during the beam movement. Secondly, the anode cover shouldn't collapse during the fast beam downward movement. Hence, during the tests the bath level change and anode cover condition were also observed. The 2nd stage of the project was to program an AET module to activate the automatic beam movement for AE termination after AE is detected. In order to make sure that AE is not to reoccur, two key procedures in the program included: (1) After downward movement, the beam was held at the lowered position for 15 seconds before it goes up; and (2) Extra fast feed of 115 liters of alumina was programmed into the AET module to increase the alumina concentration in the bath after AE is detected. As the beam moves down, the bath stirring caused by the instability during short circuiting increases the transfer of alumina into the bath under the anodes. Moreover, when beam is dropped down and ACD reduces, bath gets displaced. When beam returns up and ACD increases again, this displaced bath comes back under the anodes, and can bring with it more alumina and help to quickly terminate AE. The anode current density is reduced as more anode surface area is submerged into the bath. All these factors contribute to quick AE termination. After AE termination is confirmed, the pot control system goes back to its normal control program.

Before the AET project started in Zhongfu, AEs are killed manually. In order to determine the effectiveness of the AET program, baseline anode effect duration (AED) was compared with the AED of the test pots. The baseline AED, i.e. 28s, of the 400 kA potline was calculated using 36 pots' 3 months data from a normal operation period, i.e. excluding pot start up and amperage fluctuating period. As a result of implementing the AET program, the average AED has been reduced to 13s, achieving 53% reduction compared to the baseline AED. Experimental To develop and implement the new AET program, the project was conducted with five key stages. Beam movement tests in the 1st stage of the project were used to determine the particular beam positions, at which MHD instability in the pot generates self-sustained metal waving to short-circuit the anodes. As a result, the gas film at the anode bottom surface during AE is removed and then the AE is terminated.

The 3rd stage of the project was to test the newly programmed AET module. During the tests AEs were triggered for a number of pots by stopping the alumina feed to the test cells. The automatic beam movement and feeding behavior of the cells were observed after AEs started. AE durations were recorded accordingly. At this stage, an adjustment to beam upward movement was made to the AET module. Instead of moving back to its original position, the beam was returned up to a position 2-3 mm lower than its original position after AE termination. The reason is that there is some bath loss due to the downward movement of the beam. When the beam moves down, the bath level increases. When the beam goes up, the bath level decreases. However, during this process some of the bath will remain on the cell sidewall, carbon blocks and/or the bottom of the crust, resulting in the total volume of bath being reduced. To compensate for bath loss and maintain a relatively stable bath level, which is particularly important for low level bath pots, the beam's upward movement distance was adjusted to 2~3mm less than its downward movement distance. During the AET module testing, it was also observed that due to the difference in feeding hardware of different cells the extra fast feed did at times result in different feed intervals. However, all the AEs were successfully terminated with the AET module. The cell voltage stayed stable after AE termination.

The experiments were carried out on different pots with different pot conditions to ensure the results were reliable. Beams of these pots were progressively lowered, up to 20mm from original beam position, to observe how pot stability changed with different beam positions. Pot stability can be measured by "pot noise", which is due to fluctuations in anode-cathode distance (ACD). Pot noise can be calculated as the difference between the maximum and minimum pot resistance reading within a specified period of time - this is called "peak-to-peak noise" or PPN [12]. In the beam movement tests, these specified periods of time are numbered for different beam positions as "stage 1, stage 2, etc." The noisiest stage was determined based on the PPN, and the corresponding beam position is the particular position which can generate MHD instability to kill AEs. The pot noise indicator, PPN, was calculated with the following equation (1): PPN=

Max (Pot resistance of Stage n) - Min (Pot resistance of Stage n) (n=l,2,3,4)

After the success of AET module tests, the project proceeded to its 4th stage: a 7-week single pot trial. One test pot operated with the AET program integrated into its control system for 7 weeks. The purposes of the trial were: (1) to ensure that the AET module was functioning as programmed; and (2) the AET program did not conflict with the pot daily operations and other automatic control functions.

(1)

The pot pseudo-resistance, R, was calculated with the following equation (2): n

, j

N

K (onm) = v

J

Working voltage (V)-1.65V

.^.

Working current (A)

v y

—

——

(z)

The 5th stage of the project was a 4-week 72 pots trial. The new control system including AET program was installed for 72 cells. The purpose of this trial is to further verify the reliability of the AET program. Within these 72 pots, 36 pots operated with the

Wherein working voltage is the cell voltage measured as pot to pot voltage on the bus work.

320

Table 2 Beam drop distances for tested cells at different stages

AET module switched on, and the other 36 pots were operated with the AET module switched off, which means that when an AE occurs, manual kill will be carried out. The addition of the AET module switch is to give the process engineers an option to disable the AET module in case of any abnormalities on the cell. Results and Discussions (1) Beam movement test results The beam movement tests first started at a normal pot. The PPN and beam position change is shown in Figure 1. From Figure 1, it can be seen that the pot instability increased with the beam moving down. When the beam was dropped for 1520mm, there was a sharp increase in PPN. Therefore, a similar beam downward movement was tested for other pots with other conditions, as listed in Table 1.

Beam Drop Distance (cm)

Cell Number

Stage 1

Stage 2

Stage 3

Stage 4

5045

0

5

20

0

5051

0

10

20

0

5056

0

12.5

17.5

0

5059

0

10

20

0

5069

0

10

20

0

Π î

U.3

i

\

1

- 0 . 4 ° n 3 0 U.3 im u

1 0.2 z

Q.

°- 0.1 -

k HΦ

1

/ t

Ë

k

*~A fc-*

- * - C e l l 5045 1

*Y

\ < **,

-*-Cell5051 - * - C e l l 5056 HiHCell5059 — C e l l 5069

t

u n ~ Stage 1 Stage 2 Stage 3 Stage 4

CL Q_

Figure 2. Noise change for different cells during beam movement test

0.10

2 3 Stage

From these tests, it was observed that 15-20mm of downward beam movement resulted in significant instability in most test pots during the beam movement tests. However, for termination of an AE, the anode beam should not need to move down to the same extent. The instability already present in the cell during an AE would likely mean that only 10-15mm beam movement would be required to generate the same level of metal waving to shortcircuit the anodes and kill an AE.

i

PPN «A-Beam Drop

Figure 1. Pot instability change with beam position Table 1 Conditions of test cells for beam movement tests Cell Number

Cell Condition

5045

High bath level

5051

With carbon dust, skimmed before beam movement test

5056

High metal level

5059

Low metal level

5069

With carbon dust, not skimmed before beam movement test

During the beam movement tests, cells with different conditions responded similarly to beam down movements, i.e. no cover collapse, no bath spillage, with bath level returning to normal after the beam was returned to its original position. These results indicated that beam downward movements of up to 20mm could potentially be used to kill AE at Zhongfu smelter's 400kA prebake cells. (2) AET module tests An automatic AET module was programmed based on the beam movement tests, which activated downward beam movements of 10-15mm to kill an AE. Also extra 115 liters of alumina feed of was triggered by the module to restore alumina levels in the bath, terminating AE and also preventing AE from reoccurring. Figure 3 is the Screenshot of one AET module test on a pot before, during and after AE with the AET module functioning.

Figure 2 shows the PPN change with different beam down distances. The beam downward movement distance of each test cell at different stages is shown in Table 2.

From Figure 3, it can be seen that when AE occurred, the beam automatically lowered to kill the AE and went back to a position reaching the set voltage of the cell. Meanwhile, feeding was activated. The test had an AED of only 5s and demonstrated that the AET module functioned as programmed.

The test pots had highest PPN values at stage 3, as shown in Figure 2, indicating the cell had the highest instability at stage 3.

321

t » « 0*12 51 *m *12 1&kk i ;:*ta** «is.ss» o T W O . #*4474.n

Current

Automatic beam | up Cell voltage

Figure 3 AET module functioning as programmed (4) Group pots trial After the 7-week single pot trial, the AET module was integrated into the control system of 72 pots. For the 36 pots with the AET module enabled, AEs were successfully killed during the trial period with an average AED of 13 s, a performance which matched the single pot 7-week trial. The average AED of the other 36 pots with the AET module disabled, i.e. AEs were killed manually, was 17s, 4s longer than those pots with the AET module enabled. No conflicts with daily operations were observed in the 4 weeks among the 72 pots. This again proved that the AET program can reduce the AED at the 400kA prebake cells. It was also observed that the average AEF of the test pots was not changed after implementing AET program compared to the baseline AEF.

Altogether seven AET module tests were carried out during the 2nd stage of the project. It was confirmed that the AET program was able to kill AEs quickly and safely. Results of the tests are listed in Table 3. Table 3 AE durations (AEDs) during AET module testing Test 7 2 4 6 5 1 3 Number AED(s) Average AED(s)

25

13

19

35

9

5

32

20

(3) Single pot 7-week trial After stage 2 was finished, the AET module was integrated into the control system of one 400kA pot in Zhongfu. The single pot trial ran from July 6th, 2010 to August 19th, 2010. Nine AEs were recorded during this period. The average AED was 13s. No conflicts between the AET module and daily operations were observed during the trial period.

Conclusions A fast anode effect termination program has been developed and tested at 400kA prebake cells. After implementation of the new AET program, the average anode effect duration was reduced to 13s, achieving 53% reduction compared to baseline anode effect duration of 28s. The results reveal that the new anode effect

322

l l . J . Marks et al. "Factor Affecting PFC Emissions from Commercial Aluminium Reduction Cells, Light Metals, 2001, 295. 12. Alton T. Tabereaux, Presentation at workshop in Zhongfu Smelter, November, 2009. 13. International Aluminium Institute, Results of the 2009 Anode Effect Survey, Report on the Aluminium Industry's Global Perfluorocarbon Gases Emissions Reduction Programme (July 5, 2010)

termination program has a huge potential to reduce anode effect duration in Chinese aluminium smelters, which accounted for 13 million tonnes (MMt) of annual aluminum production during 2009 [13] and most of which still rely on "manual kill" to stop anode effects. Implementation of similar automatic AET programs at all Chinese manual kill smelters have the potential to reduce the annual PFCs emissions by about 4.6 MMt C02-eq. [13] and would contribute to overall improvements in energy efficiencies in the primary aluminum production process worldwide. Acknowledgement The authors thank the Institute for Governance and Sustainable Development, the Henan Zhongfu Industrial Company, Limited, China Non-Ferrous Metals Industry Association (CNIA), Northeastern University Engineering and Research Institute (NEUI) for their collaboration in this endeavour. They also thank Sally Rand of United States of America Environmental Protection Agency, Xue Guohui from the R&D Centre of Zhongfu Smelter and Liu Rui from CNIA. Financial assistance was provided by the United States of America Department of State Bureau of Oceans, Environment and Science Office of Global Change. References 1. United States Environmental Protection Agency, Office of Atmospheric Programs, Global Mitigation of Non-C02 Greenhouse Gases (2006). 2. Forster P. et al., Changes in Atmospheric Constituents and in Radiative Forcing. In: Climate Change 2007: The Physical Science Basis. Contribution of Working Group I to the Fourth Assessment Report of the Intergovernmental Panel on Climate Change (2007). 3. Markus Gräfe, Greg Power and Craig Klauber, The AsiaPacific Partnership: An Important New Initiative for a Sustainable Alumina Industry, Light Metals 2009, 5 4. EPA 43-R-03-006- May 03. "Protocol for Measurement Tetrafluoromethane (CF4) and Hexafluoroethane (C2F6) Emissions from Primary Aluminium Production". US Environmental Protection Agency (Washington) and The International Aluminium Institute (London). 5. Warren Haupin, Edward J. Seger, Aiming for Zero Anode Effects, Light Metals 2001, 329 6. P. Homsi, M. Reverdy, the Reduction of PFC Emissions from Electrolysis Cells, Proceeding of 6th Australasia Aluminium Smelting workshop, 1998, 691 7.Thonstad, J., Utigard, T. A., Vogt, H., On the anode effect in aluminum electrolysis, in TMS. 2000. 8. Carlo Braga, et al. Faster Anode Effect Kill, Light Metals 2007, 417. 9. Pablo Navarro et al. A New Anode Effect Quenching Procedure, Light Metals 2003, 479. 10. Alton T. Tabereaux, Maximum Anode Effect Voltage, Light Metals 2007, 405

323

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SUSTAINABLE ANODE EFFECT BASED PERFLUOROCARBON EMISSION REDUCTION Neal R. Dando1, Lise Sylvain2, Janice Fleckenstein1, Ciro Kato3, Vince Van Son4 and Laura Coleman5 1 Alcoa Technology, 100 Technical Drive, Alcoa Center, PA 15069-0001 USA 2 Alcoa Canada Primary Products, 1 Boulevard des Sources, Deschambault, Quebec GO A ISO Canada 3 Alcoa Pocos de Caldas, Rod. Pocos-Andradas, Km 10, Pocos de Caldas 37719-900 Brazil 4 Alcoa, 2300 North Wright Rd, Alcoa, TN 37701 USA 5 Alcoa Global Primary Products, 6641 West Broad Street, Suite G100, Richmond, VA 23230 USA

An increasing number of smelters opt for plant-specific measurements of PFC emission factors, since these factors more accurately represent the true emission rates for that particular plant, as opposed to IPCC default Tier 2 coefficients. To facilitate this process, Alcoa has a long-established (> 15 years) program in place to ensure that PFC emissions are measured in a comprehensive manner at each Alcoa smelter on a recurring basis. As a result of this program, Alcoa has the largest Tier 3 data base in the aluminum industry. Alcoa's Tier 3 coefficients have also been used to improve the accuracy of the default Tier 2 coefficients for multiple cell technologies.

Abstract Commercial aluminum reduction is a significant anthropogenic emission source of the perfluorocarbon (PFC) gases carbon tetrafluoride (CF4) and hexafluoroethane (C2F6). Given the high global warming potentials of these two gases (6500 and 9200 respectively), concerted efforts are underway to monitor and reduce emissions of PFCs from aluminum smelters. Alcoa has been performing plant-specific perfluorocarbon (PFC) emission testing at operating smelters for over 15 years and has the largest data base of Tier 3 coefficients in the aluminum industry. From 2005-2009, Alcoa smelters have reduced anode effect (AE) based PFC emission intensity by 0.48 T C02e/T Al produced, resulting in an absolute annual reduction of 2.3 million tonne CO^ over this same time period.

Most aluminum companies have initiated voluntary programs for actively reducing PFC emissions4 and all modern pre-bake smelters have implemented automated methods for terminating anode effects. These methods vary by cell technology however all have the same goal of minimizing the cumulative anode effect minutes per cell-day at the operating location. Perhaps more importantly, smelter pot feeding and monitoring programs are continually optimized to minimize the frequency of anode effect events.

These systematic, sustained PFC reductions were the result of 1) continuous attention to detail and process optimization at operating locations 2) consistent sponsorship and interest from all levels of the organization 3) comparative monthly performance reporting on anode effect performance, 4) internal workshops regarding shared AE reduction efforts and 5) production technology changes. This paper will discuss examples of these key enablers for sustaining AE-based PFC reductions of 15.9 million tonnes from Alcoa smelters over the period 1990 to 2009.

The smelting operations of larger, established aluminum companies typically encompass different smelting cell technologies, each typically having significantly different alumina feeding and control practices. This situation precludes a "one-size fits all" approach to anode effect reduction programs. The purpose of this report is to discuss contributing factors that led to a systematic, sustained reduction in annual PFC-based CO^ emission of 15.9 million tonnes from Alcoa smelters over the period 1990 to 2009.

Introduction Perfluorocarbons (PFCs) are known greenhouse gases with exceptionally long atmospheric lifetimes.1 These trace gases are linked to global warming due to their strong infrared absorption in the upper atmosphere. The two most common PFC gases emitted to the atmosphere are carbon tetrafluoride (CF4) and hexafluoroethane (C2F6). Aluminum smelting is considered to be the largest anthropogenic source of these PFC emissions worldwide.2 PFC emissions from aluminum smelters occur during a transient process condition known as an anode effect (AE), which occurs when the alumina (A1203) level in the bath drops below - 1 % , the cell voltage rises, and the bath itself begins to react.3

Discussion In 1995 January, Alcoa entered into a five-year, voluntary agreement with the U.S. Environmental Protection Agency (EPA) known as the Voluntary Aluminum Industrial Partnership (VAIP),. The EPA-VAIP was a cooperative effort to measure and reduce the principal PFCs, CF4 and C2F6j produced in the U.S. Aluminum Industry. The original VAIP program goal was to reduce US PFC emissions from aluminum smelting by 30 to 60%, from 1990 levels, by the year 2000. The actual performance achieved over this time period (by all US partners) was a 57% reduction in PFC emissions.5 The PFC reduction observed over this same period for Alcoa smelters worldwide was 58%, from 17.7 to 7.4 million tonne C02e> as shown in Figure 1.

Estimates of a smelter's PFC emissions (kg PFC/tonne Al produced) can be calculated by multiplying the total number of anode effect minutes (AE minutes/cell-day) observed by a PFC slope coefficient [(kg PFC/ tonne A1)/(AE min/cell-day)]. The PFC slope coefficients for (CF4) and (C2F6) can be estimated using Tier 2 default coefficients from the good practice guidelines of the Intergovernmental Panel on Climate Change (IPCC)4 or they can be calculated from plant-specific measurements of PFC emission rates (Tier 3 coefficients).

Initial successes in anode effect reductions (relative to 1990 levels) were easier to achieve, owing to pre-existing practices that

325

year locations that were the "most improved" and that had the lowest AE minutes/cell-day were recognized to acknowledge their accomplishment.

used anode effects themselves as a simple process control tool for de-mucking pots and validating feed control program with respect to alumina concentration. An early key enabler was education on how modern pot control strategies and measurement technologies could effectively eliminate the need for deliberate, or scheduled, anode effects as a process control tool.

A 0.5 million tonne reduction in AE-based C0 2e was observed during the first year of the One Million Tonne Challenge program, from 4.1 MT (2005) to 3.6 MT (2006), as shown in Figure 1. During the second year of the program (2007), aggregate PFC performance remained essentially static and in response a worldwide PFC reduction workshop was held at an Alcoa smelter. Representatives from all Alcoa smelters participated in three days of meetings to: 1) 2) 3) 4)

Figure 1. AE-based PFC CO^ performance of Alcoa smelters, where the Y-axis in million metric tons (MT). Over the next 5 years, from 2000 to 2005, Alcoa smelters continued to reduce anode effect-based PFC emissions by an additional 45%, from 7.4 to 4.1 million tonnes CO^, as shown in Figure 1. By 2005, however, it became apparent that additional, systematic gains in PFC reductions were going to require a renewed focus and additional attention to assure continued performance in line with Alcoa's sustainability goals.

Share "local successes" or "best practices" regarding process or work practice optimizations to reduce anode effect minutes. Share "local challenges" regarding AE minutes/cell-day reductions Learn more about AE root causes, and R&D based developments for improving feed control and anode kill routines. Participate in like-technology breakout sessions to identify additional tools/options for continuing to reduce AE minutes/cell-day at their locations.

Table 1 lists 19 Alcoa aluminum smelters in eight countries currently in operation. Alcoa smelting technologies encompass point-feed prebake (PFPB), side-work prebake (SWPB), vertical stud Soderberg (VSS) point feed vertical stud Soderberg (PFVSS) and horizontal stud Soderberg (HSS). These cell technologies require significantly different local approaches to achieve and sustain reductions in AE minutes/cell-day. Table 1. Operating Alcoa Smelters

1 Million Tonne Challenge In 2005 Alcoa challenged its global network of aluminum smelters to reduce annual CO^ emissions from anode effects by an additional one million tonnes. The One Million Tonne Challenge initiative was sponsored and monitored by top Alcoa executives

Country Australia

Baie Cornea u

PFPB & VSS

Becanour

PFPB

Descahambault

PFPB

Iceland

Fjardaal

PFPB

Italy

Portovesme

PFPB

Lista

PF-VSS

United States

326

PFPB

VSS

Spain

In addition, root causes for performance concerns as well as planned countermeasures were solicited from locations that significantly exceeded their target. At the end of each calendar

PFPB

PFPB

Norway

1) Locations that established new performance benchmarks 2) Locations that met or beat their target 3) Locations that met or beat their target for 4 or more consecutive months

Portland Point Henry Sao Luis

Canada

A One Million Tonne Challenge Team was charged with establishing AE minute/cell-day targets for each plant as well as tracking and globally communicating individual plant and aggregate performance against targets on a monthly basis. Key performance metrics and milestones were noted:

Technology |

Pocos de CaIdas

Brazil

The One Million Tonne Challenge was based on the following premise: if each smelter sustainably closed 90% of the gap between 2005 annual AE minutes/cell day performance and each plant's own best monthly performance in 2004, a collective reduction of one million tonnes of C0 2e per year could be achieved.

Smelter

Mosjoen

PFPB

Aviles

PF-VSS

La Coruna

PF-VSS

San Ciprian

PFPB

Intalco

SWPB

Massena

PFPB & HSS

Mount Holly

PFPB

Warrick

PFPB

Wen a tehee

PFPB

1

A PFC Best Practice Team was formed to assist locations that consistently missed their monthly performance targets - or who requested assistance to enable them to push past "plateaus" in local performance. The team was composed of "core" and "invited" members wherein the invited members were typically Technical Managers from like-technology smelters, relative to the location being visited.

An additional one million tonne reduction in AE-based CO^ emission was achieved in 2009, with respect to a 2008 baseline (2.8 vs 1.8 MT). The net AE-based CO^ emission reduction performance over the period 2005-2009 was 2.3 million metric tonne, or 56% (4.1 vs 1.8 MT). The systematic reductions in AE-based C02e, discussed above, offer a significant contribution to the total direct C0 2e from Alcoa's worldwide operations, shown in Figure 2. From a 1990 baseline, direct CC>2e has been reduced by 43%. Approximately 80% of the CO^ reduction from 1990 to 2009 is based on AEbased PFC reductions from smelting operations.

A PFC Best Practice Team would typically spend a week at a given smelter performing several tasks that can generally be listed as: 1) 2) 3) 4)

Review plant operating history and key performance indicators with plant personnel Audit anode kill practices on the plant floor Walk the floor during normal operation and anode change (or stub change) to observe plant practices Inspect pot conditions (metal level, feeder condition, anode cover practices, etc).

At the end of the week, the Team would make a presentation to the plant operating team with a list of specific, cost-effective, prioritized recommendations or action items to enable further reductions at that location. A time-line would be agreed upon for implementation. 1990

One of the most readily available opportunities for reducing AE frequency is to use the potroom computer system to report pots exhibiting higher than average AE's per shift (or day). AE action sheets can be used to document (and develop Pareto charts of) the root causes for atypical AE behavior and serve as the basis for implementing countermeasures to systematically reduce the number of high AE frequency pots. Point feeder issues tend to dominate the Pareto analyses developed by this approach.

4) 5) 6)

2006 Year

2007

2008

2009

Given significant financial constraints regarding the world economy and more specifically the aluminum industry over recent years, it is interesting to note the relative changes in aluminum production, CO^ intensity and mass CO^ emissions (the latter two being from AE-based PFCs). Figure 3 shows the relative performance of these metrics from 2005 through 2009, where CC>2e intensity is expressed at T CC>2e/T Al. As indicated in the data shown in Figure 3, over the period from 2005-2009, Alcoa achieved a 56% reduction in AE-based C02e, while only a 7% reduction in aluminum production occurred over this time. The large 56% mass-based CO^ reduction and the 54% CO^ intensity reduction are both due to the significant, sustained AE minute/cell-day reductions achieved at Alcoa smelters over this period.

Additional systematic, sustained PFC reductions were enabled by several projects and practice optimizations, such as:

3)

2005

Figure 2. Total direct C02e and AE-based CO^ emissions from Alcoa operations (in million tonnes).

At side break smelters (prebake or Soderberg), variations in crust breaking/feeding tend to cause similar variations in cell AE frequency. These variations can be often be tracked to differences in crust breaking equipment (stroke length) or breaking practice (full side vs partial side).

1) 2)

2000

Eliminating root causes for long duration anode effects, Use of predictive tools to alarm pot room operators to take preventive actions before anode effect occurrence, Technology retrofits - from Soderberg side break technology to point feeder technology, Revised pot start-up procedures with more uniform, higher temperature target pre-heating practices to improve pot stability during first hours of pot operation, Technology-specific countermeasures for variations in alumina properties, Improved standardized procedures to minimize the impact of power outages and load modulation.

Alcoa Smelters - C02e from PFCs » A I Prod

The impact of the PFC Reduction Workshop and PFC Best Practice Teams was evident in 2008 performance, as shown in Figure 1, wherein a 1.3 million tonne reduction in AE-based C0 2e emission was achieved, with respect to a 2005 baseline (4.1 vs 2.8 MT).

IMC02e

-ilrlntensity

Figure 3. Aluminum production, C0 2e mass emission and C02e intensity.

327

Systematic reductions in either AE frequency, AE duration or both contribute to an overall reduction in PFC-based CC>2e intensity. Table 2 shows the production-weighted AE frequency and AE duration for Alcoa smelters in 2005 and 2009. Significant reductions were achieved in both metrics over this period; however the major lever contributing to an overall C0 2e intensity reduction was AE frequency. As reported previously, reductions in anode effect frequency can offer a linear route to CO^ reductions while a focus on fast anode effect kills (to minimize duration) can reach a point of diminishing returns, given that PFC emission rates (per second of anode effect) are highest during the initial AE onset.

2.

R. Zander, S. Solomon, E. Mahieu, A. Goldman, C. Rinsland, M. Gunson, M. Abrams, A. Chang, R. Salawitch, H. Michelsen, M. Newchurch, and G. Stiller, "Increase of Stratospheric Carbon Tetrafluoride (CF4) based on ATMOS Observations from Space," Geophysi. Res. Lett, v. 23 1996, 2353-2356.

3.

K. Grojtheim and H. Kvande, eds., Introduction to Aluminum Electrolysis, (Dusseldorf: AluminiumVerlag), 1993.

4.

Protocol for the Measurement of Tetrafluoromethane (CF4) or Hexafluoroethane (C2F6) emissions from Primary Aluminum Production, April 2008, http://www.epa.gov/highgwp/aluminumpfc/documents/measureprotocol.pdf

5.

Voluntary Aluminum Industrial http://www.epa.gov/highgwp/aluminumpfc/accomplishments.html

Table 2. Production Weighted AE Frequency and Duration Year

AE Frequency Prod Weighted

Calculated AE Duration Prod Weighted

(#/pot day)

(mins)

2005

0.51

1.58

2009

0.29

1.15

% Reduction

44%

27%

1

6.

7.

Conclusions The sustained reduction of AE minutes/cell-day at operating aluminum smelters requires consistent sponsorship by leadership, the coordinated collaboration of Like Technology Teams, Best Practices Teams, R&D personnel and - most importantly - the continuous efforts of potroom floor personnel to identify, implement and transfer tools and enabling options for achieving and sustaining reductions in AE minute/cell-day performance. In the early stages of anode effect reduction programs, significant reductions could be achieved by ending past practices of using anode effects as a process control tool. The magnitude of the reductions already made and sustained will make achieving additional reductions far more challenging. Given the range and age of smelting technologies in operation, new options for continuing to push to lower AE minutes/cell-day need be explored on a plant-specific basis. At Alcoa, the sustained, combined efforts of monthly performance tracking, periodic PFC workshops, inter-plant best practice teams and potroom floor personnel have allowed for a systematic AEbased reduction from 17.7 to 1.8 MT CO^ on an annual basis. Further reductions from existing pot operations will require even more intense effort and innovation than have been expended to date.

Acknowledgements The authors would like to acknowledge the significant efforts of our smelter operating personnel to achieve and sustain these CC>2e reductions and Alcoa for permission to publish this work. 1.

References A. Ravishankara, S. Solomon, A. Turnipseed and R. Warren, "Atmospheric Lifetimes of Long-Lived Halogenated Species," Science, v. 259, 1993, pp. 194199.

328

Partnership,

J. Marks, A. Tabereaux, D.Pape, V. Bakshi, and E. Dolin, "Factors Affecting PFC Emissions from Commercial Aluminum Reduction Cells," Light Metals, 2001, pp. 295-302. N. Dando and W. Xu, "Root Causes of Variability Impacting Short Term In-Plant PFC Measurements," Light Metals 2006, 189-194.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

THE INITIATION, PROPAGATION AND TERMINATION OF ANODE EFFECTS IN HALL-HEROULT CELLS TMS1, Gary Tarcy2 and Alton Tabereaux3 TMS (The Minerals, Metals & Materials Society); 184 Thorn Hill Rd.; Warrendale, PA 15086, USA 2Alcoa Technical Center, 100 Technical Drive, Alcoa Center, PA 15069, USA

lr

Keywords: l Alcoa Inc; Consultant 1)

Abstract Anode effects in Hall-Heroult cells have been the subject of multiple investigations and studies. The current state of the knowledge is fairly well advanced and there is very little discrepancy or controversy with respect to many of the phenomena associated with anode effects. Included in this is the belief they are: 1) AE are the predominate emitter of PFC into the atmosphere 2) AE are triggered by low alumina concentrations near the anode surface 3) Short circuiting of at least part of the anode cathode inter electrode gap is required to terminate an anode effect. This paper will cover some of less discussed aspects of anode effect including: 1) the initiation at a single random anode in the circuit; 2) propagation to multiple anodes until the whole circuit is on anode effect and; 3) reason it is necessary to short circuit the anodes to terminate anode effect.

2) 3)

4)

Introduction

Figure 2 below shows the changes in C0 2 and CO concentrations measured off of a single anode as the pot approaches anode effect (2). The measurements were made by drilling a hole and inserting a steel pipe completely to the bottom of the anode to assure a minimum contribution from the Boudouard reaction. Infrared was used to measure both the CO and C0 2 concentrations. The significance of this measurement with respect to anode effect initiation is that there is a change in the primary reaction at the anode as the anode effect is approached. The usual reaction:

The Initiation and Propagation of Anode Effects The initiation of anode effect is known to be caused by the depletion of alumina at which point a critical current density is reached and the pot then goes into anode effect. Although not fully confirmed, the most respected correlation for the critical current density has been published by Piontelli (1). Figure 1 depicts how the Piontelli correlation predicts the critical current density to vary for a typical pot. The critical current density becomes smaller (easier for the cell to exceed and thus have an anode effect) for either lower alumina concentrations or cooler temperatures.

A1203 + 3/2C * 2A1+3/2C02

1

1

1

1

(1)

which is thermodynamically favored but kinetically hindered due to the difficulty in breaking carbon to carbon bonds get replaced with the thermodynamically favored but usually kinetically hindered reaction

Critical Current Density vs. Alumina Concentration for Different Bath Temperatures 3 I

Change in the electrolysis reaction from being a predominately C0 2 producing reaction to more of a CO producing reaction. The CO production result is twice as much gas. Dewetting of the anode as the alumina becomes lower thus having less real anode surface available for electrolysis Different anodes draw different amounts of current and the alumina concentration becomes more non-uniform in concentration in the pot as the pot gets closer to anode effect. Resulting in the anode effect usually being initialized at a single anode and then propagating to the rest of the anodes. The overall voltage on the pot begins to increase due to the increased bubble coverage on the pot and the change in the anode reaction to a more kinetically unfavorable condition.

A1203 + 3C -»2 AI + 3CO

1 Temperature T3

(2)

The low alumina concentration in the pot is the cause for the shift in reactions. Although under most circumstances it is more difficult to break the carbon-carbon bonds the increasing lack of oxygen bearing anion leads to the competing reaction becoming more favorable. Finally at anode effect due to even further depletions of alumina approximately 16% of the gas produced from the anode becomes CF4 and this gas is evolved from the bottom of the anode. The balance of the gas at anode effect is essentially CO and this is produced off the sides of the anode. 1

2

3 4 5 Alumina concentration [weight percent]

6

7

Figure 1. Critical current density vs. alumina concentration for different bath temperatures Mechanistically this correlation can be explained by:

329

Figure 4 shows the position of the under loading anode where the anode effect was first detected. The study was on 13 pots over a two week period. All positions were responsible for the initiation of the anode effect there is no pattern by which the positions trigger an anode effect. In other words the position effect is essentially random.

Figure 2:Changes in C 0 2 and CO concentration off a single anode as Anode effect is approached

Figure 4. Distribution of the first unloading anode with respect to position in the pot

Figure 2. Changes in C0 2 and CO concentration off a single anode as a Anode effect is approached. Figure 3 below shows the onset of anode from an Alcoa P-225 pot. Due to the special design of the pot every anodes current is measured and recorded by the computer every second. The figure shows the usual sequence of events. Prior to the onset of anode effect (in this example it is approximately 3 minutes before) one of the anodes begins to under load with respect to the rest of the group due to the change in reaction cited above. The CO producing reaction in the Hall-Heroult case requires more energy to run and produces more gas thus insulating the anode that is running out of alumina at the fastest rate. This shifts the amperage (in this example 50% of the amperage) that the under loading was carrying to the other anodes in the circuit. The overall low alumina concentration in the pot then leads to more anodes reaching the critical current density In the end the anode effect eventually spreads in a cascading series phenomenon as the current density increases significantly on the remaining anodes until all anodes in the cell are on anode effect.

Figure 5 shows the distribution of the "warning time" from the under loading anode prior to anode effect. The signal detection criteria shows that 95% of the anode effects can be detected in sufficient time (30 seconds prior to anode effect) that action should be able to be taken to prevent anode effects.

F^

>8fcre Anode Effect I

Figure 5: Distribution of the Early Warning of Anode effects 10 i

Individual Anode Readings Aproaching and During Anode Effect

.01

.1

1

5 10 2030

50 7080 90 95

99

99.9 99.99

Percent

It is clear that anode effect starts at a localized location probably due to a local low alumina concentration under some anodes. This is caused by a combination of a local low superheat, insufficient transport of alumina underneath the anode, or carbon dust blocking the surface of the anode. The location of anodes with the lowest alumina was shown above to vary in the cells in essentially a random manner.

rloading 10

12

14

16

18

20

Figure 3. Individual Anode Current Readings Approaching and During Anode effect

Modeling results by Feng et.al (3) demonstrate that constraining the bath flow underneath anodes by reducing the ACD results in greater extremes of alumina concentration. It was also reported

330

anode beam (with all anodes) are moved down to decrease the anode to cathode distance (ACD) to cause electrical short-circuits between anodes and aluminum waves. Anodes effects can often be killed quickly <5 seconds using this practice.

that there is a stagnant zone in the ACD under the middle of each anode because the alumina in that area is being consumed at a faster rate than it is being replaced during the under-feed period. This seems to indicate that the anode effect initiation may even be more localized than even a single anode and could start in even localized areas of the single anode.

In addition to the rapid feed of alumina to bath by point feeders during the anode effect, the increase in anode immersion and the intense bath circulation produced by the metal wave causes the liquid bath to come into contact with the anode cover which is a source of alumina all around the pot. Another source of alumina is by some sludge dissolution that is removed by the metal waves from the cathode surface [9]. The intense bath circulation due to metal waves and high bath temperature (>1000 °C) quickly increases the bulk alumina concentration in the cell bath and importantly makes it more homogeneous throughout the cell.

Moxnes et.al. (4) reported that the alumina feed was changed from a "flat, equal feed in the five point feeders to an equal or "flowadaptive" feed among the five feeders in the SU4 potline at Sunndal, Norway. The alumina distribution in the cells was made more equal by increasing the alumina feed at the two feeders nearest the ends of the cells and decreased the alumina feed in the feeders near the center of the cell. As a result the number of anode adjustment, anode spikes and deformation decreased. It is further anticipated that the more uniform alumina distribution in the cell will result in decreasing the anode effect frequency as well as decreasing sludge residing in the cathode.

The isolating layer of PFC gas is removed by the combination of short-circuits on anodes, and the decrease in anode current density due to the deeper anode immersion when the ACD is reduced on all anodes.

Finally, one more interesting observation. There is a small fraction of the time when there is no warning at all that an anode effect is about to occur. If this observation is coupled with the reality that the only actions that can be taken to terminate the anode effect is to lower the anode beams and feed alumina. Lowering the anode beams is the fastest possible countermeasure that could be employed to prevent anodes but this action will take at least several seconds to be effective. Using a 30 second recognition time to react to the anode effect we can calculate the absolute minimum anode effect frequency that we can get if we were 100% successful in preventing every anode effect we observed with the anode under loading technique at least 30 seconds prior to the full anode effect occurring. The calculations are only for the Alcoa P-225 situation and may not extrapolate to all the other technologies but we have calculated the absolute minimum anode effect frequency to be 0.002 anode effects per pot day. This is both encouraging and discouraging. It is encouraging because it points out there remains plenty of room for further improvement, but it is discouraging because it also means that a true zero rate frequency may not be possible with the technology we know of today.

Conventional methods of extinguishing anode effects by electrical shorting were not effective in a drained TiB2-G cathode cell as it did not have an aluminum metal pool (5). All attempts to extinguish anode effects by electrical shorting the pilot cell failed The only effective method of extinguishing anode effects during drained operation involved the momentary interruption of power to the cell. Impact of Short-Circuiting During AE Termination on PFC Emissions Lower PFC Emissions From Soderberg Cells Marks (6) reported revised Tier 2 PFC slope coefficients that demonstrate that the emission rates for PFC gases is substantially less for Soderberg cells, (0.092) compared with prebake cells (0.143-0.272) during anode effects. When the pot is on anode effect a high fraction of the electrical current is conducted from the sides of the anode, because the isolating gas layer considerably hinders the flow of current though the bottom surface. This produces an important increase in the horizontal current densities and therefore of the MHD instabilities of the pot during the anode effects. The anode effect is inherently unstable in Soderberg cells as the bath immersion on the sides is only about one tenth of that for prebake cells. The pot develops a wave in the bath-metal interface that provides local short-circuits to the Soderberg anode. Tabereaux et.al. and Marks et.al. (7) reported that intermittent short-circuiting observed in Soderberg cells is linked to a decrease in PFC emissions.

The highly irregular current distribution after the anode effect is due the after effect oscillations in the metal pad surface. Terminating Anode Effects To sucessfully terminate (or kill) anode effects in operating cells requires a short circuit to the molten aluminum metal. Once the short-circuit occurs in cells the anode effect is killed instantly. •

•

Anode effects are killed by first adding alumina to the bath of cells and then manually inserting a wood pole under anodes into the metal, resulting in a large expulsion of combustion gases that causes molten aluminum to splash upward and short-circuit with anodes. This manual process usually requires minutes, and even longer in some circumstances

The emission rate of CF4 gas from the Soderberg cell shown in Figure 6 is ~46% less than prebake cells that have continuous emission with minimal short-circuits. The lower rate of PFC emissions due to electrical shorting agrees with the lower values determined for the Tier 2 PFC Slope factors for Soderberg cells (0.092) compared with center-break and side-break prebake cells (0.14 to 0.27)

Anode effects are typically killed in prebake cells using automatic anode effect termination programs in the computer system by: 1) rapidly adding a sufficient amount of alumina in order to increase the alumina content in the bulk bath all around the cell and, 2) the

331

circulation: Fast down move of anode typically results in one small emission peak of PFC gases during the anode effect (9). Conclusion: Anode effects are initiated by the local depletion of alumina at a single anode resulting in a shift in the primary gas producing reaction. This shift leads to production of twice the amount of gas for the same amount of current which leads to a decrease in the amount of current carried by this anode. The load is shifted from this anode to other anodes and the trend continues until the whole cell comes to anode effect. The anode effect must then be terminated by s short circuit resulting in breaking of the PFC insulating film. The usual method of short circuiting also results in partial immersion of the anode cover into the liquid bath along with increased stirring of the bath from the high degree of instability. This raise the alumina concentration overall and diminishes any localalize alumina depleted regions of the cell. References

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Figure 6. Electrical Shorting During Anode Effect of VS Soderberg Cell Differences in PFC Emission Rates in Prebake Cells The major difference in PFC emission coefficients (determined by the "slope" and "anode overvoltage" methods) between individual prebake smelters especially those using the same technology, is largely due to differences in the anode effect kill strategy. The emission of PFC gases stop, or are greatly reduced each time anodes are short-circuited with metal. Thus the manner and timing in which the anodes are lowered to cause short-circuit with the metal pad waves in order to kill the anode effects has a great influence on the rate of PFC generation in prebake cells (8). Automate Downward 8eam Movement * Uitrs-Fast Alumina feeding -.

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1.

Piontelli, R., Mazza, B. and Pedeferri, P. (1965). Ricerche sui fenomeni anodici nelle celle per alluminio. Metallurgia Italiana, 57(2), 51-69.

2.

Internal Alcoa Letter A. J. Sharkins to Rolf Rolles 1979-1226.

3.

Y. Feng, M.A. Cooksey, and M.P. Schwarz, "CFD Modelling of alumina Mixing in Aluminum Reduction Cells", Light Metals 2010, pp.455-460.

4.

B. Moxnes, A. Solheim, M. Liane, E. Svinsas and A. Halkjelsvik, «Improved Cell Operation by Redistribution of the Alumina Feeding", Light Metals 2009, pp.461-466. G.L. Fredrickson,"1999 Pilot Cell Operations - Final Report", (DOE Project # DE-FC07-97ID13567), March 2000.

5.

6.

Marks, J., "Method for Calculating the PFC Emissions From Primary Aluminum Production," Light Metals 2006, pages 185-188.

7.

Marks, J., Tabereaux, A.T., Pape, D., Bakshi and E. Dolin, "Factor affecting PFC Emissions From Commercial Aluminum Reduction Cells", Light Metals 2001, pp.295-302.

8.

A. Tabereaux, "Maximum Anode Effect Voltage", Light Metals 207, pages pp.405-410.

9.

P. Navarro, Gregoric G., Cobob O., and Calandra, A., "A New Anode Effect Quenching Procedures", Light Metals 2003, pp.479-486.

Figure 7. Anode effect Killed in Prebake Cell by Electrical Shorting with Fast, Aggressive Down Moves. Slow AE kill with anode pumping by slow multiple cycles of reducing and increasing the ACD until strong short-circuit with metal waves: Slow pumping of anodes results in a nonlinear decreasing emission of PFC gases during the anode effect. Faster AE kill with modified anode pumping utilizing faster, deeper ACD down moves until strong short-circuit with metal waves: Faster pumping of anodes results in a less irregular emission of PFC gases during the anode effect. Fast AE kill with large, aggressive ACD down moves causing fast, strong short-circuit with metal waves as shown in Figure 7: Deep fast down move of anodes results in one small linear peak emission of PFC gases during the anode effect. 4.

Fast AE kill with a smaller ACD down move until MHD instability occurs and a self-sustained wave develops that causes instant short-circuits on anodes and intense bath

332

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

TOWARDS ELIMINATING ANODE EFFECTS Ali. AI Zarouni1, BJ. Welch2, Maryam Mohamed Al-Jallaf1, A.Kumar1 ^ubai Aluminium Company, PO Box 3627, Dubai, U.A.E 2 Welbank Consulting Ltd, PO Box 207, Whitianga 3542, New Zealand. Keywords: Anode Effect, Anode Current Density, Incipient AE Logic potline was 0.31 AE/cell/day and this has been brought down to -0.05 AE/cell/day by a series of measures.

Abstract It has been established that the atmosphere contains appreciable amount of per fluorocarbon (PFC) gases. These are powerful greenhouse gases with extremely long lifetime. Although these gases are being used in the semiconductor industry, it has been established that aluminium smelters are their main source. Emissions of PFC Gases from the aluminium industry have therefore become an environmental issue. The PFCs are emitted from an aluminium reduction cell when it is on an 'anode effect'. The exact nature of the onset of an anode effect is still shrouded in mystery. However, an astounding reduction in number of anode effects has been achieved by understanding, attributing and implementing a strict process control regime to eliminate assignable causes.

Figure 1. Potline 5B

Dubai's progress in reducing anode effect frequency to less than 0.05 AE/pot-day in a sustained manner in poline 5B has been discussed and presented.

The Anode Effect During normal electrolysis the produced anode gas is a mixture of carbon oxides, with C0 2 being the dominant product formed according to the overall reaction,

Introduction

2A1203 + 3C = 4A1 + 3C0 2

The IAI survey [1] shows an overall 86% reduction in the emission rate for PFCs per tonne of aluminium produced between 1990 and 2007, equating to a 74% reduction in total emissions as C0 2 equivalent. The reduction in emission rate is despite the doubling of metal production over the same period.

Normally this reaction proceeds at an anode potential of A13+/A1 of -1.6 volt, with the increase in potential above the minimum (~ 1.18V) being due to both concentration and reaction polarisation [2, 3]. The exact magnitude of the anode potential is a complex function of the anode carbon quality, the electrolyte temperature, the dissolved alumina concentration, and operating current density.

Industrialization and rapid economic growth are major contributing factors to the increase in greenhouse gas emissions, global warming and climate change. The declining rate of PFC emissions is the result of the industry's efforts to reduce the frequency, and the duration of the anode effects in pot line cells. This has been further enhanced through the use of alumina point feeding systems and computer feed control programmes that reduce anode effect frequency. The ongoing phasing-out of older technologies and their replacement with more modern technology, wherever economically justified, is also assisting in decreasing relevant emissions.

Modern smelting operation is moving towards favouring lower dissolved alumina concentration (for better control and resulting efficiency) and higher anode current densities (for increased productivity) and this increases the concentration polarisation, and hence anode potential. The next possible electrochemical reaction product has been shown to be COF2 [4, 5] according to the overall reaction, A1203 + 3C + 2A1F3 = 4A1 + 3COF2(g).

Dubai Aluminium Company (Dubai) continues its untiring efforts in reducing PFC emissions. It commissioned Potline 5B in the year 2007, figure 1. The potline houses D20 cell design. The potline amperage was gradually raised from 240 to 260 kA. The cells operate on an anode current density of 1.045 A/cm2 which is understandably on the higher end in the industry. It was therefore a very challenging and daunting task to lower the AE frequency in this potline. Initially, the anode effect frequency in the

Thermodynamic data suggest this should occur at an anode potential of A137A1 -1.7 to 1.8 V, depending on the alumina and aluminium fluoride concentrations [6]. Coincidentally the magnitude of increase is consistent with the working range of control limits " resistance" increase during under and over feed strategy.

333

The exact mechanism of the above reaction, and the overall kinetics is less certain, but the presence of 3COF2(g) [4]. and its electrochemical formation [5] has been established in carefully controlled laboratory research. It is clear that passivation of the electrode surface plays a role, but the cause and form that it takes is a subject of debate.

Issues Impacting AE Frequency Following factors can all contribute to the onset of an anode effect,

Figure 2 shows the controlled voltage sweeps of a small cell at different alumina concentrations [7]. The electrode process changes through concentration polarisation and subsequently after a limited amount of electrolysis passivation causes the current to drop to almost zero.

I Λ

•

Addition of alumina too late or too slowly, i.e. below the limit sustainable by the operating anode current density.

•

On occasions, the anode current density exceeding the critical current density.

•

Maintenance issues within the cell, e.g. blocked feeder holes, empty ore bins, hindered transfer of alumina into the feeder hole.

•

Delayed sensing of the voltage rise.

•

Insufficient depth of immersion of anodes (very low bath level), anode burn off, etc.

The potline work pattern is a 4-shift work cycle - metal tapping idle anode setting idle shift. An analysis of the anode effects in D20 cells was performed as under: •

Applied Potential

(V)

Figure 2: Potentiodynamic voltammogram of a carbon anode electrolysed in cryolitic electrolyte, (a) =0.75 wt% A1203, (b) =1.75 wt% A1203, (c) =2.75 wt% A1203, (d) =4.55wt% A1203. From Haverkamp[ 6]

The majority of anode effects occurred during anode setting operation[9], figure4. The anode setting activity temporarily increases anode current density probably to the critical density threshold which triggers an anode effect. AEF vs Potline Work Schedule

Under the approximately constant current conditions of an operating smelter, the highly resistive nature of the passivating film causes an ohmic increase in cell voltage (often manifested by arcing) as the electrolysis continues, figure 3. Tap

Idle 1

Work Schedule

Figure 4. AEF vs Work Schedule The alumina feed cycle is super fast (SF)^over feed (OF) base feed (BF)^under feed (UF). The super fast feed is for a very short duration. The anode effects were concentrated in the over feed window, figure 5. AEF vs Alumina Feed Windows

Figure 3. The rapid rise in voltage following the onset of an anode effect Measurement in operating cells has also demonstrated that an anode effect often starts in a region of a cell causing a decrease in the current of the affected anode and a corresponding rise in the current of the others [8]. This implies that cells have concentration gradients, or varying current densities at different zone.

BF SF UF Work Schedule OF- over feed, BF - base feed, SF - super fast, UF - under

Figure 5. AEF vs Alumina Feed Windows

334

and to the difference in gas coverage of the anode surface.

This means that the demand feed was not always able to efficiently react in time to prevent an anode effect. Hence digital signal filtering becomes a contributing

Line 6 : AEF vs Noise Activations 0.140

It was established in the other potlines that potline amperage increase had an adverse effect on AE frequency [9], figure 6. The amperage increase has pushed the anode current density and also the rate of alumina depletion.

0.120

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0.40

AE Frequency

Figure 8. Impact of Noise Activations on AEF Actions & Results 2001

2002

2003

2004

2005

2006

First, the operating practices were revisited and fine tuned. Later, logic to 'kill' an incipient anode effect was developed and introduced.

2007

Year

Figure 6. Impact of Line Load increase on AEF.

• The anode setting and anode covering practice were designed to minimise spillage, figure 9. The controls were specifically targeted to permit a maximum amount alumina feeding from point feeds.

The Dense Phase System for transporting alumina to the cell hoppers primarily operates on high pressure, low velocity concept. A drop in pressure was found to have a profound impact on hopper filling and consequently on the anode effect frequency.

The changes included auto anode mark transfer - this minimized subsequent anode adjustments to achieve an optimum current distribution and also minimized the need to redress the anodes. The anode top cover thickness and the composition of the cover material were monitored closely - this helped in minimizing variations in the 'anode top heat losses'.

Pencilling of crust breaker tips lead to insufficient alumina quantity reaching the bath. Bath build up on the crust breaker hindered the delivery of the prescribed alumina dose at the right time as illustrated in figure 7.

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Figure 7. Bath Build Up on Crust Breaker Tip.

The bath chemistry controls were optimised and designed for cells to operate within a narrow superheat. The bath chemistry control strategy was followed up closely and would take into account the natural cell variations, figure 10.

Unexpected variability in alumina delivery from the conveying and delivery system that is installed in some of the Dubai potlines. A typical problem was an unannounced massive surge in the delivery of alumina fines to a cell or a small group of cells. Analysis of anode effects in relation to 'fines' on a sample basis confirms this, figure 8.

The operation at the end of an underfeed would provide a minimum of overfeed alumina dumps so that the risk of muck formation was minimised.

A weak correlation was observed between AE frequency and other parameters such as bath chemistry and bath temperature. An unstable cell can trigger anode effects at higher alumina concentrations [10]. This is attributed to difference in alumina concentration gradients between the bulk of the bath

335

L5B Correlation - Bath Temperature & %AIF3 12.0

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After implementing the 'Incipient AE logic', the potline operated on 'zero' anode effects on many days. When examining the potline operating data, the days having AEs were generally preceded by significant interruptions to the alumina feeding, usually associated with compressed air pressure. The issues have since been addressed.

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Despite the current density remaining at 1.045 A.cm"2, current efficiency was maintained at 95.4 - 95.5% level throughout, figure 12. This has been possible due to excellent cell stability and is also a confirmation of the robustness of the cell design.

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Year 2010: Current Efficiency (%) 97.0

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96.5

• The next stage was to observe the anode effects and feeding characteristics. The resistance curves were carefully analysed and inferred. A programme was developed to recognise and 'kill' an incipient anode effect. It involved lowering the anodes to short them with metal, breaking the crust and feeding alumina at a rapid rate. The success rate in the 'incipient AE quench logic is normally greater than 90%.

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Year 2009

Feb

Mar

Apr

May

Jun

Jul

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Sep

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The results were encouraging and brought down the anode effect frequency by nearly 78%, table 2. Parameter

2010

Figure 11. Line Load vs Anode Current Density

Process variations through normal operations and control were minimised; where the cell got out of the expected control band, diagnostics were performed prior to taking extreme steps. The principle was to operate the cells on very similar anode-to-cathode distance table 1. The poor performing cells were diagnosed in terms of the voltage drop (or resistance) in different segments, current distribution, and metal pad reserve; corrective actions applied as appropriate. Cull | i | I ioiBii.E.vi.ne'/iiny.'(»(».( &>»>ft«( « c W ^ l Ac Mt <»if IM 1>.W

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After Year 2010

No/ cell/ 0.22 0.069 0.31 0.19 day AEDurn Sec. 83 47 48 33 Table 2: AE data - before and after implementing 'incipient AE logic' AE Freq.

Line load in potline 5B was increased gradually from 240 to 260 kA, figure 11. Further amperage increase was limited by the rectifier limitations. The anode size remained unchanged throughout. As a consequence, the anode current density increased to 1.045 A.cm 2 . Therefore the objective of being able to operate on a low anode effect frequency was even more challenging.

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Year 2009 250.9

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336

6.

Dorreen, M.M.R., Chin, D., Lee, J., Hyland, M.M. and Welch, BJ. "Sulfur and fluorine containing anode gases produced during normal electrolysis and approaching an anode effect", Light Metals 1998 pp 311-316.

The amperage in potline 5B has been raised gradually from 240.0 kA to 260.0 kA. Consequently, anode current density increased from 0.965 to 1.045 A.cm"2. This has made the task of reducing anode effects even more daunting. Compressed air pressure plays an important role in controlling anode effects. The mechanical aspects (crust breakers, feeders, etc) of a cell shall and will continue to interfere with smooth functioning of demand feed. The above challenges make it practically impossible to achieve zero anode effect frequency with our installed hardware.

7.

Haverkamp R G. PhD thesis "Surface Studies and Dissolution Studies of Fluorinated Alumina" University of Auckland, (1992).

8.

J T.Keniry, GC.Barber, M. P.Taylor and B J.Welch, Digital Processing Of Anode Current Signals: An Opportunity For Improved Cell Diagnosis And Control Lisht Metals 2001 pp 1225.

9.

A. Kumar, A. Zarouni and M. Al Jallaf, Initiatives to Reduce SEF at Dubai, Light Metals, pages 259 - 262.

However, a 78% decrease in anode effect frequency could be achieved by tracking the voltage rise and killing an incipient anode effect. The anode effect frequency could be lowered from 0.31 to -0.069 AE/cell/day.

10. J. Thonstad, et al, 'Aluminium Electrolysis', pp 201 206.

Concluding Comments Reducing anode effects is a challenge facing aluminium smelters to minimise wastage of energy and reduce green house gas generation.

The impact of amperage and anode current density increase was kept within the heat balance limits by monitoring the superheat and energy management while ensuring minimum sludge formation as indicated by changes in the cathode voltage drop, cell instability values and alumina content in the bath. Productivity gains have been achieved by maintaining the current efficiency as a consequence of better controls. All this was achieved whilst maintaining an enviable metal purity of better than 99.90% Aluminium. Further work is in progress towards an ambitious target of <0.01 AE frequency. Acknowledgement The authors gratefully acknowledge the guidance, support and encouragement of Dubai management in carrying out this project.. References 1.

International Aluminium Institute, Results of the 2007 Anode Effect Survey, pp 1-6.

2.

Richards, N.E. and Welch, B.J. "Anode Overvoltage in Electrolysis of Cryolite-Alumina Melts", Extractive Metallurgy of Aluminium. Vol. 2, Aluminium (Ed. G. Gerard) pp 15-30 (Interscience, New York) (1963).

3.

Welch, BJ. "Tech. of Electrolyte Reduction of Alumina by Hall-Heroult Process: I. A Voltage Analysis Under Conditions of Varying Alumina Concentrations". Proc. Aust. I.M. & M., No. 214, pp 119 (1965).

4.

Dorreen M M S. PhD thesis "Cell Performance and Anodic Processes in Aluminium Smelting Studied by Gas Product Analysis". Univ. of Auckland, (2000)

5.

A. J. Calandra, C. E. Castellano and C. M. Ferro, 'The Electrochemical Behaviour of Different Graphite / Cryolite Alumina Melt Interfaces Under Potentiodynamic Perturbations" Electrochimica Acta, Vol 24,1979, pp 425-437.

337

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CORRELATION BETWEEN MOISTURE AND HF FORMATION IN THE ALUMINIUM PROCESS Camilla Sommerseth1, Karen Sende Φsen2, Thor Anders Aarhaug2, Egil Skybakmoen2, Asbj0rn Solheim2, Christian Rosenkilde1 and Arne Petter Ratvik1 *Dept. of Materials Science and Engineering, Norwegian University of Science and Technology, NTNU, NO-7491 Trondheim, Norway 2 SINTEF Materials and Chemistry, NO-7465 Trondheim, Norway Keywords: HF, moisture, aluminium process the work was to increase the knowledge concerning where hydrogen fluoride emissions escape from the bath, and the correlation with feed cycle of the cells. An HF/H20 NEO Lasergas II Single gas monitor instrument was used to measure the concentrations of HF and H 2 0 in the cells. A portable "sniffer line" was used to suck gas from various positions in the electrolysis cell. These positions included inside the duct, above the crust, the feeder hole and an open hole in the crust. HF emissions as well as H 2 0 concentration were measured at all these positions. The main objective of this work was to obtain quantitative measurements of the HF levels above holes in the crust and inside the duct.

Abstract Hydrogen fluoride (HF) emission to the working atmosphere is still a problem in the aluminium industry. Moisture in secondary alumina fed to the cell and humidity in the ambient air reacts with fluorides in the bath and fluoride vapours to form hydrogen fluoride. The relation between the various sources of water and the resulting HF emission is still not well understood. In this work, industrial measurements have been done to determine where HF escapes from the bath. The quantities of HF and moisture at the specific sites have also been determined. Measurements were done in the duct during normal operation as well as during anode change, above the feeder hole and above an open hole in the crust. A strong correlation between feed cycle and HF levels was measured. Increased HF emissions were also recorded during anode change.

From the measurements in the duct, average values of HF are found. When measuring over the feeder hole and over other holes in the crust, more knowledge concerning how the hydrogen fluoride evolution is distributed over the cell is found. This way it can be established how moisture in the alumina fed to the cell affects the HF formation. Two main sources of water contribute to the HF formation. These sources are structural hydroxyl in the primary alumina and moisture in the air sucked into the cells [4'5'61.

Introduction HF emissions to the environment have been reduced considerably since the introduction of the dry scrubber in the 1980s. However, emission to the working atmosphere is an issue that remains. Emissions to the working atmosphere mainly happen during work on the electrolysis cells, for example during changing of anodes, tapping, etc. when the covers are removed.

Experimental Setup and Procedure The experimental setup was similar to the one used at SU4, Hydro in May 2010 [3], except for the following adjustments: The drumtype gas flow meter and the cover near the junction between the probe and the tubing were eliminated. Figure 1 shows a sketch of the experimental setup t3]. It was possible to manually switch the ventilation system on the cell from normal to forced gas suction, and this way the ability to manipulate the suction rate on each cell was present. Measurements could be done in the duct on each individual cell relatively close to the crust itself. The magnetic field was moderate in this area of the pot room.

Generation of HF takes place when fluorides in the bath or in the vapour phase react with moisture according to Reactions 1 and 2 below [t\ \ AlF,(diss) + H 2 0(g) = 1 Al203(diss) + 2HF(g) NaAlF4(g) + H 2 0(g) = \ Al203(diss) + 2HF(g) + lNa 3 AlF 6 (g)

(l)

(2)

This work has been done to establish knowledge about the different factors that affect the formation of HF and gain a better understanding of where it evolves in the cells. Different technologies for producing aluminium have different challenges concerning HF emissions, and it is therefore interesting to compare the different technologies. In this paper, it is distinguished between HF formation and emission in accordance with Haupin and Kvande [5]. Formation is the total HF formed by reaction with the different moisture sources. Emission is the amount of HF released to the atmosphere, from both the cell and the fume treatment plant.[2]

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Tuneable Piece for diode laser \ suction of 4 (measuring "false air" Inert gas cell 10 cm) inert gas into outer out from circuit outer circuit^ j

emitter

Vacuum pump

The work was carried out at Alcoa Mosj0en, Norway, in June 2010, to create a basis of comparison with the measurements already done at Hydro Sunndal SU4, Norway[3]. The objective of

Figure 1: Experimental setup of the equipment used at Alcoa Mosj0en [31

339

Earlier experience from measurements using the laser equipment showed that is sensitive to the magnetic field. If the magnetic field is too strong the unit will stop working. Hence, it should be placed in an area of moderate magnetic field. The computer used has a flash disk which makes it suited for bringing into a pot room with a magnetic field. The laser instrument is made to withstand a moderate magnetic field and still giving reliable results. The laser used was an HF and H 2 0 Lasergas II Single Gas monitor from NEO. The PFA tubing between the laser and the probe was flexible and measurements could be done in different areas of the cell. The pump helped sucking the gas from the measuring site through the probe, the tubing and into the laser. The total volume flow through the measuring system was 10-20 L/h. Before entering the laser, the gas passed through a filter to prevent the laser measuring cell from being filled with dust. Nitrogen was purged through the laser measuring cell to eliminate humidity from the ambient air to affect the H 2 0 measurements. Measurements were recorded every ten seconds. The pump, probe, filter and nitrogen were standard equipment at Alcoa Mosj0en. The laser instrument was calibrated in the laboratory before going to Alcoa Mosj0en. Known amounts of moisture entered the laser instrument and the moisture level recorded was compared with the theoretical value. The comparison showed good compliance with the theoretical and recorded values. Measurements were done in the duct during constant feed rate with and without forced suction (open and closed damper). The duct measurements were partly done to verify the system. Subsequently measurements were carried out above the feeder hole and finally over an open hole in the crust.

Figure 2: Probe inside the duct. Notice how the damper can be opened and closed manually.

Results Duct Measurements The probe was placed inside the duct as shown in Figure 2. It was made sure that there were no leakages in the system and that all valves and tubes were reliable. In Figure 2, also notice the possibility to manually open and close the damper while working on the cell. Measurements were done with forced suction (open damper) as well as without forced suction.

2,3

700

2,2

600

2,1

500

2

£ 400

According to various literature data [4' 5], the hydrogen fluoride evolution rate is in the order of 20-35 kg F/t Al. Depending on the gas suction rate, this corresponds to an HF concentration in the duct in the range 200-400 ppm (volume). Figure 3 shows the results from the duct measurements.

1.9 9,

E

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S 300 200

Forced over-feeding

100

A section of Figure 3 is enlarged in Figure 4 to show the correlation between moisture and HF, as well as the response in HF and H20 concentrations after feeding of alumina.

Time

Measurements in the duct were also done during anode change, and the results are shown in Figure 5. Measurements of HF concentrations in the pot room during the same time interval are shown in Figure 6.

1,8 ^ 1,7 1,6 1,5 1,4

Figure 3: HF and H 2 0 measurements in the duct showing variations in the cell when the damper is open and closed. The lower points on the HF graph represents the time just before the feeder fed the cell.

340

Measurements above the Feeder Hole The results from the measurements above the feeder hole are shown in Figure 7, and a photo of the feeder hole is placed in Figure 8.

E

400 A 25000

Probe is moved closer and closer to feeder hole. Maximum value of HF measured to 20201 ppm

2.5

20000 H

HA i\ I xM An oxide layer 1/ \ i s

u. 15000 H

ä 10000 HF H20

5000

Feeding

covers the feeder hole stopping the Qas from emitting

Probe above crust, approaching feeder hole .

Figure 4: A section of Figure 3. This shows how the concentrations of HF and H 2 0 vary with feeding of alumina. The dashed lines indicate when feeding of the cell took place.

3

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1.5: 1 Feeder down, removes the oxide layer

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Figure 7: Measurements done above the feeder hole 500 450 400 350 £ 300

Probe placed in duct, damper open

V

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Bath is covered with alumina

2,5

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E 250 a. f s 200 150 - Four covers Changing of anodes are removed starts, crust is removed 100 from the cell 50 15:5 0:00

3

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16:11:36

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16:18:48

16:26:00

Time HF

H20j

Figure 5: Effect of anode change.

Figure 8: Feeder hole. mm

t&m

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mm

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itm

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Measurements above an Open Hole in the Crust

mmmo Figure 6: Data given by Alcoa Mosj0en of HF emissions to the pot room the 30th of June 2010. There is an apparent peak of HF emission at approximately 16.00-17.00 (this time interval is pointed out by the black arrow) showing that more HF is emitted to the pot room during anode change.

Measurements were done above an open hole in the crust, and Figure 9 shows a photo of the probe above this open hole. Figure 10 shows the results from the measurements over the open hole in the crust. Four vertical lines indicate when the feeder went down into the bath.

341

When the damper was closed the measured levels of HF was as high as 590 ppm. This is a higher value than earlier reported [4'5'61. However, since the gas suction rate inside the duct was not measured, the dilution factor is not known, so the measured and earlier reported values are not directly comparable. Different plants have different technologies and routines, and this might lead to the deviation in predicted value of HF. When operating under forced suction, the gas becomes more diluted and the concentration of HF drops. Also in Figure 3, a correlation between feeding of alumina and HF and H 2 0 levels are shown. A pattern of "peaks" and "valleys" is observed. The "valleys" in the HF graph occur just before the cell was fed with alumina. After feeding the cell the concentration of HF increased considerably, reaching a maximum value (a peak) and then started decreasing. This is shown in more detail in Figure 4. There is a short response time before the concentrations of HF and H 2 0 start increasing after feeding of the cell. This response time is most likely due to delay in the laser instrument, due to the time it takes to move the gas through the probe and the tubing into the laser and/or due to the reaction rate between moisture and A1F3 in the bath. Hence, the formation of HF is almost immediate after the feeding of the cell.

Figure 9: Probe held above an open hole in the crust.

Figure 5 shows the results from the measurements during anode change. These measurements were done to verify what happens with the duct concentrations of HF during anode change. In these measurements forced suction was introduced (by the operators as part of their anode change routine), cell covers were removed. When the anodes are removed a large surface area of the bath is exposed to the ambient air. When the covers are removed, part of the HF escapes into the pot room (as shown in Figure 6), without reaching the duct and hence the dry scrubbers. Also, when the crust is removed more HF is formed by hydrolysis due to the large surface area of the bath being exposed to ambient humidity. When the bath is covered with alumina to create a crust and the damper is closed, the cell returns to normal. The HF concentration increases in the duct and also in the pot room.

HF H20 Feeding

Measurements above the Feeder Hole The results from the measurements above the feeder hole are shown in Figure 7. The figure shows maximum values of HF of about 20000 ppm. At Hydro Sunndal the maximum values of HF measured above the feeder hole was only approximately 10000 ppm[2]. One problem of measuring over a feeder hole is that the probe has to be removed from the hole every time the feeder feeds the cell. Variations in HF levels according to feed rate are not obvious from Figure 7 in the same way as it was observed in the duct. The photo in Figure 8 shows that the feeder hole is very narrow with only a small surface area of the bath exposed to ambient air. When comparing the measured HF levels at Alcoa Mosj0en with the measured levels at Hydro Sunndal, it is obvious that the measured level at Mosj0en was significantly higher than at Sunndal. One reason might be the different shapes of the feeder holes at the two plants. At Mosj0en the feeder holes are narrow, giving a very concentrated flux of gas upwards from the hole. At Sunndal the feeder holes are more open and less narrow, giving a gas that is quickly diluted with the ambient air. This makes it difficult to "pick up" the HF gas at Sunndal with the probe used during the measurements, and prevents an exact comparison of the real variations between the different plants. Further studies of the effect of physically changing the feeder hole geometry in one cell

Figure 10: Results from measurements done above the open hole in the crust. The vertical lines indicate feeding of the cell. Discussion Duct Measurements Measurements in the duct were performed to create a basis for understanding how feed rate, feeder cycle, anode changes and various air suction rates affect the HF concentration in the cell. The measurements in the duct demonstrated a clear correlation between HF evolution and feeding of alumina. The results from the duct measurements are shown in Figure 3. The figure shows time intervals of different feed rate and with and without forced suction. It is worth noting that the concentration of HF decreased when the suction rate increased due to increased dilution of the raw gas. In Figure 3, the time intervals of tracking and overfeeding are shown. Tracking of the cell means that the cell is underfed. At over-feeding the concentration of HF increased.

342

may be useful for determining the reliability of the feeder hole measurements. This was not done in this work.

Conclusion In these measurements it has been established a strong correlation between feed cycle and HF evolution from the bath. It can be concluded that most of the HF evolution happen through open holes in the crust, both feeder holes and other holes in the crust. A maximum value of about 27000 ppm of HF was measured through an open hole in the crust. It can also be established a correlation between moisture and HF levels and they follow each other like "mirrors". The results indicate that the shape and diameter of the feeder holes can be a factor that decides how much HF it is possible to measure with this method.

Measurements above an Open Hole in the Crust Measurements were also done over an open hole in the crust. An advantage of measuring above an open hole versus the feeder hole is that it was not necessary to remove the probe every time the feeder fed the cell. In this way it was possible to determine whether or not it was necessary to be close to the feeder hole to detect any correlation between HF, moisture and feeding rate. The results given in Figure 10 show clear indication of the variations in HF and H 2 0 concentrations during feeding of the cell. This is regardless of where the probe is located as long as it is above an open hole in the crust. The maximum level for HF was measured to be almost 27000 ppm which was significantly higher than the maximum value measured over the feeder hole. The fact that it was not necessary to remove the probe from the hole every time feeding took place, is believed to contribute to the higher maximum value measured. Another observation made was a change in flame colour in the hole during the cycle between every feeding. After feeding, the flame became bluish, turning more and more yellow until a new feeding was done.

Table 1: Summary of the results found in the different measuring campaigns. Location Above crust, far away from feeder hole Above open feeder hole Duct Above closed feeder hole Just above bath, open tap hole In open feeder hole just above bath Above open feeder hole Above open hole in crust Duct (open damper) Duct (closed damper)

Moisture Measurements and Moisture Importance An observation that is worth noting is that the water level in the cell in general follows the HF levels. This is found both when measuring in the duct and above the open hole in the crust. However, this is not the case for the measurements done over the feeder hole as shown in Figure 7. In these measurements the moisture and the HF graphs show a negative correlation. This has to do with the fact that every time the feeder went down, the probe had to be removed from the hole. Also the feeder generated a lot of dust and this might give the apparent "drying" after feeding of the cell. The measurements indicate that the moisture content of the alumina is an important factor contributing to the HF formation. Also, it seems likely that the HF adsorbed on the alumina in the dry scrubber, is released during feeding. This may, however, be determined by comparing primary alumina feeding with secondary alumina feeding.

Range 5-10 ppm*

Comments Oct2008, SU4

Reference [1]

5-200 ppm*

[1]

3000-5000 ppm*

Oct2008, SU4 Oct 2008, SU4 Oct 2008, SU4 May 2010, SU4

9000-10000 ppm*

May 2010, SU4

[1]

10000-20000 ppm* 15000-26000 ppm* 300-500 ppm* 400-600 ppm*

June 2010, Mosj0en June 2010, Mosj0en June 2010, Mosj0en June 2010, Mosj0en

This work

200-400 ppm* 50 ppm

[1] [1]

"

m

This work This work This work

* Fluctuates according to feed cycle

Acknowledgement

Discussion of the Method

The present work was funded by Hydro Aluminium, Alcoa Norway, Alstom, The Norwegian Research Council, and the other companies involved in the KMB ROMA project. Permission to publish the results is gratefully acknowledged. First a great thanks to Ellen Myrvold for welcoming us to Alcoa Mosj0en, helping us with equipment that we needed and arranging everything for us. Thanks to Asbj0rn Kj0nnas at the laboratory at Alcoa Mosj0en for helping us with equipment and answering all the questions we had. Thanks to Ingar Solberg for replying to our numerous e-mails with questions, Sjur Dalsbotten for information about the cells and last a thank you to all the workers in the electrolysis hall for showing consideration to us while we were working in the hall. Thanks are also due to Chris Torjussen and his colleges at NEO for valuable advice in connection with the laser measurements.

A filter was used to prevent dust from the crust and the bath to reach the measuring cell of the laser device. The laser instrument needs a transmission above 70 % and dust in the measuring cell will rapidly decrease the transmission. When measuring at Alcoa Mosj0en, the filter was changed several times. It was not possible to know exactly when the filter needed changing on site. An issue that has been considered, but not investigated is the possibility of the filter to adsorb HF. Tests on this matter can be a source of further work. Summary of the Measurements Table 1 gives a summary of the measured HF levels, both at Alcoa Mosj0en and at Hydro Sunndal.

343

References 1. 2.

3. 4.

5. 6.

K.S. Osen, A. Solheim, C. Rosenkilde and E. Skybakmoen, "The behaviour of moisture in cryolite melts", Light Metals 2009, pp 395-400 M.L. Slaugenhaupt, J.N. Bruggeman, G.P. Tarcy, N.R. Dando, "Effect of open holes in the crust on gaseous fluoride evolution from pots", Light Metals 2003, pp 199204 K.S. Osen, T.A. Aarhaug, A. Solheim, E. Skybakmoen, C. Sommerseth, "HF measurements inside an aluminium electrolysis cell", Light Metals 2011, this issue W. E. Wahnsiedler, R. S. Danchik, W. E. Haupin, D. L. Brackenstose and J. W. Colpitts, "Factors Affecting Fluoride Evolution from Hall-Heroult Smelting Cells", Light Metals 1978, pp 407-424 W.E. Haupin og H. Kvande, "Mathematical Model of Fluoride Evolution from Hall-Heroult Cells", Light Metals 1993, pp 407-411 M. Hyland, E. Patterson and B. Welch, "Alumina Structural Hydroxyl as a Continuous Source of HF", Light Metals 2004, pp 361-365

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PARTICULATE EMISSIONS FROM ELECTROLYSIS CELLS Heiko Gaertner1, Arne Petter Ratvik1, Thor Anders Aarhaug2 1

NTNU, Department of Materials Science and Engineering, NO-7491 Trondheim, Norway 2 SINTEF Materials and Chemistry, NO-7465 Trondheim, Norway Keywords: Aluminium, Off-gas, Particulates, Fluorides, Emissions the chiolite composition, contributes about 35 % of the total fluoride emissions. The coarse faction > 6 μπι, consisting of alumina, carbon and frozen cryolite droplets, represents about 20 % of the total fluorine emissions. Exact particle size distribution was difficult to determine due to the many different components and the wide range of particle sizes.

Abstract In the dry cleaning of the exhaust gas from the aluminium cells impurities are accumulated in the finer fractions of secondary alumina from the dry scrubbers. The present work describes new methods for the determination of dust composition, aiming at increasing the understanding of the effect of cell operation on the amount and the composition of dust in the fume. New and advanced analysis methods are used to characterize a broad specter of emissions. An Electrical Low Pressure Impactor is used to sample and analyze the dust from the cells. The equipment enables real-time particle size distribution analysis of 12 particle classes in the range 30 nm - 10 μιη. The size classified samples are analyzed by means of SEM/EDS and XRD to determine the characteristic chemical composition of the different fractions. Understanding the evolution, evaporation, and condensation of particulates in the cell emissions under different operational conditions may facilitate new standards for environmental friendly and energy efficient high amperage electrolysis cells.

In 2000 Hyland, Welch, and Metson [8] reviewed the current knowledge on fluoride and sulfur emissions from aluminum reduction cells. The emissions were classified in two main categories; gaseous emissions as HF, CF4, C2F6, S0 2 and COS and particulates as Na3AlF6 (cryolite), Na5Al3F14 (chiolite), NaAlF4, A1F3 and CaF2. The following mechanisms are suggested: - Vaporization of electrolyte and subsequent condensation as fine particles - Mechanical entrainment of liquid electrolyte or fines from crust cover material - HF generation due to reaction of electrolyte with hydrogen from the anodes, vapour or particulates with moisture from air and/or alumina

Introduction

Based on models assuming that particulates are mainly condensation and hydrolysis products from evaporated NaAlF4, according to Equation 1, different reactions are considered [9,10,11]. The gaseous NaAlF4 may disproportionate to solid chiolite, Na5Al3Fi4, and A1F3 according to Equation 2 and/or undergo hydrolysis as described in Equation 3.

An understanding of formation, evaporation and condensation of gas compounds in the aluminium electrolysis cell and the off-gas duct system under various operational conditions and current densities are of interest both for setting new standards for environmental friendly production and for efficient operation of high productivity electrolysis cells. It is commonly accepted that beside the bath chemistry the operational practice has a considerable effect on the gas composition and condensation products in the off gas [1,2]. Large current density increases may only be realized by increasing the heat losses to maintain the heat balance of the cell. As most of the surplus heat has to be released with the exhaust, it is likely that the particulates formation in the cell will be affected. A higher temperature may affect the cell superstructure and the subsequent gas treatment [3, 4], however, it may be a requirement for economical heat recovery [5, 6]. Also, cost effective C0 2 capture, if developed, requires less draught air to increase the C0 2 concentration, resulting in higher off-gas temperature. Hence, the composition of the finer particulates is of interest, both from a dust recovery perspective and scale buildup in the duct and in heat recovery units.

Na3AlF6(l) = NaAlF4(g) + 2NaF(l)

(1)

5NaAlF4(g) = Na5Al3F14(g) + 2AlF3(s)

(2)

3NaAlF4(g;s) + 3H20(g) = 6HF(g) + Na3AlF6(s) + Al203(s)

(3)

Hydrolysis of A1F3, Na3AlF6 and Na5Al3Fi4 may occur in contact with moisture in the draught air, Equations 4 and 5. 2AlF3(diss) + 3H20(g; diss) = 6HF(g) + Al203(s; diss)

(4)

3Na5Al3F14(g;s) + 6H20(g) = 12HF(g) + 5Na3AlF6(s) + 2Al203(s)

(5)

Hφflich et al. [12] collected potroom fumes from Soderberg and prebake smelters in a five stage impactor (0.18-0.35 μιη, 0.350.65 μπι, 0.65-1.2 μηι, 1.2- 3.5 μιη and 3.5-10 μιη). The electron microscopy approach to determine the particle distribution substantially underestimates the particle concentrations in the size range below 300 nm as shown elsewhere [13]. With 45-65 % particles containing O, F, Na and Al, a mixture of aluminium oxide and cryolite are most abundant in the potroom air. The ultrafine particles are considered to be condensation and hydrolysis products of vapor compounds from the electrolyte.

Fume emitted by aluminum reduction cells with prebaked anodes have been investigated by Less and Waddington [7]. The fume and dust particles sampled under varying operational conditions were examined by chemical analysis, X-ray crystallography, and optical and scanning electron microscopy. Approximately 50 % of the fluoride emissions were in form of particulates. The fine fraction < 2 μιη, consisting of condensed fluorides approximating

345

The equipment consists of a heated sampling and dilution line (Dekati ejector diluter and DAD-100 axial diluter). A cyclone collects the coarse dust particles before the diluter. The residual fine particles, D50 < 10 μιη is feed to the impactor with a flow of 10 1/minute. Dilution ratios were determined with a Portea 204M/C FTIR by comparing concentration changes in tracer gas before and after the dilution. The Dekati ELPIVI 4.0 software was used to compute the particle numbers. Particles on impactor substrates were dried at 120 °C for 24 hours before weighing and further analysis.

Large carbon particles (up to a few μιη) are most likely entrained dust from the anodes. Due to the complex chemical composition, it is assumed that agglomerates contain additional phases as aluminium oxides, cryolite, silicates and sulfides/sulfates. The authors also noticed needle-like particles. Experimental An Electrical Low Pressure Impactor, DEKATI ELPI™, is used for real-time sampling of 12 particle classes in the size range 7 nm - 10 μιη. The operating principle can be divided into three parts; particle charging, size classification in a cascade impactor, and electrical detection with electrometers. The particle collection on each collector plate stage is dependent on the aerodynamic size of the particles. Measured current signals are converted to particle number and size distributions. The size classified samples over the operational range of the impactor instrument is given in Table I.

A Testo 435, Pitot tube and Fluke Thermologger with K-type thermocouple were used to determine the duct gas velocity and temperature in the duct. An Actaris gas meter type G4 (0.04-6 m3/h) was used to monitor the gasflowwhen dust samples were collected onfilters.The dust samples were dried at 160 °C until constant weight was achieved. A LVSEM type Hitachi S-3400N and a FESEM type Zeiss Ultra 55 Limited Edition were used to perform optical and X-ray energy dispersive spectrometer (EDS) analysis of the samples. Nonconducting samples were not coated. To prevent accumulation of electrostatic charge the Hitachi S-3400N was operated in low vacuum mode at 70 Pa using the BSE (backscatter electron) detector. To achieve high resolution micrographs, the FESEM type Zeiss Ultra 55 was operated in the low voltage SE mode at 2 kV to reach a dynamic charge balance, as described by Yu et al. [14].

Table I. Size classification and mass characteristics of the Dekati ELPI™ impactor 101pm (without dilution). Inlet pressure 1013.3 mbar, temperature 21.6 °C. Stage D50 Di Number (l/cm3) Mass (μ^πι 3 ) (Mm)

13

$M

12 6.64 11 3.97 10 2.38 9 1.59 0.942 8 7 0.609 6 0.378 5 0.259 4 0.153 3 0.0908 2 0.0541 1 0.0279 Filter 0.07

(μπι)

mm

6.25 3.07 1.95 1.22 0.76 0.48 0.31 0.19 0.11 0.70 0.04 0.02 0.014

0.10 0.10 0.17 0.33 0.60 0.11 2.0 3.7 6.0 12 23 50 250

max

mm

max

2.0E+04 31 2.0E+04 10 4.5E+04 3 8.8E+04 1.4 1.6E+05 0.7 2.9E+05 0.03 9.8E+00 0.1 9.8E+00 0.07 1.7E+06 0.03 0.02 3.1E+06 6.1E+06 0.006 1.4E+07 0.002 6.9E+07 0.0004

8300 2700 810 370 180 80 40 18 8 4 1.7 0.5 0.11

A Bruker AXS D8-Focus X-ray powder diffraction (XRD) and EVA 15.0.0.0 analysis software were used to identify different crystalline phases in the filter dust samples. The laser diffraction analysis of filter samples were performed with a Mastersizer 2000 APA200 (Malvern Instruments Ltd) with a size range of 0.02 to 2000 μιη. Powder from filter samples were dispersed in water under ultrasonic agitation for 90 seconds before passing the laser beam in the closed Hydro MU dispersion loop. Particle distribution was converted from volume to number distribution.

The ex perimental setup for the impactor is shown in Figure 1.

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Results and Discussion Impactor Various impactor substrates are used to accommodate different objectives. Samples deposited on aluminum foils are used for image analysis, while polycarbonate substrates are used for chemical analysis. On aluminium, silver and polycarbonate substrates, we typically observe a shift in the impactor recordings during the initial period as shown in Figure 2.

Impactor

For the substrates capturing the coarser particles, this effect may be due to particles bouncing off the substrates or smaller particles released from agglomerates hitting the plates. As the opposite effect is observed for the substrates capturing the finer particles the breakup of agglomerates seems to be the most likely cause. The effect on calculated mass concentrations due to the initial variations is illustrated in Figure 3. Using sintered substrates with an oil film reduces this effect significantly. However, these substrates are not suited for chemical analysis as the oil film interferes with subsequent analysis.

Figure 1: Sketch of experimental setup for sampling of particulate emissions.

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Table II. Calculated weight on the impactor plates after the dry scrubber and ambient air for the given average particle sizes. Air Di Duct Ratio μιη μg/m3 (Duct/Air) μg/m3 0.044 0.41 0.0279 0.108 0.0541 0.384 0.175 0.46 1.10 0.44 0.0908 0.48 2.62 0.153 1.22 0.46 4.32 4.45 0.97 0.259 17.4 4.28 4.07 0.378 2.74 0.609 4.66 12.7 0.942 1.39 5.11 3.67 1.59 4.76 1.48 3.21 2.41 2.38 7.68 3.18 6.64 3.30 4.07 0.81 9.85 66.1 36.6 1.81 Concentration 37.2 19.8 1.88 1441 2211 Number of data points Density g/cm3 1 1

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Figure 3: Effect of initial and final period versus total calculation of particles hitting the impactor plates. Compared to electron microscopy data published by Hφflich [9], a significant higher number of smaller particles than 0.5 μιη are observed with the present impactor method, making the method well suited for extracting on-line information on the state of the particulates in the off-gas from the cell and in the potroom air. Scrubber To obtain recordings of dry scrubber performance, gas samples for one filter line were drawn from the stack between the dry and wet scrubbers. For comparison, also ambient air at the same location is measured. The results, calculated by Dekati ELPI software assuming particle density equal 1 g/cm3, are shown in Figure 4.

The measured period after the scrubber is extracted to Figure 5. Cyclic changes in the particle recordings are in phase with the filter-bag regeneration pulses. The leakage through the filter varies between 0.01 to 0.07 mg/m3 during each cycle. The coarser particles have sharper peaks than the finer particles as shown in Figure 6. This is expected as the new filter cake build-up after cleaning initially captures coarser particles more effectively than finer particles. The measurements show that it is possible to determine the effect of the cleaning cycle and the state of the filter bags. Although not investigated in the present work, it is believed that the method is well suited for improving bag performance and

347

determining optimal conditions for efficient cleaning with lowest possible dust leakage through the filter bags.

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Figure 5. Particles in the gas duct after the dry scrubber.

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Figure 6. Particle distribution on the various impactor plates over some of the bag cleaning periods; particle sizes corresponding to the impactor plates can be read from Table II.

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Duct Results from sampling of the finer fractions of the raw gas in the duct at two plants are presented in Figure 7. Subject to normal operations, the distribution of finer particles is comparable between the plants.

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348

agitation, indicating that current impactor measurements are better suited for the determination of the fines in the raw gas from the electrolysis cells. Particulates Scanning electron microscopy images, Figure 10, 11 and 12, show that the dust particles adopts numerous shapes, all from discrete spherical droplets, platelets and hexagonally shaped primary particles to agglomerates and particles embedded in a composite matrix of oblong or needle shaped fibrous framework with alumina, cryolite and sodium aluminum fluoride. Typically, many of the fibers seem to be formed at an early stage acting as condensation sites for the vapour phase forming knobbly or grapelike droplets and smaller particles.

Figure 12. Grapelike condensate structure on fibers. Based on a qualitative chemical analysis the most dominant particle group contains sodium, aluminum, oxygen and fluorine. For the particle size ranges with Di equal 60.9 nm, 942 nm, 1.59 μηι and 2.38 μιη, respectively, the atomic ratio of the elements are close to Na:Al:0:F = 1:1.1:2.2:5.6. In the pictures taken with help of the BSE detector, Figure 12, heavier elements appear brighter and stand out from the lighter surroundings. The BSE analysis shows a large variety of morphologies of coarse particles and agglomerates. Ni, Fe and S containing particles are identified due to their very bright appearance. Often the brighter particles are located with the carbon particles. Possibly, the heavier elements are abrasion products.

Figure 10. Sponge-like structure of samples with D50 = 1.590 μιη. On substrates with Di = 2.39 μηι individual particles can quite easily be identified while particulates on the impactor stages with Di < 1.59 μπι the agglomerates seem to grow together to a fabric framework. This is probably due to higher surface forces of the smaller particles, resulting in particles growing together.

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The XRD analysis of filter samples shows the dominant phases to be cryolite Na3AlF6, chiolite (Na5Al3F14) and corundum (A1203). Sulfur may be present in the form of sulfates attached to sodium and/or alumina as millosevichite, A12(S04)3, or matteuccite, NaHS04'H20, which is identified in the XRD pattern. Large carbon rich agglomerates (up to a few μπι) are observed, which is most likely entrained dust from the anodes. Soot particles are abundant in the fractions with particle sizes above 1.59 μπι. For

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Figure 11. Fabric like framework with grapes of smaller particulates attached to oblong particles/fibers of samples with D50 = 0.259 microns.

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the finer particulates with D50 below 1.59 μιη, sampled on silver substrates, the XRD shows that the dominating compound is NaAlF4. The XRD identification of NaAlF4 is in accordance with Gylseth et al. [15] and Heiberg et al. [16] who reported fibrous NaAlF4 in the potroom air. Also Kirik and Zaitseva [17] report the formation of metastable fibrous crystals and colorless NaAlF4 powder by condensing vapors arising from chiolite heated to 800 °C. The condensed NaAlF4 was stable up to 390-400 °C before disproportionating exothermally into solid Na5Al3Fi4 and A1F3 and into solid Na3AlF6 and gaseous NaAlF4 at higher temperatures. The high amount of NaAlF4 on the substrates with the finer particles indicates that most of the finer particulates are rapidly quenched to low temperatures by the draught air in the cell.

References 1. W. Wahnsiedler, R. Danchik, D. Backenstose, W. Haupin, and J. Colpitts, "Factors affecting fluoride evolution from HallHeroult smelting cells," Light Metals 1978, pp. 407-421 2. C.N Cochran, W.C. Sleppy and W.B. Frank, "Fumes in Aluminium Smelting: Chemistry of Evolution and Recovery," Journal ofMetals, Sep 1970, pp. 54-57 3. A. Sorhuus and G. Wedde, "Pot gas heat recovery and emission control," Light Metals 2009, pp. 281-286. 4. S.J. Linday, and N.R. Dondo, "Dry Scrubbing for Modern Pre-Bake Cells," Light Metals 2009, pp 275-280 5. A. Solheim, B. Moxnes, K. Vamraak and E. Haugland, "Energy recovery and amperage increase in aluminium cells by active cooling in the anode yokes," Light Metals 2009, pp. 10911096 6. M. Fleer, O.-A. Lorentsen, W. Harvey, H. Palsson and G. Saevarsdottir, "Heat recovery from exhaust gas of aluminium reduction cells," Light Metals 2010, pp. 243-248 7. L. Less and J. Waddington, "The characterization of aluminium reduction cell fume", Light Metals 1971, pp. 499-508 8. M. Hyland, B.Welch and J. Metson, "Changing knowledge and practices towards minimising fluoride and sulphur emissions from aluminium reduction cells," Light Metals 2000, pp. 333-338 9. W. Haupin, "Mathematical model of fluoride evolution from Hall-Heroult cells," Light Metals 1984, pp. 1429-1439 10. W. Haupin and H. Kvande, "Mathematical model of fluorine evolution from Hall-Heroult cells", Light Metals 1993, pp. 257263 11. B. V. L'vov, L. K. Polzik, S. Weinbruch, D. G. Ellingsen, and Y. Thomassen, "Theoretical aspects of fluoride air contaminant formation in aluminium smelter potrooms", J. Environ. Monit, 7(2005)425-430 12. B.L.W. Hφflich, S. Weinbruch, R. Theissmann, H. Gorzawski, M. Ebert, H.M. Ortner, A. Skogstad, D.G.Ellingsen, P.A. Drablos and Y. Thomassen, "Characterization of individual aerosol particles in workroom air of aluminium smelter potrooms," J. Emiron. Monit, 7(2005)419-424 13. M. Ebert, S. Weinbruch, A. Rausch, G. Gorzawski, P. Hoffinann, H. Wex and G. Helas, "Complex refractive index of aerosols during lace 98 as derived from the analysis of individual particles," Journal of geophysical research: Atmospheres, vol. 107, NO. D21, 8121, doi:10.1029/2000JD000195, pp. LAC 3, 115, September 2002 14. Y.D. Yu, M.P. Raanes and J. Hjelen, "Characterization of nonconductive polymer materials using FESEM," in Proceedings of the 17th International Microscopy Congress (IMC17), Rio de Janeiro, Brazil, 19-24 September 2010 15. B. Gylseth, O. Bjorseth, 0. Dugstad and J. Gjonnes, "Occurrence of fibrous sodium aluminumtetrafluoride particles in potrooms of the primary aluminum industry," Scand. J. Work Environ. Health, 10(6)( 1984) 189-195 16. A.B. Heiberg, G. Wedde, O.K. Bockman and S.O. Strommen, "Pot gas fume as a source of HF emissions from aluminium smelters - Laboratory and field investigations," Light Metals 1999, p 255-261 17. S.D. Kirik and J.N. Zaitseva, "NaAlF4: Preparation, crystal structure and thermal stability", Journal of Solid State Chemistry, 183(2010)431-436 18. N.R. Dando and S.J. Lindsay, "Hard Gray Scale," Light Metals 2008, pp. 227- 232

Dando and Lindsay [18] found that hard gray scale (HGS) in dry scrubbing consisted of an amorphous reaction product formed from the attrition-induced reaction of bath superfines, alumina fines and water. HGS could be artificially created by co-grinding these three components. It is postulated that the energy release by "new alumina surface" re-hydration is the principal energy driver. If any of the three ingredients was left out, scale was not formed. The present work indicates that the paniculate fines in the raw gas may possess fouling properties. Chemical composition and origin of different particle size classes will be studied further and compared to deposition mechanisms in heat exchangers to identify particle groups having an effect on both the growth and removal of depositions. Conclusions and Further Work On-line sampling and monitoring of changes in composition of particulate fines in the off-gas from aluminium cells are established and will be used in future work to compare the effects of various operational parameters. Larger numbers of particles in the sub-micron range are measured with the impactor equipment than previously reported. The SEM pictures of the size classified particulates reveal a change in agglomerate morphology for the different particle size classes. Further work will also focus more on formation mechanisms of the observed compounds and agglomeration products. Also, samples will be collected on silver substrates for better EDS and TEM analysis of the different particle size classes. Acknowledgment This study is a part of the ROMA research project with financial support from the Research Council of Norway and Norwegian aluminum and metallurgical industry. An essential part of this research is conducted as measurement campaign at industrial cells. We appreciate the support by operators at Alcoa Mosjoen and Hydro Sunndal.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INVESTIGATION OF SOLUTIONS TO REDUCE FLUORIDE EMISSIONS FROM ANODE BUTTS AND CRUST COVER MATERIAL Guillaume Girault1, Maxime Faure2, Jean-Marc Bertolo1, Stephanie Massambi1, Georges Bertran1 ^ i o Tinto Alcan L.R.F ; BP114, Saint-Jean-de-Maurienne Cedex, 73303, France 2 Rio Tinto Alcan Smelter Technology ; Centr'Alp, BP7, Voreppe Cedex, 38341, France Keywords: Spent Anode, Butt, Anode Crust Cover, Fluoride Emissions. Abstract

Emissions from spent anodes and crust cover material

For many aluminium smelters, reducing fluoride emissions is often a condition to increase production. Since they contribute up to 40% of the overall roof vent emissions, anode change operations are often targeted for improvements and specifically the emissions from the anode butts and crust cover. The Rio Tinto Alcan Research Centre, LRF, has been conducting an R&D programme over the past few years aimed at improving the understanding of the physical phenomena involved and ultimately minimising this specific contribution. On this basis, different conceptual solutions have been developed and their relative performance evaluated. These tests, associated with past experience of enclosed butt boxes and crust bins, concluded that any container receiving a hot butt and bath crust would require almost complete sealing to be effective. The next stage in this programme is engineering design with the aim of developing smelter technology with the lowest environmental footprint.

Hydrogen fluoride (HF) is produced by the chemical reaction between fluoride species and compounds containing hydrogen. In the case of spent anode and crust cover material, ambient moisture is the major source of hydrogen. Fluoride-containing species can be found in the bath crust and in the anode cover mixture on the spent anode as well as impregnated in the surface of the butt itself. Fluoride emissions from these sources are due to two mechanisms.

Introduction

This reaction is endothermic, facilitated by an external source of energy, such as the hot anode. The equilibrium constant linking partial pressures (P in atmospheres) and activities (a) of the different molecules in this reaction is as follows:

The first mechanism involves a direct reaction of ambient moisture with bath material. Humidity reacts preferentially with aluminium fluoride (A1F3) rather than any other constituent of the bath, because the equilibrium constant is much higher than for the other bath species [4]. The associated chemical reaction is: 1/3 A1F3+ 1/2 H 2 0 -> HF +1/6 A1203

As production is increased, many aluminium smelters are faced with the concurrent challenge of reducing, or at least maintaining, specific fluoride emissions in order to meet their regulatory obligations. Tighter environmental regulations will make this even more difficult in the future, and will require operations to consider investment in state-of-the-art pollution control equipment to help achieve this objective.

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Since anode change operations represent the single largest contributor to overall roof vent emissions, they are often targeted for improvements [1, 2, 3]. A significant contribution to the emission from the anode changing operation occurs when cell hoods are removed, especially during the period when bath is exposed [3]. This contribution can be minimized by reducing the total time that hoods are opened and/or significantly increasing cell exhaust flow during this operation, through the use of a dedicated Boosted Suction System [3].

The second mechanism involves vaporization of the bath material prior to reaction with moisture. At temperatures above 700 / 800°C, some bath material vaporizes, forming mainly NaAlF4 [4], Part of this vapour is hydrolysed by ambient moisture to form HF while another part condenses and forms fluorinated particulates. Because of the relatively high temperature required for such vaporization to occur, this reaction only takes place during the first few minutes following the material removal from the cell. The contribution from this second mechanism is therefore expected to be much lower compared to thefirstone.

Two other significant contributions to emissions from the anode change operation are related to the spent anodes and crust cover material after their removal from the cell. These contributions cannot be reduced by better operational practices and therefore require additional equipment to capture the emissions. Some technologies have already been trialed on an industrial scale within Rio Tinto Alcan plants but most have considerable drawbacks which have prevented their widespread deployment. In this context, Rio Tinto Alcan has been conducting an R&D programme over the past few years aimed at improving its understanding of the physical phenomena involved. Different conceptual solutions have subsequently been developed and tested. This paper presents the main conclusions to date from this work.

From the above, one would expect fluoride emissions to be influenced by temperature, and the rate at which it decreases, since all the previously described reactions are favoured, both thermodynamically and kinetically, at high temperature. Other parameters include the quantity of bath material exposed to water vapour in the ambient air as well as, to a lesser extent, the bath material composition (higher A1F3 excess results in higher activity for example).

351

its installation in the enclosure was typically 5 to 8 minutes, which is much less than what occurs during published tests carried out in industrial potlines [8]. Additional information, such as the anode location in the cell, age and spent anode mass, as well as ambient meteorological conditions, were also recorded to facilitate data interpretation. The same procedure was applied to crust bins.

Experimental set-up The characteristics of spent anode emissions from older cell technologies and the main drivers influencing these have been previously investigated [5, 6]. It was felt however, that it would be beneficial to update these studies with newer, more accurate, measurement techniques and check whether the conclusions drawn on the basis of older cell technologies are still valid for the latest generation, which are designed in particular with much bigger anode assemblies.

Spent Anodes - Baseline measurements The following graph shows the measured emission rate (in mg/s of HF) of 16 similar anode assemblies as a function of the time after the anode assembly was positioned inside the measurement enclosure.

Experimental equipment was installed at the Rio Tinto Alcan Research Centre, LRF, in Saint-Jean-de-Maurienne for this investigation. It consisted of a measurement enclosure in which a single anode assembly or a single crust bin could be placed. The enclosure stack was equipped with a continuous HF analyzer (Neo Monitor), a CO and C0 2 continuous analyzer (Testo). The gas flow rate and temperature were also continuously monitored and there was a facility to conduct manual dust sampling.

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Figure 2 - HF emission rate curves for anode assemblies The first set of measurements (A to I) showed that the gaseous fluoride emission rate had generally decreased to background level after 10 hours, though a few spent anodes still evolved measurable emissions after this period. For the second measurement campaign (1 to 7), it was therefore decided to leave the anode inside the enclosure for at least 10 hours, and to integrate the emission rate over this period to calculate total specific emission in kgFg/tAl (kilograms of gaseous fluoride emissions per tonne of liquid aluminium produced). The average cumulative emission for the first 10 hours was 0.14kgFg/tAl ± 0.06 (95% confidence interval). Extrapolating these emission rates to account for the initial 5 to 8 minutes where the emissions were not measured adds an extra 0.02kgFg/tAl to the previous contribution.

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Figure 3 shows the cumulative emission as a function of time. The results have been normalised on the basis of the cumulative emission after 10 hours and include the contribution from the first 5 to 8 minutes before the spent anodes were placed in the enclosure.

Reference measurements were first carried out to assess the characteristics of the emissions from the spent anodes and crust bins under normal conditions where there was no attempt to minimize or capture the emissions.

This graph shows that, based on the cumulative emission after 10 hours, approximately 30% is evolved during the first 30 minutes, less than 40% during the first hour and 80% during the first 5 hours.

Two measurement campaigns were carried out in 2008 (tests numbered A to I) and in 2010 (tests numbered 1 to 7), which consisted in inserting spent anodes from the LRF AP50 prototype cells into the measurement enclosure immediately after their removal from the pots. The time between the anode removal and

352

linear relationship. Consequently, emission rates can be described by the following equation: Emission rate (mg /s) =

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Each experimental curve, representing the emission from a single spent anode, could therefore be characterized by these two parameters. The following graph provides examples of actual and corresponding modelled curves to illustrate the good fit between the two.

Manual dust sampling was also carried out. This sampling showed that the cumulative particulate fluoride emission was less than O.OlkgFp/tAl during this period and thus makes only a minor contribution to the total fluoride emission. Total dust emissions represent approximately 0.02kgPt/tAl. Chemical analysis of the dust showed that approximately 50% of the bath vapours evolved is hydrolysed to HF, the rest condensing as fluorinated particulates. Spent anodes - Discussion The overall average cumulative emission is consistent with previously published information, though it is on the high side [1, 2, 3, 5, 6]. When compared to existing information, the curve of cumulative emission versus time also has a slightly different shape, as represented below [5,6].

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On this basis, it is then possible to investigate what influences the two parameters A and B. Interestingly, the parameters were found to be uncorrelated. Both parameters exhibited high variability, with the standard deviation representing approximately 40% (for 1/A) and 30% (for \IB) of their corresponding averages.

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The following graph represents HF emissions, CO and C 0 2 emissions as well as the heat exhausted at the stack from a typical example (trial number 6). Three phases can be identified: A first pt|f§e, lasting for approximately 30 minutes, during which HF emissions decrease very rapidly and the exhaust heat increases (probably as the measurement enclosure heats up).

Compared with the published information, in our experiments a greater part of the total emission occurred at a later stage. We believe that this is due to the much bigger anode assemblies that have been used for these tests. The higher thermal inertia of the larger spent anodes results in a slower cooling rate, contributing to maintaining emission rates for a longer duration. The data analysis then focused on investigating the drivers for the large variability. Plotting l/(emission rate) versus time results in a

353

A second phase of approximately 1 to 2 hours, when the exhaust heat and CO emissions remain constant. During this period, HF emissions reduce at a slower rate compared to the first phase. A third and final phase, characterised by a slow decrease in the HF and CO emissions as well as the heat dissipation. 80

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A high variability in emissions from spent anodes has been observed. Both the initial emission rate as well as the rate at which the emissions decrease are very variable. A statistical analysis has concluded that the different carbon and cover masses do not explain the variability of these two parameters. This is thought to be due to the relatively small range of masses explored (similar anode sizes; the standard deviation of the carbon and cover masses represented 11% and 14% of their respective averages). Comparing the emission rate curves from smaller anode assemblies does indeed suggest that bigger sizes result in an extended period of emissions. Ambient humidity does not seem to influence emissions, probably because water vapour is already in large excess even in the driest conditions (range explored: from 4.1 to 6.8 g/kg averaged over the 10 hours period). This high variability implies that it is necessary to conduct a high number of experiments when measuring these emissions.

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As expected, the emission rate decay (A) appears to be related to the kinetics of the spent anode cooling. A longer cooling time results in a slower decrease in the emission rate. It is thought that the carbon oxidation participates in maintaining the spent anode surface temperature at a high level, slowing down the emission decreasing rate. Crust cover material - Baseline measurements While the R&D programme primarily focused on the contribution of the spent anodes, a few tests were carried out to estimate the emissions from the crust cover material, using the same procedure. Bins containing crust cover material (typically around 700kg), corresponding in size to the skimming from a single anode cavity, were inserted inside the measurement enclosure within minutes of the material being extracted from the cell.

The factor A, representing the rate at which emissions reduce with time, has a stronger impact on the overall variability than B (75% of the variability is explained by A). This factor is related to the thermal behaviour of the spent anode and how quickly it cools down, which is thought to be itself, to a degree, influenced by the anode combustion. For a given anode size, qualitative parameters, such as the spent anode shape and cover integrity, are suspected to be the most important factors.

Two typical examples of the emission curves obtained are shown in Figure 7, as well as their corresponding modelled curve built on the same principle as the one used for the spent anodes. Similar to the spent anodes, there is a good fit between the experimental and modelled curves. The emissions from the crust bins exhibit lower initial rates and decrease more rapidly than for the spent anodes. This results in lower emissions for the crust bins. Total emissions, integrated over a period of 10 hours, are estimated to be slightly less than O.lOkgFg/tAl.

Solutions for reducingfluorideemissions from spent anodes and crust cover material Once this preliminary analysis of reference spent anodes and crust cover material had been completed, the experimental set-up was used to test different conceptual solutions. Enclosed spent anode boxes

First lessons from the baseline measurements

This concept aims at minimizing the spent anode exposure to both ambient moisture and oxygen from the air. It was assumed that the efficiency of this type of equipment depends on its sealing level. Such a concept is not new and has already been tested and implemented by Rio Tinto Alcan and others [7, 8].

Emissions from spent anodes and crust cover material have been estimated to be around 0.16kgFg/tAl and 0.10kgFg/tAl respectively. For both, the contribution of particulate fluoride is minimal, estimated to be less than O.OlkgFg/tAiisk is to be noted that, under normal operating conditions, a portion of these emissions escapes from the potline roof vent whilst they are cooling on-line, with the rest being emitted from the anode cooling building.

In order to prove the concept, a prototype box was built, with the objective of being close to 100% tight. Gaps between the cover and the tray, as well as around the anode stem were manually sealed using rock wool. This prototype is shown in Figure 8.

354

These tests concluded that a container receiving a hot spent anode would require almost complete sealing to be effective. This was deemed to be impractical in industrial conditions, as it would be difficult to maintain good sealing in the long term.

The same protocol as the one used for the baseline measurements was applied; spent anodes were placed in the prototype box immediately after their removal from the cell, and the box was itself placed in the measurement enclosure shortly after.

Covering the anode with a granular material An alternative concept to the enclosed anode box was therefore considered, which consisted in burying the spent anode in a granular material available in the smelter, such as crushed bath, cover mix or alumina. Proof-of-principle trials confirmed the potential of this solution. Spent anodes were placed in a tray and completely covered with approximately one tonne of crushed bath material before being inserted into the measurement enclosure as in Figure 9. The tests were repeated five times and the emissions were very low and repeatable, averaging 0.05kgFg/tAl (±0.01), excluding the contribution from the first minutes, a performance similar to that obtained with the first prototype "fully" sealed anode box.

Figure 8 - Sealed anode box used for the proof of concept test

Figure 9 - Spent anode covered with cover mix Following these initial successes, tests using cover mix and alumina were conducted. While they confirmed that using cover mix was also very efficient in reducing emissions (nine tests averaging 0.06kgFg/tAl), they indicated that alumina cannot be used. Covering the hot butt with alumina resulted in geysers being formed because of this material's fluidizable nature. This in turn generated high levels of dust emissions and poor efficiency in reducing HF emissions.

Six trials were conducted with this prototype, and the results were very repeatable. Emissions were very low, averaging 0.04kgFg/tAl (±0.005), excluding the contribution from the first minutes. Although the tightness was considered as being close to 100% tight, it was not sufficient to fully counteract the internal pressure building related to the gas generation and heat. Following these encouraging preliminary trials, it was decided to test an industrial box based on the same principle.

Additional tests were organized to estimate the sensitivity of the cover quality on emissions. The quality of the cover was visually evaluated as being either "good" (the spent anode was completely recovered) or "medium" (there were still a few visible parts). The following graph represents typical emission curves for one reference / uncovered spent anodes, spent anodes with "good" and "medium" cover.

A new concept was designed, based on Rio Tinto Alcan's past experience in the Alma smelter where several concepts had been tested in the past. Particular attention was paid to the sealing system around the anode stem. Another measurement campaign of six tests was then organised, based on this prototype. The corresponding average emissions and variability were very close to the spent anodes baseline measurements. Different modifications of the box were made, providing only marginal improvement to its efficiency. It could be seen that the buoyancy driving force of the hot spent anode was so strong that emissions would escape from any small gaps in the box.

355

Conclusions The results of these tests have improved our understanding of the behaviour of spent anodes and crust cover material emissions. Spent anode and crust bin emissions have been estimated to be around 0.16kgFg/tAl and O.lOkgFg/tAl respectively. It is thought that this contribution could theoretically be reduced to less than 0.05kgFt/tAl through the implementation of an efficient fume capture. On this basis, different conceptual solutions aimed at reducing this contribution have been tested.

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The sealed box concept provided good results. However, a container receiving a hot spent anode requires almost complete sealing to be effective. This was deemed difficult to achieve and maintain in industrial conditions, and it was therefore decided to investigate alternative solutions. On the other hand, a simple cover on the crust bin is sufficient to be effective.

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Testing the covering of the spent anode with a granular material has been successful, but similarly to the sealed anode box, it requires very good coverage to reduce emissions significantly.

The conclusion is similar to that drawn from the anode box test. For the concept to be efficient the spent anode needs to be completely covered. Spent anodes that are incompletely covered exhibit similar emissions to the baseline values.

The next stage in this programme is engineering design, with the objective of developing a viable industrial solution based on these concepts.

In practice, the spent anode could be recovered with granular material by using, for example, autonomous pallets, equipped with reservoirs containing the appropriate material, or alternatively using the pot tending crane.

Acknowledgments The present work is fed by and takes into account results from previous Rio Tinto Alcan tests and evaluations, not published externally, and carried out by the Arvida Research and Development Center as well as by the Pacific Technology Center.

Enclosed crust bin Installing a simple cover on the crust bin significantly reduces its emissions. A few tests conducted with such covers in place concluded that emissions below 0.04kgFg/tAl can be achieved (compared with a reference value of 0.10kgFg/tAl, refer to Figure 11). Whereas it was necessary to have complete seal on the spent anode box to achieve good performance, the crust bin cover efficiency is less sensitive to sealing.

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References [1] Nilton F. Nagern et al; Understanding fugitive fluoride emissions at the Alumar; Light Metals 2005; 289-292. [2] Elaine Y.-L. Sum et al.; Understanding and controlling HF fugitive emissions through HF monitoring and air velocity characterisation in reduction lines; Light Metals 2000; 357-363.

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[3] Michel Meyer et al; Development of a jet induced boosted suction system to reduce fluoride emissions; Light Metals 2009; 287-292.

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[4] Warren E. Haupin et al.; Mathematical model of fluoride evolution from Hall Heroult cells; Light Metals 1984; 1429-1439.

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[5] T.A. Lowe et al.; Fluoride evolution from spent anodes; 73rd Annual Air Pollution Control Association Annual Meeting; Montreal, Canada; June 22-27, 1980.

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- ftp* f"!;'-f

[6] Warren E. Haupin et al.; Fluoride evolution from spent anodes; Light Metals 1981; 941-951.

8:00

Time (hours) REF1

REF2

Covered crust bin 1

Covered crust bin 2 |

Figure 11 - HF emission rates of covered crust bins

[7] Jean-Pierre Gagne et al; New Design of Cover for Anode Trays; Light Metals 2006; 213-217. [8] Jean-Pierre Gagne et al; Update on the evaluation of the HF emission reduction using covered anode trays; Light Metals 2010; 291-294.

356

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PFC Survey in Some Smelters of China Wangxing Li, Qingyun Zhao, Shilin Qiu, Shuchao Zhang, Xiping Chen Zhengzhou Research Institute of CHALCO, Zhengzhou, China, 450041 China National Engineering & Technology Research Center for Aluminum, Zhengzhou, China, 450041 Keywords: PFC emission, reduction cells, anode effects, smelters Abstract

Measurement Strategy

PFC survey was performed by Zhengzhou Research Institute of CHALCO (ZRI). Results of the first PFC survey from several different Chinese smelters are discussed in this paper. This survey included normally-operated, newly-started, power-limitation and new control-mode potlines. The measurement equipment used in the survey is MG2030 FTIR which comes from America MKS Instruments. Gas sampling was directly pumped out of main exhaust duct before stack and transmitted into FTIR by Teflon tube. Fugitive PFC emission was not measured. The calculation of PFC amount was based on accepted IPCC Tier 3 method. Total PFC emission was calculated from duct PFC emission divided by collection efficiency. Newly-started, new control-mode and power-limitation potlines emitted much more PFC than normallyoperated ones. PFC emission from power-limitation potline was the highest among seven potlines. Three of the seven potlines have PFC emission lower than original IAI 2010 target. Two potlines have PFC emission close to Australia 2006 average level. The lowest amount of PFC emission is 0.21 tC02-e/t-Al.

MG2030 FTIR The measurement equipment is MG2030 which comes from America MKS Instruments. Preparation and training of the FTIR were finished in early 2008. The FTIR and its pump were assembled separately. Sampling pump is assembled together with gas filters and gas temperature controllers (see Figure 1). They are placed into wooden boxes and suitcases for shipment.

Introduction Reduction of PFC emissions has been a priority for the global aluminum industry since the industry became aware of the climate change impact of the two PFC compounds, tetrafluoromethane (CF4) and hexafluoroethane (C2F6) emitted during anode effects. Primary aluminum production in China has increased more than four times over the past 10 years. The fast-expanding of Chinese aluminum production is largely due to investment in new capacity, employing PFPB technology with the lowest PFC emission. Now China is the single largest primary aluminum producer and PFPB technology has been spread to all of its smelters along with extinguish of Soderberg technology cells. The International Aluminum Institute (LAI) conducts annual survey of global primary aluminum producers on anode effect (AE) performance from which PFC emissions are calculated according to Intergovernmental Panel on Climate Change (IPCC) Good Practices methodology (1). The participation of IAI AE survey has for a number of years remained around the 90% mark except China. Since China continues to make a larger share of global Al production, the total survey participation rate falls annually (2). IAI is making efforts to seek cooperation with the Chinese aluminum industry in order to improve the accuracy of the emissions projection. At the same time, as part of the APP (Asia Pacific Partnership for Clean Development and Climate), Chinese government began to pay more attention on PFC emission from aluminum industry. Thus, Zhengzhou Research Institute of Chalco (ZRI)/National Engineering and Technology Research Center for Aluminum decided to investigate PFC emission from Chinese smelters. A FTIR was bought from MKS Instruments in 2008, at the same time a research team was set up in ZRI.

Figure 1-MG2030 FTIR Sampling First survey included seven potlines which belong to several different smelters of China. PFC measurements were finished by ZRI team from June to September in 2008. During measurement, gas samples were directly pumped out of main exhaust duct and transmitted into the FTIR through Teflon tube while fugitive gas emission was not measured. Sampling location was usually chosen on the horizontal duct between dry scrubber and chimney stack (see Figure 2).

Figure 2 - Sampling Location

357

PFC survey was performed according to the Protocol for PFC Measurement published by IAI/USEPA in 2008. PFC emissions were calculated according to accepted IPCC Tier 3.

Data Collection The MG2030 was monitored with a laptop. Sampling gas was heated to 150°C and passed continuously through FTIR gas cell. PFC spectra are collected, transformed into digital signals and plotted vs time by special software in the monitoring laptop. PFC concentration was recorded by the laptop at intervals of 1-60 seconds (see Figure 3). There would be a PFC peak when an anode effect occurred. Some gases such as HF, S0 2 and H 2 0 may bring some negative effect on PFC measurement. But in this survey these gases were not removed from gas sample in order to learn their concentration fluctuation.

Code A

Table I - Status of Potlines in the Survey Total cells Annual aluminum 1 Line including production, kilotons current, kA 182 261 124.830

B

335

30

25.345

C

239

180.325

D

202

286 204

107.610

E

295

276

209.923

F

215

263

150.015

G

180

256

121.545

Survey Results According to IPCC Tier 3 methodology, total PFC emission from seven potlines was calculated and compared with overseas smelters (see Table II). Table II - PFC Emissions from Measured Potlines Overseas and indoors PFC emission PFC emissions, 1 tC02-eq/t-Al 0.32 Measured PFC emission from Potline A Measured PFC emission from Potline B 0.66 Measured PFC emission from Potline C 1.33 (newly-started) | Measured PFC emission from Potline D 2.38 (newly-started) | Measured PFC emission from Potline E 2.45 (*power- 1 limitation) 1 Measured PFC emission from Potline F 1.24(2new control- 1 mode) | Measured PFC emission from Potline G 0.21 | Australia average PFC emission in 2006 0.30 Original IAI PFC target in 2010 0.99

Figure 3 - PFC Time Lines Calculation Methodology The survey requests AE performance data from the measured potline, usually reported by smelters directly to ZRI research team. The key data, required to estimate potline or facility performance including anode effect performance and related data for calculating total PFC emissions, are as follows: - Primary aluminum production (MP) during measurement, in tonnes of molten metal tapped from cells; - Anode effect frequency (AEF), the average number of anode effects occurring per cell day; - Anode effect duration (AED), the average time in minutes of each anode effect; - Collection efficiency, percentage of exhaust gas collected in main duct. The anode effect performance data above allows for the calculation, by IPCC Tier 3 methodologies, to obtain facilities' total CF4 and C2F6 emissions, and then tonnes of C0 2 equivalent emitted per tonne of aluminum produced through applying each PFC's GWP (Global Warming Potential).

Power limitation: The potline was running at lower line current 295kA (normal line current is 3l5kA) because of insufficient power. New control-mode: The potline was changing its control software which has different feeding intervals from original mode.

It can be learned from Table II that potline A and potline G have good AE performance and their PFC emissions are as low as Australia average value in 2006. But PFC emissions from powerlimitation, new control-mode and newly-started potlines are much higher than normally-operated ones (3). Potline F in powerlimitation emitted the highest PFC. Three of the seven potlines have PFC emission lower than original IAI PFC reduction target in 2010 (2).

PFC Survey

PFC Concentration Curve

In this survey, in order to get information as much as possible, smelters in different region were chosen. They locate in Northwestern China, South-western China and Central China, respectively. Seven potlines were selected including different situations such as normally-operated, newly-started, new controlmode and power-limitaion potlines (see Table I). Newly-started potlines were those which suffered heavy ice-rain in the beginning of 2008.

During measurement, when anode effects occurred, there would be peaks on the PFC concentration curve. Peaks for CF4 and C2F6 always arose at the same moment (see Figure 4 and 5).

358

20 15

lA

£

10

II

b

-d 0

5

10

L , 0

15 20 25 30 35 40 AS 50 55 60 65 70 75 80 85 90 95

5

10

15 20 25

d u r a t i o n /k

40 45 50

65 70

55

75 80 85 90 95

Figure 5 - C2F6 Concentration Curve (also fourteen peaks)

Figure 4 - CF4 Concentration Curve (fourteen peaks)

Measured potlines

30 35

d u r a t i o n 7k

Table III - PFC Emissions and Continuous PFC in the Survey A B D E C

F

G

Line current, kA

182

335

239

202

295

215

180

Including cells

68

13

36

95

45

53

39

86.29

72.22

72.11

72.07

71.85

73.03

90.14

AE CF4 emission rate, kg/t-Al

339,594 0.0227

101,358 0.070

189,938 0.162

430,198 0.199

291,666 0.0628

260,619 0.135

197,598 0.0249

Non-AE CF4 emission rate, kg/t-Al

0.0191

0.0163

0.011£

0.1121

0.2575

0.0271

0.0026

Total CF4 emission rate, kg/t-Al

0.0418

0.0863

0.1739

0.3111

0.3203

0.1621

0.0275

54.26

81.13

93.17

63.96

19.61

83.28

90.70

45.74 Normally operated

18.87 Normally operated

6.83 Newly started

36.04 Newly started

80.39 Power limitation

16.72 New controlmode

Measurement duration, hr Tapped metal during measurement, kg

The ratio of AE CF4, % The ratio of non-AE CF4, % 1 Potlines running status

9.30 Normally 1 operated |

including cells: Cells in PFC measurement.

Measured potlines Line current, kA Measured flow rate, NnrVh Calculated flow rate, Nm3/h Measured uncertainty, % Remeasured flow rate, Nm3/h

Table I V - Measured and Calculated Flow Rate A B C D 202 182 335 239 275178 81481 161357 507476 302616 177432 89018 492728 9.97% -2.91% 9.25% 9.96%

/

/

/

/

E 295 408742 344830 -15.64% 380704

F 215 273964 252046 -8.00%

/

G 180 186261 176634 -5.17%

/

mode were included in the survey. Satisfactory results were obtained from the first survey. PFC emission was obtained by calculation from anode effect performance and related data. Influences of aluminum production, collection efficiency, and especially duct flow rate were considered when PFC emissions were calculated. In this survey, duct flow rate was measured by Pitot tube. Carbon balance calculation was carried out in order to prove accuracy of measured ductflowrate.

Observation of Continuous PFC Emission In this survey, low concentration of PFC emission not related to anode effects, named as non-AE PFC, was found (see Table III). Non-AE PFC rapidly attracted interests of ZRI and IAI experts. Non-AE PFC has characteristic of continuous emitting and low concentration, also called as continuous PFC emission. What cause non-AE PFC emission? Why potline F in power-limitation running emitted a large portion continuous PFC, but potline G in normally-operated gave off a small percentage continuous PFC emission?

Carbon Balance Calculation During aluminum electrolysis, following anode consumption, carbon changed into gaseous carbon compounds such as C02, CO, COS, CFL^ CF4 and C2F6. Total carbon amount from these compounds should be equal to net carbon consumption per tonne aluminum. This carbon amount can be calculated based on duct flow rate and gas concentration. If the calculated value is close to measured value, the measured duct flow rate is reliable, or the duct

Discussion PFC Survey plan was mapped out by ZRI on the base of experiences from overseas research work on PFC from primary aluminum industry. The first survey was performed under the urgent need for reduction of PFC emission. Different abnormal statuses such as power-limitation, newly-started and new control-

359

flow rate should be remeasured. The formula for carbon balance calculation is as following. C = 12 X [(MCO2/44)+(MCO/28)+(MCOS/56)+(MCHV16)+

MCF4/88]+(Mc2F6/138)

Here, C - Total carbon amount from gaseous carbon compounds, kg; MCo2> Mco> Mcos, MCH4, McF4 and Mc2F6—Amount of C0 2 , CO, COS, CH4, CF4 and C2F6,which can be calculated from measured duct flow rate and gas concentration, respectively, kg. Duct flow rates of seven potlines were measured and calculated (see Table IV). D and G potlines had lower uncertainty, but potline E had a higher uncertainty and its duct flow rate was given a remeasurement. Conclusions PFC emission from Chinese normally-operated smelters is similar to that of Western ones. The lowest PFC emission is 0.21 tC02e/t-Al. The potlines operated in various abnormal situations emitted much more PFC than normally-operated ones. Powerlimitation was the worst circumstance which caused the most PFC emission. Potlines in newly-started and control-mode changing period released much PFC emission as well. There is still a way ahead for PFC survey and investigation in China although some work has been done by ZRL Continuous PFC emission will be focused on in the future surveys. Factors causing continuous PFC emission will be fully investigated. Acknowledgement This work is supported by The National Natural Science Foundation of China (No. 50974127) and The National Key Technology R&D Program (No. 2009BAB45B03). References 1. Protocol for Measurement of Tetrafluoromethane (CF4) and Hexafluoroethane (C2F6) Emissions from Primary Aluminum Production. U.S. EPA (Washington, D.C.) and IAI (London, U.K.), April 2008 2. Results of the 2008 Anode Effects Survey, International Aluminium Institute. Available at http://www.worldaluminium.org, September 2009 3. Results of the First Survey in CHINALCO Smelters, Zhengzhou Research Institute of Chalco, May 2009

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CONSIDERATIONS REGARDING HIGH DRAFT VENTILATION AS AN AIR EMISSION REDUCTION TOOL Stephan Broek1, Dr. Neal R. Dando2, Stephen J. Lindsay3, Alain Moras4 ^atch Ltd; 2800 Speakman Drive; Mississauga, Ontario L5K 2R7, Canada Alcoa Technical Center, 100, Technical Drive, Alcoa Center, PA, 15069-001, USA 3 Alcoa, 300 N. Hall Rd. MS S-29, Alcoa, Tennessee, 37701-2516, USA 4 Alcoa, Aluminerie de Deschambault, 1 Boulevard Des Sources, Deschambault, Quebec, JOA ISO, Canada 2

Keywords: primary aluminium smelter, potroom ventilation, air emissions, high draft, dual duct, gas collection efficiency into immediate contact with humidity from the ambient air. This creates an emission that is estimated to be 0.05 kg TF/tonne. Lastly, the excess hot crust/bath material that is removed from the cell also comes in contact with the air. This is perhaps a less intense emission source, but still provides for approximately 0.03 kg TF/tonne.

Abstract High draft ventilation is an effective technique for reducing emissions from electrolysis cells while panel covers are removed to perform maintenance. In recent years, many new smelters have implemented high draft ventilation as one of the tools to further reduce air emissions from potrooms. In this paper the principles of high draft ventilation are discussed followed by a presentation on its impact on smelter performance. Practical observations are provided concerning the implementation of high draft ventilation in greenfield and brownfield smelters.

After the introduction of GTCs, stack emissions have been reduced to a very low level, assuming that the evolved fluoride load to the GTC is maintained below 85% of the saturation level of the alumina [1]. Therefore, in order to achieve further reductions the next most productive target is the emission from cells during anode changes. Here is where high draft ventilation is most effective.

Introduction Primary aluminium smelter operations continuously seek improvements for higher production efficiencies, for lower costs, to achieve better working conditions and to reduce emissions. This paper is focused on reduction of fluoride emissions from potrooms, more specifically, on the use of high draft ventilation in cells. All values in kg Total Fluoride per tonne aluminum

While there are different approaches to high draft ventilation, in this paper we focus on the so-called 'dual duct' type of high draft ventilation system. This system uses a separate (thus 'dual') but smaller diameter ventilation duct that runs parallel to the main ventilation duct. Depending on the situation, the extra flow from the high draft system is taken to the GTC or (in case there is no spare capacity in the existing GTC) taken to a smaller, separate GTC.

°-°6 \

For each project it is recommended that during the planning phase several, often plant-specific, options are evaluated first. This is what happened in the early phase of the addition of Line 3 of the Alumar smelter in Brazil. Here a total of five options were developed and for each option comparisons were made of the total cost to install and the cost to operate. Alcoa operates three smelters with high draft ventilation systems (Deschambault, Alumar and Fjarφaäl) while a fourth smelter (the Ma'aden JV) is in the design phase. While it is an accepted practice to include high draft ventilation in the design of new smelters, very few existing smelters have actually been equipped with a high draft ventilation system. The smelter in Deschambault is a rare example of where a full dual duct high draft system has been retrofitted during full production.

Figure 1 - General distribution offluorideemissions in a modern prebake smelter High draft ventilation is a tool to reduce emissions into the potroom and it all starts at the electrolysis cell. See Figure 1. When a cell operates normally and all panel covers are closed, background fluoride escapes into the potroom through gaps and other means (0.04 kg TF/tonne). Most fluorides are actually contained by the ventilation system and transported to the gas treatment center. There the HF is scrubbed and the particulate fluoride caught on the filters. The residual fluoride in the stack is approximately 0.06 kg TF/tonne. In the potrooms the temporal fluoride emissions peak during anode changing. Opening the covers to access the bath and anodes causes a loss of draft and the average contribution of these emissions are 0.24 kg TF/tonne. Then, the spent anodes, covered with hot anode crust, can come

Principles of High Draft Ventilation In principle, high draft is the same as normal draft except for the flow rate of air drawn from a cell, which is a factor X higher. Under normal draft a particular vacuum is maintained under the closed hoods. This is sufficient to keep the gas collection near 99 percent during periods of no open cover panel activity.

361

Q ^

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P
A OH

H 9 P

All covers closed

Figure 2 - Indicative pressure profiles during normal draft and all covers closed

0P

T

The issue is that once covers are removed or doors are opened that the vacuum changes such that the tipping point from vacuum to positive pressure moves upwards inside the superstructure so that more of the cell gases can escape through the opening. In Figure 2 the normal situation is sketched while Figure 3 shows the effect of opening one or more covers (on one side only).

Using these equations, engineers can design the ventilation system for normal draft and for high draft ventilation. The impact of this is summarized in Figure 4 where point 1 represents the case when all covers are closed and point 2 the case when 3 covers are removed. From the data in Figure 4 it is clear that under normal draft the gas collection efficiency is reduced once covers are opened. Under high draft the point where the collection efficiency starts to deteriorate is pushed away to the right allowing up to 3 covers, for example, to be removed with minimal impact on cell losses. Please note that sometimes high draft ventilation is also referred to as 'double suction' ventilation. While some applications have shown that the high draft ventilation flow is twice the normal ventilation flow, one has to be careful that the high draft ventilation depends on a number of factors and needs to be evaluated case by case. For the purposes of this paper we shall refer to the use of 'high draft ventilation'.

Two or more covers removed

Figure 3 - Indicative pressure profiles when 2 or more covers are removed After a cover is removed the pressure inside changes, especially around the area that is now open. Two important effects occur: First, the buoyancy of the cell gases is now strong enough that it overcomes the internal pressure and drives the gases outside the cell, and second, the natural draft of air around the cell (and from floor vents) creates a stack effect that further pulls the cell gases from the cell into the potroom [2].

Normal Draft

To prevent the cell gases from escaping from the cell when one or more covers are opened the profiles must be restored so that the cell remains under a light vacuum. Dernedde [3] defined the flow under these conditions as the minimum ventilation flow ÖÌ where the flow is just enough that all cell gases remain captured. From data taken from tests on a prebake cell, Dernedde subsequently developed a mathematical model that presents a relationship between the actual ventilation flow and the minimum required ventilation flow. 100 r,^

xo.50

V

Normal Draft

In this equation the minimum ventilation flow ÖÌ is a function of how the cell is configured. Dernedde evaluated the draft under different cell configurations vs. the buoyancy of the cell gases and derived the following relationship for QM'. /

M

v

°

'

(P-CP-T

Normal Draft

^ High Draft

Next, two important questions must be answered in the design phase:

= gas collection efficiency, % = Applied ventilation flow, Nm3/s = Minimum ventilation flow for E = 1, Nm3/s

E Φ ÖÌ

High Draft

Figure 4 - The gas collection efficiency as function of the applied ventilation flow

50

-((V) -<) With:

= Minimum ventilation flow for E = 1, Nm3/s = Orifice coefficient = Open area, m2 = Heat released from the crust, W = Vertical height of one cover plate, m = gravity constant, 9.81 m/s2 = density of air, kg/m3 = heat capacity of air, J/kg.K = temperature of the air into the hood, K

•

For how many removed covers must the ventilation flow be able to provide for 100 percent gas collection?

•

How many cells (per GTC) will be under high draft ventilation at the same time?

With this, the high draft ventilation of a single cell is configured plus the complete system now starts to takes shape. The answer to question 1 will result in the design flow for high draft ventilation. A margin for error is needed (this is not exact science - the engineer will apply a safety factor) so that the design flow under high draft conditions exceeds the minimum calculatedflowÖÌ-

NO.33

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362

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Question 2 is a very practical question. Operators often need to split their activities and movements between cells and in some occasions they will work on one cell and open the covers on the cell that is next in line. In some cases this can be as many as three cells that have a number of covers removed. In this case the answer is three. In any case, this answer leads to the total flow under high draft conditions that need to be taken to the GTC either through the dedicated high draft duct or through the main duct.

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f

Dual Duct High Draft System Design The design for a high draft ventilation system in a new smelter is different from the design of a system in an existing smelter. Implementation in an existing plant is a considerably more complicated exercise. This is further explained in the following sections.

h | 1

| I

I ' · ]

Figure 7 - High draft boosterfans at the Alcoa Fjarφaδl smelter (Photo courtesy of Solios Environnement)

New aluminum smelters (Greenfield application) In this case we refer to the installation at Alcoa's Fjarφaäl smelter in Iceland. For this smelter the emissions limits are very low (<0.35 kg TF/tonne Al) and the high draft ventilation is used here as a tool to achieve consistent low emissions from potrooms.

In order to create the high ventilation flow a lower negative pressure must exist in the parallel duct. Detailed pressure loss calculations will determine what this pressure is but it is inevitably different from what the main fans deliver. Therefore dedicated booster fans are used (See Figure 5 and Figure 7). Each has double the capacity for high draft ventilation so that the second booster fan can be stopped for maintenance, for instance. This way there is sufficient redundancy while capital costs are optimized. In a new design the extra ventilation flow is incorporated in the total flow that is treated in the GTC. Sufficient filter compartments are provided and the main exhaust fans are sized properly to work optimally in the design points. Existing aluminum smelters (Brownfield application) There are several considerations for a smelter to install a high draft ventilation system:

Figure 5 - Schematic of one section of high draft ventilation in a new smelter Because the ventilation air flow from a cell in this system is redirected to the parallel duct, a diverter type of valve is installed (Figure 6). When the operator engages the high draft ventilation mode (also referred to as maintenance mode) this valve moves its position and enables the parallel duct to ventilate the pot gases at the high rate.

Normal Draft

•

Tighter regulatory standards or working within a unit mass per unit time permit restriction

•

Creating better working local conditions in the potroom during anode changes

•

A considerable reduction of total fluoride emissions is required

•

Over time several line amperages increments have occurred, thereby increasing open pot work and emissions into the potroom so that more flow under normal draft is required.

One of the major constraints is that the GTCs are likely to run at 100 percent of their capacity. It is not a common practice that GTC operations are adjusted to changes in the potline such as line amperage increases. In other words, the normalized ventilation flows remain unchanged over time. However, higher amperages generate higher heat losses to the ventilation air and the average gas temperatures increase to a point the GTC operation and performance are negatively impacted. Higher temperatures push the air-to-cloth ratios and can also cause increased fluoride emissions from the GTC.

High Draft

Figure 6 - Diverter damper

363

Benchmarks without High Draft Ventilation

1 f 1

The case for high draft ventilation, in specific using a dual duct system, has been put forth as a means to achieve lower levels of fluoride emissions primarily during anode change. Many smelters in the world regard this technology to be the price of admission to attain total fluoride emission rates of less than 0.40 kg TF/tonne. However, this is not necessarily the case. Emission rates below 0.30 kg TF/tonne Al have successfully been attained at some Alcoa smelters without the use of high draft ventilation.

$M

Π I

TOTAL FLUORIDE EMISSIONS Without High Draft Ventilation 0.50

BLfLej

EBFS

Figure 8 - Dual duct high draft ventilation at Alcoa Deschambault with at therightthe smaller, dedicated GTC to handle the extra flows •

Figure 9.

If there is no room to accommodate the flow requirements for high draft ventilation then an alternative solution must be found. One solution that has successfully been applied is to add a small GTC besides the exiting GTC, as shown in Figure 8 and

2

3

4

5

6

7

8

Annual Periods (Years)

Figure 10 - Total fluoride emissions without high draft ventilation 1

/4 Potroom

As illustrated in Figure 10, the total fluoride emissions (combined pot room roof emissions and the GTC stack emissions) can be sustained at, or below, 0.30 kg TF/tonne as an annual average. In this case it should be noted that hot butts have been removed from by transport vehicles to an area of fume capture and treatment for more than 20 years.

parallel duct

- diverter damper

There are multiple factors that contribute to this success. Dedication and commitment of the work force is essential. This includes the proper management of draft on pots using a single duct system with dampers at each cell. When the operator puts the pot into a status that requires covers to be removed, the system automatically goes into the high draft mode for that pot. After a certain number of minutes the system times out and is returned to the normal damper position. Figure 9 - Schematic of one section where high draft ventilation is applied in a Brownfield application

The dual duct system provides significant improvement of capture of emissions from the pot as compared to single-duct dampered systems that rely upon manual placement into the proper drafting position. The greatest shortcoming of this approach is that pots that are left in the improper draft position can greatly detract from overall emission performance of other pots and the potroom.

What needs special attention is the challenge of installing a high draft system while a potline is in operation. Each cell needs a tiein by removing a piece of existing ductwork and replacing this with a diverter damper. This operation needs to be done very quickly. During the change over all draft is lost and emissions escape freely. However, this installation must also be executed very carefully because of the imminent electric dangers that exist while working on parts that connect directly to the superstructure. This meant that during the project in Deschambault special installation tools needed to be developed to safely install all different shaped ducts and pieces without any injuries.

The location that achieved the impressive result shown in Figure 10 has taken extra measures to assure that the overall system is tightly sealed to prevent in-leakage between the main fans and the reduction cells. They also place strong emphasis on regular checks of the system balance of flows from each cell.

364

What is most significant to the steady decline in the emissions baseline observed in Figure 10 has to do with how normal, or low, draft is managed. As equipment, pot covers and work practices were improved it became possible to reduce the normal drafting of the cell to levels below what is typical for modern technology. In effect this is a reduction in Derneddes's open area factor, A, that is used to estimate minimum cell ventilation.

FLUORIDE EMISSION BENCHMARKS

When combined with the other efforts to prevent in-leakage and to balance the system, this allows a pre-existing system to deliver greater suction and exhaust flow when it is needed at a cell. Exhaust rates may then be increased by >50% at every reduction cell when it is needed during anode change or other operations. B

This more efficient use of exhausting is not unique to this location, but it is somewhat uncommon in our industry. There are multiple advantages to be gained by placing focus on low draft that is indeed low and high draft that is at a maximum when needed. Most of these advantages are in costs, but there are process control benefits as well.

C

D

E

F

G

Benchmark Smelting locations

Figure 11 - Totalfluorideemissions from benchmark smelters Benchmark performance beyond this graph has been recently identified at 0.16 kg TF/tonne Al and at 0.13 kg TF/tonne Al at one smelter that had exceptionally good performance over a nine month period.

With this example given it is possible to reset some of the common perceptions around dual duct systems. They do open the door to lowering rates of fluoride emission. However, the absolute threshold is closer to <0.30 than 0.40 kg TF/tonne Al with a single duct system.

Figure 11 gives examples of fluoride emission performance that can be achieved using high draft ventilation systems at smelters with extremely demanding emissions limits and air quality standards. It will also apply to mega-smelters that strive to have no significant impact upon sensitive vegetation beyond the buffer zone around the facility.

Benchmarks with High Draft Ventilation Examples have been given in this paper on the reduction of fluoride emissions that are achievable with the installation of high draft ventilation systems. These systems can greatly reduce emissions provided they are designed, maintained and used properly.

So far, the data presented in this paper shows overall results of what smelters can achieve. However, there was a rare occasion where it was possible to capture the overall impact of a dual duct ventilation system. This is associated with the start-up of the dual duct system in Alcoa's Deschambault smelter during the fall of 2002. Here it was possible to collect totalfluorideemissions data from roof monitors prior to and after the start-up. This data is shown in Figure 12. Please note that during 2001 the plant also started using enclosed pallets to transport spent anodes and bath. This lowers the HF part of the total fluoride emissions and is included in the trend in Figure 12.

Residual fluoride emissions will primarily exit the pot room roof during anode changing, which is shown in Figure 1. If the gas treatment center (GTC) is not followed with wet scrubbers there will also be a few one-hundredths of a kilogram per ton of emissions from the stacks of these systems as well. Figure 11 shows what might be expected after the installation of high draft ventilation and other world class technologies and practices are implemented and maintained. These are annual average values for the best smelters in the world in the year 2008. Most, but not all, of these top performing smelters use dual ducts.

r~~~

TOTAL FLUORIDE EMISSIONS Before & After start-up of High Draft Ventilati 0.50

\ % V \ \ % % %u V %0 % V

365

Figure 12 - Fluoride emissions recorded on roof monitors before and after the start-up of the high draft system in Alcoa Deschambault

For locations with dual ducted high draft ventilation systems the practical limit has been demonstrated to be <0.20 kg TF/t Al.

It can be clearly seen that the total roofline fluoride emissions have been reduced after start-up of the high draft system. Moreover, the variability in the emissions (outside of variations in emissions due to the seasonal variation in ambient conditions) was greatly reduced once the high draft system was on-line. This is another important aspect for operations.

Current world benchmarks for annual averages of total flouride emissions from the pot room roof plus dry scrubbing systems typically in cooler climates, are in the range of 0.15 to 0.20 kg TF/t Al. Less than 1% of our industry currently attains this standard of performance. Acknowledgements

The geographical location of a smelter will impact any seasonal trends observed in monthlyemission performance. Deschambault has cooler, shorter summers than a smelter in the Middle East, for instance. The seasonal trending shown in Figure 12 clearly reflects this impact.

The authors would like to acknowledge the cooperation of many people from Deschambault, Fjarφaäl, Alumar and Hatch in the preparation of this paper. We further wish to acknowledge the sponsorship of Alcoa, Inc. and Hatch Ltd. for their dedication to attaining high environmental standards at specific operating locations and in our industry as a whole.

Future considerations High draft systems are being used but there is still limited long term experience available. Some of the current experience is that to keep the full benefit of the system one has to maintain it well. Furthermore, improvements are needed to increase the consistency. For instance, flow detection should be considered in all high draft connections to ensure the right amount of draft is established. This is not always evident.. Also the position of the valve should be monitored more closely.

Stephen Lindsay wishes to acknowledge Dr. Margarita R. Merino (Ph.D. - Florida State University) - for her encouragement, dedication and support. References

Other considerations relate to the increasing temperatures of the ventilation air. The use of high draft ventilation is also a good tool when both temperatures and emissions need to be reduced. Presently, temperature reduction is not a key design factor but can be in future designs of systems for existing smelters. Conclusions Dual duct-based high draft ventilation systems present an important, technology based option that, when combined with practical measures such as enclosed spent butt and bath pallets, enable smelters to achieve world-class levels (~ 0.2 kg TF/tonne Al) of fluoride emission performance. These systems are rapidly becoming cost-of entry mandates for greenfield smelters. It is a tool that can be considered in existing smelters but then, based on the experience gained from the Alcoa Deschambault project, planning and installation must be done very carefully. At this point it is not certain how reductions of the last 0.2 kg TF/tonne Al will be addressed or achieved. If pot tending machine technology is somehow modified to capture fumes during work activities it will very likely be tied in to the gas treatment systems or use portable, crane-mounted "capture stations." While "triple duct" technology is unlikely, additional ducting from gas treatment centers to capture and control point sources of fugitive emissions (other than from cells) may be the way of the future beyond dual duct technology. Cost-effective methods for continuously monitoring the status and "health" of the ducting systems presents an option for minimizing periods of sub-standard ventilation flow. The practical limit for existing locations installed with single ducted systems is <0.30 kg TF/t Al.

366

[1]

Stephen Lindsay and Neal Dando, "Dry Scrubbing for Modern Pre-Bake Cells," TMS Light Metals, 2009, 275-280

[2]

Morten Karlsen, Victoria Kielland, Halvor Kvande, Silja Bjerke Vestre, "Factors influencing cell hooding and gas collection efficiencies", TMS Light Metals, 1998, 303-310

[3]

Edgar Dernedde, "Gas Collection Efficiency on Prebake Reduction Cells", Am. Ind. Hyg. Assoc. Journal 51(1): 4449 (1990)

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

Cells Thermal Balance SESSION CHAIR

Bernard AUais Rio Tinto Alcan France

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INCREASING THE POWER MODULATION WINDOW OF ALUMINIUM SMELTER POTS WITH SHELL HEAT EXCHANGER TECHNOLOGY Pascal Lavoie1, Sankar Namboothiri1, Mark Dorreen1, John JJ Chen1, Donald P Zeigler2 and Mark P Taylor1 1 Light Metals Research Centre, The University of Auckland, Private Bag 92019 Auckland, New Zealand 2 Alcoa Primary Metals, Aluminerie Deschambault, 1 Boul. Des Sources, Qc, Canada, GO A ISO Keywords: power modulation, shell heat exchanger, sidewall heat transfer Abstract With power prices constantly rising, and varying aluminium prices requiring operating flexibility, the financial incentive for smelters to adopt a power modulation strategy is becoming larger. However, the power modulation window, in which a smelter can safely operate its reduction cells, is limited. The Light Metals Research Centre has developed the Shell Heat Exchanger (SHE) technology for controlling the heat dissipation from aluminium smelting pot shells. By varying the air flow through the SHE, the heat removal from the shell can be increased or decreased as desired, doubling the previous power modulation window or allowing power modulation with minimal disturbance to the pot thermal balance.

Figure 1. EEX hourly power price and volume for the week of September 6 2010

This paper presents experimental results from LMRC's test facility, which show the shell temperature response when the SHE is operated in cooling or insulating mode. Steady state thermoelectric model results for these operating scenarios are also presented, outlining the impact on ledge thickness and other pot operating conditions.

Effect of Power Modulation on Cell Operations As described by Eisma and Patel [4], the dynamic response of the electrolysis cell to a reduction of 10 to 12% in power input starts with a sudden drop in temperature, followed by a linear drop in both temperature and liquidus over the following 24 hours, resulting in ledge freeze and reduction in liquid bath volume. It was also noted that the dynamic response returning to the original power input was slower, in regards to ledge melting and bath volume.

Introduction Need for Power Modulation A continuous increase in energy cost is changing the cost structure of smelters. Previously, energy accounted for less than 30% of production costs, recent figures show it amounts to 40%, or more in many smelters [1]. This forces smelters to change their operating strategy by finding alternative ways to reduce energy consumption in order to remain financially viable. Reducing the anode-cathode distance is a mechanism of choice, helped by the re-emergence of drained and other cathode technologies, especially in China [2]. Another mitigating strategy increasingly being considered and implemented is power modulation. This is mostly advantageous for plants requiring the acquisition of electricity on the spot market where a large shift in price is observed on a daily, or other periodic basis, or those able to obtain better contractual rates by agreeing to modulation. Figure 1 shows spot price on the European market for the week of September 6, 2010 [3].

The long term effects of power modulations on the electrolysis cells can be quite damaging. Operating a power modulation scheme over extended periods can lead to severe damage of cathodes due to the solidification of sludge on the surface following the mechanisms described by Taylor et al if the surface temperature remains below 945 to 950°C [5]. However, with adequate understanding of the dynamic heat balance of the cell and proper adjustment to the modulation scheme, day-night cycles of up to 10% power (amperage) variation can be operated indefinitely without observable damage to cathodes. In any case, the power modulation window, i.e. the range of power inputs where a given cell will operate sustainably, is very limited by the heat loss rate of cells, and their re-heating characteristics. Increasing the Power Modulation Window with Shell Heat Exchangers Power modulation introduces the opposite thermal challenge compared to capacity creep. In modulation the heat balance must be altered to either increase the heat generation or decrease dissipation to avoid excessive ledge freeze. By varying the heat dissipation from the pot shell, it is possible to adjust the dynamic

369

response of the pot to a change in power input, thereby enabling a larger power modulation window.

alternate SHE model, driven by a blowing device (such as compressed air or powered fan) produces similar heat transfer. These SHE options have been plant tested and the choice of suction or blowing mode should be based on specific smelter design and operating considerations such as cost of installation.

The Light Metals Research Centre (LMRC) at the University of Auckland has developed a technology capable of providing both controlled cooling and insulation to sidewalls using heat exchangers, installed on-line, with variable air flow. Moreover, the hot air from the heat exchangers is collected and removed from the potroom. The heat content of the air at 150-200°C can be recovered and re-used in various low grade thermal applications including desalination and refrigeration. The performance of LMRC Shell Heat Exchanger (SHE) using compressed air was previously published along with its potential application during capacity creep [6]. The present paper reports the performance of SHE powered using an extraction fan and fitted to the sidewall of a full scale cell demonstration model. The expected benefits of power modulation for smelters from SHE technology are also discussed using 3D thermo-electric modelling of a mid-range amperage reduction cell.

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Experimental Shell Heat Exchangers Figure 2 shows a schematic of one of the SHE designs developed by LMRC. The shell heat exchanger is comprised of two parts; an exchanger body and an exchanger inlet and outlet. The exchanger body incorporates vortex generators (not shown in the Figure) to enhance the heat transfer. None of the other cell heat exchangers reported to date has made use of turbulence promoters to increase the heat transfer rate. Although shell fins are used extensively to increase heat dissipation by increasing the transfer surface, they can restrict convective flow on the shell face and reduce radiative heat transfer by reflecting to each other at higher temperatures. The SHE however, makes use of proprietary turbulence promoters which increase the heat transfer rate when compared with the rates calculated using standard engineering design equations. More details are given below.

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Figure 3. SHE operation showing air flow path Sidewall cooling Demonstration Facility The details of the LMRC sidewall cooling demonstration models were discussed previously [6]. The facilities essentially represent two 3-cradle pot shells of 350 kA technology facing each other. The sidewalls are heated by electrical elements mounted inside the sidewall and are thermally insulated on the inner sidewall face to ensure almost all heatflowsthrough the shell. Experimental Procedure Four shell heat exchangers, (two of 140 mm wide and two of 95 mm wide) were mounted to the central inter-cradle space on one pot shell. This has given 77% area coverage of the shell within the inter-cradle space. Air flow through the exchangers was powered by a centrifugal fan located outside the demonstration model. Ducting from the fan was fitted to the exchanger outlets to provide the suction and remove the hot air. Figure 4 shows the actual arrangement of SHEs mounted on the central inter-cradle space of the demonstration model.

Figure 2. Schematic of Shell Heat Exchangers (SHE) in the demonstration model inter-cradle spacing Figure 3 illustrates in broad terms the mode of operation of extraction fan powered SHEs. The vacuum generated by the fan draws air into the shell heat exchanger inlet from the outside vicinity. The air flows up the gap between the steel shell and the heat exchanger and then exits via the opening at the top of the exchanger into the extraction duct which is connected to the fan. The cell wall is cooled while the exchanger air is heated. The

Figure 4. SHEs mounted using a simple one-piece hanger in central inter-cradle space of the demonstration model

370

The pot shell was initially heated to a steady state temperature with zero air flow through the SHEs. Once the required temperature was attained on the shell, air was admitted through the SHEs to cool the central inter-cradle space and the same flow rate was maintained until a new steady state value was attained. Figure 5 shows a typical shell temperature recording of the experiment. The following parameters were recorded during the experiments: • Temperature at various locations on the inter-cradle space with and without SHEs. • Flow rate, pressure and temperature of the extracted air, measured at the fan exhaust. • Velocity and temperature of the hot exit air from the SHEs (measured in the duct, approximately 1 m from the exit of SHEs). • Ambient temperature

Shell Temperature Reduction vs Vertical Position 355 SCFM -»-234 SCFM -«-130 SCFM - -39SCFM

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Shell Temperature vs Vertical Position

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Figure 6 compares the shell vertical temperature profile at various air flow rates with and without SHEs installed while Figure 7 shows the corresponding temperature reduction. When no forced air extraction is applied to SHEs (0 scfm), they act as insulators, locally increasing the average shell temperature by up to 75°C. When cooling was applied, a peak sidewall temperature reduction of up to 200°C was obtained in this case. The sidewall cooling was controlled by adjusting the extraction fan speed. The temperature of the air measured at 1 m from the outlet of the SHEs was between 170 to 200°C. This corresponds to a heat content of 9-10 kW per cradle position. -»-234SCFM

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The effective shell to air heat transfer coefficient for the SHE was calculated for various air flow rates by Equation (1). This is based on the heat gained by the exit air from the shell heat exchanger and considering only the shell surface covered by the shell heat exchanger as the heat transfer area.

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The shell to air heat transfer coefficient is an important parameter in the sidewall heat transfer circuit as it dictates the overall sidewall heat flow and hence the ledge thickness of high amperage aluminium cells. The effective shell to air heat transfer coefficient ranges from 25 to 30 W/m2.K for a typical pot shell approaching 350°C [7]. The LMRC shell heat exchanger, depending on the air flow, alters the shell to air heat transfer coefficient to adjust the sidewall heat flow.

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Shell Temperature °C

Figure 6. Vertical temperature profiles at various air flow

371

represent different air flow situations. In the second case 10 W/m2.K was used to represent an insulating, minimum flow situation. Two SHE sizes were also modelled, representing 25% and 80% surface coverage of the pot side shell (above the cathode surface level).

*μ is the viscosity of air, kg/m.s *CP is the specific heat capacity of air, J/kg.K *K is the thermal conductivity of air, W/m.K * calculated at mean air temperature inside the rectangular duct

Modelling results

Figure 8 shows the heat transfer coefficient at different air flow rates (air flow quoted per cradle position of the demonstration model). The overall heat transfer coefficient increased with increasing flow rate. Note that the heat transfer coefficients measured for the SHE are 1.5 to 2.5 times higher than that predicted for forced convection through a duct of same dimensions. This increase is due to turbulent promoters mentioned earlier. The increase in heat transfer coefficient compared to forced convection is consistent with the previous findings [9, 10].

The key inputs and outputs of the model are shown in Table I and the predicted ledge profile in Figure 9. Table ] . Key modelling input and output

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Base Case 222 4.14 948 0 35

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BC2_001 BC2_002| 200 200 4.82 4.82 948 948 25% 80% 10 10 |

KEY OUTPUTS Base Case Avg Bath Temp (°C) 958.2 10.2 Superheat (
BC1 964 16 404.1 638.1 162.3 340.4 0.7

BC1J001 BC1 002 961.9 962.2 13.9 14.2 329.3 274 515 429.4 161.2 160.5 339.9 340 1.6 2.1

BC2 954.4 6.4 259.9 381.4 157.4 321 9.5

BC2_001 BC2_002 954.7 954.8 6.7 6.8 288.7 382.3 435 568.1 157.8 158.9 321.3 321.5 7.8 3.8 I

VOLTAGE BREAKDOWN Base Case 1.76 Cell EMF 0.544 Anode V Drop 3.315 Bath V Drop (Including EMF) 0.491 Cathode V Drop 4.350 Model Voltage (Excl. extern.)

BC1 1.76 0.593 3.389 0.528 4.51

BC1 001 BC1 002 1.76 1.76 0.592 0.593 3.392 3.395 0.529 0.53 4.513 4.518

BC2 1.76 0.482 3.363 0.443 4.288

BC2 001 BC2_002| 1.76 1.76 0.395 0.481 3.460 3.370 0.414 0.443 4.269 4.294 |

HEAT GENERATION Reaction Voltage (V) Ohmic Generation (V) Heat Generated (kW)

Base Case 2.022 2.328 516.8

BOI 2.022 2.488 597.1

BC1J001 BC1J)02 2.022 2.022 2.491 2.496 597.8 599.0

BC2 |BC2_001 BC2_002| 2.022 2.022 2.022 2.266 2.247 2.272 | 453.2 I 449.4 454.4 |

HEAT LOSSES (kW) Loss Rod/Yoke Loss Cover Total Loss Shell Loss Shell Side Loss Shell Bottom Loss Collector Bars Total Heat Loss (kW)

61.6 173.4 215.9 174.5 41.4 61.9 512.8

65.7 187.5 271.4 229.0 42.4 68.4 593.0

KEY INPUTS Line Current (kA) ACD (cm) Liquidus (°C) SHE coverage (%) HTC in SHE (W/sq.m.k)

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Figure 8. SHE effective heat transfer coefficients as a function of air flow rate To determine an approximate value for effective shell to air heat transfer coefficient in insulation mode, experiments were conducted by supplying a small but measurable amount of air to the shell heat exchanger (10 SCFM per cradle position). This small amount of air supplied did not change the vertical temperature profile from a no-flow situation. The effective heat transfer coefficient under this condition was calculated using Equation (1) and found to be between 7 and 10 W/m2.K (as shown in Figure 8).

65.6 182.6 278.5 236.6 42.0 66.6 593.4

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56.9 159.4 177.5 137.0 40.6 54.5 448.4

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A thermo-electric 3D slice model of a mid-amperage cell was built to illustrate how the heat balance of a cell can be manipulated by controlling the heat transfer coefficient at the shell to increase the power modulation window. Although the steadystate nature of the model does not show the short term dynamic response of the cell (in the first few hours), it has proved useful in indicating the effect of SHE operation on the sidewall heat flow and ledge thickness over a number of days.

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Taking a base case where the cell would normally operate at 222kA, with a voltage of 4.36V, alternative cases were ran at 200kA and 240kA, both at increased ACD, with and without SHE. Various heat transfer coefficient inside the exchanger were applied. In the first case 80 and 150 W/m2.K were used to

It can be seen that the use of heat exchangers at low flow partially mitigates the increase in shell temperature caused by the higher power input in the cell, while providing an increase in ledge thickness (BCl and BC1_001). The redistribution of the heat balance, the increase in ledge thickness and the reduction in

372

superheat offer an opportunity to increase the amperage further. Increasing the heat transfer coefficient in the SHE to 150 W/m2.K (BC1_002) results in a further increase in ledge thickness. Simulating the opposite situation, where the power input to the cell is reduced by 22 kA with an increase in ACD (BC2), the ledge pinch point increases to 9.5 cm and the superheat is reduced to 6.4°C. Introducing the SHE with a heat transfer coefficient of 10 W/m2.K (BC2_001) results in a small increase in superheat (0.4°C) and a retraction of the ledge, especially at the pinch point (-1.7 cm). It is important to note that the top of the ledge remains unchanged, contacting the anode side. Unlike the higher amperage case, a further reduction of the heat transfer coefficient is difficult without endangering the SHE in its present configuration. However, the heat transfer through the side can be further limited by increasing the surface coverage of the heat exchangers (BC2_002). This results in the ledge retracting almost to the Base Case position, effectively doubling the power modulation window on the low side to a reduction of 40kA. Higher SHE coverage would also increase the heat recovery potential, requiring lower flow for a given cooling at high amperage and increasing heat grade. Note that in all cases, the heat balance is adjusted mostly between the sides and the top of the cell, leaving the ledge position on the cathode surface without significant changes. Figures 10, 11 and 12 show the predicted impact of the shell heat exchangers on the sidewall heat flux, the bath superheat and the ledge thickness.

Scatterplot of Superheat (°C) vs Heat Generated (kW)

150, 25% coverage coverage HTC=80, 25% coverage

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500 525 550 Heat Generated (kW)

Figure 11. Impact of SHE configuration on superheat Scatterplot of Ledge Pinch Point (cm) vs Heat Generated (kW) No SHE

i6 HTC=80, 80% coverage HTC=150, 25% coverage-

HTC=10, 80% coverage

HTC=! No SHE 500 525 550 Heat Generated (kW)

Figure 12. Impact of SHE configuration on ledge pinch point It can be seen that the heat flux through the sidewall is manipulated by the Shell Heat Exchanger configuration. This results in significant changes of the ledge profile and superheat inside the cell. Higher heat flux and ledge thickness can be achieved with increased air flow through the SHE or by increasing the area covered by the exchangers but the effect on the ledge is small. Although the heat loss with the SHE in insulation mode changes only through coverage, the impact on the ledge thickness is large.

Figure 10. Impact of SHE Configuration on Side Heat Loss

In contrast, the impact on the bath superheat appears to be limited in insulation mode, while the effect in cooling mode seems constant and attributed to the presence of the Shell Heat Exchangers, with no further decrease in superheat at higher flow or coverage. Conclusions This paper demonstrated that the Shell Heat Exchangers designed by the Light Metals Research Centre, through air flow variation and the use of proprietary turbulence promoters, are effective in manipulating the heat transfer coefficient at the pot shell,from10 to more than 180 W/m2.K. With SHE's operated in insulating mode, an average shell temperature increase of up to 75°C was achieved. When cooling was applied, a peak sidewall temperature reduction of up to 200°C

was obtained and the cooling was controlled by adjusting the extraction fan speed. The temperature of the exit air measured was 170 to 200°C, corresponding to a heat content of 9-10 kW per cradle position. The heat transfer coefficients measured for the SHE were 1.5 to 2.5 times higher than that predicted for forced convection through a duct of same dimensions, due to the turbulence promoters used in the SHE. SHE operation enables a fast readjustment of the cell heat balance by controlling the heat loss from the shell, suitable to counterbalance major power input variation to the cell. It results in an increased window of power modulation, in which a given electrolysis cell can operate stably over time. The power modulation window was shown to be doubled with the use of the shell heat exchangers The next step of work in regards to the operation of SHE for power modulation involves confirmation of the benefits in operating cells and a study of the slowed dynamic response when using Shell Heat Exchangers. Acknowledgement The authors would like to acknowledge the essential contribution of David Cotton and Dr. Ronny Etzion to the experimental work, as well as Dr. Jianning Tang to the thermo-electric modelling. References [1] AME Smelter Cost Curve, 2009 [2] Feng N., Tian Y., Peng, J., Wang Y, Qi X., Tu G., "New Cathodes in Aluminum Reduction Cells", Light Metals, 2010 [3] European Energy Exchange, www.eex.com [4] Eisma D., Patel P., "Challenges in Power Modulation", Light Metals,2009 [5] Taylor M.P., Liu X., Fräser K.J. and Welch B.J., The Dynamic and Performance of Reduction Cell Electrolytes, Light Metals 1990 [6] Namboothiri S., Lavoie P., Cotton D. and Taylor, M.P., "Controlled Cooling of Aluminium Reduction Cell Sidewalls Using Heat Exchangers Supplied with Air, Light Metals 2009 [7] Haugland E., Borset H., Gikling H. and Hoie H., "Effects of Ambient Temperature and Ventilation on Shell Temperature Heat Balance and Side Ledge of an Alumina Reduction Cell", Light Metals 2003 [8] Perry R.H., Green D.W. and Maloney J.O., "Perry's Chemical Engineering Handbook", 6th ed. 1984, USA: Mc GrawHill chemical engineering series [9] Sutherland W.A., "Improved Heat Transfer Performance With Boundary - Layer Turbulence Promotors, Int. J. Hear Mass Transfer Vol. ICI, pp. 1589-1599, Pergamon Press Ltd. 1967 [10] Hung Y. H. and Lin H.H., "An Effective Installation of Turbulence Promoters for Heat Transfer Augmentation in a Vertical Rib-Heated Channel, International journal of heat and mass transfer Vol. 35. No. 1, pp. 2942, 1992

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

New approaches to power modulation at TRIMET Hamburg 1

Till Reek1 TRIMET ALUMINIUM AG, Niederlassung Hamburg, Aluminiumstrasse, 21129 Hamburg, Germany Keywords: Power modulation, Energy price, energy models, side ledge measurements Due to liberalization of the energy market in Europe, it is virtually impossible to realize energy tariffs that are comparable to most of the other production location for aluminium in the world. In addition, in Germany several taxes and levies were introduced to advocate renewable energies, resulting directly in a distortion of competition [1]. Due to the effect, that most of the renewable energies are non-base load suppliers of electricity - such as wind or solar power - the prices are also affected by the actual generation of renewable energy. The availability of wind energy is highly inconsistent and unfortunately - not related to the actual demand. Figure 1 illustrates the variability of energy from wind generation. In 2006 in peak times up to 16,000 MW was generated from wind in Germany, however, most often the energy amount generated did not surpass 1,000 MW. This reflects directly onto the actual energy prices as power suppliers have to keep spare capacity to be able to still deliver energy on demanded, independent of the availability of wind or solar energy.

Abstract TRIMET ALUMINIUM AG acquired the Reynolds cell technology P19 smelter in Hamburg in 2006 after a complete shutdown. The 135 ktpa smelter was restarted successfully in 2007 and has been in continuous operation since. The increasing spread of the energy price during night and day time, as well as long term price difference have lead to novel approaches to decrease the average energy price by operating the smelter with non linear energy input. Theoretical calculations were done to estimate the maximum energy difference of various cell states. Experiences were made of the impact on performance vs. modulation range. This paper presents the theoretical background as well as practical data of a smelter operating with power modulation. Introduction TRIMET acquired the facilities of the Hamburg smelter in November 2006. The plant was shut down by the end of 2005 by the former owner HAW. TRIMET restarted the plant in fast order, operating at full capacity of 270 pots at 175 kA at Christmas 2007. In 2008 operation was very smooth; however an increasing power price as well as a huge spread of prices over the day presented an incentive to operate at non continuous current. For the 2nd half of 2008, all potlines were operated with power modulation. Due to the economic crisis in 2009, production was cut to 50%. During that time, modulation was stopped as energy prices had leveled out at a comparatively low level. In expectation of increasing energy prices and increasing price spreads, work continued to set up all systems for power modulation. After restarting to full capacity at the end of 2009 and early 2010, modulation was restarted in one potline to continue optimizing procedures.

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Incentives for power modulation

In addition, with a trading platform as the European Energy Exchange (EEX) which was formed in 2002, prices tend to reflect the generation cost of the last MWh instead of a market average. Since the founding, energy is traded for each individual hour of the day, fixing a price between a fixed generation capacity and a variable energy demand. As energy demand peaks during daytime, prices increase for these hours as well.

Customarily, aluminium smelters are built in locations where there is constant and moderately priced electrical energy available. As electrical energy is the largest part of the production cost, such a development is unsurprising. Following this strategy, aluminium smelters were built in central Europe in the early 70s. When the first oil crisis struck in the mid 70s, construction was stopped and expansions curtailed. With increasing energy prices and increasing energy demand in Germany since then, several smelters were closed down over the past 10 to 15 years. The remaining ones are facing an increase in energy price, which is pushing them to the brink of profitability, especially as they have to compete with modern, large scale green field projects as they are constructed in the Middle East or Asia. The only advantage older smelters have is that their construction cost is already amortized and they don't have to pay capital costs for the initial investment.

Figure 2 shows the average energy price per hour in June 2008. Price varied from 40 €/MWh during night time to 150 €/MWh during midday on average. The peak price in that month was 213.08 € on June 25th between 11:00 and 12:00 am[4]. Even considering aluminium prices listed at the LME between 2,900 US$ and 3,100 US$ for this month, producing aluminum when the energy price is repeatedly above 100 €/MWh or 155 US$/MWh is not feasible.

375

The aim of process control is to drive the cells back into their respective target range and keep them there. With increased effectiveness of process control, the distribution of cell states represented as bath temperature - has narrowed, leading to a better operation t31. Relating a smaller variability of cell states to an actual range for optimum target bath temperature, results in a variance of cell states that still leads to optimum performance.

i

2

3

i

S

6

7

8

9

10

- ~Jun-08

ii

12 13

14 15

S6 1?

18

19 20 21 22 2Ú

Materia 1

2*

AvgJune 2008

Amount [kgl

Temp. [°C]

Energy contained rkWhl 4,438 5,282 2,829

Anode 13,230 850 c 935 Cathode 14,000 c Ramming 935 7,500 c paste Collector Fe 12,400 750 1,603 bars Na3AlF6 890 Side ledge 5,000 1,623 3,344 Al 955 Metal pad 8,500 Bath 6,000 Na3AlF6 Cryolite 4,860 955 2,467 A1203 150 955 88 955 690 408 A1F3 CaF2 955 300 129 Sum 22,210 Table 1: Temperatures and amount of materials contained in a cell at955°C

Figure 2: Average energy price variation in June 2008 as listed on the EEX[4] Scope of power modulation Due to the incentives presented earlier, some smelters have to adopt a strategy to cope with power modulation in order to remain profitable. Even though most of the industry still tries to operate with a current as constant as possible, there are incidents which indicate that a certain degree of modulation to the power input is in fact the norm. For example, line amperage drops if there in an anode effect on the line. Depending of the length and cell voltage of the anode effect, the decrease of line current can exceed 10 kA and last for several minutes. This happens routinely and it is not known to have a negative impact on the rest of the pot line. The most severe case of power modulation is the line shutdown to cut out a cell for relining or to cut a cell back in line. Measures are taken to reduce the time needed for that operation, but a shutdown of a line for 10 to 15 min is quiete common.

Materia 1

Amount [kg]

Temp.[°C ]

Energy contained [kWh] 4,507 5,467 2,928

It might be a stretch to call this power modulation. However applied differently, it very well falls into this exact field. It is common practice in Germany that in times of grid instabilities, the grid operators, after consulting with the smelter management, may throw the smelter off the grid as a substitution for cutting in a peak load power plant to stabilize the electrical grid. Also, over the past years, a practice called peak shaving as been used repeatedly. This is practiced mainly during summer, when energy demand around midday peaks and prices at the EEX soar as well. In consultation with the grid operators, the smelters cut out one or several lines to reduce their electricity uptake for one hour. The energy is sold back to the grid operates. Smelters have been able to generate a profit of several tens of thousands Euros on each occasion.

Anode C 13,230 860 955 Cathode C 14,000 Ramming 955 C 7,500 paste Collector 850 Fe 12,400 2,236 bars Na3AlF6 Side ledge 2,375 920 793 Al Metal Pad 8,500 975 3,399 Bath 8,625 | Na3AlF6 975 Cryolite 7,485 3,877 975 150 89 A1203 690 975 413 AIF3 CaF2 300 975 131 Sum 23,841 Table 2: Temperatures and amount of materials contained in a cell at975°C

Looking at the cell side of power modulation, it is a well know fact, that cells can attain different energy states due to various process conditions. The control of a cell always work to keep the cell within a very limited range of bath temperature und composition. In some instances, the control even works towards a set, single value temperature. However, looking at publications over the last years, there is no general agreement on the optimum temperature. Typically, the target temperatures in the range of 955 to970 °C are cited t5], depending on cell technology, process control and operating philosophy. If the bath is lithium modified, temperatures as low as 945 °C are quoted. The general agreement is that cells perform optimum in that range and efficiency decreases if you operate at bath temperatures considerably below or above.

If accepting temperature swings between 955 °C and 975 °C, assumptions can be made to the energy differences between these cell states. Table 1 and Table 2 state different temperatures for various materials located in the cell as well as the energy needed to heat the material to the given temperature and - if applicable melt it. Temperatures for certain materials are a collection of cell investigations, such as bath and metal inventory determinations, temperature gradient measurements in the cathode lining, cell construction data and extrapolation. To simplify the calculation, materials which have a definite temperature gradient - such as current collector bars, anodes and cathodes - are still treated as having a average, uniform temperature. Care is taken for the

376

collector bars to avoid temperature swings beyond recrystallisation temperature. Absolute temperatures are not as important as the temperature variation between the different states.

Trails Using the data from the delta energy calculation, a model was set up to calculate the energy demand of the cell operating over a range of amperages. This model assumed a constant heat loss as a first step on simplification. The model includes variables for the flowing cell parameters: Line amperage Cell resistance Alumina addition Anode consumption Parameters related to anode change Heat loss through off gas

The change in liquid bath mass is the result of the melting of side ledge. In this scenario, the AlF3-content dropped from 13 % to 8 %, which is a common occurrence for the stated temperature change. Figure 3 shows the contained energy graphically. The major contribution to storing energy is from melting side ledge and increasing the liquid bath mass during the high energy period. It highlights, that cells operated on power modulation are very susceptible to hot or red side walls. If the cell has been driven to maximum amperage, this danger is increased manifold. Due to constrains mainly in the anode bake furnace and also in anodic current density, the TRIMET Hamburg P19 cell is operated with a wide side channel and sufficient side ledge.

With this model, modulation schemes were set up to stay within the parameters defined previously, namely, aiming for a midterm energy balance that is similar to the one the cell operated at continuously. In anticipation of the effect of the modulation on efficiency, only a fraction of the calculated theoretical energy difference was reached. The cell was operated keeping the resistance set point constant, aiming for an unchanged anode cathode distance. However, due to cooling effects during the modulation cycle ACD was effectively squeezed. This method is the most severe with regards to cell condition; however, it is also the most effective to reduce absolute energy consumption for the reduction line.

* Side ledge ft Additives

| 1

M Na3A!F6(!) ♦ Collector bars

The first scheme was then put into operation on one reduction line, keeping the second line operating at constant amperage as a reference for general performance. As both lines are supplied with the same alumina as well as the same anodes, both featuring the same pot design and operating practice, the difference in performance can be attributed to power modulation.

♦ Metall pad

c

« Ramming paste

1

• Cathodes ♦ Anodes

cell at 955 *C

The modulation scheme was designed to have a maximum energy difference of 490 kWh, or roughly 30 % of the theoretical value. The scheme consisted of increasing the current up to 185 kA during night time. To release this energy during the daytime the current was lowered to 160 kA for 4 hours. To avoid negative impact of excessive cooling, as was experienced during the modulation in 2008, the scheme operated at a slightly increased energy input compared to stable operation.

cell at 975 X

Figure 3: Comparison of two different energy states of a reduction cell For the cell as it is installed in Hamburg, a theoretical limit of 1.6 MWh was calculated for the energy difference that can be stored in the cell. Power modulation should aim for two goals: 1.

2.

The line operated at this modulation strategy for 2 month. To be able to evaluate the performance after only 2 month, several cells were set up as reference pots. Measurements were done to evaluate the cells in detail. To ensure that production figures are indeed representing the actual cell performance, metal inventory test were made and the cells were tapped separately for the whole period. It was proven that the cell performance corresponded to reduction line performance with regards to current efficiency and energy consumption. No impact of this modulation scheme on reduction line efficiency was found.

The average power introduced into the cell shall average to an energy input that is similar to the energy input with which the cell was operated previously on continuous amperage. This should keep both, energy needed for metal production as well as heat loss in mind. The maximum deviation from continuous power input shall not exceed the theoretical maximum.

The assumption is that your process efficiency will decrease significantly if you exceed these boundaries. In 2008 TRIMET Hamburg operated on a modulation scheme that was significantly below nominal energy input. During this time, significant losses in current efficiency where recorded. Not all cells can be operated at exactly that temperature swing, but are scattered close to an average. Thus efficiency will decrease significantly as one approaches or exceeds this limit as an increasing fraction of cells is beyond .the optimum control range.

Bath inventory changes and side ledge measurements Side ledge measurements were done on selected test cells during the modulation. There are several reasons for measuring the thickness of the side ledge during different energy states. Firstly, the measurement was used to validate the change of bath inventory during the modulation. Secondly, it was used to check if there is a change in the shape of the side ledge, e.g., moving from side ledge to bottom freeze or similar. Lastly, during the high

377

energy phase, there is little side ledge left. If there are one or more problem spots where the side ledge melts first, exposing the underlying SiC bricks, a serious impact on cell life is expectable.

Anode 18 Anode 1? Anode 16 Anode IS Anode 14 Anode 15 Anode 12 Anode 11 Anode 10

For measuring side ledge thickness, a device was used that lowers a turnable hook into the bath. The hook is turned until it touches the side ledge on the inside of the cell. By measuring the angle by which the hook can be turned until it touches solid ledge, the thickness of the side ledge can be calculated. The hook is lower in 50 mm steps to measure the thickness over the whole height of the liquid phases. Measurements points were taken in front of all 18 anodes. The device is reverenced to the outer edge of the deck plate. Figure 4 shows the measuring device in detail. Figure 6: Side ledge profile of the downstream side of cell 331 after the high energy period Changes in side ledge thickness were found to be more pronounced on the downstream side of the cell. In general, the side ledge was thinner here as well. This is in agreement with the flow patterns within the cell. Anode 18 Anode 17 Anode 16 Anode 15 Anode 14 Anode 13 Anode 12 Anode 11 Anode 10

Figure 4: Measuring device for measuring side ledge Measurements were done on test cells at the beginning and at the end of the period, during which the line operated with modulation. During two months of operation with power modulation, a trend of decreasing side ledge thickness in the metal pad was found. This was attributed to the slightly higher energy input than during continuous operation. At the end of the modulation period, the thickness of side ledge was still within acceptable limits. Anode 18 Anode 17 Anode 16 Anode 15 Anode 14 Anode 13 Anode 12 Anode 11 .Anode 10

Figure 5 and Figure 6 show the thickness of the side ledge from the cathode surface upwards. As can be seen by the rather high values in general, the cell operates with a substantial side channel, enabling a thick side ledge that can be used for power modulation. On average, the cell melted 5 cm of side ledge moving through a 24h modulation cycle with a neutral energy balance. Assuming that a similar amount of side ledge melted on the short sides of the cell, this amounts to close to 2 tons of material. To melt this much, 315 kWh are necessary. 64% of the energy introduced during the high energy period was used to melt side ledge. This is a significantly higher percentage than was estimated in the theoretical energy difference. Bath inventory tests were also done, but they were not conclusive. The time delay between the addition of tracer material and the final sampling proofed to be too long to result in a bath inventory that corresponds with the energy state of the cell. Results and future work

Figure 5: Side ledge profile of the downstream side of cell 331 after the low energy period

A test campaign was conducted at TRIMET Hamburg, operating one reduction line with a power modulation ranging from 160 kA to 185 kA. The maximum energy difference stored in the cell amounted to 490 kWh. This is about 30 % of the theoretical energy difference that a cell of this type can hold. It was found that with this modulation, there was no effect on current efficiency, compared to a second line of the same pot type operated on the same site.

With side ledge measurements a substantial change in bath volume was proven. No critical spots were found where the side ledge had molten completely. However, the change in ledge thickness was higher than anticipated. Further tests with different modulation schemes will be conducted to investigate the correlation between power modulation and melting and freezing of side ledge. Also, all models used assume a constant heat loss during power modulation. Exemplary heat flux measurements and off gas temperatures show that this is evidently not true. Due to the melting and freezing of side ledge and top crust, heat flux and heat loss change. This has to be incorporated into models tracking the energy state of the cell. To increase the range of power modulation beyond what is currently possible, further steps have to be taken that allow a modulation of the heat loss directly. For the heat loss through the off gas tests will be conducted to vary the suction rate of the dry scrubber. For the heat loss through the side wall, installations that regulate the air flow along the side wall have to be designed and tested. References [1] Kruse, H.; German power market - Final countdown for the aluminium smelters, proceedings 8th Austalasian Aluminium smelting conference 2004 [2] Pieper, T.; Fleckenstein, H.; Rosen, M.; Das Wohl und Wehe des Windes, EMW, Release 01-2007 [3] Iffert, M.; Rieck, T.; White, P.; Rodrigo, R., Kelchtermans, R; Increased current efficiency and reduced energy consumption at the TRIMET Essen smelter using 9 Box Matrix Control; Light Metals 2003 [4] www.eex.com [5] Utigard, T.A.; Why best pots operate between 955 and 970 °C; Light Metals 1999

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SOME ASPECTS OF HEAT TRANSFER BETWEEN BATH AND SIDELEDGE IN ALUMINIUM REDUCTION CELLS Asbjorn Solheim SINTEF Materials and Chemistry; P.O. Box 4760 Sluppen, NO-7465 Trondheim, Norway Key words: Bath; sideledge; heat transfer; superheat

average superheat, and local variations in the heat transfer coefficient (or superheat) may give local variations in the ledge thickness. The main resistances against heat flow are represented by conduction through the sideledge and by heat transfer (convection and radiation) between the potshell and its surroundings. To ensure a stable and thick enough sideledge, cooling ribs at the potshell could be a solution in some cases.

Abstract A literature review concerning the heat transfer coefficient between bath and sideledge (h) is given. Normally, the heat transfer is controlled mainly by the circulating bath motion due to gas drainage into the peripheral channel. After the introduction of slotted anodes that direct the gas towards the centre channel, it is likely that h will be determined by natural convection in some cases, and an equation is suggested to take this into account. The coupling between heat and mass transfer during melting and freezing of sideledge was studied in a numerical model involving multicomponent diffusion. During freezing, the concentration of bath components other than cryolite is higher at the ledge surface than in the bulk of the bath. The surface temperature of the ledge varies with the rate offreezingand melting, in such a way that the variation of the ledge thickness becomes slower than thought earlier.

The bath heat transfer coefficient determines the rate of melting or freezing of the sideledge in periods when the cell is not at thermal balance, i.e., during melting or freezing of sideledge. If the heat fluxfrombath to sideledge is different from the heat flux into the ledge, that difference is used for melting orfreezing.However, the sideledge has a chemical composition that is different from the bath. Therefore, melting or freezing of sideledge must involve mass transfer between the bath/sideledge interface and the bulk of the bath, and consequently, the bath composition facing the ledge is not equal to the bulk composition. This means that the assumption of a ledge surface temperature equal to the liquidus temperature of the bulk bath is strictly true only when the sideledge is at thermal balance, i.e., when neither freezing nor melting takes place.

Introduction No other material than frozen cryolite can withstand the harsh conditions encountered in aluminium reduction cells, comprising oxidizing gases, a cryolite-based electrolyte (bath) kept at high temperature, and reducing molten or dissolved metal. To maintain such afrozenlayer, the cells are designed to have a high heat flux through the sides. Besides protecting the sidelining, the freeze (sideledge, ledge) acts as a thermal buffer byfreezingand melting.

The present paper provides a general description of heat transfer between bath and sideledge. The main topic, however, is the problem of predicting the rate offreezingor melting in a situation with coupled heat and mass transfer.

The heat transported from the bath to the surface of the ledge is commonly calculated by the well-known equation q = h-(T b -T liq )

2

[Wm· ]

Heat Transfer at Steady State Literature Data

(1)

Some literature values [M3] for the heat transfer coefficient are given in Figure 1. The figure shows a considerable scatter, partly because the data were obtained by a variety of methods, including calculation and modelling, experiments in physical models, and measurements in cryolitic melts. Newer data, from the 1980s on, show higher values than older data, which is harder to explain. One fact that to be pointed out however, is that modelling, as well as interpretation of data obtained in physical models, requires a correct value for the thermal conductivity of the bath. New data on the thermal conductivity was published by Khokhlov et al. [13] in 1998. Their data showed a more than two-fold increase above earlier values. In fact, the heat transfer coefficients given by Khokhlov et al. [13] (item 13 in Figure 1) emerged from a reinterpretation of Solheim and Thonstad's experimental data obtained in a physical model [7]; the original interpretation is

where h is the heat transfer coefficient [Wm^K"1] and T is the temperature; subscripts "b" and "liq" indicate the bulk of the bath and the liquidus, respectively. Eq. (1) explicitly assumes that the ledge surface temperature is equal to the liquidus temperature of the bulk of the electrolyte. The heat transfer coefficient can be controlled by forced convection, or - probably less recognized by natural convection. When calculating the total heatflowfromthe bath and out through the sides, it becomes obvious that the bath heat transfer coefficient represents only a minor part of the total heat resistance between the bath and the surroundings. Therefore, it can be argued that the average value of the heat transfer coefficient between the bath and the sideledge is not important for the average ledge thickness. However, the average heat transfer coefficient determines the

381

shown as item 7 in Figure 1. The semi-empirical correlations given by Khoklov et al. [13] were later fully confirmed by CFD modelling [14]. It can, therefore, be assumed that the heat transfer coefficient between bath and sideledge is in the order of 1000

1500

x

1000

I

I 500

1 2 3 4 5 6 7 8 9 10 11 12

Korobov&Yanko(1967) Haupin(1971) Aral & Yamazaki (1975) Arita, Urata, & Ikeuchi (1978) lkeuci& Arita (1978) Peacey&Medlin(1979) Solheim & Thonstad (1983) Ohta & Matsushima (1984) Taylor & Welch (1985) Fräser, Taylor, & Jenkin (1990) Gan& Thonstad (1990) Warczok, Utigard, & Descleaux (1992) 13 Khokhlov et al. (1998)

*«

•2

"CD

4

*· 1960

* 1970

4,5

Nu

Ψ12

¥· i 13

1990

μ<^ρ

Ra =

pApgL 3 C p ëμ

0.31 "natural = 3 0 7 - ( Δ Τ ) '

2000

(4)

(5)

Figure 2 gives a graphical representation of Eqs. (2), (3), and (5). The heat transfer coefficients due to forced convection and due to natural convection are not additive; one of them will usually dominate. A simple way to account for this is to merge Eqs. (3) and (5) on the form

Forced Convection. The heat transfer coefficient is normally thought to be governed by gas induced circulation in the peripheral channel. The intensity of the convection increases with increasing rate of gas drainage into the channel (G, [mV1!!!"1]), increasing anode immersion depth (H, [m]) and decreasing anodeledge distance (L, [m]). Based on Solheim and Thonstad's [7] data for different part of the ledge, later re-interpreted by Khokhlov et al. [13], the following expression for calculating the average heat transfer coefficient can be suggested, [Wm-2K"']

Pr =

With data taken from Table I below, Equation (3) can be replaced by the much simpler expression

•8

Functional Dependencies

m"

hL. ë '

The transition from laminar to turbulent natural convection normally takes place at Ra » 1 109. Equation (3) is valid both in the laminar and the turbulent region, however[15].

Figure 1. Literature values of the heat transfer coefficient between bath and sideledge in aluminium cells.

31480

(3)

where L is the characteristic length (the bath height, taken to be 0.2 m), ë is the thermal conductivity of the bath, μ is the dynamic viscosity, Cp is the heat capacity of molten cryolite, p is the density of the bath, and Δρ is the density difference between the ledge surface and the bulk (calculated as 0.927 ΔΤ [19]). The numerical values for the physical data used in the present paper can be found in Table I in the following.

f

1980 Year

0.387 -Ra 1/6 9/16 4/9 [l + (0.492/Pr)9/16 ]

where Nu is the dimensionless Nusselt number, Ra is the Raleigh number, and Pr is the Prandtl number, Nu

Φ10

0.825 +

1500

1000

(2)

I

For an anode with dimensions 1.6 x 0.7 m and carrying 10 kA at 960 °C, the rate of gas drainage will be G = 5.7 10"4 mYW 1 , provided that the drainage is uniformly distributed along the anode periphery. Taking H = 0.15 m and L = 0.20 m, the heat transfer coefficient becomes 888 Wm"2K_1. This number may be somewhat higher due to magnetohydrodynamic flow along the channel[7]; on the other hand, it is known that more gas is drained into the centre channel than into the side channel. The latter is particularly important when using drained cathodes.

500

*

4

Natural Convection. Natural (free) convection is set up when there are density differences between the liquid facing the surface and the bulk. The density difference may be due to temperature gradients, gradients in chemical composition, or both. The following relationship was derived by Churchill and Chu (recommended in Welty et al.[l5]),

Q^-104

6 /m3s-1m-1

Figure 2. Graphical representation of the heat transfer coefficient between bath and sideledge, calculated from Equation (2) (forced convection) and Eqs. (3) and (5) (natural convection, Eq. (5) probably not visible).

382

h = ({31480(GH/L)0-46}n + {307-AT0-31M

Liquid

(6)

where the exponent n is a large number; e.g., n = 20 as used in Figure 3. A situation with high superheat and thin ledge (large L) will favour the domination of natural convection.

Bath (cryolite + X)

Solid

Pure cryolite

X

1200

Concentration gradient ofX

* 1000

I

iß

800

Eq. (6), n =

600

Eq.(5), Eq.(5),

AT^15°C AT=10°C

Eq.{5),

AT=5°C

Figure 4. Schematic illustration of fluxes at the surface of the ledge during crystallization of pure cryolite, using the moving surface of the solid as the reference frame. The total (global) flux corresponds to the rate of solid cryolite formation. The total flux of other components than cryolite is zero.

$ 400 Eq- (2)

The rate offreezingand the total molar flux Ntot are related by

200

N„ = ENi =

_i_

0

2

4

6

8

Jfre(

Ó Ì ΐ χ i(s)

[molm s J

(8)

where Mj and xi(s) are the molar weight and the molar fraction in the ledge, respectively, of component i. The positive molar flow direction is from the ledge towards the bath, whereas negative values ofj f r e e z e means that the ledge is melting away.

^■104/m3s1m-1

Figure 3. Transition between natural and forced convection (see text).

For a given rate of freezing the individual molar fluxes are given by the composition of the ledge, Dynamic Sideledge Behaviour

Nj = N t o t .x i ( s )

Theory

Turbulent Heat and Mass Diffusivities. In the bulk of the bath, the heat and mass fluxes depend exclusively on convection. Within the boundary layer at the sideledge, the convection decreases gradually and finally becomes zero at the ledge surface, at which point the transport depends only on chemical diffusion and heat conduction. Following the treatment in a previous paper [21], this can be described by using turbulent diffusion coefficients,

Basic Principles. The sideledge and the bath have different chemical composition. If the sideledge is formed at near equilibrium conditions, it will consist of nearly pure cryolite. This implies that any changes in the ledge thickness must be accompanied by mass transfer as well as by heat transfer (in fact, the terms "precipitation" and "dissolution" could be used, just as well as "freezing" and "melting"). During freezing ("precipitation"), a bath constituent that is not incorporated in the ledge will be enriched at the ledge surface, and it has to diffuse back to the bath. The basic principle was described earlier by Solheim and Stoen [20], and it is illustrated in Figure 4.

1 •(xiNj-XjNi) j=l tot^ij

For mass: Dy = D c ( y )+D t

(10)

Forheat: a = a c + a t

(11)

where a c = ë/ρΟ ρ

where Dc is the chemical diffusion coefficient, Dt is the turbulent diffusion coefficient, and a are the corresponding thermal diffusivities of the bath.

Multicomponent Diffusion. The bath is a multicomponent mixture containing NaF, A1F3, A1203, and CaF2. The diffusion phenomena must therefore be described using the Stefan-Maxwell equations (e.g., seeref. [15])

Vxi = Ó c

(9)

The turbulent quantities are zero at the ledge and increase with increasing distance from the ledge (y). Using boundary layer theory, it can be shown that Dt = <xt = C y 3

(7)

(12)

where C is a constant depending only on the degree of convection (i.e, a function of the Reynolds number). The mass and heat transfer coefficients, respectively, can then be calculated by

where x, is the molarfractionof component i, c^ is the total molar concentration [molm"3], Dy (= D^) is the binary diffusion coefficient for the system i-j [m2s ], and N is the molar flux [molm'V1].

k (i) = 0.827 C 1 / 3 - D ^

383

(13)

and

Results

h = 0.827 C 1/3 pC p a;2/3

(14)

Temperature Gradients. Some calculated temperature gradients are shown in Figure 5. During freezing, A1F3, CaF2, and A1203 become more concentrated near the ledge, and the liquidus temperature decreases. During melting, the concentrations of bath constituents other than cryolite decrease, and the liquidus temperature increases correspondingly. It follows that the "superheat" measured in specialized in situ sensors [25] in no way should be taken to represent the difference between the bath temperature and the surface temperature of the sideledge. Unexpected high or low measured "superheats" may indicate that the sideledge is either melting away or freezing during the time of measurement.

1

where k^ is the mass transfer coefficient [ms" ] for species / in cases where the overall ("global") flux may be neglected. Eq. (14) is derived in the Appendix. It should be noted that the transfer coefficients are proportional to D2/3 or ot2/3 at a given degree of convection (expressed by C), which is in accordance with boundary layer theory. Numerical Treatment Main Features of the Model. The numerical calculations were aimed at estimating the bath composition and the local liquidus temperature at the surface of the ledge at a given rate of freezing or melting. The calculations were performed in a spreadsheet, where the region from the ledge and 15 mm into the bath was divided into 1000 equally sized elements. The bath composition was fixed (at "bulk composition") 15 mm from the sideledge.

970

The numerical procedure involved the shifting between "old" and "new" values of the molar fractions in the discretised version of Equation (7). The procedure produced constant molar fractions after only about 5 iterations, and the operation was performed in an Excel spreadsheet. Assumptions and Input Data The physical properties for the bath and the diffusion coefficients are summarized in Table 1. The most uncertain data are probably the diffusion coefficients. For the system NaF-AlF3-Al203(Sat) Burgman and Sides [22] found effective diffusion coefficients for A1F3 close to 1 10"8 mV1, depending on the NaF/AlF3 molar ratio. The value 1.5 10"9 mV 1 for A1203 in cryolitic melts is based on work by Gerlach et ah [23] and Thonstad [24]. The values for binary systems including CaF2 are little more than surmises.

Local liquidus temperature

950

Limit for Formation of Pure Cryolite. If one assumes that pure cryolite is the only solid phase formed, the area close to the ledge will be supercooled if the rate of freezing is high enough. This is related to the fact that the thermal boundary layer (which determines the course of the local actual temperature) is thicker than the diffusion layer (which determines the course of the local liquidus temperature). The supercooling will increase with increasing distance from the surface, which favours dendritic crystal growth. In that case, pure cryolite will not longer be formed, since bath will be "trapped" between the dendrites. The criterion for formation of pure cryolite is, obviously,

Table I. Physical data used in the present paper. The bath contains 12 wt% excess A1F3, 5 wt% CaF2, and 3 wt% A1203. The liquidus temperature is 955.2 °C [17] and the superheat was assumed to be 10 °C. The melt was treated as a mixture of NaF (component 1), A1F3 (2), A1203 (3) and CaF2 (4).

D13 D,4

D23 D24 D34

Value 2.52 10° 2087 1883 502.3 0.795 1.0 10°

Ref. 17 19 18 18 13 22

1.5 10"9

23,24

3.0-10"9 1.5 10 a 3.0 10"a 1.5-10*

2.0

Figure 5. Temperature gradients close to the ledge. Grey curves - during melting (-1.96 mm/h), black curves - during freezing (0.84 mm/h). It was assumed that the sideledge consists of pure cryolite.

The heat flow through the ledge (qO was fixed at 8000 Wm"2, referred to the ledge surface area. The value of the "convection constant" C in Eqs. (12-14) was fixed at 365 m'V1, which gives a heat transfer coefficient of 800 Wm"2 and s a superheat of 10 °C when the cell is at thermal balance. The density of the sideledge was fixed at 2900 kgm"3.

Parameter Viscosity [kgrrf's"'] Density [kgm"J] Specific heat capacity [Jkg^K1] Heat of melting [kJkg1] Thermal conductivity [Wm^K"1] Diffusion coefficients [mV 1 ] D12

0.5 1.0 1.5 Distance from ledge / mm

dy

y=0

dT,liq dy

(15) y=0

-

With the present data and assumptions, the onset of dendrite formation was found to take place at afreezingrate as low as 1.13 mm/h (3.28 kginV).

-

Composition of Freeze. Dendritic crystal growth during rapid freezing was simulated by adjusting the flow (N) of each bath constituent in such a way that dTb/dy = dTliq/dy (see Eq. (16)); at

23,24

384

the same time, there should be agreement between the rate of freezing and the heat flow. From the calculated flow, the composition of the ledge could be computed, as shown in Figure 6. The cell is at thermal balance when the superheat (defined as the difference between the actual temperature and the liquidus temperature for the bulk) is 10 °C. If the superheat is above 7 °C, pure cryolite, or a solid solution containing small amounts of A1F3 and CaF2, is formed. If the superheat drops below 7 °C, the contents of A1F3, CaF2, and A1203 in the ledge start to increase abruptly. It is interesting to note that the change from "almost pure cryolite" to "almost bath composition" takes place in a very narrow temperature range, which may explain the alternating distinct bands of needle-shaped crystals and disordered or even porous structures observed in samples of industrial sideledge [20].

where the function f, which comprises the bath and ledge compositions, must be determined by numerical treatment. By replacing ΤΗò by T* in Eq. (16), followed by inserting Eq. (17) into (16), one obtains q s i-h(T b -T l i q )

Jfre«

where

AH m CF

Qp _ i

|

(18)

Hxmelt> xsolid)

In cases where the bath and the sideledge have the same composition, the correction factor CF will be unity and Eq. (16) applies, since there will be no concentration gradients near the ledge. A large number of numerical calculations were made in order to assess CF. A reasonablefitwas obtained by using the function

T

Jfre«

q s i-M T b- T iiq)

[kgm-Y1] (19)

AH m . 1 + 0.2 (X bath -X ledge c

c

where c are concentrations in weight percent of other bath components than cryolite in the bath and in the ledge. Figure 7 shows a comparison between Eq. (16) and Eqs. (18)-(19), using randomised variables (A1F3, CaF2, and A1203 concentrations in bath and ledge, heat flow, heat transfer coefficient, and superheat). As can be observed, Eq. (19) generally agrees much better with the numerical data, especially at low and moderate rates, and it predicts lower rates offreezingand melting than the traditional Eq. (16). This agrees with the work by Rye et al. [25], who studied the ledge behaviour in industrial cells; citation: "Results indicate that the melting of the frozen ledge [during anode effect] may be less severe than predicted by thermal models.'"

Figure 6. Composition of the sideledge as a function of the superheat. Open symbols - no solid solubility of CaF2 in cryolite, filled symbols - assumed 2.5 wt% solid solubility of CaF2 in cryolite. It was assumed that the solid solubility of A1F3 is 1 wt%. Rate of Freezing and Melting. The rate offreezingand melting is normally calculated by Jfre«

q s i- h -(T b -T l i q )

1

[kgm-Y ]

-

'•a

E Φ) -*

(16)

• •

• Traditional equation (16) o Revised equation (19)

100

0

\

·+ s

-

W^

C^s

where qsi is the heat flow from the ledge surface and out of the cell, and AHm is the heat of melting. As will be clear from the treatment above, however (e.g., Figure 5), the surface temperature equals the liquidus temperature only at thermal balance.

(TJ

analyti

o

Jfreeze *f(xmelt, xsolid) h

·· • • • •

— ♦

1

The numerical calculations indicated that the ledge surface temperature (T*) varies linearly with the rate of freezing or melting, at least when the rates are small. Furthermore, according to the Chilton-Colburn analogy [15] the mass transfer coefficients vary linearly with the heat transfer coefficient. Consequently, one would expect that the concentration difference between bulk and surface, and hence, the difference between T* and Tliq, is inversely proportional to the heat transfer coefficient. This gives Miq

0

^^

o

-200

^

· •

-200

-150

· -100

•

I

-50

50

jfreeze (numerical) / kgm'2h'1

Figure 7. Rate of freezing or melting; comparison between Eqs. (16) and (19) and numerical results, using randomised input data. The straight line represents the "ideal" relationship.

(17)

385

Acknowledgement The present work was financed by Hydro Primary Metal Technology. Permission to publish the results is gratefully acknowledged. The author wants to thank colleagues at SINTEF and in Hydro for discussions and suggestions during the preparation of this paper. Special thanks to Nancy Holt, Hydro PMT, for initiating this work. References 1. A. Korobov and E.A. Yanko, "Pecularities in the Convective Heat Transfer by the Smelt in Aluminium Converters of High Capacity", Publication of the VAMI Institute #60, Leningrad, 1967. 2. E. Haupin, "Calculating Thickness of Containing Walls Frozen from Melt", J. Metals 23 (July), 41 (1971). 3. Arai and K. Yamazaki, "Heat Balance and Thermal Losses in Advanced Prebaked Anode Cells, Light Metals 1975, Vol. 1, p. 193. 4. Y. Arita, N. Urata, and H. Iteuchi, "Estimation of Frozen Bath Shape in an Aluminium Reduction Cell by Computer Simulation", Light Metals 1978. Vol. 1, p. 59. 5. H. Ikeuchi and Y. Arita, "Treatment of Freeze-Melt Boundary Layer in Aluminium Reduction Cell", Yoyven 21 (2) 215 (1978). 6. J.G. Peacey and G.W. Medlin, "Cell Sidewall Studies at Noranda Alumium", Light Metals 1979. p. 475. 7. A. Solheim and J. Thonstad, "Model Experiments of Heat Transfer Coefficients between Bath and Side Ledge in Aluminium Cells", J. Metals 36 (3) 51 (1984). (Also in Light Metals 1983. p. 425. 8. T. Ohta and T. Matsushima, "Thermal Analysis of Soederberg Pots", Light Metals 1984. p. 689. 9. M.P. Taylor and B.J. Welch, "Bath/Freeze Heat Transfer Coefficients: Experimental Determination and Industrial Application", Light Metals 1985. p. 781. 10. K.J. Fräser, M.P. Taylor, and A.M. Jenkin, "Bath Heat Transfer and Mass Transport Processes in H-H Cells, Light Metals 1990. p. 221. 11. Y.R. Gan and J. Thonstad, "Heat Transfer Between Molten and Solid Cryolite Bath", Light Metals 1990. p. 421. 12. A. Warczok, T. Utigard, and P. Descleaux, "Heat Transfer Between Molten Cryolite and Solid Phase", Savard/Lee International Symposium on Bath Smelting", Montreal, Quebec, Canada, October 18-22,1992 (Proceedings, p. 325). 13. V.A. Khoklov, E.A. Filatov, A. Solheim, and J. Thonstad, "Thermal Conductivity in Cryolitic Melts - New Data and Its Influence on Heat Transfer in Aluminium Cells", Light Metals 1998, p. 501. 14. K. Bech, S. T. Johansen, A. Solheim, and T. Haarberg, "Coupled Current Distribution and Convection Simulator for Electrolysis Cells", Light Metals 2001. p. 463. 15. J.T. Welty, C.E. Wicks, and R.E. Wilson, Fundamentals of Momentum. Heat, and Mass Transfer. 3rd edition, John Wiley &Sons, 1984. 16. T. Herzberg, K. Torklep, and H.A. 0ye, "Viscosity of Molten NaF-AlF3-Al203-CaF2 Mixtures", Light Metals 1980. p. 159. 17. A. Solheim, S. Rolseth, E. Skybakmoen, L. Stoen, Δ. Sterten, and T. Store, "Liquidus Temperatures for Primary Crystallization of Cryolite in Molten Salt Systems of Interest for the Aluminium Electrolysis", Met. Trans. B, 27B 739 (1996).

18. M.V. Chase et al., JANAF Thermochemical Data (J. Phys. Chem. Ref. Data, Vol. 14, Suppl. 1, 1985). 19. A. Solheim, "The Density of Molten NaF-LiF-AlF3-CaF2A1203 in Aluminium Electrolysis", Aluminum Transactions 2 (1) (2000), p. 161. 20. A. Solheim and L.I.R. Stoen, "On the Composition of Solid Deposits Frozen out from Cryolitic Melts", Light Metals 1997. p. 325. 21. A. Solheim, "Crystallization of Cryolite and Alumina at the Metal-Bath Interface in Aluminium Reduction Cells", Light Metals 2002. p. 225. 22. J.W. Burgman and P.J. Sides, "Measurement of Effective Diffusivity in Hall/Heroult Electrolytes", Light Metals 1988. p. 673. 23. J. Gerlach, U. Hennig und H.-D. Pφthsch, "Zur Auflφsungskinetik von Aluminiumoxid in Kryolithschmelzen mit Zusätzen von A1203, A1F3, CaF2, LiF oder MgF2, Erzmetall, 31 496 (1978). 24. J. Thonstad, "Chronopotentiometry Measurements on Graphite Anodes in Cryolite-Alumina Melts", Electrochim. Acta, 14 127 (1969). 25. K.Δ. Rye, T. Eidet, and K. Torkiep, "Dynamic Ledge Response in Hall-Heroult Cells", Light Metals 1999. p. 347.

Appendix: Derivation of Heat Transfer Coefficient from Turbulent Thermal Diffusivity The heat flow in the boundary layer next to the sideledge can be expressed by q

= -pCp(ac+at)-y-

[wm- 2 ]

(Al)

where q is formally negative when heat isflowingfromthe bath to the sideledge. Assuming that the turbulent thermal diffusivity close to the sideledge varies as cct = C y 3

[m 2 s _1 j

(A2)

the above equations combine to give

dT =

5

J!>_

Where

β = ί^-1

(A3)

C.pC p ί3 + y 3 [c] The temperature profile near the ledge can be found by integrating Eq. (A3)fromy = 0 to a distance yfromthe ledge surface,

v.—f-ff#4V -L.tan-i fiZzg. + _ * _ [ V3 l V 3 ί J 6Λ/3[ For very large y, we obtain Tb-Ti = - ^ * 3V3 pC p a? / 3 .C 1 / 3

(A5)

and the heat transfer coefficient becomes h =

^ \ r

= 0.8270pCpac2/3C1/3

(A6)

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Towards a design tool for self-heated cells producing liquid metal by electrolysis Sophie Poizeau, Donald R. Sadoway Department of Material Science and Engineering, Massachusetts Institute of Technology, Cambridge, Massachusetts 02139-4307, USA Keywords: finite element modeling, Hall-Heroult cell, electrolysis, sideledge, turbulent flow, heat transfer coefficient duction by molten oxide electrolysis (MOE) [2], the wellAbstract studied Hall-Heroult process serves as an excellent startAs part of an effort to assess the technical feasibility of ing point. MOE has become the focus of attention for producing metals by molten salt electrolysis, a design tool its potential for carbon-free iron making as well as in situ is under development for the purposes of estimating the generation of oxygen from lunar regolith [3]. To design threshold cell size and current for self-heating operation. a self-heating cell, a tool is needed to optimize parameTo make the model broadly applicable to the production ters such as the size of the cell, the current density, the of different metals, two major issues must be addressed. interelectrode gap and the composition of the electrolyte. First, accurate values of the heat transfer coefficient are The first interest would be to obtain a temperature prorequired in order to model the position of the ledge. In file that provides a similar crust to what is achieved in the Hall-Heroult cell, the heat transfer coefficient is de- Hall-Heroult cells. To validate this design tool, we would termined experimentally from industrial operation, an ap- compare its solution for the bath temperature of a known proach that is not possible for a cell that has never been process, namely Hall-Heroult. built. Second, thorough treatment of transport phenomena in the cell involves solving the equations for liquid Existing models: operating tools adapted t o t h e thermal management of t h e Hall-Heroult cell and gas flows simultaneously; however, the methods used to model the turbulent flows in the Hall-Heroult cell are usually not well coupled. Introduction

As seen on Figure 1, the Hall-Heroult cell is composed of three main parts: the carbon anode at the top, the electrolytic bath in the middle and the aluminium pool at the bottom. The electrolyte contains the alumina dissolved into enriched cryolite. Because of the difference in density between the liquid aluminium and molten electrolyte, the liquids are self-segregated: the aluminium produced goes to the bottom of the cell. Carbon dioxide bubbles are produced at the electrolyte/anode interface. Frozen electrolyte forms a sideledge against the walls. There is an optimum thickness for the sideledge: too thin fails to protect the wall against chemical attack; too thick reduces useful electrolyte capacity. This equilibrium can be optimized through determination of the temperature profile. The electric current imposed across the cell has two main functions: producing aluminium but also keeping the liquids in the cell in the molten phase at 960°C. The temperature is not homogeneous throughout the cell: the temperature of the sidewalls varies from 100 to 300°C, the lowest temperatures being observed at the bottom and top of the cell, the highest temperature at the electrolyte level [1]. In the process of designing a new process for metal pro-

conduction

convection

(bubble evolution + magnetohydrodynamics Joule heating convection (magnetohydrodynamics) conduction

conduction ^ V · . radiation (+ convection)

Figure 1: heat transfers at steady state In this part, the authors will focus on the modeling work reported in the literature to evaluate to which extent it could be transposed to other chemistries. The heat generation (grid pattern on Figure 1), the heat flows inside the cell and finally the heat loss (dashed line) at the boundaries will be evaluated. The ledge position is closely

387

of the velocity [11]. Even if the authors do not seem to agree on the value and the dependence of this coefficient, their results match very well to the reality. Indeed, the authors were able to tune their functions. However, be1. Heat generation cause of this essential tuning, it makes those models nonJoule heating Heat is produced in the Hall-Heroult cell transferable to other chemistries for which the cell does through Joule heating. The two parameters that deter- not exist. Severo [13] used the law of the wall determined mine the importance of heat creation are the electrical experimentally by Kader [14] in the case of a temperaresistivity and the current density. The electrolyte is the ture field developed in a turbulent boundary layer along poorest conductor compared to the other cell components a smooth wall. This complex law has the advantage of where current flows (through the anodes and the cath- being transferable to other chemistries. However, it used ode). Indeed, the conductivity of the electrolyte is about the temperature of the wall and the temperature near the 200 S/m [4] in a pseudo steady state, while the conducti- wall, whose determination was left unexplained. vity of aluminium is 4 10 6 S/m [5] and the conductivity of The metal-ledge interface has not been studied as excarbon in the anode and cathode is about 10 4 — 10 5 S/m. tensively as the bath-ledge interface and is usually treated Therefore, the interelectrode gap determines the amount similarly to the bath-ledge interface, with a heat transof heat produced and must be kept constant during the fer coefficient. However, as Solheim mentioned [6], the operation of the cell to keep the temperature constant. sideledge is separated from the metal pool by a thin bath The interelectrode gap is usually 4 to 5 cm large to en- film that does not have the same composition as the bulk sure that the fluids are at 960°C. of the bath. Its temperature is therefore not fixed at a Sideledge formation The sideledge formation is impor- melting point temperature. In all the simulations pretant to be taken into account to know its position and sented [11; 9; 13], the heat flux determined to get the where the thermal properties of the bath are discontinu- position of the ledge is the product of the heat transfer ous. The sideledge present in the Hall-Heroult cell is cru- coefficient and the superheat temperature, defined as the cial to the good functioning of the process for two main temperature difference between the temperature of the reasons: the sideledge (1) protects the sidewall from the bath and the melting point of the bath. It is not clear combined action of high temperature, oxidizing cell gases, how the superheat can be used to describe an interface molten aluminium, and the fluoride-based bath and (2) that is not at a defined temperature. limits heat losses. The composition of the sideledge is 2. Heat distribution inside the cell not uniform from the top of the electrolyte to the botInside the cell, the heat is distributed everywhere by tom of the aluminium pool. Indeed, at the bath level, the conductive heat transfer, as well as convective heat transsideledge is due to the solidification of the bath, and the fer in the molten aluminium and cryolite. interface is defined by the melting point of the bath, while at the metal level a thin film of bath saturated with alu- Convection This is the dominant heat transfer mode in mina of the order of millimeters covers the sideledge [6]. the electrolyte and liquid aluminium. The heat motion in The mechanism of formation of the top and the bottom the case of the Hall-Heroult cell is due to: - natural convection via buoyancy force. ledge is completely different, even though heat transfer - forced convection, via gas bubble flow in the cryolite, coefficients are usually used in both cases. or magnetic forces in fluids that are electrically conducThe literature is prolific concerning heat transfer coeffitors (molten aluminium and cryolite). To calculate the cients for the bath-ledge interface. Chen [7] listed twelve velocity, the heat equation has therefore to be coupled different papers, giving heat transfer coefficients in the range of 116 to 1820 W / m 2 K . The main reason for this with the Navier-Stokes equation. The forced convection diversity is the dependence of the heat transfer coefficient contributions are developed below.

related to the consumption of additional heat to go from liquid to solid state (or vice versa), and will be treated with the heat generation.

on experimental conditions. For this reason, authors have developed experimental functions depending on those experimental conditions. Solheim [8] and Khoklov [9] recognized the impact of the gas flow rate, the immersion depth and the anode sideledge distance on the heat transfer coefficient. Nazeri [10] and Dupuis [11] used a function of the bulk fluid velocity in the immediate vicinity of the ledge, depending linearly [11; 12; 10] or on the square root

Gas bubbles Anode gas evolution in aluminium reduction cells is an important driver of electrolyte flow and alumina mixing. In the cryolite, the values of the velocity due to gas bubbles and of the velocity due to the magnetic forces [15] are similar, while the orientation is different. Therefore it is also necessary to calculate the impact of the bubbles for the sideledge determination.

388

In the literature, several methods are used to determine the velocity of the fluid. Solheim [16] determines the mean velocity of the bubbles in a water-air full scale prototype, fitting a dimensionless equation with the data. The result outlines the role of all the physical parameters, such as the distance from the liquid metal, the dynamic viscosity and the slope of the anode. These data however do not indicate the direction of the flow and therefore could not be used as base flow in the solution of the global velocity in a full computational model. The bubbles can be tracked individually. Johansen [17] undertook this first in the case of bubble stirred ladles. He treated the problem from a hydrodynamic view, considering two phases, the electrolyte as a primary phase, introducing the k-e method for turbulent flow in Fluent, and the bubbles as a secondary dispersed phase with an approximate drag coefficient for spherical bubbles. In this method, turbulent viscosity is expressed as a function of kinetic energy and the dissipation rate of the turbulence energy, two dependent variables adding two additional transport equations. The bubble drag term is characterized by a relative Reynolds number and a drag coefficient. Fräser [18] reused this method in 2D, with a slightly different drag force coefficient in the case of the Hall-Heroult cell, as well as Purdie [19] in 3D, who did not specify the drag force coefficient he used in Fluent. In those two last cases, the maximal velocity calculated was 44 and 33 cm/s respectively. Solheim [15] also used a two-phase approach to model the fluid flow, using the k-e turbulence model as well and a slightly different drag force term from Fräser and Johansen. For an inclination angle of 2° he calculated a maximum velocity of 5.8 cm/s away from the center of the anode, which is consistent with experimental data for Hall-Heroult cell, and a maximum velocity of 30 cm/s. In all those methods, only two phases were considered: the gas and the electrolyte. The liquid metal is not considered. Indeed, this Lagrangian tracking method is only suitable to two phases in the way the equations are defined. Magnetohydrodynamics Magnetohydrodynamics constitute the major cause of fluid flow in the molten aluminium since no bubble flow is active there. They are also a major contribution of the fluid flow in the cryolite. Due to the fact that the anode area is smaller than the cathode area, the current diverges in the cell and creates a high magnetic field. It is usually on the order of 1 0 - 2 T. Because the molten aluminium and the cryolite are both conducting fluids, the combination of the current and the magnetic field causes electromagnetic forces, also called Lorentz forces. Those

forces are responsible for the metal and bath flow, as well as for the metal-bath interface deformation. The typical flow velocity of the cell is 0.1 m / s while the interface is about 8 cm higher in the middle than on the sides of the cell [20]. Two main challenges need to be overcome to get a proper velocity profile inside the cell and be able to get a right convective heat flow in a finite element simulation. The first one is the determination of the moving interface between two fluids, while the second one is the treatment of turbulent flows. Simulation of t h e moving interface To model the moving interface, several methods are used as found in the literature. If a fixed mesh is used, the interface is the line where the volume fractions of both fluids are equal. This line does not go through any particular nodes of the mesh. It has a finite width, along which the properties change continuously from the cryolite to the aluminium phase. For instance, most software has the ability to use the so-called method of phase field. This method solves the Cahn-Hilliard equation and is usually used to model phase separation. Severo [20] used a similar method provided in Ansys CFX. Not only such methods are computationally expensive, since the mesh at the interface has to be very small, but Severo noticed that his results were not good compared to other methods if the air gap on the top on the cryolite is not included. But this increases of course the complexity of the problem, since two interfaces and three fluids have to be taken into account at the same time. If a moving mesh is used, the interface is located at the common boundary of two finite elements/volumes, one having the properties of cryolite, the other having the properties of aluminium. This method is usually used in combination with the shallow layer approximation, also called St Venant approach [21; 22; 20], which is justified in the case of the Hall-Heroult cell. Indeed, the vertical dimensions of the fluids are considered small compared to the horizontal dimensions of the cell and the interface wave amplitude is small compared to the depth. The St Venant approach neglects the vertical dependency for the velocity while approximating the pressure by the hydrostatic pressure [22]. It gives fast results consistent with experimental ones [20]. Treatment of t h e turbulence in aluminium and cryolite flows As seen above, some artifacts are usually necessary to be able to get a solution in modeling when turbulent flows are present, increasing the viscosity of the fluid to take

389

into account the effects of the turbulence. We saw in the case of gas bubble flow that the cryolite flow could be solved using k-e method. Severo tried to use this in the case of a three-phase (aluminium, cryolite, air on the top) model and it led to an underestimation of the displacement of the interface by 4 cm (50%). Indeed, the k-e method often shows poor agreement in the description of rotating flows dominated by body forces, which is the case here. Therefore, other methods are used to treat the turbulence of the flows in the case of the Hall-Heroult cell. The most common way to model turbulence is by assuming that it is purely diffusive. The turbulent viscosity is considered to be a constant and called the eddy viscosity. This viscosity is usually around 0.5-1 Pa.s in the case of the Hall-Heroult cell in both fluids [20]. This represents 1000 times more than their real viscosity (about 0.8 mPa.s for aluminium and 3 mPa.s for the bath). Severo [20] and Wahnsiedler [23] used this method. The turbulent viscosity is determined by tuning the model. The last method used is the k-ω turbulence method [24], in which the transport of the dissipation per unit turbulent energy is modeled instead of the dissipation itself, as it is the case in the k-e method. The results [21] are consistent with those from Severo [20] for the interface shape. The velocity profile is similar to those from other simulations [25; 23] with values around 10 cm/s. To conclude this part on magnetohydrodynamics, the floating grid seems to be preferable for models already complex enough, and the k-ω method to treat the turbulence would allow the user not to depend on measured quantities. However, the combination of magnetohydrodynamics with bubble flow is challenging, since the models for bubble flow are using Lagrangian tracking of bubbles, a gas-liquid phase model and the method to deal with the turbulence is k-e, that does not work well for magnetohydrodynamics [20].

Other n e e d s for other problem: design tool Experimental tuning In this article, the authors have pointed out several coefficients that are determined experimentally: the heat transfer coefficient for the ledge, the superheat (difference of temperature between the bath temperature and the melting point of cryolite), the drag coefficient due to bubble flow, increased viscosity of the fluids to model the turbulence of the flows. Those parameters, such as the heat transfer coefficients and the superheat necessary for the sideledge determination, depend on the configuration of the cell [26], and are therefore unknown for unbuilt cells with different chemistries. This shows the difference in the requirements to get a good design tool versus operating tool. Fully coupled model to get a temperature profile The authors have shown that to evaluate the convective heat transfers correctly, the most problematic features are the velocity (involving bubble and magnetohydrodynamic flow) and the determination of the ledge. Even if each part can be solved separately, the combination of those solutions may not give a reasonable answer. Indeed, the experimental constants used to solve those problems are usually experimentally determined while other phenomena are active, even if those phenomena are not modeled simultaneously. For instance, for most models of the ledge, no bubble flow is involved in the simulation [9], while the bubble flow (rate but also direction) has a large impact on the shape of the ledge [8]. This means that the bubble flow contribution is included in the heat transfer coefficient to get results consistent with experimental data [9]. Therefore, if the bubble flow would be added on top of this heat transfer coefficient, the contribution of the bubbles would be overestimated.

Moreover, the methods used to model one feature are not easily compatible with other methods. For instance, The heat losses in the Hall-Heroult cell are typically the determination of the velocity of the bath flow would distributed equally through the top and sides of the cell, need at least the treatment of four phases at the same and a small amount through the bottom (less than 10%). time: the carbon dioxide bubbles, the cryolite, the aluThe heat can be lost at the boundaries through convec- minum, the sideledge (which could be divided in two), tive or radiative heat transfer. Radiative heat transfer is and eventually the gas on top of the channels. This is a the main factor of dissipation of energy, since the emis- huge challenge for the existing software. Another examsivity of the sideblocks is about 0.8 and the tempera- ple is the treatment of the turbulence of those flows. The ture gradient is large. Haugland [1] calculated that, for bubble flow is usually solved using the k-e method, while a wall temperature of 350°C and an ambient temperature this does not work for the magnetohydrodynamics [20]. A of 30°C, radiative heat is equivalent to convective heat common way to solve turbulence in molten cryolite due with a transfer coefficient of 23 W / m 2 K , while the heat to both bubble flow and magnetohydrodynamics needs to transfer coefficient due to pure convection is 5 W / m 2 K . be found. 3. Heat losses

390

Other approach for the sideledge determination The sideledge at the level of the electrolyte is due to the solidification of the bath, and the interface is defined by the melting point of the bath. In the approach used in the literature reviewed, a heat transfer coefficient and/or the superheat are used as parameters to evaluate the heat needed by the ledge to melt and the position of the ledge. Also, two phenomena happen at the liquid-solid interface: dissipation of energy and zero velocity in the solid phase. While the dissipation of energy can be modeled in the current literature of the Hall-Heroult cell models by those heat transfer coefficients, this method alone does not treat the zero velocity in the ledge. Another approach that could allow the model to depend less on measured coefficients in the context of generalization to other chemistries would be to treat this problem as a phase transition from solid to liquid. Several methods are used to model phase transition, treating the two phenomena (latent heat effects and zero velocity) at the same time. To treat the energy dissipation/production at the interface, authors specialized in phase transition use mostly the enthalpy source method [27; 28], where the latent heat is incorporated in the heat equation as a source term, or the heat capacity method, where the latent heat is incorporated into the heat capacity [29]. To treat the zero velocity inside the solid phase, most authors assume a very large viscosity inside the solid phase [30; 29], or an artificial force that opposes the force applied to the fluid that provokes its movement [27; 28]. The usual drawback to those methods is that they require a really fine mesh at the interface, but they could be an interesting alternative to the empirical methods used in most articles reviewed.

transfers were varied but without much impact on the solution since the radiative heat transfers were dominant, which was consistent with Haugland [1]. In this model, the variation of the properties of the bath with temperature (to model the ledge vs. liquid electrolyte) was taken into account using a step function. The temperature of the liquid phases is obviously not correct, but the temperature of the solid phases is not too far from realistic temperatures. In particular, the outside temperatures are about 75-100°C as can be seen in Figure 2.

Figure 2: Steady state temperature with internal conductive heat transfers only Natural convection in liquid aluminium

Preliminary results A first model of the Hall-Heroult cell was built following the recommendations of the first parts of the article, using Comsol Multiphysics v3.5a. The goal of this preliminary model was to determine the limits of simple modeling to determine the temperature inside the electrolyte in pseudo-steady state using only measured properties of the separate components used to built a cell in the industry. A 2D model was built to simplify the situation. Internal heat transfers limited to conduction In this first model, emphasis was placed on the impact of the boundary conditions. The heat source was determined by Joule heating, while the sinks were determined taking into account radiative and convective heat transfers. The outside temperatures to solve the boundary conditions were 160°C for the top of the cell, 60°C on the sides and 30°C at the bottom. The convective heat

0

0.2

0.4

0.6

0.8

1

1.2

1.4

1.6

1.8

2

M n:

'

72.055

Figure 3: Steady state temperature with convection in aluminium As a next step, the natural convection in the liquid alu-

391

minium was added. It is interesting to notice that even if the dominant convective flows due to forced convection were not present, the temperature of the aluminium phase homogenizes quickly (about 100 s), reaching a steady state after 500,000 s. The temperature in the aluminium is about 880°C. There are too many missing contributions (particularly the convection in the electrolyte) to draw conclusions about temperature discrepancies. Limitations of this preliminary model The limitations of the preliminary model were recognized when couplings of fluid flows or of phase transition with convection were tried using Comsol. When a phase field method was used to model the liquid-liquid interface, the number of elements was too large to get a solution on a large scale while on a smaller scale the proportion of the volumes were not conserved. When the liquid-solid phase transition was coupled with convection, it was not possible to reproduce the simulation of the melting of noctadecane as published successfully in the literature [31] and known to be consistent with experiment [32]. Acknowledgements The authors thank Dr. Alton Tabereaux and Dr. Adam Powell for their advice in this work, supported by the American Iron and Steel Institute and NASA. References [1] Haugland, E., Borset, H., Gikling, H., and Hoie, H. (2003) In Light Metals : pp. 269 - 76. [2] Sadoway, D. R. (1995) Journal of Materials Research 10(3), 4 8 7 - 4 9 2 . [3] Sibille, L. (2009) In 47 th AIAA Aerospace Sciences Meeting : pp. Paper AIAA 2009-659. [4] Hives, J., Thonstad, J., Sterten, A., and Fellner, P. (1994) In Light Metals : pp. 187 - 194. [5] Brandt, R. and Neuer, G. (2007) International Journal of Thermophysics 28(5), 1429 - 1446. [6] Solheim, A. (2006) In Light Metals : pp. 439 - 43. [7] Chen, J. J. (1994) JOM46(11), 15 - 19. [8] Solheim, A. and Thonstad, J. (1984) Journal of Metals 36(3), 5 1 - 5 5 . [9] Khokhlov, V. A., Filatov, E. S., Solheim, A., and Thonstad, J. (1998) In Light Metals : pp. 501 - 506. [10] Nazeri, H., Utigard, T., and Desclaux, P. (1994) Light Metals pp. 543-559. [11] Dupuis, M. and Bojarevics, V. (2005) In Light Metals : pp. 449 - 54. [12] Dupuis, M. (2002) In CIM : Canadian Institute of Mining, Metallurgy and Petroleum. [13] Severo, D. S. and Gusberti, V. (2009) In Light Metals : pp. 557 - 562.

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Kader, B. (1981) Int. J. Heat Mass Transf (UK) 24(9), 1 5 4 1 - 4 . Solheim, A., Johansen, S., Rolseth, S., and Thonstad, J. (1989) J Appl Chem 19(5), 703 - 712. Solheim, A., Johansen, S., Rolseth, S., and Thonstad, J. (1989) In Light Metals : pp. 245 - 252. Johansen, S. and Boysan, F. (1988) Metallurgical transactions. B, Process metallurgy 19(5), 755 - 764. Fräser, K., Taylor, M., and Jenkin, A. (1990) In Light Metals : pp. 221 - 226. Purdie, J. (1993) In Light Metals : pp. 355 - 360. Severo, D. S. (2008) In Light Metals : pp. 413 - 418. Bojarevics, V. and Pericleous, K. (2009) In Light Metals : pp. 569 - 574. Gerbeau, J.-F. (2006) Mathematical methods for the magnetohydrodynamics of liquid metals, Oxford University Press, . Wahnsiedler, W. (1987) In Light Metals : pp. 26 287. Bojarevics, V. (2004) Metallurgical transactions. B, Process metallurgy 35(4), 785 - 803. Evans, J. (1981) Metall. Trans. B, Process Metall. (Ί7&4;ΐ2Â(2),353-60. Dupuis, M. and Haupin, W. (2003) In Light Metals : pp. 255 - 62. Brent, A., Voller, V., and Reid, K. (1988) Numer. Heat Transf. (USA) 13(3), 297 - 318. Voller, V. and Prakash, C. (1987) Int. J. Heat Mass Transf. (UK) 30(8), 1709 - 19. Morgan, K. (1981) Comput. Methods Appl. Mech. Eng. (Netherlands) 28(3), 275 - 84. Ma, Z. and Zhang, Y. (2006) International Journal of Numerical Methods for Heat and Fluid Flow 16(2), 204 - 225. Cao, Y. and Faghri, A. (1990/08/) Trans. ASME, J. Heat Transf (USA) 112(3), 812 - 16. Okada, M. (1983) In ASME-JSME Thermal Engineering Joint Conference Proceedings, volume 1, : pp. 281 - 288.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

HEAT RECOVERY FROM ALUMINIUM REDUCTION CELLS Yves Ladam \ Asbjorn Solheim2, Martin Segatz3, and Odd-Arne Lorentsen 4 1

2

SINTEF Energy; P.O. Box 4761 Sluppen, NO-7465 Trondheim, Norway SINTEF Materials and Chemistry; P.O. Box 4760 Sluppen, NO-7465 Trondheim, Norway 3 Hydro Primary Metal Technology; Koblenzer Str. 122, D-41468 Neuss, Germany 4 Hydro Primary Metal Technology; P.O. Box 1128, NO-3905 Porsgrunn, Norway Key words: Heat recovery; Electricity Production; Energy Consumption; Productivity

get rid of the heat by active cooling of the cell. On the other hand, lowering the energy consumption requires lower heat loss. This requires a better insulated cell which gives higher temperatures in the lining, and active cooling may be necessary to protect vital components. As will be substantiated in the present paper, heat recovery and utilisation of heat may be instrumental in enabling reduced energy consumption as well as further capacity creep.

Abstract About half of the energy spent in aluminium electrolysis is lost as heat. A preliminary study concerning the possibilities of recovering part ofthat heat was carried out, primarily focusing on electrical power production. The three main heat sources (cathode sides, anode yokes, and gas) were combined in different ways, using different types of power cycles. The potential for electric power production is significant (up to 9 percent of the total consumption). The two most promising families of power cycles appear to be 1) distributed open Brayton cycle based on a turbo charger and 2) centralised power production with a Rankine cycle. The temperature and amount of heat available in the anode match well with the heat from the sides, while the potential of integrating the flue gas is limited. The main aims in energy recovery may be increased productivity or reduced energy consumption, which gives different strategies for heat collection.

Even though there is no tradition for heat recovery and utilisation of heat from aluminium electrolysis cells, some patents have been filed. Newer patents comprise cooling by placing heat exchangers in the cell wall [1'2] as well as active cooling of the anode yokes l3]. The application of a heat exchanger in the dusty and contaminated raw gas from the cell has also been addressed [4]. Moreover, systems for driving the thermal medium through the heat exchanger [5] and utilisation of the heat [6] have been patented. However, a large amount of development work is required to realise a cell with significant heat recovery.

Introduction

Although all systems for heat recovery and utilisation will have some flexibility, the system must be designed to fulfil a primary purpose. The present work is based on a preliminary study where two extreme cases were considered; either aiming for maximum production, or aiming for minimum energy consumption. We focus on principles, challenges, and possibilities in heat collection from the cells, as well as utilisation of the recovered heat. Estimates concerning necessary infrastructure and cost are not part of the present work.

Since the invention of the Hall-Heroult process in 1886, there have been steady and long-lasting trends towards larger electrolysis cells, increased current efficiency, and lower specific energy consumption. Although the main principles in the process remain unchanged, it is fair to say that the difference between the first commercial cells in the 1890s and today's standard is comparable to the development of automobiles in the same period. However, during the last few decades, the trends towards higher current- and energy efficiencies have been gradually replaced by a trend towards higher productivity; in particular, by amperage increase in existing potlines (capacity creep). Besides reflecting the economical benefits by capacity creep, this also illustrates the increasing challenges in further reduction of the specific energy consumption. However, there is no reason to believe that the energy costs will become lower in the following decades. Therefore, it will become increasingly attractive to reduce the net amount of energy used for aluminium production by utilising the waste heat for electricity production or other purposes.

Methods and Equipment for Heat Utilisation Thermodvnamic Cycles Thermodynamic cycles can be used to convert heat into work. The cycle always comprises compression, heating (internally or externally), expansion, and heat rejection. This can take place in three or more steps. A thermodynamic cycle is categorised according to how the different steps are performed (adiabatic, isobaric, etc.), and usually named after its inventor. The most relevant cycles for utilisation of heat from aluminium cells are the Brayton cycle and the Rankine cycle. Principal sketches of these cycles are shown in Figure 1.

In any new cell designs, as well as when introducing modifications in existing electrolysis cells, the energy balance is of primary concern. This involves maintaining the protective layer of frozen bath at the sidelining. Increasing amperage inevitably leads to more ohmic heating, and sooner or later, one will face the scenario that the only way of achieving increased amperage is to

The Brayton cycle is a single phase gas cycle that is commonly used for high temperature applications (above 300-400 °C), like a gas turbine.

393

heat with the incoming cold working fluid. In this mode, the working fluid can attain higher temperature in the cell (constant mass flow), or the mass flow can be increased at constant outlet temperature from the cell. Both solutions give higher efficiency, as compared to a simple open loop configuration.

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The closed loop arrangement is illustrated in Figure 2c. In this case, the working fluid is not in contact with the cell. An intermediate closed loop recovers heat from the cell, and in this way, many working fluids can be considered. In this arrangement, the power cycle does not need to cool down the heat transfer fluid as much as possible, since the non-recovered heat is brought back to the cell. This allows for a better fit between the heat source and the power cycle. If the heat is to be transported a long distance, it may make sense to have an additional closed loop between the heat recovering loop (inert gas) and the working fluid. This intermediate loop can be based on thermal oil, which has high volumetric heat capacity, leading to small dimension piping and easy application of thermal insulation.

4 Entropy (b)

Figure 1. Schematic representation of two thermal cycles, a) - the basic Brayton cycle (dotted lines are isobars, cycle comprises gas only), b) - the basic Rankine cycle. 1 - compression, 2 - heating, 3 expansion, 4 - heat rejection, cp - critical point. The Rankine cycle implies a phase change. The dominating working fluid is water. However, for low temperature heat sources (below 300-400 °C), or for small installations (less than 3MWe) other media better adapted to the heat source can be used[7]. There is an intense research activity in this field, particularly on the choice of working fluid[8].

In general, the efficiency of electricity production by a thermal cycle increases with increasing temperature of the heat source. Still, high temperature operation is not always the most economical solution, since operation at high temperature requires expensive heat media, or it may lead to degradation of the medium.

Heat Recovery Strategies

Location of Power Cycle

Depending on the heat source (amount of heat, temperature), several strategies for heat recovery can be considered, as illustrated in Figure 2.

Distributed Plant. Each cell or group of cells can have their own small power plant. The advantage is short distance between the heat source and the turbine, which gives small heat losses. The main drawbacks are higher cost and operating complexity of the plant. Assuming that there is 200 kW heat available from each cell at a temperature of 600 °C, the best power cycles will have efficiencies in the range 35-40 percent, corresponding to a potential production of up to 80 kWe per cell. Finding efficient equipment at a reasonable price will be challenging for such a small size class, even if heat from a few cells is combined.

(a) Heat exchanger (recuperator)

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Centralised Plant. At the other extreme, the heat from all the cells could be collected to serve a centralised power plant. Typical plants are organised in such a way that the flue gas from, e.g., 100 cells is treated in one dry scrubber, and the collection of heat could function in a similar manner. The potential power production is in the order of 10 MWe, and efficient components can be found for this size. The downside of this solution could be the cost and heat loss from the long pipes necessary for transporting the heat medium.

(b) -*4Power generation)-} (c)

Figure 2. Some strategies for heat recovery, a - open loop, b - open loop with recuperator, c - closed loop.

Assumptions for Power Cycle Calculations Power cycle calculation and optimisation were performed using both ProII and Excel. For ProII, the equation of state was used with access to steam tables. For the Excel model, calculations were linked to the thermodynamic property calculator REFPROP8. The following assumptions were made,

In a simple open loop arrangement (Figure 2a), the working fluid is heated in the cell, expanded to produce power, and released to the atmosphere. Basically, only air can be used in this cycle, since production of steam from water as well as the use of flammable gases will be considered too risky in an aluminium plant. The temperature after the power generation is high, and it may be possible to apply a second power cycle to recover the waste heat (dual cycle).

• • • •

In the open loop arrangement with recuperation (Figure 2b), the working fluid at the outlet from the power generator exchanges

•

394

80 percent turbine efficiency (typical literature value [9). 70 percent pump efficiency (typical literature value [9]). 10 °C minimum temperature difference in the heat exchangers. 5 °C minimum temperature difference in internal heat exchangers (recuperators). The lowest achievable condensation temperature is 20 °C.

• •

Table I. Energy balance for three cases. In the "minimum energy" and "maximum production" cases, the cell resistance was assumed to be lower than in the base case.

In calculations concerning heat recovery from the flue gas it was assumed that the lowest permissible temperature is 50 °C, due to the risk of surpassing the sulphuric acid dew point. Pressure drop in and heat loss from the piping were not accounted for.

Base case 300 4.150 94 13.16 6000 23.4 13.41

Heat Collection. Implications in the Electrolysis Cells

Amperage [kA] Cell voltage [V] Current efficiency [%] EC, cell only [kWh/kgAI] Air draught [ N m V ] Effect in fans [kW/cell] EC, including fans [kWh/kgAI] *) Heat collectedfromthe yoke [kW] Heat collectedfromthe cathode [kW] 195 140.3 Flue gas temperature [°C] 296 Yoke temperature [°C] Cathode heat exchanger temperature 600 *) Neglected pressure drop in heat exchanger

A spreadsheet-based model was used to predict the implications of heat recovery on the electrolysis process itself. Although simplified, the model takes into consideration all known couplings in the aluminium electrolysis process, and the model also comprises an integrated model of the cell superstructure, gas ducts, dry scrubber, and fans. Earlier tests have proved that the predictions in the model generally are in line with measurements as well as with Hydro's other models. All calculations shown in the present paper relate to a hypothetic modern 300 kA cell with 30 prebaked anodes. In the "base case", the cell voltage was assumed to be 4.15 V and the current efficiency was 94 percent, which gives a specific energy consumption of 13.16 kWh/kg AI.

Minimum Maximum energy production 300 340.5 3.794 4.150 94 94 12.03 13.16 2000 2000 1.2 1.2 12.04 13.17 90 165 140 290 162.8 168.2 440 146 600 300

temperature and chemical conditions, but they are permeable to sodium andfluorine-containinggases. Furthermore, ceramics are brittle, and the thermal shock resistance may be too low. On the other hand, it may prove difficult to find a metallic material that is chemically stable in such a harsh environment.

Calculations were also made for two additional cases, aiming at minimising the specific energy consumption and aiming for increased production, respectively. In the former case, the outlet from the heat exchangers should have high temperature in order to ensure high power cycle efficiency. Not too much heat should be extracted from the cells however, since reduction of the specific energy consumption will be accompanied by reduced heat loss from the cells (unless the amperage is increased at the same time). On the other hand, when aiming for increased production, the main reason for having heat exchangers is to remove as much heat as possible from the interior of the cells, but the utilisation of the heat in a power cycle or for other applications may be less important.

Maximum Production. When aiming for maximum production, the purpose with heat collection from the cathode is to obtain the highest possible heat flow through the side. This means that the sideledge should be thin (due to its relatively low thermal conductivity) and that the superheat should be high, leading to high heat flow through the ledge. A thin sideledge is also beneficial when introducing longer anodes for further amperage increase. In practice, however, it might be difficult to operate the cell with thinner sideledge than 5-10 cm, due to the trench formed at the metal-bath interface [10].

It was assumed that measures were taken to decrease the electrical resistance in the cell, allowing either decreased cell voltage or increased amperage. The main data for the three cases is given in Table I, and the reasoning behind the differences will be discussed in the following.

High heat flow and a correspondingly steep temperature gradient in the sideledge may imply that the temperature at the heat exchanger will be low. This makes it difficult to utilise the heat efficiently. On the other hand, it will be easier to overcome the material challenges. By keeping the temperature low, it is possible that the vapour pressure of fluorides (NaAlF4) will be acceptable low close to the heat exchanger. However, it has been found that the metallic potshell in aluminium cells may corrode due to the presence of fluorides [11], and it appears that systematic studies concerning the corrosion behaviour of different metals in atmospheres containing NaAlF4 have not been carried out.

Heat Collection from the Cathode The cathode sides, including the cathode current collector bars, represent about 40 percent of the heat loss from an aluminium cell. The available temperature may be quite high; this can be exemplified by the potshell temperature, which is above 300 °C and sometimes close to 400 °C. Minimum Energy Consumption. Cell operation relies on having a sufficiently thick layer of frozen electrolyte at the inner lining (sideledge). In order to reduce the heat flow, the cell must be operated at low superheat. In a modified cell with heat exchangers, the best technical solution seems to be a "hot" heat exchanger placed inside the cell close to the sideledge. By placing thermal insulation materials between the heat exchanger and the potshell, more heat will be collected. Moreover, the potshell will be protected against too high temperature. By having low superheat (difference between bath temperature and liquidus temperature) and a "hot" internal heat exchanger, it will still be possible to have a relatively small heat flux through the sideledge.

The temperature difference across the sidelining is quite small, due to the high thermal conductivity of Si3N4-bonded SiC. Therefore, it is possible that almost the same amount of heat can be collected by placing the heat exchanger directly at the outside of the potshell, as compared with location inside the lining. The temperature will be limited by the permitted maximum temperature of the potshell, however. This is not necessarily a problem when the goal is increased production. By having the heat exchangers at the outside of the potshell, they will be available for inspection and repair or replacement during normal operation, which is impossible with heat exchangers on the inside. To summarise; aiming for minimum specific energy consumption requires that the heat is collected at high temperature, in order to utilise it efficiently for power production. The heat flow out from

The main challenges with this concept will probably be related to the choice of materials. Ceramics can probably stand up to the

395

the cell should be limited, and it is necessary to operate the cell at low superheat. When aiming for maximum production, the most important issue is to get rid of the extra heat, which requires high superheat and thin sideledge, and the temperature in the heat exchanger will be lower. The difference between these extreme cases is illustrated in Figure 3. Clearly, one cannot obtain the best of both worlds at the same time.

(or trimetal) plate between the anode rod and the yoke. However, by combining increased thermal insulation with cooling of the yoke, it is possible to achieve lower heat loss as well as acceptable low yoke temperature ("cooling for protection"); see Figure 4a. In the "minimum energy" column in Table I, it was assumed that totally 90 kW is collected from the yokes at a temperature of approximately 440 °C. 1000

250 200%

Insulated yoke and anode cover material 200

400 600 800 1000 Temperature in heat exchanger / °C

50 100 150 Heat removed from yokes / kW (a)

Figure 3. Relationship between superheat, sideledge thickness, and temperature in the heat exchanger. Thermal conductivity of ledge: 1.5 Wm^K"1, bath-ledge heat transfer coefficient: 800 Wm^K"1, ledge surface temperature: 950 °C. The figure is based on the simplistic assumption of one-dimensional heat flux.

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Heat Collection from the Anode Tests with an air-cooled anode yoke was reported earlier [12]. The testing was performed with different pressure of compressed air and different air flow. The temperature of the air at the outlet from the yoke and the amount of heat removed were measured. It was found that by placing thermal insulation at the yoke/stubs and above part of the anode cover material, relatively large amounts of heat could be collected from the yoke at a reasonable high temperature (e.g., 4.3 kW per yoke at 320 °C). A simple heat flux model for the anode was made, based on the measurements. This model was used in the construction of Figure 4, which shows the heat flux from the bath through the anode and the yoke temperature as a function of the amount of heat removed from the yoke.

Non-insulated yoke and anode cover material 50 100 150 200 Heat removed from yokes / kW (b)

250

Figure 4. Yoke temperature and heat flux from the bath through the anode as a function of the amount of heat removed from the anode yokes (cell with 30 anodes), a insulated anode yokes and anode cover material (in the test, the yoke and most of the anode were covered by 50 mm rock wool mats [12]), b - non-insulated anode yokes and anode cover material, The dotted line represents the heat flux in the normal situation without heat collection from the yokes.

The main challenges in cooling the anode yoke, at least when the collected heat is utilised, are related to connection and disconnection of the anodes during anode change, avoiding leakage of the heat medium, and making constructions at the anode stem that survive through the rodding shop without substantial repair and maintenance. Minimum Energy Consumption. Although the thickness of the sideledge, and thereby the heat flux through the sides in principle can be varied, it seems likely that low energy operation also will be based on reducing the heat loss from the anode. This can be obtained by increasing the thermal insulation on top of the anode, e.g., by increasing the thickness of the anode cover material. This may give problems related to too high temperature at the bimetal

Maximum Production. The heat loss from the anode can be increased by reducing the thickness of the anode cover. This may be risky however, since increased anode airburn may occur. It is known to be difficult to apply the anode cover material uniformly, especially at neighbouring anodes with very different age, i.e., different height.

396

The heat flow through the anode can also be increased by cooling the anode yoke, provided that the yoke and stubs are not thermally insulated, as shown in Figure 4b. As can be observed, although much heat can be collected from the yoke, the yoke temperature may become too low to utilise the heat. In the "maximum production" column in Table I, it was assumed that 165 kW is removed from the yokes, producing a yoke temperature of 146 °C.

The effect of reduced draught is shown in Figure 6. The fan power is proportional to the draught raised to a power of approximately 3. The main challenges with the application of DPS seem to be related to the high temperature. It may, therefore, be necessary to take measures for thermal protection of DPS units themselves as well as other equipment at the superstructure, such as electronics and pneumatics. It should be mentioned here than during an experiment at the Reference Centre in Δrdal, 3-4 percent C0 2 was reached for a short period [14]. Recent experiments have demonstrated that cells equipped with DPS can be operated with 1.5-2 percent C0 2 in the flue gas for weeks without encountering problems, which illustrates that this is a promising technology.

The largest potential of cooling the anode might well be the possible reduction of the voltage drop by introducing new materials in the yoke. The current conducting parts of the yoke can be made of copper (even aluminium could be considered). The bimetal plate can be placed between the yokes and the stubs. The stubs can be made much shorter, and it might be possible to have a copper core in the stubs.

200

If the heat taken out from the yokes is not utilised for production of electricity, removal of heat by using a thermosiphon [13] may be possible. This device can move heat from the yoke to the top of the anode stem by circulation of C0 2 in a closed circuit (flow driven by temperature-induced density difference, i.e., no moving parts). It may also be possible to get rid of the heat by conducting it into the anode beam (which needs to be cooled to retain its structural strength and protect electronics and pneumatics in the superstructure). In this way, the previously mentioned problem with connecting and disconnecting the anode from an external cooling circuit will be eliminated.

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The flue gas (process gas and excess air) represents the far largest mass flow in the aluminium production, up to 100 tons of gas per ton aluminium. Typically, the gas temperature is in the range 120160 °C. Heat collection from the flue gas is probably the energy recovery concept that has come closest to industrial implementation. There is only a weak coupling between the gas temperature above the crust and the heat loss from the top of the cell, and therefore, heat collection from the flue gas does not influence much on the cell performance or the choice of heat collection equipment. However, it is important to remember that heat collection from the anode gives less radiation from the anode yokes and stems, and thereby, less heat in the flue gas.

1

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Figure 5. Generalised figure showing the amount of heat that can be collected from the flue gas a function of the gas flow per cell by cooling from the flue gas temperatures given in the figure (straight lines) and down to 80 °C. Upper curve - standard operation, lower curve - 180 kW removedfromthe anode yokes. 80

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Figure 5 illustrates the relationship between gas draught, temperature, and amount of heat available for collection. As can be derived from the figure, collection of 180 kW from the anodes gives only about 100 kW reduction in the heat available in the flue gas. It should also be noted that, for the purpose of heat recovery, it will be beneficial to increase the gas temperature by reducing the draught. Reduced draught also leads to reduced fan power. It may also enable future C0 2 capture, since the gas will contain more C0 2 than the present 1 percent.

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A more concentrated gas suited for future C0 2 capture. Higher flue gas temperature; easier utilisation of the heat for power production or for C0 2 capture. Substantial savings in the fan power consumption.

397

Power Production from Recovered Heat

Concluding Remarks

The performance of a large number of thermal cycles, layouts, and combinations of heat sources were studied, but only some of the most promising candidates will be mentioned here.

A heat recovery system for an aluminium plant will only be built when it is considered economically viable with a minimum of process disturbance. The complexity and the cost of a heat transporting system (infrastructure) must be carefully evaluated. The economical attractiveness could be increased by introducing other ways of utilising the heat besides electric power production, such as production of cold, drying, and desalination. Customers for the recovered heat need to be on the spot.

The cathode (cell sides and ends) is the heat source with the largest potential. Two families of power cycles appeared to be particularly interesting, •

Distributed open Brayton cycle, based on a turbo charger.

•

Closed loop heat recovery and centralised power production with a Rankine cycle. For high temperature application (e.g., 600 °C), a steam cycle with recuperation is the natural choice. For medium temperatures (e.g., 300 °C), a steam cycle with a condensation turbine may be interesting. However, good results can also be obtained with an organic Rankine cycle (ORC) based on toluene with a butane bottoming cycle (the condensation heat of toluene is recovered in a butane power cycle, which increases the total power production).

Acknowledgement The present work was mainly financed by Hydro Primary Metal Technology and partly through the project CREATIV, financed by the Research Council of Norway and Norwegian industry. Permission to publish the results is gratefully acknowledged. References 1. O.-J. Siljan, "Electrolysis Cell and Structural Elements to be Used Therein", WO/2004/083489. 2. J.A. Aune, K. Johansen, and P.O Nos, "Electrolytic Cell for the Production of Aluminium and a Method for Maintaining a Crust on a Sidewall and Recovering Electricity", WO/2001/94667. 3. B.P. Moxnes and A. Solheim, "Method and Means for Control of Heat Balance", WO/2006/088375. 4. E. Naess, T. Slungaard, O. Sonju, and B.P. Moxnes, "A Method and Equipment for Heat Recovery", WO/2006/009459. 5. H.K. Holmen and S. Gjorven, "A Method and a System for Energy Recovery and/or Cooling", WO/2006/031123. 6. J.A. Aune and P.O. Nos, "Method for Controlling the Temperature of Components in High Temperature Reactors", WO/2002/39043. 7. A.W. Crook, "Profiting from Low Grade Heat", IEE Energy Series 7, IBSN 0 85296 8355. 8. U. Drescher and D. Brüggemann, "Fluid Selection for the Organic Rankine Cycle (ORC) in Biomass Power and Heat Plants", Applied Thermal Engineering 27 (2007), pp. 223/28. 9. Y. Chen, P. Lundqvist, A. Johansson, and P. Platell, "A Comparative Study of the Carbon Dioxide Transcritical Power Cycle Compared with an Organic Rankine Cycle with R123 as Working Fluid in Waste Heat Recovery", Applied Thermal Engineering 26 (17-18) (2007), pp. 2142/47. 10. A. Solheim, H. Gudbrandsen, and S. Rolseth, "Sideledge in Aluminium Cells: The Trench at the Metal-Bath Boundary", Light Metals 2009. pp. 411/16. 11. V. V. Sharapova, "Study of the Interconnection of Aluminium Electrolyzer Life with Corrosion of its Casing" Translated from Novye Ogneupory, No. 1, pp. 39/42, January 2009, http ://www. springerlink. com/content/nx83 8151 v 13 60721 / 12. A. Solheim, B. P. Moxnes, K. Vamraak, and E. Haugland, "Energy Recovery and Amperage Increase in Aluminium Cells by Active Cooling of the Anode Yokes", Light Metals 2009, pp. 1091/96. 13. S. Gjorven, Y. Ladam, B.P. Moxnes, P. Neksä, and A. Solheim, "Method and Means for Extracting Heat from Aluminium Electrolysis Cells", WO/2010/050823. 14. O.-A. Lorentsen, A. Dyroy, and M. Karlsen, "Handling C02eq from an Aluminum Electrolysis Cell", Light Metals 2009. pp. 263/68.

No great potential was found by integration of heat from the cathode and heat from the flue gas. It is probably better to utilise the flue gas heat in a separate plant. On the other hand, heat from the anodes and heat from the cathodes could probably be integrated easily in some cases; it is always easier to integrate heat sources that are comparable in amount of heat and temperature. For the flue gas, it seems most reasonable to collect the heat in front of the dry scrubber. Power production should be based on an ORC cycle. A summary of the calculations made for the cases described in Table I is shown in Table II. As can be observed, the potential is up to about 9 percent of the total DC power consumption. Maybe surprising, the potential for power production from the flue gas (0.04-0.1 kWh/kgAl) is smaller than the savings in fan power following a draught reduction from 6000 Nm3 to 2000 Nm3 (0.24 kWh/kgAl). In addition, the higher gas temperature achieved with reduced draught opens for higher power cycle efficiency. The use of distributed pot suction may therefore have a great potential, also when C0 2 capture is not the primary focus. Table II. Upper part of table: Summary of possible power cycles and potential for energy utilisation in kWh/kgAl for the three cases described in Table I. R Rankine cycle, B - Brayton cycle, o - open loop heat recovery, c - closed loop heat recovery, M - medium temperature, H - high temperature. Lower part of table: Specific energy consumption in kWh/kgAl. Heat source Flue gas Cathode

Cathode + anode

Cycle RcM RcM RcH BoH RcM RcH+RcM BoH

Specific energy consumption Including fans, no heat utilisation Best case, max. heat utilisation

Base case 0.10 0.45 0.77 0.17

Base case 13.47 12.60

Minimum Maximum energy production 0.04 0.04 0.47 0.33 0.55 0.12 0.77 0.86 0.19 Minimum Maximum energy production 12.04 13.17 11.15 12.66

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECTS OF COMPOSITION AND GRANULOMETRY ON THERMAL CONDUCTIVITY OF ANODE COVER MATERIALS Hasini Wijayaratne1, Margaret Hyland1, Mark Taylor1, Andreea Grama2, Tania Groutso1 Light Metals Research Centre, The University of Auckland, Private Bag 92019, New Zealand department of Chemical & Materials Engineering, The University of Auckland, Private Bag 92019, New Zealand 1

Keywords: Anode Cover, Thermal Conductivity, Granulometry, Composition thermal conductivity [2]. Dependence of thermal conductivity on the bulk density of consolidated anode cover crusts was also suggested by Richards [3]. A correlation between bulk density of anode cover crust and alumina content was observed by Liu et al. indicating that density and proportion of alumina are determining factors of thermal conductivity [4]. However, relatively limited studies have been conducted on crushed bath based granular cover and consolidated anode cover crust materials [1,5]. Apelt [6] and Shen [7] obtained thermal conductivity values for crushed bath based granular cover and alumina, which were an order of magnitude lower than those measured for crust materials by Hatem et al [2, 6, 7]. Thus, it is understood that the granular cover layer is the controlling factor which determines the top heat losses of the cell. Figure 1 illustrates the different factors that affect thermal conductivity of cover material. Granulometry, or more precisely, voidage, is a major factor, as demonstrated by Shen [7]. For example, for fine crushed bath with a voidage of 45%, thermal conductivity was measured to be 0.5W/m°C while for coarse crushed bath with a voidage of 35%, thermal conductivity was measured to be 0.7W/m°C [7]. However, the effect of voidage on thermal conductivity is not simple and is affected by other factors such as particle size that contributes to the voidage. Literature related to heat transfer in packed beds, suggest that relatively large particles at high temperatures can give rise to radiation effects [8]. Specifically, Schotte [9] observed radiation effects, for lmm particles above 400°C and for lOOum particles above 1500°C. This supports studies of Yagi and Kunii, where the radiation contribution to thermal conductivity was found to be as high as 80% for large particles measured at 840°C (which was the highest temperature measurement) [9, 10]. These imply the complexity of heat transfer through powder systems of varying particle sizes and indicate the interrelated nature of particle size, voidage and packing density as well as temperature, all of which contribute to thermal conductivity. This demonstrates the importance for modern smelters for achieving the right granulometry to avoid operational problems such as crust melting from beneath due to over insulating nature of poor cover material. An example of this is shown in Figure 2. However, many smelters are equipped with autogenous mills for bath processing and these are known to produce excessively fine bath material. This has led to smelters facing challenges in obtaining the correct granulometry necessary for optimum anode cover material. This paper aims to discuss the findings of a rigorous study conducted in the laboratory to understand the effects of granulometry and composition, (which are the two main control parameters used in anode cover design) on thermal conductivity of granular anode cover materials. This paper also aims to discuss practical implications of the findings for smelters in obtaining the ideal anode cover material and provides a basis for further studies.

Abstract Thermal conductivity of anode cover material is critical in determining cell top heat loss. It has been observed that thermal conductivity of cover material is strongly dependent on packing and particle size distribution. Granular material that is densely packed (lower voidage) has higher thermal conductivity. When two sizes of spherical particles are mixed at various size ratios, the theoretical voidage can be reduced from 0.4 to 0.2-0.3. This can be applied to a particle system of crushed bath and alumina that constitute cover material. Currently, many smelters produce cover material that is too fine with high voidages. This has the effect of lowering the thermal conductivity which can cause unnecessary operational problems within the cell. Additionally, the effect of cover composition on thermal conductivity is not obvious. This paper describes studies conducted in the laboratory to understand the effects of composition and voidage on thermal conductivity of cover material. Introduction Anode cover is a crucial part of an aluminium reduction cell because of its important role in maintaining overall heat balance, protecting the carbon anode from air burn, and controlling fluoride loss. Control of anode cover material properties such as composition, depth and granulometry provide means of optimising the top heat losses in the cell. Due to the increase in line current over the years in the reduction cell technology, heat generation inside the cells has increased. This has led to higher service temperatures of both the anode assembly and anode cover material which needs to be controlled through optimising anode cover [1]. The effective thermal conductivity of cover material determines its capacity to dissipate heat through the top of the cell. It is an important property that is useful in modeling the heat flows through the top of the cell, predicting the thermal and structural stability of the cover material. For example, highly insulating anode cover material with low thermal conductivity will over heat the cover, leading to partial melting and subsequent collapsing due to chiolite phase melting inside the cover. In contrast, extremely conductive material will dissipate increased amounts of heat from the top of the cell leading to operational problems such as low bath temperature, low alumina dissolution, increased side ledge thickness leading to overall heat imbalance in the cell. Anode cover in an aluminium reduction cell consists of two distinct layers; a top layer of granular anode cover material and a bottom layer consisting of a consolidated crust formed due to exposure to heat and fluoride fumes over time. Previous studies have focused on properties of mainly alumina based anode cover materials and in particular alumina based consolidated crusts. Hatem et al. measured thermal conductivity of various alumina based crusts and alumina powders and observed that there is a relationship between crust density and

399

|

1

Raw materials composition

The voidage of each sample was calculated using equation (1) with measured values of 'as poured' bulk density (pb) and true density (pt).

Cover thermal conductivity

Raw material particle size

Voidage ·

Cover Sintering

(i)

| 1

Fresh Alumina/ Crushed bam mix

Cryolite. Chiol its. Alumina Phase Transitions Voidage

Semi-quantitative phase composition of cover and crushed bath samples were determined by X-ray Power Diffraction.

1

| Temperature |

Figure 1: Factors affecting cover material thermal conductivity.

Thermal conductivity measurements of each material were obtained using an apparatus developed by Shen based on the Fourier's law of heat conduction [7]. According to Fourier's law, this method relies on establishing a temperature gradient across the sample of interest and measuring the heat flux required to maintain that gradient at steady state. As long as uni-dimensional heat transfer is ensured, this enables the calculation of thermal conductivity of the sample. For a system where one dimensional heat transfer occurs in the radial direction, heat transfer can be described as shown in equation (2):

, dT dr

(2)

Where; qr is the heat flux at steady state, ë is thermal conductivity of material and dT/dr is the temperature gradient across a known radial thickness of the sample. Figure 3 shows a schematic of the apparatus which is of a concentric cylindrical setup. It consists of an inner heating core with a smaller diameter (d) and an outer cylindrical heating shell of larger diameter (D). A radial temperature gradient of approximately 60°C was set up between the inner heating element (TH) and the outer heating element (TL) during experimental runs. The granular material to be measured was packed inside the cylindrical cavity between the two heating surfaces as shown in Figure 3. At steady state, measurements were recorded and results obtained were used in calculating the thermal conductivity of the material (ë) according to equation (3) which is the integrated form of the heat transfer equation indicated by equation (2).

Figure 2: Crust collapsed from beneath due to over insulating nature of fine cover [1]. Experimental Materials Industrial granular anode cover material and processed crushed bath from two different smelters, referred to as Smelter 1 and Smelter 2 respectively were used for experiments during this study. Methods Both anode cover and processed crushed bath bulk samples as received from the respective smelters were size fractionated into 'coarse', 'intermediate' and 'fine' size fractions. The size fractions were defined as coarse being above 4mm, intermediate being between l-4mm and fine being below 150μπι. The Smelter 1 bulk sample was more homogenously mixed with fine material whereas Smelter 2 crushed bath bulk sample was predominantly coarse material. The fraction between 150μπι and 1mm is recognized to be part of the intermediate size fraction, however has not been used in this particular set of experiments. This is because the objective of this study was to establish distinct size fractions, which would give distinctly different thermal conductivity values.

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thermal conductivity is reasonable and the variation between measurements is in the range of 0.1% to 2.0%. It is clear from this plot that all samples have very similar thermal conductivity, despite the differences in composition of these materials. Table I shows that the density and voidage values of the fine materials from both smelters lie in the range of 59 to 62% with alumina voidage being slightly higher. Since, all four samples have particle sizes below 150microns, radiation effects cannot be expected at the measurement temperature range, as suggested by Schotte [9]. Therefore, the corresponding voidages can be expected to be insulating. This corresponds well with the thermal conductivity data for these materials which fall almost on the same line in Figure 5.

(3)

It is to be noted that, all thermal conductivity data presented in this paper are plotted as relative (normalized to coarse crushed bath) and not as absolute data. Results and Discussion The purpose of this study was to determine the effects of particle size distribution, voidage and composition on thermal conductivity of granular material. Figure 4 shows the semiquantitative phase composition results obtained for both bulk and size fractionated samples from each smelter. Alumina composition is also indicated as a comparison for the fine materials. Samples labeled as 'cover material' was received already blended with alumina, whereas samples labeled 'crushed bath' was received unblended with alumina. All samples shown in Figure 4 were used 'as received', apart from sizefractionatingand no further blending of alumina was performed in the laboratory.

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Alumina Smelter 1 Interm Cover Material

In general, composition of crushed bath and anode cover varies with size fraction. For example, alumina is concentrated in the fine fraction in the cover material, which is to be expected. In contrast, coarse materials tend to have larger quantities of cryolite and chiolite phases and only trace levels of transition alumina (which were not quantifiable by X-ray diffraction).

Crushed Bath 1 Smelter 1 - Coarse Cover Material Smelter 1 - Bulk Cover Material Smelter 2 - Bulk

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Phase composition data plotted in Figure 4 indicates that the four fine materials have quite different compositions. Smelter 1 fine cover material has higher quantities of cryolite and transition alumina compared to Smelter 2 fine cover material. Smelter 2 fine crushed bath and fine cover material have similar quantities of chiolite, however varying amounts of corundum and other bath phases.

Intermediate Material (Between 1- 4mm)

Figure 5 shows the thermal conductivity data plotted for the four fine samples including multiple data points for alumina showing the reproducibility of multiple measurements. Reproducibility of

Figure 6 shows the thermal conductivities of both intermediate size samples, with data for Smelter 2 fine crushed bath also plotted for comparison. The thermal conductivities of the intermediate materials are notably higher than the fine materials,

The cover and crushed bath intermediate fractions differ in both composition and voidage. In particular, amounts of cryolite and corundum are significantly different when comparing the two samples. In addition, Smelter 1 sample has a lower voidage due to its relatively high packing density.

401

as might be expected on the basis of lower voidage of the intermediate materials. In addition, it is probable that radiation effects begin to make contributions to the thermal conductivity at higher temperatures at this intermediate particle size [9], as can be seen by the higher slopes of the thermal conductivity lines, compared to the fine material. The two intermediate samples have similar thermal conductivities; however, Smelter 1 cover material has a slightly higher thermal conductivity than Smelter 2 crushed bath sample. This can be expected because of the relatively lower voidage and higher bulk density of Smelter 1 sample.

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Figure 7: Relative thermal conductivity data obtained for Smelter 1 and 2 coarse materials together with Smelter 1 intermediated and fine cover data plotted for comparison. Bulk Material Initial phase compositional analysis performed on bulk materials before size fractionating also show differences in phase composition. This is to be expected as Smelter 1 bulk sample is a cover material which includes blended smelter grade alumina while Smelter 2 bulk sample consists of pure crushed bath with only trace levels of transition alumina. Bulk density and voidage values listed in Table I also indicate clear differences between the samples, with Smelter 1 sample having a lower voidage. This is to be expected as Smelter 1 sample included noticeably more fine material homogenously mixed with coarse material which contributes to the higher packing density. This relates to the two particle packing theory which indicates how voidage can be significantly reduced by mixing two particles of different sizes in specific size ratios as shown by Figure 8 [5, 11].

Figure 6: Relative thermal conductivity data obtained for Smelter 1 and 2 intermediate materials as well as Smelter 2 fine crushed bath plotted for comparison. Coarse Material (Above 4mm) The coarse fraction of both smelter samples have notable differences in phase composition as can be seen from Figure 4. Smelter 2 coarse crushed bath sample has a larger sum of bath phases and less corundum compared to Smelter 1 coarse sample. According to Table I, bulk density and voidage values for the two samples are also different as the calculated uncertainty based on instrumental uncertainty for each voidage value is ±0.5%. However, it is interesting to note that the thermal conductivity values plotted in Figure 7 for these samples are nearly identical as can be seen by the overlapping data points. Again, the coarse materials have higher thermal conductivities compared to both intermediate and fine materials (plotted with dashed lines for comparison). Similarly, the rate of increase of thermal conductivity of the coarse materials with temperature is also higher than both intermediate and fine materials. This phenomenon observed even with relatively high voidages (refer coarse material voidages in Table I) suggest radiation effects contributing to thermal conductivity as demonstrated by previous workers studying heat transfer through packed beds of large particles [8-10].

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402

Table II: Density and voidage measurements obtained for mixed samples.

Bulk Density (g/cm3) True Density (g/cm3) Voidage (%)

Sample

Figure 9: Relative thermal conductivity data obtained for Smelter 1 and 2 bulk materials. Although each bulk sample has a different composition and granulometry, it is interesting to note that both samples have very similar thermal conductivities up to approximately 320°C. Above this temperature, Smelter 2 sample thermal conductivity begins to increase with respect to Smelter 1 sample. This behavior of Smelter 2 sample is very interesting as having a higher voidage compared to Smelter 1 sample; it is expected to be lower in conductivity. However, as can be seen from Figure 8, higher voidages can be expected even when the coarse particles are in high proportions (above 0.9 volume fraction) and this effect was also observed with the coarse samples. Kunii and Smith suggest that radiation effects can become significant in packed beds of relatively large particles above 900°F (approx 482°C) [8]. Shen observed the same affect (above 400°C) in particular with samples having voidages greater than 40% [7]. Therefore, it is possible that heat transfer within the thermal conductivity measuring apparatus (which is essentially a packed bed of granular particles) is affected by radiation at higher temperatures when particles are relatively large which contributes to high voidage (as in the case with Smelter 2 bulk sample which consists largely of coarse material) resulting in a corresponding increased thermal conductivity.

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ΔSmelter2-40%CCB+60%FCB (<200micron)

XSmelter2-40%CCB+60%FCB(45-150microns)

Figure 10: Relative thermal conductivity data for the mixed crushed bath samples.

Mixed Samples using Smelter 2 Crushed bath and Alumina

Data plotted in Figure 10 show interesting features. When a mixed sample has coarse bath in a proportion of 60% blended with 40% alumina, the thermal conductivity of the sample is higher than for a sample blend of 40% coarse crushed bath and 60%alumina. This can be expected due to the relatively higher packing density of the 60%coarse bath and 40%alumina sample, although both samples have the same voidage. This suggests that voidage alone does not determine thermal conductivity when the cover blend contains coarse material in the higher proportion as other factors such as radiation can contribute to thermal conductivity. In addition, it is clear from Figure 10 that the thermal conductivity of 40%coarse bath and 60%fine bath (45-150microns) is nearly identical to that of the 40%coarse bath and 60% alumina blend. However, thermal conductivity of the sample with ultrafine bath (<200micron sample) is comparatively slightly lower, as can be expected from its corresponding higher voidage. This suggests that when cover blends contain fine material in the larger proportion, voidage can predominantly determine thermal conductivity as other factors such as radiation are not as important. These results clearly illustrate that in a sample consisting of a blend of coarse and fine particles, thermal conductivity is

In order to understand the effect of cover blend composition on thermal conductivity, coarse crushed bath from Smelter 2 was blended with comparable proportions of alumina and fine crushed bath material from the same smelter. The blend compositions and corresponding density and voidage data are listed in Table II. The corresponding thermal conductivity data for these samples are plotted in Figure 10. It is to be noted that the sample labelled with adjacent notation (<200micron) is a sample that contain large quantities of fine material (in particular -45microns particles) in the fine portion. In order to see the effect of removing this ultrafine portion from the fines, the same composition blend was prepared by removing the 45micron fraction. This is distinguished by the label (45150micron) adjacent to the sample name (See Table II and Figure 10).

403

determined by granulometry rather than by composition. This means that in a blend of cover material, coarse crushed bath can be blended with either alumina or fine bath as long as each of these fine material have the same particle size distribution. In either case, thermal properties will be very similar. However, the relative proportions of coarse to fine material blended in the cover will determine to what extent, the corresponding voidages created within the material will aid heat transfer. This suggests that particle size has a more pronounced effect on thermal conductivity than composition of granular anode cover materials. However, as the granular cover sinters due to exposure to heat and fumes within the cell over time, phase transformations occur dynamically changing the structure of the consolidating crust. In addition, when temperatures exceed 600°C, cover components such as chiolite phase may melt and as a result properties of the material would radically change.

of granular cover materials. It is clear that, granulometry or particle size distribution has a more significant effect on the effective thermal conductivity of cover mixtures than composition. In general, higher proportions of coarse bath in a cover mixture increase the effective thermal conductivity whereas a higher proportion of fine material (either alumina or bath fines) lowers the effective thermal conductivity. Controlling the granulometry of cover material, in particular the fines generated is found to be critical as it contributes to lowering the overall thermal conductivity of the cover. This has very important implications for smelters in that the amount of alumina blended with cover material needs to be controlled (as it contributes to the proportion of fine material), especially in the case of crushed bath from autogenous mills which produce excessive fines. These findings provide a firm basis for further studies such as, determining the critical amount of alumina required for structural strength in cover and evolution of thermal properties of consolidated crushed bath based crusts over time.

Implications for Smelters The trends observed with the laboratory studies conducted, have relevant implications for smelters. More thermally conductive cover material can be prepared using coarse bath in a higher proportion in the cover mixture, with a lower proportion of fine material to give higher packing density and reduced voidage. This supports a theoretically based conclusion reached by the authors earlier [5]. In this case the fine fraction can be made up of either alumina or fine bath. Increasing the proportion of fine material in the blend will contribute to lowering the thermal conductivity and ultrafme fine bath material aids further deterioration of thermal conductivity. These findings are significant for smelters as this indicates the importance of controlling granulometry of cover material and in particular the generation of fines. Uncontrolled fine fraction in cover material will have the effect of decreasing thermal conductivity, especially when more fine material is added in the form of alumina without consideration of how much fines is already present in the base material. However, there is likely to be a 'critical' proportion of smelter grade alumina that is necessary in cover material, in order to ensure sufficient structural strength in the resulting consolidated crust. Determining the ideal proportion of alumina to be blended with cover was outside the scope of the current study since this requires study of the higher temperature behavior of the crust material over time. Separate studies are needed for understanding this issue and the findings of this paper will be a firm basis for further studies in this regard. Although the loose cover material layer is the controlling factor that determines immediate heat losses from the cell, it is also important to understand the thermal properties of consolidated crusts over a longer period of time; in particular modern day crushed bath based crusts which have not been widely studied. The key to understanding the thermal conductivity of consolidated crusts lies in understanding the evolution of voidage and phase transformations during the consolidation process, which need to be further investigated. These will contribute to the ongoing development of the understanding necessary for anode cover material design for modern smelters.

Acknowledgements We would like to acknowledge the contribution of Rio Tinto Alcan in supplying industrial materials for carrying out laboratory experiments for the purpose of this study. References 1. Taylor, M.P., et al. "The Impact of Anode Cover Control and Anode Assembly Design on Reduction Cell Performance", in Light Metals: TMS (2004), 199-206. 2. Hatem, G., et al. "Thermal Conductivity of Some Alumina Powders and Synthetic Hall-Heroult Crusts", in Light Metals : TMS (1989), 365-370. 3. Richards, N.E. "Anode Covering Practices", in 6th Australasian Aluminium Smelting Workshop. (1998). 4. Liu, X., M.P. Taylor, and S.F. George. "Crust Formation and Deterioration in Industrial Cells", in Light Metals:TMS (1991), 489-494. 5. Taylor, M.P. "Anode Cover Material - Science, Practice and Future Needs", in 9th Australasian Aluminium Smelting Technology Conference. (2007). 6. Apelt, T., "Thermal Conductivity of Granular Anode Cover Materials from Industrial Reduction Cells",(BE Undergraduate Thesis, Dept. of Chemical Engineering, University of Queensland, 1992) 7. Shen, X., "Top Cover and Energy Balance in Hall Heroult Cells". (PhD Thesis, Dept of Chemical & Materials Engineering, University of Auckland, 2006). 8. Kunii, D. and J.M. Smith, "Heat Transfer Characteristics of Porous Rocks". AlChe Journal 6 (1) (1960), 71-78. 9. Schotte, W., "Thermal Conductivity of Packed Beds". AlChe Journal, 6 (1) (1960), 63-67. 10. Yagi, S. and D. Kunii, "Studies on Effective Thermal Conductivities in Packed Beds". AlChe Journal, 3 (3) (1957), 373-381. 11. Yu, A.B., R.P. Zou, and N. Standish, "Modifying the Linear Packing Model for Predicting the Porosity of Nonspherical Particle Mixtures". Industrial & Engineering Chemistry Research, 35 (1996), 3730-3741.

Conclusions Laboratory studies have been conducted in order to understand the effects of granulometry and composition on thermal conductivity

404

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Restart of 300kA Potlines after 5 Hours Power Failure Xinliang ZHAO1, Jitai YAN1, Bingliang GAO2, Hua HAN1, Jie LIU1, Jiuyi XIAO1, Jianxun QIAN1 Delai WANG1 1. Keao Aluminium Company, Zoucheng, China 273515 2. Northeastern University, Shenyang, China 110004 Keywords: restart, 300kA potlines, power failure, aluminum reduction cell metal and bath to the maximum extent possible, reducing anode changing schedule and other related activities, prior to switching off the lines.

Abstract A Chinese reduction plant has an installed capacity of 140 kt of metal per year and employs 180 pots with line current of 300kA. In May, 2008, power failure happened. The situation lasted 5 hours. The bath temperatures fell to 900°C below. Power was switched on after recovering the power system. At the initial stage of restart the anode effect frequency was high. The voltage and amperage of lines fluctuated severely, which caused higher anode effects. Rectiformer failure would have occurred in such a situation. We finally solved the problems and led the lines to the normal status. The paper discusses the strategies adopted; restart operations and the technological parameter normalization during restart.

Preliminary work Fortunately, the plant has an emergency power plant, which can supply power for automation system and lighting system. After blackout accidentally, the operations are largely manual with very little automation. Some preliminary works were carried out: • Maintenance of cranes, electric panels, alumina feeding system, anode adjusting system and automation system; • Arrangements for adequate quantities of pot isolation materials like shunt plates, clamps and accessories were also made; • Raw materials like cryolite, aluminum fluoride; • Communication check between potrooms and automation station. • 100 piece of anode blocks

Introduction The Hall-Heroult process for the production of aluminum based on the electrolysis of alumina dissolved in cryolite, using carbon anodes and cathodes, was introduced in 1886 and is still the most important process for the production of the metal. The process has a strict requirement for power system. For a potline with current input of 160 kA and over, it is difficult to drive the cells to normal status if some kinds of power failures occur. These failures include power blackout or power limit. If blackout duration is longer than 3 hours or limit duration longer than 4 hours at input current not satisfying the thermal requirement, some serious problems will happen during the restart of the potlines and lead to cells shutdown if wrong strategy for restart is carried out.[1]

During the blackout, the strategy adopted includes following aspects: • Profile control of frozen shell/sludge • Bath temperature control and monitor • Anode control Profile control of frozen shell. After power failure, sludge (cryolite and alumina crystals) formed gradually with decreasing bath temperature and precipitated at the bottom of the cell. The sludge solidification may bring on the poor profile of the frozen ledge. For the sake of forming good profile of the frozen ledge, some operations were carried out: • Switch automatic mode to manual mode in order to avoid immoderately feeding after recovering the power failure. • Adjust anode-cathode distance gradually in order to avoid poor profile formed underneath the anodes which may bring on short circuit between anodes and cathodes.

Comparing with the start-up of a new cell, the different features of a cell being restarted include the following[2]: • the cathode base is more rigid; • there is more metal in the cell during early operation; • there will be more sludge/muck in the cells because of collapse of crust and bath during start-up. The consequence of these differences is that the probability of forming cracks in the cathode blocks increases significantly. Liquid aluminum can penetrate into the cathode block more readily in restarted cells. Thus the metal purity will go down more rapidly in the future operation. The pot life will be short due to cathode deterioration.

Bath temperature control and monitor. After power failure, all operations were terminated. The bath temperatures were measured hourly. All cover plates were put on the pots, and fire holes were sealed with cryolite and crushed bath; anodes were covered with crushed bath and alumina at higher thickness; alumina was manually removed from alumina hopper; dry scrubbing system was stopped. As a result, the bath temperatures decreased slowly as expected, see figure 1 and table 1.

A Chinese reduction plant has an installed capacity of 140 kt of metal per year and employs 180 pots with line current of 300kA. In May, 2008, power failure happened suddenly. The situation lasted 5 hours. The bath temperatures fell to 900°C below and electrolyte almost solidified. This type of situation was never faced by us. According to some literaturesI3], some shutdowns were planned unlike the emergency on hand at this smelter. Planned shutdown involve tapping of the

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very difficult to push the amperage beyond 120kA. After restart, the situation of anodes current distribution was improved (see table 3). However, most pots were not normalized.

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Fig. 1. Bath temperature evolution after power failure. Table 1. Bath temperatures for some pots 113# 112# lh 962 958 2h 939 935 3h 912 918 4h 895 901

115# 951 930 921 885

Table 2. Anodes current distribution during restart A2 A3 A4 Al A5 A6 A7 A8 3.2 3.7 2.6 1.5 0.8 3.9 7.6 11 Bl B2 B4 B5 B6 B7 B8 B3 5.3 2.1 3.6 2.9 3.1 5.8 10.2 8

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Table 3. Anodes current distribution after restart. Al A2 A4 A5 A6 A7 A8 A3 2.4 3.2 4.2 3.8 3.7 3.8 2.8 4.0 B4 Bl B2 B3 B5 B6 B7 B8 3.8 4.0 3.3 4.4 L2 3.8 3.7 3.6

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Side ledge. During restart, the average temperatures of steel shell were never over 400°C (see table 4). It indicated that the side ledge was not damaged.

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Table 4. The average temperatures of steel shell before and after restart. /°C Tap End Duct End Downstream Upstream Before 332 334 376 367 Restart After 341 380 350 330 Restart

Anode control. The anodes were put down at a rate of 2 cm/h if bath temperatures were higher than 930 °C. In the case of 930 °C below, the anodes were moved downward at a rate of 3 cm/h till contacted with the liquid metal.

Operating parameters. 15 days later, the cell voltage was controlled at 4.13-4.15V with average cathode voltage drop of 330mV. Anode effect frequency was around 0.2. The metal purity was greater than 99.7%, which indicated that the forming cracks in the cathode blocks were not occurred significantly. After 15 months, the average current efficiency was about 92.5%. The lines come to normal operating levels.

Restart Operations After recovering the power failure, the pot lines were switched on. Initially, anode effects occurred for the most of the cells, and some cells were in the status of short circuit. The line voltage built up reached up to 840 V, and the amperage was very low, less than 120 kA. The pot lines were running at a very poor status and continued to deteriorate.

Summary

To overcome this, some strategy were carried out: • Killed anode effects as soon as possible; • Stop feeding alumina; • Stop anode changing operation and metal tapping

This was the first experience of restarting lines with current input of 300kA after 5 hours blackout. It is important that the aluminum industry should operate on steady power. Such accident and restart require lot of efforts and have long term impact. However, such restarts enrich our knowledge.

Meanwhile, anodes were leveled slowly, and cell voltages were kept at 5 to 6 V. Cryolite was gradually added to increase the bath level. Once electrolyte started melting and bath level was built up to some extent, the voltage was slowly raised to about 10V and this increased thermal input to the pots and thereby helping in melting of cryolite and bath.

Acknowledgements The authors are grateful to the management of Keao Aluminium Company for granting permission to publish this paper. Thanks are also due to our colleagues involved in the restart operations.

After restart operation, bath temperatures gradually reached up to normal level, mainly in the range of 950-965 °C.

References

Extra cell voltage during the early operatine period. During the initial stage of normalization, the voltage was set at 4.3V over. This extra voltage is associated with the heat-up, melting of bath and is necessary for removal of sludge.

1.

Z. Qiu. Principle and Application of Aluminum Electrolysis, (in Chinese), the China Mining Press, 1998.

2.

BJ. Welch and K.Grojtheim, "Considerations of the Impact Regular Shutdown has on the Operation of Aluminium Smelters," Light Metals 1988, pp 613-616.

3.

U.B. Agrawal, G.D. Upadhyay and C.W.Deoras. "Restart of lOOkA VSS potlines after long shutdown" Light Metals 2002, pp 51-56

Anodes current distribution. Seen from table 2, the initial anodes current distribution was very poor, with number of pots on more than 30V. In some of the pots, arcing and voltage fluctuation were observed. Some anodes were not conducting the current. It was

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Multiblock Monitoring of Aluminum Reduction Cells Performance Jayson Tessier1, Carl Duchesne2, Gary P. Tarcy3 *: STAS, Alcoa-STAS R&D Team, Chicoutimi, QC, G7K 1H1, Canada 2 : Aluminium Research Centre-REGAL, Chemical Engineering Department, Universite Laval, Quebec, City, QC, GIK 7P4, Canada 2 : Alcoa Inc., Alcoa Technical Center, Hall Process Improvement, Alcoa Center, PA, 15069, USA Keywords: Process Monitoring, Reduction Cells, Process Control, Multivariate Statistics Abstract

Hence, following a pot excursion on CE or EC, engineers have to look at the behavior of many variables before taking corrective actions. Typically, this is achieved through the investigation of each variable using univariate statistical methods (i.e.: computation of variables average and standard-deviation or through the investigation of univariate Shewart control charts). This is time consuming and suboptimal as it does not account for the fact that reduction cells performance are resulting, from the simultaneous interaction of all these variables. Reduction cells are multivariate processes. Unfortunately, through the use of univariate statistical methods, it is impossible to capture the combined effect of these variables.

Aluminum reduction cells performance are affected by many factors. In order to efficiently understand possible causes of performance upsets, all major sources of variations have to be monitored. This implies monitoring all anode and alumina properties, as well as pot state and manipulated variables, while also taking into account pot design or integrity after start-up. Considering the high number of variables involved in such a task, this is practically impossible using typical statistical process control tools. The problem is even worst when applied on a pot basis. This paper proposes the use of multiblock PLS (MBPLS) to build a monitoring scheme on a pot basis, simultaneously taking into account the influence of alumina and anode properties, of pot state and manipulated variables, as well as the pot state following its start-up. Derived from a regression model, the monitoring policy ensures that only variations relevant to pot performance variations are highlighted.

A better approach for investigating a process upset would consist of looking at the pot status based on all variables at the same time. This is advantageous as it captures the structure or the interactions between process variables and pot performance. This is achieved through the use of multivariate statistical methods such as principal components analysis (PCA) or partial least squares (PLS). These methods enable the construction of multivariate monitoring charts and diagnosis tools while accounting for the multivariate nature of processes.

Introduction Aluminum reduction smelters typically operate a hundred to a thousand metallurgical reactors known as reduction cells or pots. Although each cell process variables are followed by process engineers on a daily or a weekly basis, key performance indicators (KPI) are typically followed on a weekly or a monthly basis. Hence, reports on current efficiency (CE) and energy consumption (EC) are investigated by process engineers and feedback corrective actions are taken, aiming at keeping cells in an optimal productive state.

This work proposes the use of these methods to develop a pot overall status or health monitoring scheme based on all available information aiming at explaining pot performance variations. Hence, the proposed monitoring strategy includes alumina and anode properties, pot state and manipulated variables, the pot initial state following its start-up and different binary variables arising from process knowledge. Doing so, a single monitoring strategy would enable to easily and efficiently monitor pot status based on all variables leading to its performance variations. Moreover, this strategy would also help diagnosing process upsets through the use of contribution plots.

However, it is generally not easy for process engineers to determine what might have been the cause of a process excursion leading to low CE or high EC. The reason is that reduction cells performance are simultaneously affected by the behavior of variables of different nature that could be grouped in different blocks. This is illustrated in Figure 1.

This paper is divided as follows. First, the drawbacks of univariate statistics and advantages of multivariate statistical analysis are illustrated. The dataset used for the model development is presented and the construction of the monitoring strategy is discussed. Finally, a case study illustrating how contributions plots can be used for upset investigation is presented. This is followed by a conclusion. Multivariate Process Monitoring Over the last decades, multivariate process monitoring has started to find its place in chemical and pharmaceutical industries, as well as within the food or the metallurgical industries. This is mainly due to the fact that more sensors, collecting data at a fast sampling rate, are now available. This resulted in the availability of massive databases for process monitoring and upsets diagnosis.

Figure 1: Variables typically involved in reduction cells performance variations.

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Nevertheless, the information enclosed in industrial databases encloses some noise and many variables are highly collinear (i.e.: correlated or dependent).

cell behaviour with respect to its bivariate nature, assuming that only bath temperature and excess of A1F3 are of importance.

Multivariate statistical methods such as PCA or PLS become handy as they are well suited to the analysis of thousands of correlated noisy variables, even possibly enclosing some missing values arising from different sampling rate or defective sensors. These methods have been used for different Multivariate Statistical Process Control (MSPC) applications [1, 2, 3].

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Recently, multivariate statistics have been used to develop different monitoring strategies for reduction cells. Majid et al. [4] used PCA on 5 minutes pot data for early detection of anode spikes and anode effect. Tessier et al. [5] also used PCA to develop a daily multivariate monitoring tool for reduction cells based on 65 variables. This later approach was used for detecting cells behaving abnormally as well as for troubleshooting.

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A limitation of univariate process monitoring can easily be illustrated using a simple case study. Daily bath temperature and excess of A1F3 were gathered for a pot over two years. These are presented in Figure 2 and 3 with their respective +/- 1, 2 and 3 standard-deviations limits (std).

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This procedure is easily extended to more than two variables and allows the computation of a single metric indicating how far a pot daily observation is from the control space. This is computed using the Hotelling's T2 statistics, which is χ2 distributed.

Figure 2. Univariate SPC for bath excess A1F3. An engineer can inspect these two graphs on a daily basis to determine if a pot is in control, based on bath temperature and excess A1F3. Engineers would flag observations based on some SPC rules [6]. However, this would not account for the fact that these two variables have a correlation coefficient -0.71. Hence, the underlying assumption for univariate SPC that variables must be independent is not met. This consequence is well illustrated in Figure 4. This figure presents bath temperature as a function of excess A1F3. The +/- 2 standard-deviations limits are illustrated by dotted lines for both variables. Independent variables would have filled the gray square formed by the dotted lines and univariate SPC would have been an efficient tool for process improvement. However, it is evident that these observations only move along a defined axis. Monitoring the behaviour of a cell based on the joint confidence interval of these two variables would be much more appropriated. This is highlighted in Figure 4 through the +/-2 standard-deviation joint confidence interval marked by the ellipse. It is seen that the joint confidence interval captures the correlation structure between the two variables and thus enable monitoring

Ti2 = (x i -x)S- 1 (x i -x) T

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where Xi is a row vector enclosing process measurements, x is a row vector enclosing process variables means or target, S is the process variables variance-covariance matrix and Tj2, a scalar, is the Hotelling's T2 statistics for the ith observation. However, due to correlations among process variables, S is often ill-conditioned and leads to poor monitoring tools. PCA/PLS Based Monitoring In a way to overcome the ill-conditioned problem, it is advantageous to use PCA or PLS to compute the Hotelling's T2 metric. PCA and PLS methods simply project process data into a sub-space (i.e.: the latent space) consisting of less variables that the original process data space. PCA aims at finding A latent variables P (JxA), each one explaining the greatest amount of

408

variance of the original data X (IxJ) that is unexplained by the previous latent variables. On the other hand, PLS aims at finding A latent variables (P), each one explaining the greatest amount of the covariance between X and a response matrix Y (IxH). The latent variables are computed in such a way that they are independent, or uncorrelated, of each other, ensuring that each latent variable captures new information unexplained by the previous latent variables. Mathematical details of PCA and PLS are presented in the literature [1, 2, 3, 7].

weight that will be used for the next tapping operation. As seen in Figure 5, small differences in metal height may lead to large differences in CE. Another source of perturbation arisesfrombath carryover. As an operator taps metal, a variable quantity of molten bath is tapped from the pot. As this bath is tapped instead of metal, some metal is left in the pot. Hence, this artificially boosts the metal height and affects CE. Unfortunately, it is not possible to quantify the amount of bath that is carried over through the metal tapping operation on a pot basis. At best, it is possible to determine the amount of bath that has frozen on the crucible lining for some group of pots. Finally, sideledge freezing and thawing also induce some artificial variations on CE. For a defined quantity of metal inventory, sideledge freezing will boost the metal height and hence artificially increase CE as the metal level will be higher. Conversely, sideledge thawing will negatively affect daily CE. These sources of noise or error will also affect EC as it is a function of CE.

The structure of the PLS model is given by: X = TPT + E Y = T QT + E T = XW* W* = W(PT W)"1

(2) (3) (4) (5)

where T (IxA) is the common latent variable space defined by the weight matrix W* (JxA) and capturing the information in X that is the most highly correlated with Y. The P (JxA) and Q (HxA) matrices contain the orthogonal loading vectors mapping the common latent variable space in the space of X and Y (models of these blocks). The PLS model residuals for X and Y are stored in E (IxJ) and F (IxH), respectively.

Hence, daily CE and EC computed from metal tap weights, on a pot basis are irrelevant for performance assessment. Typically, plant operator will average CE and EC on a monthly basis in order to smooth out variations created by the tapping table, bath carryover and sideledge freezing/thawing. 400

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The dataset for this work comes from the Alcoa Deschambault smelter, operating 264 AP-30 reduction cells above its nominal capacity. Data were collected over the complete life cycle of 31 cells. For each cell, a total of 209 preheating, start-up and early operation data, as well as alumina and anode properties and pot operating data were retrieved for complete pot life cycles. The pot related variables were retrieved on a daily basis, from the plant database while the other variables were stored as available. These data, the process descriptor data, are enclosed in the X matrix.

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Developing the Monitoring Strategy A PLS based monitoring strategy was selected for this application. On this basis, it is possible to link variations enclosed in the process descriptor matrix X with different response variables (Y). Here, the objective is mainly to track reduction cells production metrics. Hence, CE and EC are used as response variables. Therefore, these metrics were computed on a daily basis from daily tap metal weights, line current and pots volts.

Hence, data gathered over the life cycle of the 31 cells were averaged on a monthly basis, leading to 2271 observations. Therefore, X contains 2271 observations and 209 variables and Y encloses 2271 observations and 2 variables. A schematic of the data structure is presented in Figure 6. The A1203 and Anodes blocks enclose alumina and anode properties, respectively. The MV and the SV blocks encloses pot manipulated and state variables, while the PSE encloses pots start-up and early operation information. Finally, the PLV block encloses some logical (i.e.: binary) information with respect to the pot location within the potrooms.

The assumption is that, if a good statistical model can be built, it would be possible to assess pot status by simultaneously taking into account alumina and anode properties, of pot state and manipulated variables and also based on some pot specific parameters.

Process Lag and Dynamics Considerations

Time Basis Considerations

In order to cope with process dynamics, lagged version of data blocks were used. The idea was that a pot processing bad alumina or anodes, for example, for an extended period of time may be more negatively affected than a pot processing such raw materials for a short period of time.

However, it is known that CE and EC computed through metal production, on a daily basis, are noisy metrics. One reason for that is coming from the use of metal tapping tables. This particular smelter uses a tapping table, similar to the one presented in Figure 5, to determine the amount of metal that should be tapped from a pot. As an operator measures the metal pad level, the metal height indicating the quantity of metal in the pot is converted into a tap

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data, and weekly population averages are used to describe anode quality based on a limited number of analyzed core samples. A better traceability of these raw materials would certainly help explain additional variance. Furthermore, the uncertainties involved in computing, the CE and EC values due to the tapping tables and errors, as discussed earlier, may very well limit the theoretically explainable performance fluctuations.

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Figure 6: Schematic of the available data structure. Therefore, PLS models using different lags were developed and assessed. Table 1 presents statistical details of three PLS models developed for different lags. The first model (No Lag) includes only present month data, while the One Month and Two Months models include present monthly values as well as one and two months lagged versions, respectively, of the data for the A1203, the Anodes, the MV and SV. The PSE and the PLV blocks are not lagged as they do not evolve over time. This table includes; the number of variables included in X, the number of principal components used in the different PLS models, the amount of variance of X accounted by the model (R2X), the root mean squared error in calibration (RMSEC) for CE and EC and the amount of variance explained for CE and EC (R2Y). Table 1: Statistical details of PLS models.

Nbr.X Variables A R2X (%) RMSEC CE (%) RMSEC EC (kWh/kg) R2Y CE (%) R2Y EC (%)

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One Month 345 4 19 2.21 0.37 51 50

Figure 7 presents predicted against measured autoscaled CE values for the 2271 observations. This figure demonstrates that most of the predictions follow the prefect prediction line and that the prediction errors are distributed along the complete scale of measured CE. The reader should keep in mind that 2271 observations are plotted on this figure and that most of the observations are close to the prefect diagonal prediction line. 6

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From these results, it could be said that 19 to 24% of the variance of X (R2X) explains about 50% of the CE and EC variance (R2Y), indicating that most of the variance enclosed in X is not used for explaining variations in CE and EC. As seen from these results, including three months (present month and last two months) of data slightly improves the model performance as opposed to only using the present month data. The R2Y CE slightly improves from 0.51 to 0.54, while it stays constant for EC. On the other hand, the RMSEC diminishes from 2.21 to 2.14 for CE and stays unchanged for EC. Hence, it was decided to use the Two Months model. Even if this model includes more variables, while leading to slightly better prediction performances, these are data collected in the present and last two months of the pot operations and are easy to extract from the plant historian while not causing any problems for the PLS algorithm. Still, it has the advantage of highlighting recurrent problems leading to CE or EC drifts while troubleshooting.

Case Study As a statistical model enclosing most of the systematic interactions between pot process variables and performance is available, it is possible to use this model for process monitoring and diagnosis. The methodology is illustrated here using a case study.

A considerable amount of efforts were made to improve predictive ability (R2Y and RMSEC) using other ways to preprocess the data, but with no significant gain. Data quality may help explain some of the difficulties in capturing a greater percentage of variations (R2Y) in CE and EC. Alumina and anode quality variables are not measured on a pot-to-pot basis. Alumina properties are rather estimated from suppliers COA and blending

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After the first three months of operation, the monitoring strategy was applied to pot A003. Over time, on a monthly basis, process engineers follow predicted and computed autoscaled CE as presented in Figure 9. It is seen that the predicted CE does not exactly match the computed values, mainly arising from the different points discussed in the previous section. However, predictions follow the major trends and at least indicate if CE is improving or degrading. This figure illustrates a CE drop from observations (months) 26 and 27. The computed CE indicates a CE drop of 10%, probably greater than what the pot experienced as the computed CE were suspiciously high for months 24 and 25, while the predicted values indicate a drop of 4.5%, which is about twice the magnitude of the RMSECV. At -350 kA, a 4.5 % CE drop equals ~4 tons of metal and is a significant drop. After highlighting this drop, the engineer can interrogate the variables contributions [3] in order to determine which variables played a role in this significant CE drop.

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where xitj and x,.;^ are values of the / h variable for observations i and 1-7; w*aj· is the weight associated to the/ 1 variable in the 0th latent space and 52ta is the variance of the αΛ score. It basically consists of the difference between two observations of variable j , weighted by its importance in the model (w*). Dividing by the score variance gives an equal chance to each latent variable to influence the variable contribution. Note that the weights and variances in the above expression are taken from the regular PLS model with block scaling. To obtain the contribution of the variables in a particular block, one only needs to use the appropriate variables within that block (x,,/s). Variables contribution to the difference in predicted CE from observations 26 to 27 are presented in Figure 10.

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Conclusion

The contribution of variable j , to the overall movement between two consecutive observations, i and i-7, within the latent space of A dimensions is computed using the following expression:

In this paper, an efficient monitoring tool based on all variables collected around reduction cells has been proposed. This methodology, based on multiblock PLS can efficiently cope with

411

hundreds or thousands of correlated and noisy variables as it projects the pot information into a latent variable space of much lower dimensions than the original data space, while capturing most of the systematic variations. Such a model was developed based on data collected over the complete life cycle of 31 reduction cells. Information on alumina and anode properties, on pot state and manipulated variables, on pot integrity after start-up and some logical information indicating pot position in the potroom was included. The resulting model accounts for 54% of the CE variance and 51% of the EC variance. Since different sources of variations may create added noise to computed CE and EC values, the model is judged useful and enables to predict CE and EC monthly values, on a pot basis, based on process data. Following a drop of CE, or an increase of EC, for two consecutive months, it is possible to compute variables contributions and investigate what may have caused the performance upset. The reader should keep in mind that the model was built upon happenstance data (correlations) and not from a design of experiment. Hence, the model can not indicate causation. However, the lack of correlation indicates no causation. Therefore, process engineers have to use their knowledge and judgements, combined with the information arising from contribution plots, to identify possible root causes. References 1. Kourti, T., MacGregor, J.F. (1995). Process analysis, monitoring and diagnosis, using multivariate projection methods, Chemometrics and Intelligent Laboratory Systems, Vol. 28, pp. 321. 2. Kourti, T. (2002). Process analysis and abnormal situation detection: From theory to practice, IEEE Control Systems Magazine, October 2002, pp. 10-25. 3. Kourti, T. (2005). Application of latent variable methods to process control and multivariate statistical process control in industry, International Journal of Adaptive Control and Signal Processing, Vol. 19, No. 4, pp.213-246. 4. Majid, N.A.A., Young, B.R., Taylor, M.P., Chen, J.J.J., 2009, Detecting abnormalities in aluminium reduction cells based on process events using multi-way principal component analysis (MPCA), Light Metals 2009, pp. 589-593. 5. Tessier, J., Zwirz, T.G., Tarcy, G.P., Manzini, R.A., 2009, Multivariate statistical process monitoring of reduction cells, Light Metals 2009, pp. 305-310. 6. Ostle, B., Turner Jr., K.V., Hicks, C, McElrath, G.W., (1996), Engineering Statistics: The Industrial Experience, Duxbury Press, 568 p. 7. Tessier, J., Duchesne, C, Tarcy, G.P., Gauthier, C, Dufour, G., 2008, Analysis of a potroom performance drift, from a multivariate point of view, Light Metals 2008, pp. 319-324.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

Cells Technology, Development and Sustainability SESSION CHAIR

Gilles Dufour Aluminerie de Deschambault Quebec, Canada

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

HIGH AMPERAGE OPERATION OF AP18 POTS AT KARM0Y Marvin Bugge1, Haakon Haakonsen2, Ove Kobbeltvedt2 and Knut Arne Paulsen2 ^ydro Aluminium, P.O. Box 2560, NO-3908 Porsgrunn, Norway 2 Hydro Aluminium Karm0y, Hydrovegen 160, NO-4265 Hävik, Norway Keywords: High-Amperage, API8, Cell Operation, Karm0y Abstract

The first amperage increase up to 200 kA was done with some modifications of the anode size and improved cathode block quality [1,2]:

The API8 potline at Karm0y has increased the amperage to 230 kA. This has increased the productivity and reduced the specific production costs. The operation showed some lack of performance at high amperage for the old pots designed for lower amperage. The old pots had too high heat input, resulting in thin side ledge and higher Si content in the metal. To improve the heat balance the pot voltage had to be reduced for the older pots, and this resulted in reduced current efficiency (CE). The cathode life was reduced to -2250 days due to the use of graphitized cathode blocks and the increased amperage. The new cathodes designed for 230 kA show good operational results. The young pots show 94-94.5% current efficiency and 13.0 kWh/kg AI in energy consumption. Due to less bath volume in the new pots it has been a challenge to reduce the anode effect frequency.

• • •

The amperage increase from 200 kA up to 230 kA will be further discussed here. The amperage has been pushed up much the last years due to: High metal price (in periods) Power was available at an affordable price The cast house had extra capacity The rodding shop was able to rod more anodes Extra anodes were available from the suppliers The rectifiers had some surplus capacity Loss of production from the stopped S0derberg lines

Introduction Karm0y started aluminium production in 1967, and the first potlines used the S0derberg technology. To increase the production it was decided to use the API8 Prebake cell technology, and the API8 potline at Karm0y was started in 1982. Later expansions increased the number of pots to 288. Start-up in 1982 Expansion in 1987 Expansion in 1997

The higher production also resulted in a higher productivity, as the number of employees was not increased with the increased production. This was an important contribution to reduced specific production costs.

108 pots 114 pots 66 pots

However, some investments were necessary, Most investments were related to the use of the bigger anodes.

During the first 10 to 12 years the amperage was 175 kA, and then the amperage was gradually increased. From 1999 a booster rectifier was used for five pots as part of the amperage increase program [1], The experience from the booster pots and other potlines in Hydro [5,10] has then been used to increase the amperage up to 230 kA.

Main investments for increased amperage were: • Booster rectifier - capacity 20 kA • Extra transformers and rectifiers for higher amperage • Change to bigger anode stubs, almost 40000 stubs • Modifications in the rodding shop for bigger anodes • Extra cast iron capacity for the bigger anode stub holes • Modification of the potshells • Transportation equipment for the bigger anodes • New moulds for bigger anodes (by the anode supplier) • Slot production in anodes (by the anode supplier) • Forced suction in order to keep the emission limits

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Combined investments the last ten years have been almost 30 million USD. The amperage increase from 200 kA to 230 kA has resulted in about 22 kt Al/y of increased production. This means average investments of about 1350 USD/t Al. Compared to new potlines (both Brownfield and Greenfield) this cost of extra production is low. For comparison the cost for the new potline in Sunndal (Brownfield expansion) was about 3100 USD/t Al [11]. Greenfield capacity requires much higher investments. For the expansion of the potline in Slovalco the expansion costs was about 1800 USD/t [6].

1980 1985 1990 1995 2000 2005 2010

Figure 1. The amperage development at Karm0y.

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The original AP 18 anode size was increased to allow better utilization of the area inside the potshell. Later the use of thinner SiC slabs made it possible to increase the anode area further. Also the centre channel was reduced to maximize the length of the anodes. The anode length has been increased in steps up to 1630 mm so far. A smaller increase was done for the anode width. The anode width is limited by the small channels for chiseling between the anodes during anode change.

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To reduce the voltage drop from the anode stubs to the carbon anode it was important to increase the contact area by increasing both the anode stub diameter and the anode stub hole depth. The bigger stubs also resulted in a higher heat loss through the stubs, which was important for the heat balance of the pot at higher amperage. All stubs (almost 40000) were replaced in a 6-weeks period. A similar change was also reported from Tomago as part of their upgrade to AP22 technology [8]. An alternative to the bigger stubs has been to increase the number of stubs to 8 stubs. The number of stubs has been increased in other plants in Hydro [4,6].

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The last years the anode cycle has been stable at 80 shifts (26 and 2/3 days), and with the increased amperage this has resulted in a relatively thin anode butt and low gross anode consumption (510-520 kg/t AI). It has been important to have the 80 shifts anode cycle, as this fits with an anode changed every fifth shift (anode change every fortieth hour).

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Cathode improvements The original cathode used amorphous cathode blocks. During the years the graphite content in the cathode blocks was increased to graphitic qualities. These cathode blocks resulted in excellent cathode life (up to 3000 days). To reduce the cathode voltage drop and maximize the amperage graphitized cathode blocks were chosen. As a result of the higher cathode block erosion for graphitized cathode blocks the cathode life was also reduced.

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The big API8 anode has a high surface area, and the use of slots has been important to reduce the bubble voltage drop. Introduction of slots reduced this voltage drop significantly (see figure 2). Different slot designs were also tested on the booster pots. In the beginning a single slot was tested. Later up to three slots were tested, but two slots in the anodes have been used for the regular pots. So far the slot depth has been increased up to 300 mm. Theoretically the slots should be even deeper, and deeper slots have a bigger potential to reduce the bubble overvoltage compared to a third slot. The slots are produced by the anode supplier, and the slot depth has been limited by the installed equipment for slot production.

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Figure 4. Cathode life development. Comparison between graphitic and graphitized cathode blocks. Later the increased amperage has resulted in a further reduction in the cathode life to about 2250 days for the graphitized cathode blocks (see figure 5).

Figure 2. Calculated bubble voltage drops at 230 kA. To maximize the anode cycle the anode height has been increased to 610 mm and the anode weight from 1180 kg to 1450 kg during the last 10 years (see figure 3).

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Figure 7. Development in average anode effect duration.

The cathodes were also modified to higher amperage by using: • Higher and thinner SiC slabs • Longer cathode blocks (to fit with the longer anodes) • Bigger cathode collector bars (to minimize the cathode voltage drop) • Reduced bottom insulation (to increase the heat loss)

Magnetic field compensation has been tested in parts of the potline. So far the modifications in the busbar system have not resulted in increased performance, and it does not seem to be necessary for pot operation at 230 kA. Laser-guided anode setting operation was introduced to reduce the fluoride exposure for the operators and improve the anode setting accuracy. This laser measurement method has also contributed to improved safety during anode change, as no operator is required to be close to the anode for accurate anode setting.

The improved cathodes have resulted in a reduction in the cathode resistance.

2.4

Operational Results As the amperage increased, the side ledge was reduced in the oldest pots. For many of these old pots red shell sides were detected, and some pots also tapped out in the upper part of the potshell. These pots were not designed for the high amperage.

1980 1985 1990 1995 2000 2005 2010 Figure 6. Development in average cathode resistance. The 215 kA lining was, after verification in the booster section, introduced in the potline in 2002. Later the new 230 kA lining was verified in the booster section and used as standard cathode lining in the potroom from 2008. T

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Other improvements and tests The forced gas suction was introduced in the potline in 2002-2003 to reduce the emissions from the pots [3].

Figure 8. Si content in the metal increased with the amperage As a result of the thin side ledge and the high Si content, a new 230 kA lining was introduced as standard lining in the potline from 2008. For the old pots the pot voltage and the bath acidity were reduced to get a thicker side ledge, reduced Si content and avoiding tap outs.

Propane gas preheating was introduced in 2003. From then it was not necessary to have an anode effect at start-up [9]. The anode effect quenching was optimized in 2008 to reduce the duration. The average duration was then reduced from about 3 minutes down to about 1.5 minutes. This reduction has been important to minimize the emissions of greenhouse gases.

Also the CE was gradually reduced with the higher amperage.

417

This may be explained by different reasons: • Higher metal velocity - results in a higher mass transfer number for the back reaction of aluminium • Increased superheat • Less side ledge and a larger metal area • Low interpolar distance (ACD) • Reduced bath acidity for old pots Additionally it has been some anode problems, which also in periods contributed to a reduction in the CE. The average CE was reduced from about 94% in 2005-2006 to about 93%.

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Figure 11. CE development for the old and new linings. The improved CE has also contributed in lower energy consumption for the new lining, but also the old lining showed low energy consumption due to the low pot voltage. Low energy consumption has always been given high priority at the Karm0y plant. The results for energy consumption in January-September 2010 (about 30 mV busbar losses from crossovers, to andfromrectifiers and between sections is not included):

Jan 05 Jan 06 Jan 07 Jan 08 Jan 09 Figure 9. CE and amperage development in the potline.

• •

The new 230 kA lining is a significantly colder lining than the old 215 kA lining. The Si content has been low, and the voltage has been higher. This higher voltage combined with reduced cathode voltage drop has resulted in a higher anode-cathode distance (ACD). A 1 n H-. 1 U

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The bigger anodes have resulted in less molten bath in the pots. Less bath volume combined with the higher amperage results in more challenges to avoid AEs. This challenge increased for the new lining operating with a thicker side ledge - resulting in even less molten bath and more AEs compared to the old lining (figure 12). The pots with the new lining are more sensitive to low bath levels - figure 12 shows almost 0.2 AEs/day in February 2010, due to low bath levels. Another reason for the higher AE frequency for the new pots is that the metal level was also reduced for the new lining to increase the superheat. This resulted in reduced total liquid levels and reduced immersion of the breakers in the bath and more blocked feeder holes.

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The energy consumption for the 230 kA cathode lining at Karm0y is better than reported by the AP22 technology in Tomago [8] and even better compared to the reported energy consumption for the more modern AP39 technology [7].

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New 230 kA lining: 13.0 kWh/kg AI Old 215 kA lining: 13.2 kWh/kg AI

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Figure 10. Pot voltage development for the old and new linings.

The AE frequency has later been improved during 2010, partly due to better bath control and slightly increased bath levels.

The introduction of the new 230 kA lining resulted in significantly better CE compared to the old lining designed for 215 kA (figure 11). The average difference has been more than 3%.

418

References

July

Jan 09

July

Jan 10

July

1.

O. Kobbeltvedt et al., "Experience from Amperage Increase on Test Pots in the Prebake Line at Hydro Aluminium Karm0y", Eleventh International Aluminium Symposium, Trondheim - Bergen - Trondheim, Norway, September 19-22, 2001, pp. 9-14.

2.

J. Tonheim et al., "Experience with Booster Pots in the Prebake Line at Hydro Aluminium Karm0y", Light Metals 2004, pp. 191-196.

3.

J. Tonheim et al., "Environmental Challenges in the Prebake Line at Hydro Aluminium Karm0y", Light Metals 2005, pp. 283-287.

4.

B.P. Moxnes et al., "Potline Amperage Increase from 160 kA to 175 kA During One Month", Light Metals 2001, pp. 173-178.

5.

B.P. Moxnes, H. Kvande and A. Solheim, "Experience and Pitfalls with Amperage Increase in Hydro Aluminium Potlines during the last 10 years", Light Metals 2007, pp. 263-268.

6.

M. Bugge et al., "Expansion of the potline in Slovalco", Light Metals 2008, pp. 261-265

7.

O. Martin et AI., "Development of the AP39: The New Flagship of AP Technology", Light Metals 2010, pp. 333-338.

8.

L. Fiot et al., "Tomago Aluminium AP22 Project", Light Metals 2004, pp. 173-178.

9.

O.A. Lorentsen and K. Rye, "Twelve years of experience with a fully automated gas preheating system for S0derberg and Prebake cells", Light Metals 2008, pp. 1001-1005.

Figure 12. AE frequency development for the old and new linings. Conclusions The amperage at Karm0y has been increased to 230 kA. Extra metal production has been possible with small investments compared to new capacity. The higher amperage has resulted in reduced CE for the potline, mainly due to reduced performance for the pot with the old pot lining. The old lining designed for 215 kA has shown lack of performance at 230 kA. The side ledge has been too thin in the old pots, and the high Si content in the produced metal comes from the SiC sidewalls. The thin side ledge has also resulted in more red shell sides and some tap outs for the old pots. The booster pots have been used to develop and verify the new cathode for high-amperage operation. The new 230 kA lining shows good results. The side ledge situation is good, and the pots can therefore be operated at a higher voltage and a higher ACD. The higher ACD also results in more stable operation and better CE. The energy consumption shows results comparable to or even better than more modern technology. The new cathode lining therefore also shows potentials for even higher amperage than 230 kA. The amperage may be increased to 240-250 kA.

10. O.J. Siljan et al., "Amperage Increase Experience in Hydro Aluminium", Ninth Australian Aluminium Smelting Technology Conference and Workshop, Terrigal, Australia, 4-9 Nov. 2007, pp. 319-333. 11. K.0. Vee, J.A. Haugan and A.H. Hus0y, "Hydro Aluminium Sunndal Expansion Project", paper presented at the 2003 TMS Annual Meeting, San Diego, California, 2003.

The high amperage and the low bath volume have resulted in less stable bath level. This has resulted in a higher AE frequency for the new high amperage pots. The main challenge for the new pots has been to reduce the AEfrequencyat high amperage. Acknowledgements Hydro Aluminium Karm0y is acknowledged for the permission to publish the results for the high amperage operation.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ALUMINIUM SMELTER LOGISTICS - CAN THESE BRING REAL COST SAVINGS? Maarten Meijer *TMS (The Minerals, Metals & Materials Society); 184 Thorn Hill Rd.; Warrendale, PA 15086, USA Hencon B.V., de Stenenmaat 15, 7071 ED Ulft, the Netherlands Keywords: Aluminum Smelter Logistics, Pot Room Operation, Fleet Management, Schedule Optimization ABSTRACT Discrete logistics studies has proven to be a low cost tool to evaluate and implement technologies and lower the risk of investments into large scale infrastructural projects. So far this tools have not been used much with similar complexity as in other industries in order to evaluate design issues, feasibility questions and bankability. This paper describes how simulation tools do have an impact on capex and opex. What effect simulation has on optimizing interaction between smelter units and processes for additional cost savings. Finally how the use of simulation tools help to achieve "lean thinking" in new smelters or upgrades while at the same time lowering the risk of such innovations. INTRODUCTION Traditional, discrete simulation tools like Simio1, Arena" and Automodm are used to evaluate the feasibility of discrete material flows at airports, terminals, food and beverage industry, car manufacturing facilities, metal production, oil and gas, etc. In aluminum industry several studies have been done in a limited amount of time to support investment decisions on casthouse expansions, gantries cranes, warehouse optimization, plant traffic safety, etc. Although this studies have been helpful to take educated decisions on Greenfield and brownfield projects, this knowledge did not lead to frequent use of this tools as a decision making support in other projects.

Figure 1: typical lay out and schedule for a potline The room for optimization get more clear if we study the shift pattern and material flows of a modern smelter. Most smelters organize their potlines in several sections. Each section is treated in more or less the same pattern. Anode setting periods and tapping periods follow each other up, per section in a more or less regular pattern. Per section a PTM or GP crane is dedicated to the pots to assist the operators in completing the task in time. This operating regime is adapted by most smelters as the ideal and most reliable way to handle the pots in time with a reliable operator performance and relative stable line. However when amperage increased and smelters become larger, this regime start to become a critical aspect of running a smelter. For instance where the original design concentrated on one or two PTM's that could handle the pots in sequence. Nowadays the smelter is handled by a minimum of two PTM's and 1 GP to catch up with the production rate of the pot. Keeping the operating regime in sequence requires gantry cranes and after a shift is concluded the cranes are set up in the right sequence to be ready for the next shift. As a consequence fly times (the time that a crane is moving without a load) is increased and vehicle performance has decreased. On both the cranes and the vehicles OEE indexes of 60% are not unusual. Finally if we compare manufacturing results with the principals of lean manufacturing we see many cases in which waste is produced, a view examples of such waste are: • Anode's and Ladle's arriving to early at the crane or stacked in a row in the potroom to solve just in time problems between the transport cycle and handling cycle of the crane. • Uneven distributed anode change patterns or metal delivery patterns during the shift. • Not optimized metal flows to the cast house, resulting in a overflow or underflow of metal to the casthouse • Empty rides between the section and the different stations of vehicles. • Early or late anode changes to keep cells in the desired shift pattern per section.

Now with a new challenge to the Aluminium Industry where most smelters have cost reduction programs running of $100 to $200 per ton produced, discrete simulation can help to optimize Aluminium Manufacturing and turn it into a lean process. LEAN MANUFACTURING Electrolyze cells have always been studied and optimized within the industry. There is also a long tradition in this technology to use fluid dynamic calculations and test pots in order to come to a reliable and stable cell design that has been tested for many years. Typical examples of such long running research programs are: Development of the AP39: The new Flagship of AP technology and DX Pot technology powers green field expansion^ However from a discrete point of view this development programs still have room for further optimization of the material flow and a discrete approach towards anode handling and metal tapping.

421

All this items indicate there is room for improvement with regards to a lean distribution of materials into the potline. However they also indicate that the material handling system is not optimized to the knowledge collected in running stable pots at high current density. Due to the physical nature of cranes, human beings and introduction of sections handled by a dedicated team. There are two disturbing factors that are not controlled by the cell, but determined by the shift pattern. This factors are: anode setting and metal tapping.

introducing a new design in an existing smelter. The model convinced the operating staff with regards to the use and flexibility of this new design. Apart from that the model started a discussion on how this vehicle should be operated that resulted in additional functionality in order to utilize the vehicles more successfully than could have been done in their old shift pattern. The 20% reduction in OPEX, never would have been approved and achieved if the model was not build. The reason for this was that the model was the first platform in which different stake holders from, engineering, operation, maintenance and smelter management shared their ideas. The model was used as a "common language" tool, to evaluate consequences from their mutual decisions, based on their combined knowledge presented in the model.

Today's pot control systems collect enough information to plan this activities more smart and according to the demand of the pot and casthouse. However the operating regime and lack of knowledge on discrete logistics applied to our industry stop us from changing our typical section based approach.

This working method, convinced the investment team that a 20% reduction in cost was achievable, however more important once they agreed that this was the best option, it helped them to illustrate their complex task and results to top management in order to get the investment improved.

DISCRETE LOGISTICS STUDY FOR POTROOMS Light metals gives a limited number of papers that deal with using simulation tools to optimize the flow of discrete materials within a potroom. In 2009 Hydro published a study.™ In 2010 an outlook of our research program was given in another publication. vu Apart from that a number of experts within our industry are using discrete flow simulations to take educated investment decisions or improve the safety of a factory lay out. However making a potroom model is a complex task in which the detail of the model is very often limited due to the time constraints of the project at hand. Within our own research it tookfrom2007 till 2010 to upgrade the model to enough complexity to handle the material flows successfully and start using it as a base model to discuss and evaluate alternative manufacturing policies with our clients. At the moment the model is a good tool to investigate alternative scenario's for the use of vehicles in the manufacturing of Aluminium. Everybody that sees the model and use the model immediate understands its potential. However adaption's are still going on to add functions that help to evaluate lean manufacturing strategies for existing and future pot rooms.

DISCRETE MODELLING CONSIDERATIONS Like in any good model, the biggest dilemma stay's how to reflect the reality without modeling the reality in every detail. Therefore for every model we have to evaluate what level of detail is needed? Within our model we solved this issue by using a object orientated approach. This results in objects (for instance a cell) with afixedset of variables that successfully simulate a cell. This allows us to reuse information and make objects as smart as they need to be. Next to that we simplify processes as much as possible. For instance if we study vehicle movements we do not include the crane as an object, but use the cell as an item that generate the demand. This allow us to only evaluate movements on the floor and not interfere them with a complex object like a crane. VALIDATION Validation of results will always be important. This is a part where the stakeholders of our customers play the most important role. If they do not recognize their own manufacturing process in the model, the model is not suitable to run new experiments and suggest alternative manufacturing processes. This is an important validation step within our process. Without this step the model cannot be used to evaluate future processes and decide they are valuable to invest in. As a manufacturer and supplier of transport systems we know how important it is to match system performance with design calculations. Unlike funded research validation, the results are confidential in most cases.

Figure 2: a typical result of cost saving by using the model

A typical result of the model is shown infigure2, in this study we successfully reduced the fuel consumption and repair cost by

422

CONCLUSION Combining discrete modeling with the advanced models of cell simulation open up a new area of research in which the material flow can be optimized, towards the need of the individual cell. Since anode changing and metal tapping are important parameters that influence the stability of a cell, such an attempt can create a more efficient use of assist and higher productivity of the smelter. With today's model it is possible to support an investment team during the feasibility and design phase of an investment in order to come to a best practice solution that is, validated and proven making use of the combined know how of all stakeholders in the project. Discrete modeling is a tool that helps to justify or research new manufacturing methods up to a point that reaches further than gut feeling. This tool can help to find new manufacturing methods (outside pot technology research) that makefruitfuluse of the capacity of modern smelters with nameplate capacities between 150 thousand and 1 million tons of Aluminium produced a year. 1

http://www.simio.com/ http://www.arenasimulation.com/ m http: //www. automod. se/eng/home .html 1V Development of the AP39: The new Flagship of AP technology O.Martin, X.Berne, P.Bon, L.Fiot, D.Munoz, C.Ritter, R.Santerre, (Light Metals 2010, page 333-338) v DX Pot Technology Powers Green Field Expansion. A. Zarouni, M. Zelicourt, M. Jallaf, K. Alaswad, A. Kumar, A. Reyami, V. Kumar, D. Bakshi, J. Blasques, and I. Baggash (Light Metals 2010, page 339-344) Λ1 Logistic simulation of discrete material flow and processes in aluminum smelters. Anton Winkelmann, Ingo Eick, Christian Droste, Martin Segatz,.( Light Metals 2009) ™ New Logistic Concepts for 400 and 500 KA Smelters. Maarten Meijer (Light metals 2010, page 345-348) u

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SIMULTANEOUS PREHEATING A N D FAST RESTART OF 50 ALUMINIUM REDUCTION CELLS IN A N IDLED POTLINE A new soft restart technique for a potline Albert Mulder1*, Anita FolkersO, Marco A. Stam1*, Mark P. Taylor2) υ

2)

Aluminium Delfzijl B.V., P.O. Box 133, 9930 AC Delfzijl, the Netherlands The Light Metals Research Centre, University of Auckland, Private Bag 92019, New Zealand Keywords: Start-up, Process Control The tests confirmed that a 'gradual amperage increase' start-up was possible and that the required minimum line voltage and a controlled current distribution within the cathode bars could be achieved. The main focus of the start-up procedure is keeping the line voltage high enough to prevent the rectifiers from increasing the amperage above safe operating levels.

Abstract Due to the global economical crisis a significant amount of primary aluminium production capacity has been shutdown. A number of different strategies to restart idled aluminium reduction cells have been discussed in the literature [1, 2]. This paper describes the successful development and execution of the start-up of 50 cells simultaneously in one potline. The procedure is based on restarting reduction cells using a cold metal plate. Contrary to electrical preheating of new cells with use of cokes or graphite, these cells have been prepared with anodes positioned in direct contact and on top of the cold metal plate. The rate of preheating of the cells and associated melting of metal is controlled by a gradual line current increase. The actual start-up of the cells is performed sequentially by the addition of liquid electrolyte and moving the anode beam upwards. In this respect 50 cells have been preheated and restarted in 8 days.

Several methods for preheating were tested. Although these methods demonstrate more or less the same outcome in respect to the voltage characteristics per cell, the cathode current distribution between the individual collector bars shows large variations. The anode current distribution was identically sensitive to the rate at which the amperage was increased over time. Restart restrictions For the restart of Potline 1 a number of restrictions were identified that determine the characteristics of the new method. Prior to the shutdown these restrictions and associated actions were clarified.

Introduction A complete shut down and restart of a large number of cells is not a practice that a smelter carries out regularly, because of the time consuming characteristics and additional costs involved. At Aldel, the last shutdown and start-up was after a fire in the transformer building in 1977. However, due to the economical circumstances, Aldel has decreased its annual production capacity by 40% with the closure of Potline 1 in 2009.

Cell conservation The shutdown procedure was linked to the condition of each cell in order to be prepared for preheat and restart. Cells that were not considered for restart were completely drained of metal. Cells that were suitable for the new preheating and restart method were preserved with a metal pad layer of 10 cm. This was undertaken to protect the cathode surface during the shutdown period, provide a flat and conductive base for positioning the anodes during preheat, and to protect the surface during liquid bath mass transfer (slow down the rate of increase of the cathode temperature).

This capacity was partly taken into operation again by June 2010. Generally a pre-heat and start-up of a whole potline is carried out cell by cell. However, this is not possible at Aldel due to the configuration of the rectifiers. The rectifiers require a minimum voltage of 110 V, which is equivalent to 24 cells running at 142 kA. Since this is not achievable instantaneously, a new method was developed to restart one cell at a time while maintaining the total line voltage above the required level.

Rectifiers Potline 1 has five rectifiers with a capacity of 40 kA each. Four of these rectifiers were built by Siemens in 1965. The minimum voltage of these rectifiers in the lower voltage range is 110 V. If the voltage falls below this voltage, the rectifiers automatically compensate by increasing the amperage. In 2006 a new rectifier ("Ell") was added to the potline in order to secure N-l operation and to increase the line current. This rectifier operates in the range of 0-370 V.

During the design and development phase of the restart method, it was clear that preheating a large number of cells simultaneously offered the best option in terms of maintaining the required voltage and associated stability of the line current. In contrast to cold crash starts this method provides a soft and fast sequential start-up of a large number of cells (50 cells in this case).

Bath transfer

In order to develop a dedicated start-up procedure a number of industrial experiments were carried out in the remaining operating potline. In addition to the minimum required voltage, the current distribution within the cathode collector bars is of main concern.

The liquid bath transfer during start-up of the first cells must be done from Potline 2. If cells are preheated simultaneously, the rate of bath supply is obviously higher than in a normal cell by cell

425

restart. Also, the rate of bath production after start-up is much higher because of the smooth and complete preheating of the cells. With this bath production, the bath levels can increase fast creating the cells with high iron. In this respect the logistics and management of the bath transfers are of critical importance.

(given that the preheat period would extend for a number of days) and also to generate extra bath material. Figure 2 shows the set-up of a cell in Potline 2. The amount of bath material for covering cells in Potline 1 was increased in such a way that the insulation layer was not visible.

Testing In preparation of the start-up of Potline 1, a number of preheating methods were developed and tested. These investigations were conducted in Potline 2 and were focused on the identification of potential constraints in respect to the preheating and start-up. A method was developed using a predefined metal thickness and a flat metal pad to ensure good electrical contact between the anode and the metal plate. An average resistance curve for a cell was determined, which was used as the basis for the ramp-up of the line amperage and the start-up of the cells in Potline 1. At Aldel, two independent mechanisms according the overall line voltage development are important. These mechanisms are related to the preheating curve defined by the gradual stabilization of the anode and cathode current distributions, and the voltage behavior of a cell after liquid bath addition. Both mechanisms determine the balance between the rate of decrease in the overall line voltage due to preheating and the increase in the voltage due to sequential start-ups.

Figure 2: Test setup of cell in Potline 2 Preheat curve Before the start-up of Potline 1 it was very important to get a clear indication of the voltage development related to the actual current and ramp-up rate. First of all, the cells needed enough voltage to generate 110 V across the line, for stability of the line amperage at a controlled target. Secondly, it was important to understand the characteristics of the voltage development during the preheating phase. Experiments showed a fast decline in the voltage during preheating. The rate of decline depends critically on the rate of line current increase, the actual current and the initial current distribution in the cathode bars. The latter itself dependents on the rate of current ramp-up - a more even current distribution is obtained when the current is ramped up more slowly over days.

Test set-up of cell During the shutdown of Potline 1, a strong emphasis was put on the conservation of the cell condition. However, it was found that the surface of the remaining metal pad was still not smooth enough to ensure good electrical contact. Therefore the cathode surface had to be flattened out by casting additional metal into the cell. Figure 1 shows the flattened surface with a number of anodes positioned on the surface in Potline 1.

Because the experiments were carried out in a running potline, the period of the ramp-up was limited to 2 hours. It was found that several connections between the cathode bars and bus bars were lost as a consequence of rapid current increase. Due to the end-toend configuration of the potline most of the current enters and leaves the cell at the ends. However, it was not possible to predict the specific cathode bars causing most of the problems. In some cases a particular collector bar drew 50% of the total current during the ramp-up. It was found that hot cathode bars remain consistently above a line current of 40 kA. These bars had to be controlled with additional cooling above 80 kA. This phenomenon is probably due to the expansion of the cast iron in the cathode block during preheating. Figure 3 shows the preheating curve of a cell tested in Potline 2. Initially the resistance is high - equivalent to 6 V at 120 kA. During the current increase one of the cathode bars disconnected from the cell. After approximately 9 hours at full current, the voltage drops gradually to 3 V. The cell stays at this level for 2 days. The temperature in this period is equally distributed and increased in the cell. The metal reaches an average temperature of approximate 400°C before the voltage decreases to 1.5 V. At this point, the metal is liquid, after which the cell can be started.

Figure 1: Preparation of the metal surface in Potline 1. The anodes were put on top of the flattened metal surface and surrounded by hardboard. Coarse grain crushed bath material was poured between the hardboard and the remaining side ledge to protect and build new side ledge during the start-up phase, and to generate additional bath material for the start-up of other cells. On top of the anodes a thick layer of insulation material and bath material was positioned to prevent anode airburn completely

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The number of cells required for the start-up of Potline 1 were calculated based on the minimum voltage of 110 V. 84 cells were taken into the circuit of which 34 cells were short circuited. These cells have a voltage drop of 0.7 V at 142 kA. Based on the results of the tests 40 cells were required for the metal bake re-start. In order to provide maximum flexibility 10 additional cells were taken into the circuit during the final execution of the restart of Potline 1. Another 10 cells were prepared with anodes on top of the metal, but were not in circuit. These cells could be taken into the circuit if the voltage would fall below the required 110V.

Cathode and anode control at current ramp-up After 4 to 5 hours at 142 kA the anode current distribution became important. The need for inspecting the cathode current distribution was at that point not critical anymore, since it stabilizes very fast after exceeding 90 kA. Anodes that draw too much current were isolated from the anode beam and after a while reconnected again. An anode that did not draw any current was pushed onto the metal plate. Usually these anodes draw current again. During the test it became evident that preventing the anodes from heavy airburn is essential. Otherwise anodes were burned off after 3 days requiring the immediate start-up of the cell. With use of good protection against airburn of the anodes, the preheating time could be extended to at least 7 days which allows gradual heat-up of the full complement of the potline.

Bath flow logistics The availability of liquid bath was one of the major items prior to the start-up, because of the large number of cells that had to be started in a short period of time. In addition, the risk of iron contamination in Potline 2 is considered to be high. To minimize the number of cells for bath production, crushed bath material was used on top of the insulation layer as described earlier. Also, for safety reasons (to avoid transport of liquid bath over long distances) the procedure developed for Potline 1 focuses on its own liquid bath production as soon as possible.

Start-up curve As Figure 3 shows, the voltage decreases after a certain period of time. In order to keep the line voltage above the required level the cells have to be started. The start-up curve determines the actual speed in which cells should be started sequentially to keep the line voltage above 110 V in respect to the start-up of Potline 1. Figure 4 shows the start-up curve of a test cell in Potline 2.

Control of the ramp-up rate It was clear that in the initial phase of the preheating procedure, attention should be paid to the cathode current distribution of 50 cells simultaneously and therefore the ramp-up of the current should be done in a controlled manner. This prevents one or more cathode collector bars or anodes drawing too much current and consequently failing. There are only two mechanisms that can be used to prevent the break down of cathode collector flexes. These are using air lances and stop ramping up the current. Since there were no cells running in Potline 1, the current could be ramped-up very slowly, which minimized large instabilities in the cathode current distribution by providing enough time for the cathode collector bars to thermally equilibrate. The time needed for ramping up the current is limited by the airburn of anodes, but also in the decline of the total line voltage as a consequence of stabilizing cathode bars.

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Figure 4: Voltage development during start up Notice that the current in Figure 4 is decreased to 95 kA before the cell was restarted. During the start-up a number of anodes were disconnected from the beam (clad failures) and had to be replaced during the restart. The last two voltage peaks in the graph are anode effects and related to switching on the alumina feeding.

Organizational aspects of start-up It became clear that there was a need for a highly trained and specialized start-up team for the start-up of Potline 1. This start-up team consists of a number of operators, a start-up leader and a process control person. For the start-up team a dedicated control room was set up to manage the complete information flow. Each cell's status is visualized on a whiteboard.

427

During the gradual ramp-up of the line current in Potline 1 all resources available were focused on measurements of the anode and cathode current distributions. Based on these measurements the current is ramped up or actions are taken to prevent potential problems at either the cathode or anode side. After the current reaches a certain level the cells are started. From that moment on for each shift, two teams are needed to control the preheating, cell start-up and taking cells in normal operation. The first team is the start-up team led by the start-up leader. This team is responsible for the preparation and the execution of the start-up. The second team is responsible for measurements and corrections on the remaining cells under preheating condition. Secondly, this team guides the started cells to their normal mass and energy balance (bath temperature and composition, operating voltage and metal depth). The second team was led by the process control person.

The cathode and anode current distributions were measured every 2 hours. The measurements below 40 kA were all indicative, since the noise during the initial ramp-up was extremely high. Above 40 kA the cathode and anode current distribution measurements were reliable and repeatable. The target current within the collector bars was set to a maximum of 10 kA. Also, the current per anode was limited to 10 kA. Anodes, which did not draw any current, were reconnected to ensure a better contact with the cathode surface. If anodes draw too much current, the anodes were isolated from the beam. After two hours the anodes were put back in circuit again. A critical phase was reached at approximately 80 kA. As shown in Figure 5 three cells were disconnected from the preheating state. These cells became very noisy in the anode and cathode currents. The actual line current was monitored constantly. At 80 kA, it was decided to stabilize for 16 hours to redistribute the anode and cathode currents. Thereafter the current was gradually increased. No problems with the cathode current distribution were found above 90 kA and the main focus shifted towards the control of the anode current distribution. After the metal was liquefied with the anode positioned in the metal, the problems regarding the anode current distribution were gone.

Minimize interference with standard potline operation A start-up of this scope will have its impact on the performance of the other Potline 2. To minimize the interference on the operations of this line, a start-up plan was developed for the actual start-up sequence according to the geographical location of the cells in the potline. Also, the transfer of normal cells to standard operation is taken into account.

The actual line current was monitored constantly. It was found that an increased instability in the current signal could be used to predict noisy and unstable cells. This method worked very well at Aldel and the number of cells that had to be cut-off was limited to the three mentioned above.

Actual execution of start-up Potlinel Ramping up current

Starting up cells

At June 9 2010, rectifier Ell was activated and the current was ramped up very slowly to 40 kA. After that, the other rectifiers were activated and the current was increased to 120 kA. The plan was to increase the current by 5 kA per half hour. In practice the aim was to keep time flexible to give all cells sufficient time to normalize the cathode and anode current distributions.

Although the metal plate was not fully liquefied after reaching 120 kA, the cells were found eligible for start-up. The reason for starting these cells was a high level of noise and the risk of losing anodes due to too high current. After the first cells were started, 23 cells per shift were started. Finally, Aldel was able to preheat 50 and start-up 47 cells in 8 days.

Figure 5 shows the development of the overall line voltage and current during the ramp-up. The green line represents the minimal required voltage.

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possible to lower the voltage in case of high temperature due to the large amount of generated bath material by the cell.

Starting a cell Figure 7 shows the start-up voltage curve of Cell 1035. It was the first cell started via the new procedure at Aldel. The metal plate is visible liquefied after two days of preheating on full line power. The cathode side of the cells reaches a temperature of 400°C.

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The start-up was executed as follows. Under full power, three to four ladles of liquid bath were poured into the cell as shown in Figure 8. After the beam was moved upwards, bath levels went down very fast as anodes went up. The beam was moved up until the voltage went up very fast to approximately 40-50 V. After that the beam was lowered until the cell voltage was stable at 20-35 V. In order to prevent metal and bath mixing, liquid bath was poured in very slowly.

The voltage is kept high to melt the metal. In this respect the rate of melting of metal is important. After the metal melts the voltage drops very fast and the voltage needs to be controlled. When the metal was complete liquid the metal set-point of the cell is determined and the cell was ready producing metal for the casthouse after taken metal analysis. Metal/Bath height 1035

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Figure 8: Liquid bath addition to cell 1035

After 48 hours the cell was eligible for changing anodes. From that moment the cell was covered with crushed cover material and alumina on top. The set-point voltage could be adjusted based on measured temperatures to the required level.

Due to the fact that the metal was not fully liquid by the first cells the bath temperature had to be controlled very well. As shown in Figure 7 the voltage was increased several times to melt the metal. After a few days, when the metal pad was liquid, the control of the bath temperature was easier (Figure 9).

Generation of bath material The generation of bath material of newly started cells went very fast. The generation of bath material went faster then expected and the number of cells that had an increased bath height in Potline 2 could be decreased very fast. This way the number of high iron cells could be avoided.

Guiding a cell to normal operation Initially the temperature is low and therefore the initial voltage should be high or soda ash should be added. A total of 7 MT of soda ash was used during the start-up of the 50 cells. The voltage should remain on a high level to melt the metal completely. If the voltage remains too high for too long, the temperature will reach critical levels. This can cause tap-outs. Sometimes it was not

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Conclusion Preheating and fast restart of a large number of cells is a major and complex task. Thanks to extensive testing and highly motivated and knowledgeable personnel a fast start-up of 50 cells has been developed and executed successfully. A start-up rate of 1 cell per 2 hours is possible, once the gradual ramp up in line current has been completed. Several critical success factors have been identified of which the gradual increase in the line current is determined as the most important factor. Other factors are: Number of cells Critical factors on the number of cells that can be restarted simultaneously are: resources available for anode and cathode current distribution measurements, anode airburn control and the number of started cells that can be controlled at the same time during the early operation. Liquid bath transfer Especially at the start, the availability of liquid bath is another key factor. After the first cells have been started-up, fast generation of liquid bath enables fast start-up of the already preheated cells. Therefore it is critical to ensure a fast production and transfer of liquid bath. Start-up team Two separate teams that are responsible for the actual start-up and early operation respectively are needed for a good restart. A startup team that has the focus on starting-up cells and a team that is responsible for aftercare. The aftercare team focuses on control of bath levels and bringing the cells to their normal heat balance. Cell life time It is expected that cell life is increased by doing preheat before start-up. By mid November 2010 no early failures occurred at any of the 47 cells that were restarted with the new method. In retrospect it was evident that the soft restart method described in this paper resulted in improved process performance and longer cell life. Furthermore, it is possible to start-up a large number of cells because the method is easy. This raises the question about the design of a potline related to the lay-out and number of cells in respect to flexible operation. In general smelters tend to have as many cells as possible in a potline, but this makes it more difficult to frequent shutdowns and start-ups. References [1] Vaillancourt, D. and Kazadi, J., "Metal bake restart", Light Metals 2000 Metaux Legers, Edited by J. Kazadi and J. Masounave [2] Postgraduate Certificate for Light Metals Reduction Technology 2006, University of Auckland

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SWOT PERSPECTIVES OF MIDAGE PREBAKED ALUMINIUM SMELTER (Case studies of State run Smelter of National Aluminium Company Limited, Orissa, India) P.R.Choudhury1,A.K.Sharma2 Executive Director, Smelter & Power Complex, NALCO, 2Director, Production, NALCO ISmelter & Power Complex , Angul, Orissa, India, Pin 759145, 2NALCO Bhawan, P/1, Nayapalli, Bhubaneswar, India, Pin 751061 Keywords: Aluminium, Smelter, Process strength lay in its technological superiority, green field project from mining to power generation & smelting, and the elite work force in the average age group of 25-30 in the beginning years.

Abstract Aluminium smelters in nineties witnessed radical changes in technology of electrolysis, carbon manufacturing and casting. The pace of transformation in aluminium industries posed different challenges specifically to the mid-age smelters operating with relatively lower amperage. Enduring needs to match the demands of evolving technology, socio-cultural and environment issues forced mid-age Smelters to adopt appropriate strategies as long-term business imperatives. With structured approach, the threat perception could be turned into vast opportunities of innovation and improvements. Energy conservation, waste management, recycling, emissions control and customer orientation remained specific focus areas to enhance productivity and retain profitability in the backdrop of global economic recession. Process optimization and system upgradation through customised solutions, motivating and retaining high employee morale became the order of the day. The paper is a case study on the sustainable achievements of the state owned Indian aluminium major, NALCO, through well-coined strategies and pragmatic investment.

NALCO Smelter plant located at around 150 km away from Bhubaneswar, capital city of Orissa started production in 1987 with two pot lines of 240 pots each operated at approximately 180 kA for production of around 220, 000 MT of aluminium per year. In 2003, a third pot line was added with 240 pots of AP 18 technology. Production capacity enhanced to 345,000 MT of aluminium per year after commissioning of the third pot line operating at 183 kA. Additional capacity was built in anode paste plant and cast house for supporting another pot line of 240 pots. In 2009, a fourth pot line was added with 240 pots of AP 18 technology. Production capacity increased to 460,000 MT of aluminium per year after commissioning of the fourth pot line operating at 184 kA. Figure-1 shows the gradual increase in production capacity achieved during last decade.

Introduction

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The innovation in bauxite refining by "atmospheric digestion" revolutionized the era of alumina extraction from lateritic bauxite. India constituting 7.5 % of global bauxite deposit has most of its reserve in "Eastern Ghat" region. The government of India fully aware of the market potential formed "National Aluminium Company" with an initial thrust on mineral exploration and selfsufficiency of aluminium in the subcontinent. The company, which was formed in 1981 with adaptation of contemporary "state of art AP-18 technology", surged to production in 1987.

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The Company within a short spell had the enviable track record of exporting its product worldwide and has the distinction of being among the cheapest producers in the world. Company's core

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The strength, weakness, opportunity and threat of a mid age smelter is a complex business model with its diversified dimensions of risk and opportunity. While the organization has the advantages of knowledge and experience as its capital, the challenges of technological obsolescence, radical changes in the market, and reforms in the statutory & regulatory requirements are major bottlenecks in maintaining the pace of growth of the smelter. Therefore, the constraints posed by mid-age smelter offer tremendous opportunities as well as introspection to improve processes and bottom-line of the company.

Critical strategies in the areas of energy conservation through implementation of Clean Development Mechanism (CDM project), selection of energy efficient drives, improving current efficiency in potlines, lowering specific consumption, waste management, emissions control, and water resource management provided the much needed impetus to statutory and regulatory requirements. Creative & innovative pursuits like quality circles, small group activities, quality improvement projects, extensive inhouse training addressed the core issues of retaining a high employee's motivation and morale.

Smelters experience a series of challenges as they mature through the learning curve. The degree of such challenges led series of improvements, changes, innovations in every facet of smelting operation in NALCO. The distinctive changes in potlines that include operating at higher anode changing shift cycle, alumina feeding to pots, improvement in pot-lining, thermal equilibrium of pots, treatment of gases based on fluorine emission models, migration to higher kA and improved pot regulation made enormous impact to the smelter's operational efficiency. The improvements achieved in the aluminum fluoride consumption, reduction in net carbon consumption, significant improvement in pot life bears the testimony of the success through better process management.

NALCO has always been a trend setter for Indian Aluminium industry in caring for the community. Much before the word CSR was coined, NALCO has adopted many community projects like building roads and schools, providing drinking water and medical facilities in the surrounding villages matching the highest standards of corporate citizenship. Specific focus on CSR through peripheral developments brought in confidence of the society that helped the organization to excel.

Diversification, capacity augmentation and quality improvements in raw materials for carbon plant, standardization of sources, upgradation in mixer, screens, vibro-compaction in green anode manufacturing, changes of heating & regulation in anode baking, de-bottlenecking in rodding shop contributed significantly to the consistency in quality of anodes for potlines. The structured approach adopted for improvement in the green and baked anode quality through selection of sources, changes in the granulometry, improvement in kneading & vibrocompaction led to appreciable reduction in net carbon consumption. A commendable alignment with customer focus through innovative casting and downstream facilities, mechanization and technological improvements adapted in furnaces, wire rod mills, ingot casting, and addition of new product portfolios created competitive advantages in terms of cost and product quality. The capacity addition in strip casting, air slip billets, rolled product added company's new market segment in the areas of down streams.

NALCO smelter invested substantially to protect the environment by continuously upgrading the bag houses, dry scrubbing units, floor cleaning units and fume treatment centers. Like any other mid age smelter, NALCO had to address the gargantuan issue of disposal of SPL (Spent potlining) material. The solution adopted was engineered landfill and exploring the possibility of using the carbon portion with high calorific value as fuel and the refractory material as a raw material for cement industry. However, disposal of SPL still remains a challenge for NALCO as much for other smelters. Strategy followed by NALCO smelter during last decade was not only to sustain the production capacity but also to gradually increase it. This could be accomplished by converting the threat perception during global recession to challenges and resultant opportunities for growth. Increase in production capacity was achieved through continuous process evaluation, control of process inputs, and up-grading of processes and equipment etc. At present amperage increase programme has been initiated in existing smelter to migrate from 185kA to 220kA or more. Complete migration will take around 6 to 7 years and addition of 110,000 MT/Year of aluminium production capacity shall be achieved through a seamless migration.

1.

Carbon Plant

Right from the inception, anode quality was a major constraint in pot line operation. Variation in anode density was the major issue along with high rate of rejection after baking. Physical & Chemical properties of anode were also not consistent resulting in high consumption of net carbon. Variation of anode density was addressed by improving the quality of coke and pitch used for anode. Use of liquid coal tar pitch, introduction of mesophase parameters and improvement of the process at the source by the Smelter quality group improved significantly the quality of Liquid pitch. Similarly, through the bench scale and plant scale trial, source of the coke as well as the granulometry was established to meet the requirements of desired physical and chemical properties. The experimentation on raw materials blending and segregation of sources of supply established the consistency in quality parameters. Green anode plant equipment of coke fractionating circuit were upgraded to maintain consistency of coke fractions. All mechanical screens were replaced with electro-mechanical screen to have control over the screening efficiency and output. Speed variation in proportioning circuit equipments was controlled by installing variable voltage variable frequency drives to minimize the effect of grid frequency variations. Controllers of the proportioning circuit were replaced to improve regulation of the proportioning system. For accurate fines regulation in the dry mix, a weigh hopper was installed replacing the twin hoppers. A mass flow meter was installed replacing twin weigh balloons to improve accuracy of pitch regulation. Control and automation equipment of the green anode plant was replaced with new generation automation equipment with advanced user interface. Mixing tools of the paste mixer were regularly replaced to maintain mixing quality. Mechanical suspensions of old vibro compactor were replaced by pneumatic suspensions for smooth compaction. Spray nozzles were replaced in anode cooling tunnel to increase water flow and a water filter was installed to improve recycling efficiency and reduce nozzle choking. Heat regulating system of baking furnace was replaced to up grade the system for better control of the anode baking process. After installation of the new system level of baking could be effectively monitored and consistency in baked anode quality could be ensured. Twelve more sections were added to increase the number of sections from 72 to 84 in old baking furnace. Time Lag created between two

consecutive fires in four fire configuration was effectively utilised for cooling of anodes and better refractory maintenance of the furnace considering higher ambient temperature. With all these modifications substantial improvement in anode quality were achieved that led to reduction in net carbon consumption (see figure-2). At the Rodding shop, large scale modifications in the stacking crane, chain conveyor, casting zone, anode extractor, hooking & unhooking were carried out to improve the reliability of the plant. Apart from this, the bimetallic clad was replaced with trimetallic titanium clad to reduce the voltage drop.

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Reduction Area

To reduce the cell voltage drop, the first generation alteration was introduced with replacement of anthracite cathode blocks to semi graphite cathode blocks. NALCO smelter replaced all the cells having anthracite blocks to semi graphite blocks by the year 2002. Thus NALCO was able to reduce its energy consumption by reducing the cell voltage. Further to reduce the cell voltage and to obtain better process parameters NALCO smelter has upgraded its first generation pot regulation system from pot microcomputer to advanced HMI pot regulating and feeding system. The advanced regulation system helped the operating personnel to ensure effective decision making and better work organization. The distinctive feature of the new regulation system [2] encompassing "thermal balance control", "alumina feed rate", "measurement of pot resistance & rate of change of resistance" led to reduction in energy consumption and significantly improved the process and health of the pots. Based on the knowledge based information and trouble shooting guide, standard work practices

(SOP) were dynamically changed to suit the operational needs. Large scale revamping was carried out in critical pot equipment like PTMs to enhance the overall equipment effectiveness and reliability. The major modifications included replacement of the vane compressor, introduction of VFD in LT & hoist drives, metal siphoning system, and alumina feeding system. To improve the competence of the working personnel, regular interaction classes at the shop floor were organized and mentoring was initiated to guide newly recruited. This resulted in better understanding of the pot parameters by the operating personnel and they could assess the consequential impact on processes and metal quality. The consolidated approaches led to substantial reduction [1] in number of A.E./day/pot (See figure-3), decrease in average noise level/day which leads to reduction in average cell voltage and better C.E {Current Efficiency) (See figure-4). All the above improvement contributed substantially to abatement of GHG emissions.

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The regular cleaning of the cathode clad was introduced in the SOP as a lot of crushed bath and alumina covers the cathode clad and the pot shell which obstructs the heat flow from the pot shell/clad to the ambient atmosphere resulting in thermal imbalance. The pot ventilation grills were regularly cleaned to facilitate heat dissipation.

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In-house design modification carried out in stem brushing machine at Rodding shop improved the stem surface cleaning of any types of stems therefore ensuring better contact and reduction in cell voltage drop. Proper cleaning of stem surface reduced the contact voltage drop. To decrease the contact drop further between anode and pot beam, older clamps were replaced with better clamps which helped in reducing the contact drop. It also helped to decrease the noise level of the pots as well as helped in reducing the number of clad fails in pot line. Regular greasing of the clamps was introduced in the SOP which led to fewer numbers of clad fails, which ultimately resulted in reduction of average cell voltage. (See figure-5)

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Power Crisis

Aluminium smelter being power intensive, the requirement of large sustainable power source is

434

primary requisite for its survival. Despite the availability of a captive power supply system, the smelter in the recent times has been facing the problems of power interruptions. The power plant is dependent on coal supply which is controlled by the Government owned mines. To cope up with fluctuation of power, operating the pots in lower amperage as well as switching in to higher amperage as and when power is available has become part of the improvised process.

NALCO installed an additional rectifier group to strengthen the substation and planned overhaul of all the transformers to extend the reliability of the existing facility. Overhauling of twelve transformers completed in the year 2009-10 without affecting the potline operation although the substation witnessed failure of one half of a rectifier group during the year. With the experience of installing one swing rectifier group the substation engineers are planning for installation of additional two rectifier groups within the same substation area for supporting 220 KA up gradation program,

For a smelter to increase its production capacity the most critical issue is availability of required input power. Enhancement of power generation capacity becomes imperative for smelters having own captive power plant. NALCO smelter expansion was carried out in stages matching with enhancement of power generation capacity. However, with increase in generation capacity coal supply became unstable. Hence the power generation could not be ensured to maintain a consistent level of power supply to smelter. The trade off was between reducing number of pots by shutting down rurining pots or sustaining operating pots at lower amperage. Both the options were tough considering the requirement of meeting annual production target, maintaining pot parameters, minimizing effect on pot life, etc. NALCO Smelter preferred the second option to retain all the pots and reduce pot amperage. The experienced pot line operation group worked hard to set parameters for operating pots at amperage less than 175 kA and even to sustain the pots for a considerable long period of time at the lowered amperage of 160 KA. Problems related to power outage and amperage reduction like increase in instability of pots, disturbed bath chemistry, disturbed thermal regulation, fluctuating bath temperature, dusting of carbon, increased cathode voltage drop, increase in set point resistance were successfully addressed. Consistent operation of pot lines yielded optimum production quantity and smelter could add number of pots matching to increase in power generation capacity.

4.

Alumina is transported by railway network to NALCO smelter located at around 700 km away from the alumina refinery. Railway rakes are unloaded at smelter end by using air slides to belt conveyors and the alumina silos are fed trough the belt conveyors. To establish reliability of alumina supply to the pot line alumina unloading and transport system are being augmented by installation of another unloading and transportation system. Coke storage facility was enhanced by addition of one coke silo of 6,000MT capacity. Coke handling facility was retrofitted to facilitate feeding to the new silo and extraction from the new silo. However, the coke unloading and feeding system needs to be mechanized to enhance efficiency of the system. With expansion of the plant, liquid pitch storage facility was augmented by two new storage tanks of 300 MT capacity each. However, consistent pitch supply could not be maintained due to distant location of the liquid pitch suppliers. To ensure reliability of liquid pitch supply to the green anode plants two additional pitch tanks of 300 MT capacity each were constructed with in house engineering. The new unloading and storage facility is designed with complete automation for quick unloading of road tankers and ensure interConnectivity with all the existing pitch storage tanks to facilitate reliable supply of liquid pitch to both the green anode plants. NALCO uses heavy furnace oil (HFO) for its melting & holding furnaces and baking furnaces. Smelter is planning to augment its HFO handling facility to accommodate larger railway tankers for unloading and storage.

Limitation of infrastructure

Ageing infrastructure like rectifier groups of substation, alumina transportation / unloading and storage facility, coke / pitch handling and storage facility, HFO unloading & storage facility, coal handling and transportation system of captive power plant, ash disposal system were the areas of concern to sustain the smelter operation.

Matching with the expansion of smelter the generation capacity of the captive power plant was enhanced. Coal supply became a constraint as NALCO was receiving major portion of coal from state owned mines. The coal supply was augmented by introducing washed coal and imported coal. However, to find a long term solution NALCO

435

ventured into coal mining activity. One coal block is under development to start coal mining in full swing. Ash disposal is another area of concern for most of the thermal power plants. NALCO is the pioneer in implementing zero discharge technology for ash disposal among the power plants in India. Ash disposal system shall further be augmented by establishing lean slurry disposal system to fill up coal mine empty pits.

5.

Addressing Meltdown in LME Price:

Metal price was never considered as a threat for operation of NALCO smelter as the company has been extremely doing well in its financial performance from the first year of operation. From inception, NALCO maintained a robust financial performance due to its core strength of integrated operation from bauxite mining to export of aluminum metal through its captive sea port. The company ensured a dividend payment of more than 40% for all the years since 2000-01 and the dividend payment was as high as 75% in the year 2006-07. Financial performance in terms of net profit as % of sales turnover of the company is presented below for reference. (See figure-6)

Figure 6 NALCO started experiencing the effect of economic recession in third quarter of 2008-09. Smelter operation was caught between falling market price and increasing production cost. The temptation to limit number of pots in operation was also very high looking at the fluctuation in power generation capacity. It was a Catch 22 situation whether to close down some of the operating pots to reduce projected losses considering falling LME price or to avail the opportunity of adding 120 new pots under expansion for future output growth. When few of the Smelters worldwide were closing down entire potlines even Smelting operation, NALCO decided to commission the new pots and increase the production capacity to

460,000 MT, spurred by relatively low cost of production. Rigorous cost reduction measures in the areas of specific consumption of input materials including power and fuel, optimization of furnace operation, operating at higher anode changing shift cycles, recycling of waste, management of scrap generation and disposal paid off the company's risk of operating higher numbers of pots. Thrust areas were identified considering all controllable factors contributing to cost of production. Action plans were prepared and implemented in each identified area. As a result, the year 2009-10 witnessed the lowest specific consumption of major raw materials including specific power consumption. NALCO could maintain profitability due to increase in production volume although margin of profit reduced due to fall in market price.

Conclusion Aluminium industry in India is entering into a volatile phase with most of the producers adding new smelters to buildup production capacity. Although falling aluminum price in the year 2009-10 affected the in swing projects, the new smelters will start production within a couple of years. To remain competitive and retain its market share, NALCO is pursuing both brown field and green field expansion plans. The company is also aiming to be an independent power producer by 2015. This would definitely provide competitive advantage for the smelting operation. The mid age Smelter poses diversified challenges and constraints that open up new horizon of opportunities to excel. A holistic and structured approach in every facet of business dimension is needed to overcome the challenges and threat and bring in parity with new smelters equipped with high end technology & resources. The AP-18 smelter in Nalco is a paragon of case study describing the gradual journey of transformation accomplished over a period of one and half decades. References [1] - Homsi et al., Overview of process control in reduction cells and potlines (Light Metals, 2000), pp. 223-230 [2] - Sylvain Fardeau Continuous Improvement in Aluminium Reduction Cell Process Performance Using the ALPSYS ® Control System (Light Metals, 2010), pp. 495-499

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INTEGRATED APPROACH FOR SAFE AND EFFICIENT PLANT LAYOUT DEVELOPMENT Rafael L. Pires, Robert F. Baxter, Laszlo Tikasz, Robert I. McCulloch Bechtel Corp.; 1500 University Street, Montreal, QC, H3A 3S7 Canada Keywords: Layout Development, Traffic, Lean Design, Simulation the characteristics of the layout, identify roads, intersections and assign routes to traffic and calculate average inter-arrival times.

Abstract Aluminium smelter layout development usually involves dealing with an integrated mix of operations and services (i.e. smelting, carbon anode formation, metal casting, material handling, etc.). The choice of a layout can significantly impact the success of the envisaged operation, in terms of safety, life-cycle cost and environmental impact.

Further analysis using this evaluation method revealed that higher level safety and efficiency awareness was needed, and that it should be based on a better understanding of the dynamic characteristics of plant traffic flows. Discrete-event modeling (DEM) was chosen as a natural fit for the People-ProductMaterial (PPM) traffic flow problems. An example of DEM applied to traffic analysis is shown below in Figure 1 applied to a shuttle bus station during an envisaged shift change period.

This paper is a continuation of last year's publication, exploring the application of safety by design and lean manufacturing methods to the layout design of a fully integrated aluminium complex. An innovative approach, derived from lessons learnt assisting various plant layout development and resource analysis, is presented. This is an innovative approach for predicting traffic characteristics for an aluminium smelter configuration and measuring it in terms of safety, cost and environmental impact. This approach allows comparative analysis of competing layout designs and improvements to be implemented quickly to deliver a plant configuration that is Safe, Lean and Green. Throughout this paper, the design of access roads, choice of transportation modes, and planning of resource are discussed.

Introduction The Aluminium Center of Excellence (ACE) is the repository of Bechtel's institutional knowledge, technical capability, historical information and lessons learned on the design and construction of smelter projects. The mandate is to deliver value to projects by applying the above knowledge and skills focusing on sustainable design. The integrated approach presented below was developed by ACE and funded by a Bechtel internal technical research grant.

Figure 1 - Shuttle bus station at a Plant Previous case studies demonstrated the complexity of different issues that arise through, layout development. Based on DEM, a systematic approach, here referred as the integrated approach, was created. The application of this approach automates data transfer and reduces cycle time for the assessment a layout safety, efficiency and C0 2 equivalent (C02e) emissions related to the traffic network.

The choice of a layout can significantly impact an operation's long term success, both in terms of safety and its ability to compete successfully in the marketplace. In addition, investment costs associated with building a particular layout are substantial. Early development and finalization of the layout has a clear advantage from both a cost and schedule perspective to the project as compared to finalizing the layout later during project execution.

The adoption of the integrated approach has proved to deliver value by quickly simulating and analyzing a particular plant layout and assessing the effectiveness of proposed improvements. Layout Development

Layout development requires complete understanding of all operational aspects of a particular plant. The operation of an aluminium smelter relies on interactions between "customer" and "supplier" facilities to convey People, Products and Materials, with the road network having a direct impact on safety and efficiency of the operation. As a consequence, a quick and effective approach of assessing safety and efficiency of layouts early during the design phase is required.

It is well-known that a "traffic intense" plant layout not only impacts the project capital cost but more importantly it impacts decades of plant operations. The development of a safe and efficient layout requires a collection of tools and expertise in order to design out the potential for accidents, waste and uncertainties related to conveyance between customers and suppliers. Lean manufacturing and six sigma tools provide guidance to streamline layout development initiatives. Proper application of these tools during early project definition helps to "freeze" the layout early avoiding re-work, minimizing risk and providing

Typically, layout development of "traffic intense" facilities such as aluminium smelters, involve a detailed understanding of traffic interactions and behavior. A known approach is to statically map

437

The new platform automates data transfer and speeds the assessment of layouts. Figure 2 outlines the structure and the connections between the phases of layout analysis.

certainty of outcome to the design team and future plant operators that their layouts are Safe, Lean and Green. Safe - Safety by Design Safety is an integral part of corporate social responsibility. There is a clear connection between operational safety and quality achieved (6). A safe plant configuration lays the early foundations for developing an organizational culture that values behavioral based safety. Safety by Design applied to layout development has the objective of minimizing the potential for accidents. Safety-by-design reduces the likelihood of accidents over the life-cycle of a plant. As referenced above, Discrete Event Modeling (DEM) complements layout development and optimization studies. It provides a way of mimicking and monitoring the expected operation and collecting data for further analyses. A Failure Mode and Effect Analysis (FMEA) can then be applied to estimate the risk of accidents at any particular intersection in the road network.

Figure 2 Integrated Approach The integrated approach was assembled by coupling and partially integrating proven tools such as CAD drawings, Plant Basic Design Data (BDD) model, People, Product and Material (PPM) flow mapping, dynamic modeling (Flexsim® - Simulation Package), Failure Mode and Effect Analysis, and generalized Comparative Cost Function (CCF). The projected plant layout and plant operation parameters are automatically collected to and linked in a dedicated input Data Set. Then, this data set is used to specify, construct and run the plant dynamic model. The simulated plant operation is monitored, data from simulated scenarios collected and automatically exported to the same Data Set.

Overall plant layout safety is directly influenced by a variety of operational factors, such as: intersection and vehicle type, schedule and transportation mode. Intersections are designed and scored using FMEA, by severity, occurrence and detection factors that are combined into a Risk Priority Number (RPN). RPN is the resulting safety measure for any particular point of a layout. Lean - Lowest Life-Cycle Cost Design of an efficient operation requires the application of lean principles and tools, in order to identify and eliminate waste and streamline the flow of people, product and materials. In theory, efficient operations have the capability of maximizing productivity with minimum waste, effort or expense. However variations generated by transient operational conditions directly impact and reduce efficiency.

The platform provides quantitative analysis taking into account the dynamically changing conditions of the projected traffic (i.e. schedules, operational constraints, etc.). Emphasis is given to: movement of People, Products and Materials (i.e., distance traveled per vehicle/person, intersections crossings, unsafe conditions and possible bottlenecks on the road network).

DEM is used to predict the dynamic response of a particular operation including traffic and material conveyance and storage, to ensure that the proposed configuration can meet customer needs during normal, maximum and upset operating conditions. During overall layout development, DEM is used to mimic traffic flows and operations to validate the number of vehicles, load/unload stations, inspection lanes, parking lot requirements, inventory requirements, etcetera, to de-risk the overall plant layout design, offering a lower life-cycle cost.

FMEA & Comparative Cost Functions, linked to the Data Set, score the projected layout in terms of safety and efficiency. The main outputs of the platform are the factors representing risk of accidents in each specific intersection as well as an overall safety score for the layout, measures of efficiency (cost; utilization of vehicles & load / unload stations, etc.) and C02e emissions related to traffic. Moreover, the Data Set provides complete post processing of the data simulated and collected. Some of the outputs include:

Green - Environmental Impact The environmental performance of a smelter layout is influenced by the movement of People, Products and Materials. The design of the road network, conveyance between customers and suppliers, and the segregation of traffic types impact the distance driven by vehicles. As a consequence, the emissions related to traffic are also impacted. During DEM simulation, confirmation of the number of vehicles required and the distance driven per vehicle is collected, not only to understand the operational cost related to traffic, but also to estimate C02e emissions.

> > > > > > > > >

Integrated Approach

Layout Safety Score: Average RPN from FMEA Layout Efficiency Score: Operating cost, related to traffic Layout Emissions Score: C02e Emissions related to traffic RPN of each intersection Number of vehicles required per flow type Distance driven per vehicle Number of trips executed per flow type Utilization of facilities (i.e. load / unload stations) Utilization of vehicles

Optimum benefit would result from implementation of this approach early during project development.

The quick and effective development of a Safe, Lean and Green layout required the integration of the various concepts, previously mentioned, into an automated platform.

438

Figure 3 - Overall Scoring (Sample) Figure 3 shows a sample of the overall score that is generated by the platform once a few layouts are tested. A scenario could represent any modification impacting the road network such as: plant arrangement, number of vehicles, schedule, route, number of stations, etc.

Case Study The approach has been applied to an aluminium complex consisting of an alumina refinery, an aluminium smelter and an integrated rolling mill. The application focused on smelter traffic interacting with traffic from all facilities sharing the same external road network to cities and ports. The understanding of traffic patterns was crucial to validate the aluminium complex operation.

Figure 4 - Mapping offlowof PPM Once the team agreed on the list of flows and routes, the delivery schedule on each flow was defined. As part of the scheduling activity, operations such as metal tapping and anode change were executed in accordance with the expected potroom operational schedule. Annual requirements for each material were taken directly from the linked Basic Design Data model. The communication between the platform and the DEM shell is automatically executed. The discrete event model is created on the input data with no interaction required from the user. Then, the simulation was performed and predicted output data recorded.

Due to the sensitivity of the information from the above referenced project, the results presented are based on hypothetical data and plant configuration for demonstration purposes. The hypothetical results presented below focus on the approach used and the results of the analyses. The integrated approach was used to analyze the proposed layout. A summary of the key questions that the model was tasked to answers were: > > > >

Identify potential for accidents related to traffic; Identify any bottlenecks impacting operational efficiency; Validate the number of load / unload stations, lanes at the security gate and weigh stations; Validate the number of vehicles required.

A one week simulated operation was initially targeted to cover the mandate. The study started by composing a list of flows expected during a typical week of operation. Then particular routes were defined and assigned to flows. The trajectories of people, products and materials (PPM) were automatically plotted, by the platform, on the CAD drawing,

Figure 5 Example: DEM Model A typical vehicle intersection is shown in Figure 5; visualization was kept to a minimum (e.g. colored boxes, are shown instead of nice 3D objects) to keep the focus on operational problems.

The flow of PPM helps to align the project team (designers and future operators combined) on the expected vehicle movement. It raises attention to the mix of flows in any particular layout and eventually becomes the severity factor for the FMEA analysis. The flow of PPM identified potential areas of concern that would eventually be analyzed through the simulation. See Figure 4 on specifying routes (multy-colored lines) on the original CAD drawing.

Figure 6 - Example: Smelter Model

Figure 6 shows a snapshot of the smelter where various flows are active. To bridge the gap in size-differences of vehicles and the whole complex, marker strings were used to visually locate the tiny cars when the view was zoomed out.

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Utilization of trucks weigh stations; Utilization of TAC / Skim stations at the casthouse; Operating cost related to traffic.

The model validated the design of all load/unload stations, weigh stations and security gates. Figures 9 & 10 provide a sample of the results generated by the model.

Once the results are generated and post-processing is executed, an analysis follows to answer the key questions and to identify safety and efficiency issues.

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Safety Results Once the FMEA analysis was performed, a Risk Priority Number was generated for each intersection; the highest score on each is presented in a Pareto-style chart. See Figure 7 for the RPN values and Figure 8 for the location of the corresponding intersections. SAFETY- Intersection RPN 1803

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The safety result is comparative and allows the team to focus on intersections with the highest risk for accidents. Considering the case study, intersections #15, #13 and #10 scored the highest RPN's. The following solutions were recommended:

As part of the validation process, a variety of scenarios were also run to understand the impact of an accident in the road from the port to the complex or the effect of a late ship arrival followed by a very early ship arrival.

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Intersections #15 and #13 required redesign and truck deliveries were not allowed during shift change. Deliveries to the carbon plant were re-routed to avoid interferences with alumina deliveries at intersection #10.

Conclusions The integrated approach has been conceptualized, developed, tested and then applied to a particular project, validating the proposed layout in terms of safety and efficiency. This innovative platform delivered value through the application of safety-bydesign and waste elimination.

Efficiency Results

Early findings indicate that the integrated approach reduces the cycle time of aluminium smelter layout analyses by up to 50% compared to a sequential, non-integrated, approach.

The analyses on utilization focused on the following aspects: > > >

Number of lanes at the gates; Number of lanes on major roads; Utilization of load / unload stations;

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Trademarks AutoCAD is registered trademark of Autodesk group. Microsoft and Excel are registered trademarks of Microsoft group of companies. Flexsim is registered trademark of Flexsim group.

Acknowledgements The authors thank Bechtel for the funding to develop an integrated approach for plant layout development and permission to publish this paper.

References 1.

Laszlo Tikasz, C. Mark Read, Robert Baxter, Rafael L. Pires, Robert I. McCulloch; "Safe and Efficient Traffic Flow for Aluminium Smelters"; Pages 427 - 432. Light Metals, Aluminium Reduction Technology, (TMS 2010).

2.

Myer Kutz; Handbook of Transportation Engineering. McGraw-Hill.Online version available (July 2010) at: http-J/knovel. com/web/portaUbrowse/display ?_EXT_KNO VEL_DISPLAY_bookid=2534&VerticalID=0

3.

C. Mark Read, Robert I. McCulloch, Robert F. Baxter, "Global Delivery of Solutions to the Aluminium Industry". Smelter's Needs, Part I, Pag 31, 45th International Conference of Metallurgists, MetSoc of CIM, (COM 2006.)

4.

Bassem El-Haik, Raid Al-Aomar, Simulation-based Lean Six Sigma and Design for Six-Sigma, Part II, Page 71, (Wiley-Interscience, 2006)

5.

James Tompkins, John White, Yavuz Bozer, J. Tanchoco, Facilities Planning, Chapter 10, Page 639, (John Wiley & Sons Inc, 2003)

6.

Veltri Anthony, Pagell Mark, Behm Michael, Das A Ajay Data-Based Evaluation of the Relationship between Occupational Safety and Operating Performance; Journal of SH&E Research, Vol 4, Num 1; (The American Society of Safety Engineers.)

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

New Progress on Application of NEUMOOkA Family High Energy Efficiency Aluminum Reduction Pot ("HEEP") Technology Lu Dingxiong1, Ban Yungang1, Qin Junman2, Ai Zijin2 1. Northeastern University Engineering & Research Institute Co., Ltd. No.73, Xiaoxi Road, Shenhe Dist, Shenyang, Liaoning, China 110013 2. China Non-Ferrous Metal Industry's Foreign Engineering & Construction Co., Ltd. No. 10, Anding Rd. Chaoyang District, Beijing, China 100029 Key words: Aluminum reduction pot; NEUI400; high energy efficiency; new application progress Abstract NEUI is the first to have successfully developed NEUMOOkA Family High Energy Efficiency Aluminum Reduction Pot ("NEUI400 Family HEEP") Technology by incorporating the numerical simulation technology with the experiences on developing high amperage aluminum reduction pots. And the first 230kt/a potline adopting NEUI400(I) HEEP technology in China has been put into commercial operation in Aug., 2008. At present, the operating amperage of the potline is 415kA, average pot working voltage is less than 3.85V, DC energy consumption is less than 12500kWh/t-Al, and the anode effect frequency is less than 0.015effects/pot-day. On the basis of summarizing experiences on developing NEUI400(I) HEEP and analyzing the measured physical field results, NEUI has optimized the physical field of NEUI400(I) HEEP by using its proprietary "Physical Field Numerical Simulation and Analysis Software Package for Aluminum Reduction Pots". Meanwhile, NEUI has also developed some new technologies such as sub-section high level gas collection technology and compressible pot lining technology, etc. NEUI has successively developed NEUI400(II-IV) high energy efficiency aluminum reduction pot ("NEUI400(II-IV) HEEP") technologies which have been used in Linfeng Aluminum and Power Co., Ltd., Shandong Nanshan Aluminum Co., Ltd. and Jinning Aluminum Co., Ltd respectively. The design capacity of the three potlines is respectively of 250kt/a, 250kt/a, and 300kt/a. The operating amperage has respectively reached 430kA, 440kA, and 460kA. And more excellent techno-economic indices have achieved such as average pot working voltage less than 3.85V, DC energy consumption less than 12500kWh/t-Al, anode effect frequency less than 0.01 effects/pot-day and current efficiency up to 94%, etc.

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1. Introduction The aluminum output of China has been No. 1 in the world for years. And aluminum industry technologies are also galloping ahead. At present, 300-400kA family aluminum reduction pot technologies have become the main pot technology in China's aluminum industry. China has made great break-through on pot design, fabrication, reduction process and technical innovations of the aluminum reduction pot, which have brought the continuous improvement of the techno-economic indices for aluminum reduction. As an important technology supplier and service provider in China's aluminum industry, NEUI has made great contribution to the rapid development and improvement of the technological level of China's aluminum industry. By incorporating the numerical simulation technology developed by itself with the experiences on developing high amperage aluminum reduction pots, NEUI is the first to have successfully developed NEUI400(I) HEEP technology and first applied it to aluminum industry. China's first 400kA reduction potline with capacity of 230kt/a adopting NEUI400(I) HEEP technology has been put into operation in Aug., 2008, and excellent techno-economic indices were obtained. With successful application of NEUI400(I) HEEP technology, the situation has been terminated that high amperage aluminum reduction pots have to be developed by industrial test and intensify amperage step by step. The development approach has greatly reduced the cost and time for developing high amperage aluminum reduction pot technologies. At present, the actual operating amperage of the NEUI400(I) HEEP potline has reached 415kA, the average pot working voltage is less than 3.85V, DC energy consumption is less than 12500kWh/t-Al, and anode effect frequency is less than 0.015effects/pot-day. At 2010 TMS held in Seattle, NEUI has presented the main technical features, start-up

and operation conditions of NEUI400(I) HEEP which can refer to the reference [1] and [2] for detail. On Aug. 11th, 2008, NEU1400(I) HEEP technology successfully passed the appraisal during the Scientific and Technological Achievements Appraisal Meeting held by China Non-Ferrous Metals Industry Association. All experts at the meeting have the common comments that the overall technologies and equipment of NEUI400(I) HEEP has reached the world leading level. Fig. 1 and 2 are respectively the potroom and average pot working voltage of NEUI400(I) HEEP potline in Henan Zhongfu Industry Co. Ltd.

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high level gas collection, compressible pot lining and MHD stability. 2.1 Sub-section High Level Gas Collection Technology NEUI has developed the sub-section high level gas collection technology by using "Simulation Software for Negative Pressure Balance of Pot Gas Collection System" which is redeveloped by NEUI based on the CFD (computational fluid dynamics) software platform. Comparing with the traditional hooding structure, the new hooding structure was lifted above the horizontal cover board, and the new hooding structure has the following advantages: • The spaces within the superstructure and inside members can efficiently be used; • After lifting of the pot hood, not only the negative pressure generated by the hot fume can be fully utilized, but also enough operation space can be provided for process operations such as anode change, etc.; • More uniform distribution of negative pressure, flow velocity and temperature inside the hood can be obtained, which is in favor of improving the hooding efficiency. • Effective gas collection can be achieved through using the new hooding structure. Then the fume exhaust volume and the load of the scrubbing system will be reduced obviously. The simulation results of the sub-section high level gas collection structure of NEUI400kA Family HEEP are shown in Fig. 3 to 6.

Fig. 1 NEUI400(I) HEEP Potline of Henan Zhongfu Industry Co., Ltd.

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2. Recent Progress of NEUI400 Family HEEP Technology Based on overall analysis of the physical field measurement results and the operation status of the pots, NEUI has optimized NEUI400(I) HEEP by using the proprietary "Physical Field Numerical Simulation and Analysis Software Package for Aluminum Reduction Pots", which has not only further optimized the key technologies including MHD stability, thermal balance and steel structure, etc., but has also developed some new technologies such as sub-section high level gas collection technology and compressible pot lining technology, etc. By integrating the new technologies above, NEUI has successively developed NEUI400(II-IV) HEEP technologies which have already been put into commercial application respectively. The following will mainly introduce the technologies of sub-section

444

Fig. 4 Sectional View of Pressure Distribution by the Sub-section High Level Gas Collection Structure

Fig. 5 Sectional View of the Velocity Vector Distribution by the Sub-section High Level Gas Collection Structure

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Fig. 6 Inlet Streamline of the Sub-section High Level Gas Collection Structure Fig. 8 Status of NEUI400(III) HEEP When the Side Hood at the Tapping End Was Opened

The simulation results showed that the distribution of negative pressure, temperature and velocity inside the pot hood are more uniform. The total resistance of the system was only 75Pa. When the negative pressure at the gap was maintained at -20Pa, the outlet negative pressure was only required to be lOOPa, which has significantly reduced the outlet negative pressure of the pot. The hooding efficiency will be no less than 99% when the gas exhaust volume is 9150Nm3/h. The technology has already been applied to the operating potlines which adopting NEUI400(III) and NEUI400(IV) HEEP technologies, and excellent application effect was obtained.

2.2 Compressible Pot Lining Technology The compressible pot lining structure has been applied to NEUI400(II-IV) HEEP. In the design of the side lining at the lower part of the pot, the cathode paste, high strength castables, and heat resistant calcium silicate boards (or ceramic fiber boards) are organically combined to form the compressible structure, which can effectively absorb the thermal expansion stress and reduce upheaving of the cathode. When the NEUI400(II-IV) HEEP reached the start-up temperature during baking and before the liquid bath was filled in, the cathode upheaving was no more than 2cm, which was quite low comparing with the maximum 15cm upheaving of the traditional high amperage aluminum reduction pots. The application of compressible pot lining technology is benefit to prolong pot life of NEUI400 Family HEEP. 2.3 Optimization of the Magnetic Fluid Stability and Busbar Arrangement Aiming to gain more excellent distribution of the magnetic field, the non-symmetry magnetic field compensation method of "Strong Compensation at the Short Side and Multipoint Weak Compensation at the Bottom " was used for busbar arrangement of NEUI400(II-IV) HEEP. The main characteristic of the compensation is most of current on upstream will go around the end head of the reduction pot, and less current will pass through the bottom of the reduction pot. In the process of creating the electromagnetic field model, in order to be more close to the actual conditions, the whole potline was brought into the boundary conditions for calculation, and the current balance calculating

Fig.7 is the overview of NEUI400(III) HEEP potline in Shandong Nanshan Aluminum Co., Ltd. which adopted the sub-section high level gas collection technology. From the picture, we can see that the air environment inside the potroom was quite good, with almost no fume emission. Fig. 8 is the reduction pots in the potline when the pot hood at the tapping end was opened. From the picture we also can see that there was no fume escaping even when the side hood at the tapping end was opened. All above indicated that the use of the sub-section high level gas collection structure has effectively captured the fume and significantly reduced fugitive emissions, the hooding efficiency has reached 99% or above. The measurement results of negative pressure at the branch ducts showed that excellent gas collection effect can be achieved if the outlet negative pressure was kept at 100-150Pa. Comparing with the conventional outlet negative pressure of 300Pa, the operating load and cost of the scrubbing system have been reduced significantly.

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In May 2010, the research people from Auckland University came to Henan Zhongfu Industry Co. Ltd to test the control software for automatic extinguishing anode effect on NEUI400(I) HEEP potline. They found that the aluminum reduction pots still run smoothly and stably when the pot voltage drops to 3.6V. MHD stability of the improved NEUI400(II-IV) HEEP is more excellent than NEUI400(I) HEEP. The high MHD stability of NEUI400(IIIV) HEEP was also proved through the liquid aluminum pouring quantity during the start-up, when 5-10t liquid aluminum was filled in, the pots showed rather high stability. The more advanced MHD stability design created favorable conditions for NEUI400(II-IV) HEEP to run under even lower pot working voltage and higher amperage.

results were used as the busbar current data, for the iterative calculation of the magnetic field and the current balance. The calculation results indicated that no matter under normal operation condition or during short circuit period, the current ratio between upstream and downstream, current deviation ratio of each side, current density in each single flexible busbar, as well as the difference between current inside the riser busbar and the theoretical value, etc. can be all controlled within the reasonable range. For the vertical magnetic field value, no matter the maximum, average, or changing gradient, are more ideal than that of NEUI400(I) HEEP. Even when the potline's amperage was intensified by 15%, the busbar was still not overloaded, which indicated that the busbar system was more safe and reliable.

3. Expanded Application of NEUI400kA HEEP Technology

In Jul. 2010, Materials & Metallurgy School of Northeastern University tested the MHD stability of NEUI400(III) HEEP and 300kA reduction pots in Shandong Nanshan Aluminum Co., Ltd. Fig. 9 is the working site for this job. Fig. 10-13 show the measured anode current fluctuation of two type pots. The measured results indicated that at the same location (center or corner), the anode currentfluctuationof NEUI400(III) HEEP was obviously lower than that of 300kA family aluminum reduction pot, which indicated that the interface stability of NEUI400 Family HEEP has been greatly improved than that of 300kA family aluminum reduction pot.

The optimized NEUI400(II-IV) HEEP technologies have been respectively applied in Linfeng Aluminum Co., Ltd, Shandong Nanshan Aluminum Co., Ltd., and Jinning Aluminum Co., Ltd. Table I gives the detail. Table I. NEUI400(II-IV) HEEP Potlines in Commercial Operation Current Intensity Capacity (kt/a) of Potline (kA) Remarks Item ^ ^ Design Actual Design Actual Without the subsection high level NEUI400(II) gas collection HEEP potline technology; The of Linfeng construction of 400 440 250 275 Aluminum the potline began in Jan. 2008, and and Power the potline was Co., Ltd put into operation in Aug. 2009. The construction NEUI400(III) of the potline HEEP potline began in Apr. of Shandong 2008, and the 400 430 250 270 Nanshan potline was put Aluminum into operation in Co., Ltd. Nov. 2009. The construction of the potline began in Nov. 2009, and the potline was put into operation in NEUI400(IV) Jun. 2010. HEEP potline At present, the 400 460 300 350 of Jinning operating current Aluminum intensity is the Co., Ltd. highest and the capacity of the single potline is the largest in China.

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3.1 NEUI400(II) HEEP potline NEUI400(II) HEEP technology has been used in the 250kt/a potline of Linfeng Aluminum Co., Ltd. The new technologies such as the double-anode, pipe truss girder superstructure, and compressible pot lining, etc. have been used in the design of the

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Fig. 13 Anode (A10, corner) Current Fluctuation of 300kA Family Aluminum Reduction Pot

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The construction of the potline began in Jan. 2008, and been put into operation in Sep. 2009. Now, the amperage of the potline is 440kA, average pot working voltage is less than 3.8V, and DC energy consumption is less than 12500kWh/t-Al. Fig. 14 is the picture of NEUI400(II) HEEP potline of Linfeng Aluminum Co., Ltd.

2010. The potline also adopted the sub-section high level gas collection technology and compressible pot lining technology. The busbar has further been optimized in order to further increase the current intensity. At present, the actual operating amperage of the potline has reached 460kA, and the actual capacity has reached 350kt/a. It is the only aluminum reduction potline in China which has the largest single potline capacity and the highest operating amperage. Additionally, average pot working voltage is stabilized under 3.9V within three months after it was put into operation. Fig. 16 is NEUI400(IV) HEEP potline of Jinning Aluminum Co., Ltd.

Fig. 14 NEUI400(II) HEEP Potline of Linfeng Aluminum and Power Co., Ltd.

Fig. 16 NEUI400(IV) HEEP Potline of Jinning Aluminum Co., Ltd.

pot. Besides, in order to further improve MHD stability and intensify the current of a reduction pot, the busbar arrangement has been further optimized.

3.2 NEUI400(III) HEEP Potline NEUI400(III) HEEP technology has been used on the 250kt/a potline of Shandong Nanshan Aluminum Co., Ltd. The construction of the potline was started in Apr. 2008, and the potline was put into operation in Nov. 2009. The double-anode, triangle plate anode lifting device, as well as the sub-section high level gas collection technology and compressible pot lining technology, etc. newly developed by NEUI have been used on this potline. Now, the amperage of the potline is 430kA, average pot working voltage is less than 3.95V, and current efficiency is up to 94%. Fig. 15 is the picture of NEUI400(III) HEEP potline of Shandong Nanshan Aluminum Co., Ltd.

4. Excellent Operating Indices After NEUI400(II-IV) HEEP potlines were put into commercial operation, all potlines have achieved excellent techno-economic indices. Refer to Table II for detail. Table II. Main techno-economic Indices of NEUI400(II-IV) HEEP Anode 1 Average Potline DC energy Effect Date of Pot Consumption Frequency Amperage Potline Operation voltage (kWh/t-Al) (effects/ (kA) (V) pot-day) NEUI400(II) HEEP potline of 440 08. 2009 <3.80 <12500 <0.01 Linfeng Aluminum Co., Ltd. NEUI400(ni) HEEP potline of Shandong 11.2009 430 <3.95 <12500 <0.01 Nanshan Aluminum Co., Ltd. NEUI400(IV) HEEP potline 06. 2010 460 <3.90 of Jinning Aluminum Co., Ltd

-

Fig. 15 NEUI400(IU) HEEP Potline of Shandong Nanshan Aluminum Co., Ltd.

3.3 NEUI400(IV) HEEP Potline NEUI400(IV) HEEP technology has been used for 300kt/a potline of Jinning Aluminum Co., Ltd. The construction of the potline began in Jun. 2009, and the potline was put into operation in Jun.

447

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From Table II, we can see that with the NEUI400 HEEP technology, the amperage of the operating potline has reached 430-460kA, the amperage is increased by up to 15% comparing with the designed value, which means the actual capacity is increased by up to 15% too. It also indicated that NEUI400 Family HEEP can surely run stably at the current density above 0.8A/cm2. This has also laid the good foundation for the development of NEUI500 Family HEEP.

Utill Sep. 2010, four potlines adopting NEUI400 Family HEEP technologies have been put into commercial operation, and the total actual capacity of them has reached l,150kt/a. Now, six NEUI400 Family HEEP potlines with the total design capacity of l,100kt/a are under construction, and will be put into operation successively by the end of 2010. Besides, there are six NEUI400 Family HEEP potlines with the total design capacity of l,230kt/a are under detail design. So, it is obvious to see that NEUI400 Family HEEP technologies have become the main aluminum reduction pot technology in China's aluminum industry.

incorporating the numerical simulation technology developed by NEUI with the experiences, NEUI is able to provide customers with further optimized NEUI400 family HEEP technologies or even higher amperage HEEP technologies and equipments, as well as all-around technical support and service, etc. Acknowledgement The project is jointly completed by NEUI's aluminum reduction technology R&D team. The project is organized by Mr. Lu Dingxiong, Mrs. Mao Jihong, Mr. Qi Xiquan and Mr. Yang Qingchen, etc., who are also the technical advisors of the project. The development of the busbar arrangement and MHD stability technology is mainly completed by Mrs. Mao Jihong and Mr. Mao Yu; The thermal balance calculation of aluminum reduction pots is mainly completed by Mr. Ban Yungang; The ventilation structure design and simulation are mainly completed by Mr. Chen Gaoqiang and the design and simulation of the pot superstructure are mainly completed by Mr. Dong Hui and Mr. Huang Kuisheng. Thanks for the technical data they have provided for the project.

5. Comprehensive Evaluation and Conclusions As the physical field are further optimized, NEUI400(II-IV) HEEP showed more stable than NEUI400(I) HEEP. All the NEUI400(IIIV) HEEP potlines are running stably at the amperage above 430kA. Even when the amperage of the NEUI400(IV) HEEP potline in Jinning smelter reached 460kA, the reduction pots still run stably and efficiently. The sub-section high level gas collection technology and the compressible pot lining technology which are newly developed by NEUI have been applied on NEUI400(II-IV) HEEP, and gained very good effect. The sub-section high level gas collection technology has helped to realize a more effective capture of the fume and reduce the fume exhausting volume, the hooding efficiency can reach 99% or above. The application of the compressible pot lining technology has greatly decreased the cathode upheaving during pot start-up. The measured upheaving was less than 2cm, which is in favor of prolonging the pot life.

Reference [1] Qi Xiquan, Liang Xuemin, Lu Dingxiong etc. "Successful Commercial Operations of NEUI400 Potline", Light Metals 2010, TMS (The minerals, metals & materials society) 2010,359-363. [2] Yungang Ban, Xiquan Qi, Yu Mao etc. "Baking Start-up and Operation Practices of 400kA Prebaked Anode Pots", Light Metals 2010, TMS (The minerals, metals & materials society) 2010,369-373.

The pipe truss superstructure developed by NEUI continues to be used on NEUI400(II-IV) HEEP. The advantages of the structure including the high strength, less material consumption, good ventilation and heat radiation effect, as well as easy fabrication, installation and maintenance, etc. are brought into full play during their actual application. Comparing with NEUI400(I) HEEP, the optimized NEUI400(IIIV) HEEP has achieved more excellent techno-economic indices, including average pot working voltage less than 3.8V, DC energy consumption less than 12500kWh/t-Al, anode effect frequency less than 0.01 effects/pot-day, and the current efficiency up to 94%, etc. The optimized NEUI400(II-IV) HEEP technologies have made great contribution to the energy saving & emission reduction target of the China's aluminum industry. Now, NEUI has become an important technology supplier and service provider for the aluminum industry both home and abroad. Based on the specific conditions of each project and the needs of the customers, by following the development mode of

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Improving Current Efficiency of Aged Reduction Lines at Aluminium Bahrain (Alba) Khalil Ghuloom1, Abdulla Habib1, K.S.R. Raghavendra *, Hasanain Hassan1 1 Aluminium Bahrain (Alba) - Metal Production Group Keywords: Current Efficiency, Alumina Feeding, SMART Centres Abstract The side worked end-on-end pre-bake anode technologies installed during 1970's at Alba achieved a current efficiency of 90%. During the early nineties, the potlines were retrofitted from side break to point fed cell arrangement, along with an increase of 7% and 10% in line current. The retrofits included anode gas collection system, a changed anode setting pattern, and installation of alumina and aluminium fluoride feeders controlled using individual cell controller, which increased the current efficiency to 92.5%. During the last 6 years, improvements done in the alumina feeding control, thermal control and stability of cell operations has increased the current efficiency to 94.5%. Close monitoring and follow up of quality of work and key parameters on daily basis by the new employees forum introduced at Alba in 2004 called SMART Centres (See My Actions Reflect Targets) has also contributed towards this improvement and maintaining it.

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The side worked end to end pre-baked anode technology cells were progressively installed in Alba pot lines 1-2 and pot line 3 during 1971 - 1982\ Pot lines 1-2 and line 3 were initially operating at around 100 kA and 115 kA respectively achieving a current efficiency of around 87%. The performance was gradually improved and by 1991, pot lines 1 & 2 and line 3 were operating at 112 kA and 123 kA respectively, with the current efficiency averaging around 90% in both lines 1-2 and line 3. In the period 1992 -1995, the three pot lines were retrofitted and changed over from side break to point feed cells, which included installation of alumina and aluminium fluoride feeders controlled using individual cell controller, gas collection system and changing anode setting pattern. After retrofitting, the line current was gradually increased to 120 kA (lines 1 &2) and to 138 kA (line 3) as shown in figure 1. The line current in lines 1-2 was further increased gradually from 120 kA to 125 kA during 2001 to 2004 and maintained at 125 kA till now.

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As a result of the change over from side break to point fed technology, the current efficiency improved and maintained at around 92.5% till year 2004. During 2005 to 2010, significant improvement in current efficiency has been achieved as shown in figure 2. Following changes introduced and improvements done has lead to increase in the current efficiency from around 92.5 to 94.5%, which are discussed in this paper.

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Figure 2; Current Efficiency Trend in Reduction Lines 1-3 Modified Thermal Control Model Reviewing the trend of bath temperature and bath acidity of the individual pots over a prolonged period (120 days) indicated 2 to 3 swings in bath temperature and bath acidity and such swings were even observed while analysing the average data on potroom basis also. This was mainly due to the wide variation in the daily aluminium fluoride addition from 0 to 60 kg/day/cell compared to average requirement of 17 kg/day/cell. The aluminium fluoride addition control strategy used originally is represented in figure 3 by the solid line graph and at the centre point the base amount of 17 kg/day is added. In this strategy there was over emphasis on

Introduction of modified Thermal Control Model. Design changes and Voltage review. Introduction of new Alumina Feeding Model. Improvement in quality of work Introduction of new employee form called SMART Centres.

449

correcting the high bath temperature by excessive aluminium fluoride addition and expecting fast response. This approach has following errors. • It does not allow for the operating band that is inherent in the process (the oval of the Figure). • It assumes constant bath volume. • It assumes the difference in temperature is due solely to aluminium fluoride concentration.

measurements. However it enabled a beneficial increase in anodecathode distance. To improve the superheat and to reduce the voltage fluctuation of individual pots over long period, the target resistance was increased by around 0.3 micro ohms (around 40 mV) without significantly affecting the overall voltage of the pots. SILICON CARSOE SIDEWALL

Rarely any of these assumptions are valid and there was a need for major change in the philosophy as follows:

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• Allow for the dead band as indicated by the dashed line graph in figure 3. (This illustration is schematic only, the actual algorithm is non-linear but similar.) • Reduce the maximum and increase the minimum additions. • Reduce the slope of the curves for less aggressive additions.

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Alumina feeding was based on an adaptive feeding model. The characteristic nonlinear relation between the alumina concentration and the pot resistance was modeled by an adaptive linear model where the estimated slope of resistance versus alumina concentration (parameter bl) gives information about the concentration of alumina. This information was used to control the alumina supply to the pot. For the feeding control to be successful it was necessary to maintain stable pot resistance and the only "variability" in the pot was supposed to be the variation in alumina concentrations, due to the different feeding modes introduced. However, alumina concentration control was exposed to uncertainties because of

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Figure 3: Original and modified A1F3 addition strategy. The aim for these changes was to stabilize the variation in fluoride addition and enable better cell control since deviations outside the control band are usually due to deviations in other operational routines. Even though it is not expected to give significant direct improvements in cell performance, it is definitely an enabler for better control. It makes it easy to diagnose the causes of deviation from normal behavior in cells. Thus enabling to attack the root cause of the problem of high bath temperature, rather than trying to correct it through aluminium fluoride addition.

Variation in alumina dump weight (assumed constant alumina dump weight) Variation in transfer of alumina from hopper to bath (assumed 100% transfer) Solubility of alumina (assumed 100% solubility). Alumina always does not completely dissolve due to higher alumina dump weight of 1.6 kg and high number of rapid alumina dumps at the end of tracking (25 dumps) and after anode effect (60 dumps) Feeding from other sources (sludge / crust and anode top cover etc.)

Design Changes & Voltage Review of the Individual Pots: Following design changes (figure 4) gradually introduced since 2004 has resulted in a total voltage saving of more than 100 mV and has contributed towards stable side ledge and improved cell voltage stability. -

To ensure that sludge build up was prevented? every 24 hours the pots were set on tracking mode. Analysis of Alumina Feeding Irregularities and Feed Cycles

Increasing the depth of the anode stub hole from 100 to 120 mm Introducing single slotted anodes of depth 150 mm. Increasing the collector bar cross section by 28%. Composite high thermal conductivity sidewall

These savings in individual voltage components could not be realized in terms of reduced cell voltage due to heat balance and superheat requirements, which were confirmed by the

450

The total number of feed cycles typically was found to vary from 6 to 8 per day, which was considered low and could not enable good clean periods to occur for good control. The average overfeed time was in the range of 60 to 90 mins with a maximum limit of 120 mins. This was effectively increasing the alumina concentration by about 1.1%, which was high. Normally

it should be targeted to increase by around 0.6% to get to the start of the flattening of the resistance versus alumina curve and avoiding entering the insensitive rising section on the alumina rich side.

Change Over To Fixed Alumina Overfeed Strategy An attempt was made in year 2005, to change over from adaptive alumina feeding control with a sampling time of 5 minutes to fixed duration over feeding - alumina feeding strategy with a sampling time of 3 minutes in line 1-2 pots and 4 minutes in line 3 pots.

Ratio of underfeed to overfeed was extremely irregular which was due to the back feeding and sludge formation.

Implementing a new alumina feeding strategy in the existing 20 years old control system hardware, with limitations in the communication system network between the pot controller and the central computer was a challenge. Keeping this in mind a simple fixed over feeding alumina feeding strategy was chosen in preference to the advanced alumina feeding strategies evolved and adopted in modern aluminium smelters. Due to the hardware limitations in the communication system the alumina feeding control sampling time could not be reduced below 3 minutes in line 1-2 pots and below 4 minutes in line 3 pots.

The Figure 4 shows considerable variation in under feed times resulting only in 8 feeding cycles in 24 hours.

Figure 5 : Resistance, alumina feeding and instability trend graph.

Fixed duration over feeding strategy followed consists of over feed sequence for a definite period, followed by an under feed for a shorter period and subsequently putting the pot on tracking to reach a fixed alumina concentration level in bath before starting a new feeding cycle. This fixed duration over feeding strategy provided easily adjustable parameters to control the alumina concentration in the bath at the desired level and there by reducing the anode effect frequency. Different setting of the fixed duration over feeding strategy were tried and optimised at 45 minutes over feeding at 150% feed rate, followed by an under feed for 20 minutes at 60% feed rate and tracking at a feed rate of 11%. This proved to be very successful in reducing the anode effect frequency by more than 50% and improve the alumina concentration control in the cell as evident from the increase in the feeding cycles per day from around 6-8 cycles to around 13 cycles as indicated in Figures 7, 8 and 9.

Figure 6 : Resistance, alumina feeding and instability trend graph.

Figure 7 : Resistance, alumina feeding and instability trend graph

Figure 4 : Resistance, alumina feeding & slope value trend graph The Figures 5 and 6 presented below shows only 6 feeding cycles and unusual resistance slope curves at certain times.

From this detailed review it was clear that the number of feed cycles needed to be increased from 6 - 8 cycles per day to around 12-16 cycles per day by reducing the overfeeding duration to less than 60 mins. The sampling time for the alumina feed control logic to be reduced from the original 5 mins to improve the sensitivity of the control system. The optimum value of the sampling time was to be determined taking into account the existing limitations in the hardware of the communication system network between the individual pot controller and the central pot line computer. Figure 8 : Resistance, alumina feeding and instability trend graph.

The change over to the new fixed overfeed alumina feeding strategy to achieve these above goals are discussed below.

451

some pots having alumina content in the range of 3.5 to 5.0 % for few hours. Repeated tracking feature of the new fixed alumina overfeeding strategy ensures that these instances of higher alumina concentrations are brought down to the desired range with in few hours ( 1 - 2 cycles). Thus the risk of sludge formation and sludge build up was greatly reduced. The above observations confirm that the new fixed overfeeding duration alumina feeding programme is working good. Improvement in quality of Work through SMART Centre Figure 9 : Resistance, alumina feeding and instability trend graph. Other Modification to the Original Alumina Feeding Strategy Anode effect Termination Feeding Logic: The alumina rapid feeding after anode effect (60 dumps / 90 kgs in 7 mins) was high. This was a greater mass than could possibly dissolve and it was contributing to sludge build up. The number of alumina shots fed rapidly after anode effect was reduced from 60 to 20 after conducting trials and establishing that there were no repeat anode effects in the short time afterwards which would otherwise indicate insufficient addition of alumina. This also contributed improved general feeding control and reduced the chances of build up of sledge in the pots. Tracking Termination Feeding Logic:

Employees in Alba is working based on the concept of SMART i.e. S see M my A action R reflect T targets. The concept of smart centre is to converts strategic and business plans into easily understood and measurable actions on the shop floor to enable and encourage teams to share, participate, innovate, solve problems and effectively plan together. Smart centers had positively impacted on the people and their commitment, ownership and performance that lead to increased productivity. The importance of teamwork was practically emphasized through the team building training. There was a need to have a way through which shop floor workers are brought on board and engaged in the decision making process and hence the idea of implementing the concept of the SMART PROCESS was explored.

The alumina rapid feeding at the end of tracking (25 dumps / 35 kgs in 3 mins) was also considered high which was reduced to 15 dumps and subsequently to 7 dumps. Thus reducing the risk of sludge formation. Anode Change Feeding Logic: To increase the alumina concentration just before anode change and reduce the risk of anode effect during anode change due to colder bath and higher current density, alumina feeding logic was modified to add 15 shots of alumina at the normal overfed rate prior to the anode setting. This modification reduced the anode effect at the time of anode changefrom0.04 to 0.02. Performance of Modified Alumina Feeding Strategy Control of Alumina Content in bath: In order to assess the effectiveness of the fixed overfeed alumina feeding strategy in maintaining the alumina concentration in the desired range of 1.8 to 2.8%, extensive study of alumina content in bath was carried out in a group of pots over a period of 24 hours by actually measuring the alumina content in bath samples taken every 15 mins. Alumina concentration change was found to be broadly matching with the type of alumina feeding i.e, increase in alumina concentration during overfeeding and decreasing in alumina concentration during under feeding and tracking. In most of the pots the alumina content in bath was being controlled at the desirable range of 1.8 to 2.8 %. However, some times the high self feeding during activities or repeated anode moves, was found to completely mask these correlations and pushed the alumina concentration to high and outside the range values as observed in

The teams meet at the start of every shift for a 15 minutes meeting to discuss pertinent issues for the day. During the meeting, they updated and monitor their performance using daily and weekly charts. Once a week the team meets to discuss their general performance, evaluate their suggestion and resolve problems encountered. They take time to celebrate success and discuss improvement plans. The team was empowered to put solutions to their problems themselves and follow it up even with other teams or sections. As result of tight relationship built between the management and employees, the gap is closing and links established for common understanding and agreement to resolve issues related to employee's welfares, safety and meet the company corporate targets. Implementation of the changes was helped by the motivation of the smart centre groups.

Average Instability Trend

Performance Improvement with modified design & control logic The overall performance of the pot line before and after implementing the modified lining design, thermal control and alumina feeding strategy is summarised in Table 1 and Figure 2, 10,11 & 12. Reduction Line 3 | Reduction Lines 1-2 Year 2010 Year 2010 I Year 2004 Jan - Oct Year 2004 Jan - Oct Before the After the Before the After the UNIT change change change change PARAMETER 124.4 124.8 137.5 138 kA Line Current 4.67 4.69 4.66 V 4.66 Cell Voltage 94.7 94.7 92.6 92.8 |Current Efficiency 14.77 15.00 kwh/kg AI 14.95 14.70 Specific Energy 0.18 0.19 0.39 Anode Effect Frequency 0.33 1822 1828 2381 |Average Age of failed cells days 1769

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Table 1: Comparison of the overall performance of the Pot line 12 and Line 3 before and after implementing the modified lining design, thermal control and alumina feeding strategy.

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After implementing the modified alumina feeding strategy, the anode effect frequency across lines 1, 2 and 3 has been reduced from level of around 0.40 to 0.15 and correspondingly reducing the PFC emissions by more than 50%. The collective effect of improvements in alumina concentration control, thermal control, lining design and quality of work is reflected in the improvement in the current efficiency across lines 1-3 of the order of 2.0%. This high level of performance is achieved and being maintained with active participation of employees through SMART Centers.

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References

Figure 10 : Trend graph of anode effect frequency 4.75

[1] K.S.R.Raghavendra, Reducing anode effect frequency in aged pot lines by changing alumina feeding & control parameters and overhauling of breaker feeder system, Ninth Australasian Aluminium Smelting Technology Conference and Workshop 4-9 November 2007 pp 149-161.

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[2] Abdulla Habib, K.S.R. Raghavendra, Barry Welch, Improving heat dissipation and cell life of aged reduction lines at Aluminium Bahrain (ALBA), Light Metal 2010 pp xxx - xxx.

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[3] Lynette wilsenach, Customised Integrated Process for ALBA SMART Process combined with Teambuilding Activities may 2004. [4] Hasanain Hassan Al Ojaimi, EMPLOYEES BEHAVIOUR IN IMPROVING ROOF EMISSION IN REDUCTION POTLINES 13, Ninth Australasian Aluminium Smelting Technology Conference November 2007.

Figure 11 : Trend graph of Cell Voltage.

[5] Hasanain Hassan Al Ojaimi, 1st Smart Center Annual Review 2009.

453

ALBA

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Development of NEUI500kA Family High Energy Efficiency Aluminum Reduction Pot (¹ÅÅÑ') Technology Lu Dingxiong, Mao Jihong, Ban Yungang, Qi Xiquan, Yang Qingchen, Dong Hui Northeastern University Engineering & Research Institute, Co., Ltd. No.73, Xiaoxi Road, Shenhe Dist., Shenyang, Liaoning, PR China 110013 Keywords: Aluminum Reduction Pot; NEUI500kA Family; High Energy Efficiency; Technology Development structure, etc., have been improved obviously. Since the first NEUI400(I) HEEP potline with the capacity of 230,000t/a was put into operation in August, 2008, NEUI400(II-IV) HEEP potlines with the capacity of 250kt/a, 250kt/a and 300kt/a have been put into production successively in Linfeng Aluminum and Power Co., Ltd, Shandong Nanshan Aluminum Industry Co., Ltd and Jinning Aluminum Co., Ltd, and more excellent techno-economic indices have been achieved.

Abstract Based on the successful development and application experiences of NEUI400kA family high energy efficiency aluminum reduction pot ('NEUI400 family ΗÅÅΡ') technology, through the physical field measuring & analysis and with the proprietary 'Physical Field Numerical Simulation and Analysis Software Package for Aluminum Reduction Pots', by continuously following its own high amperage reduction pot development mode of incorporating the numerical simulation technology with experiences, NEUI has developed the NEUI500kA family high energy efficiency aluminum reduction pot ('NEUI500 family HEEP') technology. Comparing with NEUI400 family HEEP, the simulation results showed that the technical parameters that affecting the MHD stability, current balance, pot lining structure, gas collection system and the superstructure etc. are more excellent.

1.2 Verification of the Physical Field of NEUI400 Family HEEP by Measurement After the commercial operation of NEUI400 family HEEP, Materials and Metallurgy School of Northeastern University respectively measured and analyzed the physical field of NEUI400(I) HEEP in Henan Zhongfu Industry Co., Ltd. and NEUI400(III) HEEP in Nanshan Aluminum Industry Co., Ltd. in Jan., 2009 and Jun., 2010. The comparison results showed that the physical field of NEUI400 HEEP was in compliance with their numerical simulation results, which verified the advancement and reliability of NEUI400 family HEEP technology.

1. Introduction In order to meet market demands both home and abroad, based on the successful development and application of NEUI400 family HEEP technology, NEUI has proceeded develop NEUI500kA family HEEP technology by continuously following its own high amperage reduction pot development mode of incorporating the numerical simulation technology with experiences.

1.3 Operating Parameters of NEUI400 Family HEEP All NEUI400 family HEEP potlines have achieved excellent techno-economic indices after being put into commercial operation. Table I is a typical comparison between the operating parameters and design parameters of NEUI400 family HEEP.

In what follows, is an introduction about NEUI400 family HEEP, and the following is an introduction about the key technologies of NEUI500 family HEEP.

Table I. Comparison between Design Parameters and Operating Parameters of NEUI400 Family HEEP Design parameter Operating parameter 400±10% 400±15% Amperage (kA) Anode effect frequency <0.05 <0.015 (effects/pot-day) 4.06 Average cell voltage (V) 3.85 DC energy consumption <12900 <12500 (kWh/t-Al) 99 Hooding efficiency (%) >99

1.1 Development and Application of NEUI400 Family HEEP NEUI has created a high amperage aluminum reduction pot development mode of incorporating the numerical simulation technology with accumulated experiences. Assisted by the 'Physical Field Numerical Simulation and Analysis Software Package for High Amperage Aluminum Reduction Pots' redeveloped by NEUI based on the software platforms including ANSYS, MHD and CFD, etc, and the expert team which has experienced the whole development course of the high amperage aluminum reduction technologies in China, NEUI was the first in China to have developed the NEUI400 family HEEP technology. Comparing with the 300kA family reduction pots, key technologies of the NEUI400 family HEEP technology including MHD stability, thermal balance, ventilation system and steel

From Table I, it is included that the main techno-economic indices of NEUI400 family HEEP have fully reached or even exceeded the design parameters after being put into operation. 1.4 Amperage Intensifying of NEUI400 Family HEEP Table II showed the present operating amperage intensity of NEUI400(I-IV) HEEP.

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Fig. 1 showed that the current distribution of the cathode at upstream and downstream of NEUI500 family HEEP was relatively uniform, with the maximum deviation ratio of 0.88%.

Table II. Present Operating Amperage Intensity of NEUI400(I-IV) HEEP Present Operating Time for NameofPotline Operating Amperage Intensity(kA) NEUI400(I) HEEP in Henan 08,2008 415 Zhongfu Industry Co., Ltd. NEUI400(II) HEEP in 08,2009 440 Linfeng Aluminum and Power Co., Ltd NEUI400(III) HEEP in 430 Shandong Nanshan 11,2009 Aluminum Co., Ltd. NEUI400(IV)HEEPin 06,2010 460 Jinning Aluminum Co., Ltd.

For the busbar arrangement design of NEUI500 family HEEP, the asymmetric magnetic field compensation method of 'Strong Compensation at the Short Side and Multi-point Weak Compensation at the Bottom' was used. The main characteristic of the compensation was most of current on upstream go around the end head of the reduction pot, and less current pass through the bottom of the reduction pot. During the modeling of the electromagnetic field, the whole potline has been brought into the boundary conditions for computation in order to be more close to the actual conditions, and the busbar current parameters are taken from the results of the current balance calculation. In addition, the magnetic field and the current balance are given interative computation to improve the accuracy.

From table II, it is known that the actual operating amperage intensity of NEUI400 family HEEP potlines have completely exceeded 400kA. Take NEUI400(IV) HEEP potlines in Jinning Aluminum Co., Ltd as an example, the present operating amperage intensity has reached 460kA, the reduction pots still run stably and efficiently. This has laid a good foundation for the development of NEUI500 family HEEP technology.

Fig. 2-4 are respectively the simulation results of the magnetic field distribution along X, Y and Z direction in the molten aluminum section.

2. Development of NEUI500 Family HEEP Technology Based on the successful development and application experiences of NEUI400 family HEEP technology, NEUI has developed the NEUI500kA family HEEP technology by continuously following the high amperage reduction pot development mode of incorporating the numerical simulation technology with accumulated experiences. For the NEUI500kA family HEEP technology, the simulation results of characteristic parameters including the busbar arrangement & MHD stability, thermal balance, pot ventilation and superstructure, etc. are more excellent. 2.1 Optimization of Busbar Arrangement and MHD Stability During the development of NEUI500 family HEEP technology, following requirements shall be met: (1) The busbar voltage drop shall be reasonably chosen. The busbar consumption shall be reduced as much as possible; (2) Minimize the voltage drop difference between upstream and downstream cathode busbar to realize balancing current distribution at both sides; (3) The busbar shall not be overloaded during normal operation or under short circuit status. Refer to Fig. 1 for the busbar current balance calculation results of NEUI500 family HEEP.

Fig. 2 Magnetic Field Distribution along X Direction inside NEUI500 Family HEEP

Fig. 3 Magnetic Field Distribution along Y Direction inside NEUI500 Family HEEP *5iw

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stability design guaranteed the stable and efficient operation of NEUI500 family HEEP under low voltage. 2.2 Better Thermal Balance Design When the amperage of a reduction pot is increased, the sidewall area per unit cavity volume is reduced, and the unit area heat lost on the sidewall will be increased. Besides, comparing with NEUI400 family HEEP, the anode current density of NEUI500 family HEEP was increased obviously, which will increase the heat generation of reduction pots. Accordingly, the heat lost of each area will also increase. Therefore, relevant measures must be adopted in order to obtain a better thermal balance. (1) Favorable Conditions for Realizing Better Thermal Balance One of the good conditions to realize better thermal balance is the lower pot voltage, and then the thermal input will be reduced. It is benefit to realize ideal thermal balance and regular pot cavity.

Fig. 4 Magnetic Field Distribution along Z Direction inside NEUI500 Family HEEP

Table III is the comparison of the characteristic parameters including the magnetic field, current balance and busbar consumption, etc. between NEUI500 family and NEUI400 family HEEP.

The better MHD stability design of NEUI500 family HEEP provides favorable conditions for the stable and efficient operation of NEUI500 family HEEP under low voltage. The actual operation of NEUI400 family HEEP, has shown that all reduction pots can run stably and efficiently under the working voltage of 3.75-3.9V. As the characteristic parameters of MHD stability of NEUI500 family HEEP are more excellent, which are sufficient enough to assure the stable and efficient operation under low voltage, and accordingly, in favor of realizing better thermal balance. (2) Regulating Measures for Realizing Better Thermal Balance • Regulating Structural Parameters ® Increase the pot cavity depth The increase of the pot cavity depth can lower the heat flux of per unit pot cavity sidewall area, which can achieve the purpose of increasing heat lost. © Increase dimensions of the stub and the collector bar The increase of the stub dimension and collector bar dimension will reduce their own resistance, reduce the heat generation, and increase the heat radiation volume. (3) Reduce the width of the side block, and adhesively bonded with the pot shell Reducing the width of the side block will be in favor of reducing the thermal resistance increased due to the pot shell deformation. The adhesive bonding between the side block and the shell internal wall will further reduce the thermal resistance, which will guarantee a smooth heat passage. @ Strengthen the heat radiation potential of the pot shell Single strap or no strap and low width pot brim were used in shell design to strengthen ventilation. Besides, radiating ribs will be added to strengthen the heat radiation of the side shell. • Low resistivity and high thermal conductivity cathode blocks were selected The low resistivity and high thermal conductivity cathode blocks can not only increase the heat lost volume of the pot shell opposite to the cathode block, which will lower the temperature of the pot shell below the cathode block and make the temperature drop

Table III. Simulation Results of Relevant Characteristic Parameters of the Busbar Systems of NEUI400 Family and NEUI500 Family HEEP NEUI500 NEUI400 1 Parameter Unit Family HEEP Family HEEP 97.07 79.59 Gs |Bx|ave 12.58 10.89 Gs |By|ave 3.14 3.16 Gs |Bz| ave 19.87 22.52 Gs |Bz| max 3.09 3.77 Gs |Bz|avel 3.75 2.63 Gs |Bz|ave2 3.08 3.16 Gs |Bz|ave3 2.47 3.06 Gs |Bz|ave4 3.23 Gs/dm 3.76 Bz Max. gradient Total voltage drop of 211.40 mV 202.56 busbar Current ratio at 49.56 49.56 % upstream Current deviation ratio 0.739 0.605 % at upstream Current ratio at 50.44 50.44 % downstream Current deviation ratior 0.256 0.229 % at downstream 1 Voltage drop during mV 420.5 350.52 short circuit period 1 Current deviation during short circuit 5.30 5.20 % period 1 Max. current density 0.81 during short circuit A/mm2 0.83 period 1 Aluminum consumption kg/A 0.1311 0.1351 1 per unit current

Table III showed that the maximum value, average value and changing gradient of the vertical magnetic field of NEUI500 family HEEP were better than that of NEUI400 family HEEP. In addition, aluminum consumption per unit current for NEUI500 family HEEP was somewhat less than that of NEUI400 family HEEP, which indicated that the busbar system of NEUI500 family HEEP was more economic. Even when the potline current was intensified by 10%, the busbar was still not overloaded, which suggested that the busbar system has high security. Better MHD

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gentle, but also can lower the voltage drop of the cathode itself, which is in favor of realizing better thermo-electric balance. • Choose Reasonable Process Parameters Proper technical process conditions are chosen for the thermal balance design of NEUI500 family HEEP. The effective measures such as the selecting the bath which is of low electrolysis temperature and proper superheat, maintaining a lower anode effect frequency, maintaining proper height of liquid aluminum pad, lowering the thickness of the cover material to increase the radiation at the top, as well as selecting of reasonable gas exhaust volume, etc. to keep a good and stable heat balance (3) Application of Compressible Pot Lining Structure Technology The compressible lining structure developed by NEUI was used at the end of cathode block in NEUI500 family HEEP, i.e. the cathode paste, high strength castables and high temperature resistant calcium silicate board (or ceramic fiber board) are organically combined to form the compressible structure. The structure was equivalent to a stress buffering area between the cathode block and the pot shell, which can effectively absorb the cathode swelling stress, reduce the bottom upheaving, and prolong the pot life. (4) Thermo-Electric Field Simulation Fig. 5 and 6 are respectively the temperature contour and temperature isotherm of NEUI500 family HEEP.

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It can be seen from the simulation results (Fig. 4 and 5) that NEUI500 family HEEP can maintain a rather ideal temperature distribution, each isotherm in the cathode lining was located in the rational location, the temperature of each area was within reasonable range, which indicated that the design of the pot lining structure was reasonable. (5) Comparative Analysis of Heat Emission Distribution Refer to Table IV for the heat lost distribution comparison between NEUI400 family HEEP and NEUI500 family HEEP.

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Table IV. Thermal Balance Comparison between NEUI400 and NEUI500 Family HEEP NEUI500 Family NEUI400 Family HEEP HEEP —^Technology Simulation Measured 1 Simulation Result Result Result Parameter ^ ^ ^ 1. Heat input (excluding reaction energy — — — consumption) 100% 100% 2. Heat loss 100% 48.1% 52.6% 2.1 Heat loss of anode 46.5% 26.4% of which: packing material 21.8% 24.9% 18.9% Stub 18.4% 17.8% 7.3% 6.3% 5.4% Rod 47.4% 53.5% 51.9% 2.2 Heat loss of cathode of which: 27.4% 28.9% 42.8% pot shell 17.0% 18.8% Cradle 4.6% I Collector bar and flexible 7.6% 5.7% 3.56% 3.Heat unbalancing degree 1.90% 2.61%

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Table IV showed that the unbalance between the heat generation and the heat loss of NEUI500 family HEEP is 1.90%, which is within the allowable computing error. Based on the heat radiation characteristics of NEUI500 family HEEP, after corresponding thermal balance regulating measures were taken, the heat radiation ratio of the cathode is somewhat increased comparing with that of NEUI400 family HEEP, which strengthened the heat radiation of the pot shell opposite to the cathode block, and the anticipated target for heat balance has been achieved. The potlines which are adopted with NEUI400 family HEEP have realized good thermal balance after being put into commercial

458

Operation. The measuring results and the simulation results are almost the same. For NEUI500 family HEEP which was developed with the same development mode of NEUI400 family HEEP, corresponding thermal balance regulating measures are taken to adapt to the change of the heat radiation characteristics of the reduction pot. The simulation results indicated that the thermal balance design of NEUI500 family HEEP is also very good. 2.3 High Efficiency Ventilation Technology The sub-section high level gas collection technology was also used for the ventilation structure design of NEUI500 family HEEP. Based on the characteristics of NEUI500 family HEEP, the ventilation structure has been optimized to further increase the hooding efficiency. (1) Advantages of the Sub-section High Level Gas Collection Technology The sub-section high level gas collection technology developed by NEUI has following advantages: • The pressure distribution inside the hood is uniform, which is good for increasing the hooding efficiency. • The negative pressure caused by the hot fume itself is fully used, which is in favor of the effective gas collecting, the gas volume per ton aluminum production is reduced and the operating load of the scrubbing system is reduced. • Spaces within the superstructure and the pipe truss girder are fully used. • The overall structure is compact, which provides enough spaces for the process operations such as anode change and tapping, etc. (2) Application Effect of Sub-section High Level Gas Collection Technology for NEUI400 Family HEEP This technology has been applied to NEUI400(III) and (IV) HEEP. The operation results suggest that the hooding efficiency has reached over 99%, which provide favorable conditions for achieving the environmental indices including the total fluoride discharge of no more than 0.6kg/t-Al and dust emission of no more than 1.0kg/t-Al. (3) Development of the Sub-section Gas Collection Structure of NEUI500 Family HEEP Based on the application experiences on the sub-section high level gas collection technology for NEUI400 family HEEP, the subsection high level gas collection technology for NEUI500 family HEEP was optimized according to its own characteristics. The 8subsection high level gas collection structure was adopted and each individual area can be regulated independently. (4) Simulation of the Gas Collection System The ventilation structure for NEUI500 family HEEP was simulated by using the 'Physical Field Numerical Simulation and Analysis Software Package for Aluminum Reduction Pots' which is redeveloped based on the CFD software platform. To be more close to the actual conditions, the whole potroom was brought into the computing system, and the boundary conditions were extended to outside the potroom. Refer to Fig. 7-9 for the simulation results.

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Fig. 7 Temperature Field of the Section

Fig. 8 Inlet Streamline Diagram

Fig. 9 Velocity Field Section

The simulation results indicated that the fume exhaust system of NEUI500 family HEEP has realized the uniform distribution of the negative pressure, velocity and temperature, etc. Comparing with traditional ' V shape low level gas collection structure, this structure has significantly reduced the resistance of the system, reduced the fume exhausting resistance by 80Pa, reduced the fume exhaust volume for per ton aluminum production by 25%, which has reduced the operating load of the gas treatment center, and lowered the power consumption of gas treatment center to less than 150kWh/t-Al. 2.4 Optimized Design of the Superstructure

The superstructure with the pipe truss girder developed by NEUI has been used for NEUI500 family HEEP, and the design has been optimized based on the increased span and increased load of NEUI500 family HEEP. (1) Advantages of Pipe Truss Girder Superstructure Advantages of pipe truss girder superstructure are as follows: • The strength-steel consumption ratio is high; • Good ventilation and heat radiation effect; • Simple structure and esthetic appearance; • Easy to be fabricated, installed and maintained; • Good for the arrangement, safe and stable operation of the servo system. (2) Simulation Results The simulation results of the pipe truss girder of NEUI500 family HEEP were shown in Fig. 10-11.

The simulation results showed the ratio between the strength and steel consumption of the optimized superstructure is higher and can meet requirements under different working status. Comparing with solid web girders, the steel consumption of the support structure is reduced by 30%, which is more economic. 2.5 Auxiliary Technologies and Equipment Besides developing key technologies for NEUI500 Family HEEP, NEUI has also developed following auxiliary technologies and equipment: • Potroom high efficiency ventilation technology; • Intelligent control technology for aluminum reduction; Powder material conveying technology; High efficiency rectification and power supply technology. 3. Comprehensive Analysis and Conclusions On the basis of the summarizing the successful development and application experiences of NEUI400 family HEEP, NEUI continues to adopt the high amperage reduction pot development mode of incorporating the numerical simulation technology with experiences to develop the NEUI500 family HEEP technology. The simulation results of NEUI500 family HEEP are more excellent than that of NEUI400 family HEEP, which guarantees core technologies for achieving the anticipated techno-economic indices in Table VI. Table VI. Anticipated Indices of Nl5UI500 Family HEEP after being Put into Operation Item Anticipated Value Amperage (kA) 500±10% <0.015 Anode effect frequency (AE/pot day) Average cell voltage(V) 3.85V DC current consumption of per ton <12500 Aluminum (kWh/t-Al) Hooding efficiency (%) >99 Pot life (day) >2000

Fig. 10 Axial Force Distribution of Members (kN)

Meanwhile, NEUI has also developed auxiliary process and equipment technologies of NEUI500 family HEEP. Acknowledgement The project is jointly completed by NEUTs aluminum reduction technology R&D team. The project is organized by Mr. Lu Dingxiong, Mrs. Mao Jihong, Mr. Qi Xiquan and Mr. Yang Qingchen, etc., who are also the technical advisors of the project. The development of the busbar arrangement and magnetic fluid stability technology is mainly completed by Mrs. Mao Jihong and Mr. Mao Yu; the thermal balance simulation of aluminum reduction pots is mainly completed by Mr. Ban Yungang; the ventilation structure design is mainly completed by Mr. Chen Gaoqiang and the design and simulation of the pot superstructure is mainly completed by Mr. Dong Hui and Mr. Huang Kuisheng. Thanks for the technical data they have provided for this paper. Reference

Fig. 11 Maximum Stress Distribution of Members Table V. Simulation Results of the Superstructure of NEUI500 Family HEEP Parameter Value

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CURRENT EFFICIENCY FOR ALUMINIUM DEPOSITION FROM MOLTEN CRYOLITE-ALUMINA ELECTROLYTES IN A LABORATORY CELL Geir Martin Haarberg1, Joseph P. Armoo1, Henrik Gudbrandsen2, Egil Skybakmoen2, Asbj0rn Solheim2, Trond Eirik Jentoftsen3 department of Materials Technology, Norwegian University of Science and Technology, NO-7491 Trondheim, Norway 2 SINTEF Materials and Chemistry, NO-7465 Trondheim, Norway 3 Hydro Aluminium, NO-6884 0vre Δrdal, Norway Keywords: Aluminium, current efficiency, current density [email protected] typically in the range from 0.7 - 0.8 A/cm2 in modern industrial cells.

Abstract The current efficiency with respect to aluminium can be as high as 96 % in modern Hall-Heroult cells.

An important feature of the process is the fact that aluminium dissolves in the molten electrolyte, which is a general phenomenon taking place when a metal is in contact with a molten salt containing the metal cation or other species of the metal [3]. In molten cryolite based electrolytes dissolved Na must be considered in addition to dissolved Al. A small but significant activity of sodium is established at the metal/electrolyte interface due to the following equilibrium:

The loss in current efficiency is strongly linked to the fact that aluminium is soluble in the electrolyte. In addition the presence of dissolved sodium must be considered. The back reaction between dissolved metals and the anode product is responsible for the major loss in current efficiency. The rate of the reaction is controlled by diffusion of dissolved metals (Al and Na) through the diffusion layer.

Al + ?>NaF = ?>Na + AlFs

A laboratory cell was used to determine the current efficiency for aluminium during constant current electrolysis. Standard conditions were NaF/AlF3 (CR=2.5), A1203 (sat), 5 wt% CaF2 at 980°C and 0.85 A/cm2.

It is known that the subvalent species A1F2" is formed as well as dissolved Na, the latter being responsible for a small contribution to electronic conductivity. Solubility studies have been carried out in laboratory experiments, and under industrial operation the metal solubility is -0.06 wt% Al. The solubility decreases by increasing content of A1F3 and decreasing temperature. Reliable data for the metal solubility have been published by 0degärd et al. and Wangetal. [4,5].

Current efficiencies ranging from - 93 - 96 % were obtained. The current efficiency was found to increase slightly by increasing cathodic current density. Increasing amounts of excess A1F3 gave higher current efficiencies. Additions of LiF in the range from 1 2 wt% did not affect the current efficiency significantly at high current density.

The loss in current efficiency with respect to aluminium is mainly due to the so called back reaction between dissolved Al and the anode product according to:

Introduction Aluminium is produced by the Hall-Heroult process, where the overall primary cell reaction is: 2.»„„

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Alumina is dissolved in a molten fluoride electrolyte based on cryolite (Na3AlF6) and containing some A1F3 as well as CaF2 [1]. Modern cells are operating at ~ 955 - 965 °C. The current efficiency with respect to aluminium can be as high as 96 % and the corresponding energy consumption may be -14 kWh/kg AI in cells running at ~ 300 kA or higher. The annual production of primary aluminium was about 38 million metric tons in 2007, making it the most important electro winning process [2]. Although the basic principles of the original process remain unchanged, significant technological developments have taken place. The main improvements have been related to environmental issues and controlling the high induced magnetic fields. The productivity of the process can be increased by increasing the current in existing cells, and such efforts have been implemented in industrial cells in recent years. Another possibility is to increase the current density. The cathodic current density is

Sterten presented a theory for the mechanism for the loss in current efficiency for aluminium deposition in the industrial process [6]. Sterten and Solli and coworkers presented experimental results and model calculations for the current efficiency based on laboratory studies [7-10]. The effects of electrolyte impurities and electrolyte composition were included in their investigations. Realistic values for the current efficiency were obtained. The variation of the current efficiency with respect to current density, electrolyte composition and temperature was found to be closely linked to the concentration of dissolved metal. It is established that the back reaction takes place outside the diffusion layer next to the cathode/electrolyte interface. The rate

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of the back reaction is independent of current density, so the current efficiency with respect to Al should increase by increased current density. However, transport phenomena in the electrolyte in the cathode diffusion layer lead to more complex relationship between loss in current efficiency and cathodic current density [1]. The rate of the back reaction can be expressed as follows: vB = k' (dc/dx) = k' (c%)

(4)

where k' is a constant including the diffusion coefficient, c° is the saturation concentration of dissolved metal at the cathodeelectrolyte interface and ä is the diffusion layer thickness. Li et al. [11] recently presented a revised model for calculating the current efficiency based on local variations in cathodic current density related to the three-phase flow of electrolyte at the cathode interface. The average current efficiency for a typical 300 kA cell at 960 °C and CR=2.3 was estimated to be 95.0 %.

Figure 1. Current efficiency for aluminium deposition as a function of cathodic current density in molten Na3AlF6-Al203 (sat) with excess A1F3 corresponding to CR (molar ratio of NaF/AlF3) being 2.5 and 5 wt% CaF2 at 980 °C.

Experimental A laboratory cell similar to that of Sterten and Solli was used to determine the current efficiency for aluminium during constant current electrolysis [7-10]. The current efficiency was calculated from Faraday's law by weighing the amount of deposited aluminium. The electroysis time was 4 hours for each experiment.

Figure 2 shows the effect of varying electrolyte composition (A1F3 content) on the current efficiency for aluminium. Increasing amounts of A1F3 gave higher current efficiency. This is also in agreement with previous investigations carried out at lower current densities.

The electrolysis cell was placed in a closed furnace with dry argon atmosphere. A graphite crucible with a sintered alumina lining served as the container for the molten electrolyte. A steel plate was placed at the bottom of the crucible acting as the cathode. A graphite anode was immersed about 4 cm into the electrolyte and placed about 4 cm above the steel cathode. The influence of changing the cathodic current density and electrolyte composition including excess A1F3 and LiF additions was studied. The electrolyte was saturated with respect to alumina. Standard experimental conditions were Na3AlF6-Al203 (sat) with excess A1F3 corresponding to CR (molar ratio of NaF/AlF3) being 2.5 and 5 wt% CaF2 at 980 °C and 0.85 A/cm2. Results and discussion Liquid aluminium was deposited on the solid steel cathode plate. Steel was used to achieve good wetting of liquid aluminium, which does not wet the solid graphite. Current efficiencies ranging from - 92 - 96 % were obtained. The measured current efficiency versus current density is given in Figure 1. Good agreement with literature data was found at "normal" current densities. The current efficiency was found to increase slightly by increasing cathodic current density. This is in accordance with theory, since the main reason for loss in current efficiency is the back reaction between the primary products. The slight increase in current efficiency at high current densities is likely to be due to the fact that the transport phenomena at the cathode boundary layer cause a higher metal solubility, thus giving a slightly increased rate of the back reaction at increasing current density. However, the main reason for increased current efficiency at high current densities is due to the fact that the back reaction is almost independent of the current density.

Figure 2. Current efficiency for aluminium deposition as a function of the electrolyte composition in molten Na3AlF6-Al203 (sat) at 980 °C and 1.5 A/cm2. The effect of LiF additions of 1 and 2 wt% did not give any significant change of the current efficiency, which was determined to change from 94.2 % to 94.4 % and 94.0 % respectively. This is not in accordance with literature, where Dewing [12] and Tabereaux et al. [13] reported that the loss in current efficiency in industrial cells increases somewhat with increasing LiF content. However, theoretical considerations may suggest that the current efficiency should not depend on relatively small additions of LiF. Additions of LiF in industrial cells may

462

12. E.W. Dewing, Metallurgical Transactions B, 22B, 177 (1991).

lead to other changes of the electrolysis operation, so that industrial data for the current efficiency may not be reliable. The presented laboratory results for the current efficiency may not be directly applicable to industrial operation. Industrial cells have some constraints with respect to the design which is largely given by the energy and heat balance. An increase of the current density may not be feasible because it will imply a decrease of the interpolar distance in order to reduce the heat production.

13. A.T. Tabereaux, T.A. Alcorn, and L. Trembley, Light Metals 1993, The Metals Materials and Minerals Society, p. 221.

Conclusions The current efficiency with respect to aluminium was found to increase slightly by increasing the cathodic current density in the range from 0.85 - 1.5 A/cm2 during galvanostatic electrolysis in molten cryolite based electrolytes. The current efficiency was found to increase by increasing the content of A1F3 in the electrolyte. Additions of 1 and 2 wt% LiF did not cause any significant change of the current efficiency. Acknowledgment Support from Hydro Aluminium is gratefully acknowledged. References 1.

J. Thonstad, P. Fellner, G.M. Haarberg, J. Hives, H. Kvande, and Δ. Sterten, "Aluminium Electrolysis. Fundamentals of the Hall-Heroult Process", 353 pages, Aluminium-Verlag, Düsseldorf, 2001.

2.

U.S. Geological Survey, http://www.usgs.gov/

3.

M.A. Bredig, in "Molten Salt Chemistry", M. Blander, ed., Interscience, New York, 1964, p. 367.

4.

R. 0degärd, Δ. Sterten, and J. Thonstad, Light Metals 1987, The Metals Materials and Minerals Society, p. 389.

5.

X. Wang, R.D. Peterson, and N.E. Richards, Light Metals 1991, The Metals Materials and Minerals Society, p. 323.

6.

Δ. Sterten, J. Appl. Electrochem., 18, 473 (1988).

7.

Δ.Sterten, P.A. Solli, and E. Skybakmoen, J. Appl. Electrochem., 28, 781 (1998).

8.

P.A. Solli, "Current Efficiency in Aluminium Electrolysis Cells", Dr. ing. thesis, Department of Electrochemistry, Norwegian University of Science and Technology, Trondheim, Norway (1993).

9.

Δ. Sterten and P.A. Solli, J. Appl. Electrochem., 25, 809 (1995).

10. Δ. Sterten and P.A. Solli, J. Appl. Electrochem., 26,187 (1996). 11. J. Li, Y. Xu, H. Zhang, and Y. Lai, "An inhomogeneous three-phase model for the flow in aluminium reduction cells", Int. J. Multiphase Flow (2010), doi:10, 1016/j.ijmultiphaseflow.2010.08009

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

Improvement in Cell Equipment and Design SESSION CHAIR

Stephan Broek Hatch Ltd. Ontario, Canada

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

RETROFIT OF A COMBINED BREAKER FEEDER WITH A CHISEL BATH CONTACT DETECTION SYSTEM TO REDUCE ANODE EFFECT FREQUENCY IN A POTROOM Jonathan Verreault1, Bernard Desgroseilliers1, Rene Gariepy1, Claude Simard1, Serge Simard \ Xavier Delcorde2, Christophe Turpain2, Simon-Pierre Dery 2 ^ i o Tinto Alcan (Arvida Research and Development Center); 1955 Mellon boulevard, P.O. Box 1250, Jonquiere, Quebec, G7S 4K8. Canada 2 ECL.100 Chalant street, Ronchin, 59790 ,France Keywords: Chisel bath contactl, Anode effect2, P155 technology3 Abstract Chisel bath contact detection is a feedback system installed on aluminum reduction cell alumina point feeders to ensure that the crust breaker (or chisel) touches liquid bath. The benefits of chisel bath contact detection on anode effect reduction are well known on AP Technology™ operating with a separate system of breakers and feeders. The present paper describes the technology adaptation for a safe hot change installation of chisel bath contact detection applied to the combined breaker/feeder in a P-155 cell while, at the same time, introducing a longer pneumatic cylinder stroke. Mechanical modifications allowing the reuse of most of the existing breaker/feeder parts as well as the results obtained on four experimental cells during the development phase are presented. The retrofit procedure is also described in the context of a hot change while controlling health, safety and environment (HSE) related aspects. Introduction Background Rio Tinto Alcan is operating three smelters using the P155 technology (Sebree, USA, Grande-Baie and Laterriere, Canada). The three plants represent a total of 1200 cells using a combined breaker and feeder technology[l].

Figure 1: Operation without chisel-bath contact system Description of problem During cell operation it may sometimes happen that alumina has difficulty or is not able to reach liquid bath. This phenomenon can take place when the crust is too hard, particularly in the area of the two breaker/feeders. When this happens, the breaker is unable to maintain the opening in the crust. Bath may also accumulate on the chisel of the breaker/feeder. This occurs when the bath level is high (high pumping effect and high bath level fluctuations [3]).

Reducing greenhouse gas (GHG) emissions is part of the Rio Tinto Alcan objectives. To this end in 2006, the Arvida Research & Development Centre was commissioned to find a way to decrease the anode effect frequency (AEF) in P155 cells. The anode effects (AE) are the result of a shortage of alumina in electrolytic cells or when the alumina does not dissolve efficiently in the bath. This predominately generates two gases (CF4 and C2F6), which significantly increase GHG generation due to their high GHG intensity [2].

The combination of fixed cylinder travel and liquid height means that the chisel goes deeper into the bath and heats up with every crust-breaking operation. This eventually leads to liquid bath sticking to the chisel, accumulating into a solid mass. The mass grows at each crust-breaking operation and in the end keeps the alumina from entering the bath. Moreover, this additional mass increases the risk of mechanical failure and can even significantly reduce component life cycles. It also leads to premature chisel wear.

Operation without chisel-bath contact system When operating without chisel-bath contact system, the cell is fed with alumina using a simultaneously operated, single cylinder breaker/feeder combination. The breaker breaks the crust and creates a direct opening to the liquid bath and, at the same time, pushes in the alumina dose deposited during the previous crustbreaking operation. The fact that the opening in the crust and alumina feeding are done simultaneously in the same operation with two breaker/feeders is a particular feature of the P155 technology. A single pneumatic cylinder is used in the same operation for both tasks (Figure 1). The cylinder travels the same distance for each crust-breaking operation.

The problem can be summarized as follows: the breaker/feeders always operate in the same manner, regardless of cell conditions. They always travel the same distance, always stay the same amount of time in the down position, and always apply the same pressure, regardless of crust conditions. Furthermore, these situations raise HSE risks by increasing the number of operator interventions in problem cells. This also increases the occurrence of AEs that are directly linked to increases in GHG emissions.

467

and push down on the crust for a longer time to increase the chances of crust-breaking and enable alumina injection.

Causes of anode effects Anode effects on PI55 cells were analyzed and results show that about half of these are caused by the following: Liquid heights Blocked feed openings Frozen bath accumulated on chisel.

Another situation could occur if a voltage difference is measured even before the breaker has started to move. This is declared a permanent contact situation, and the breaker-feeder operates as if it was in the "alarm mode". A control system is needed to define which mode the breakers are in, and how each breaker-feeder should react. In short, the crustbreaking signal of the cell is sent to the control system, which activates the command to the breaker-feeders. The information from the crust-breaking operation is sent back to the system, which will creates an alarm if necessary.

All three situations lead to the same findings: alumina shortages in cells and/or inefficient dissolution. The other AEs are link with operation Choosing a Solution

Normal operation A crust-breaking signal is sent to the breaker-feeder, which travels down towards the bath. The chisel comes down and pushes the alumina into the bath. When the chisel comes into contact with the bath, the voltage difference is measured, which generates the lift command to the cylinder. During this sequence, the dosing unit comes down and deposites new alumina near the hole. This operation occurs more quickly than the standard cell operation, leading to air consumption savings [3].

Following careful analyses of potential solutions, the chisel-bath contact system used in AP Technology™ was decided upon. The main mechanical difference between the two technologies is found in the breaker/feeder system, which is divided into two separate, independent operations in AP Technology™ cells as opposed to PI55 technology. A joint partnership was set up between ARDC and ECL, who already have a tried-and-tested technology on the market, to develop an adapted chisel bath contact solution for P-155 combined breaker/feeder.

Alarm mode A crust-breaking signal is sent to the first breaker-feeder, which travels down towards the bath. The chisel comes down and pushes the alumina towards the bath. If the chisel does not come into contact with the bath or if the voltage difference is too low, the cylinder continues to travel down to build up maximum pressure over a long time period. During this sequence, the dosing unit comes down and deposites new alumina near the hole comes back up and sends an alarm to signal the situation.

Why the chisel-bath contact system? The chisel-bath contact system: Ensures that the chisel stays clean as it quickly withdraws from the bath. This eliminates the accumulation of solidified bath on the chisel and, consequently, cuts down on anode effects. Increases the potential of opening the feed hole in the crust as the full power of the cylinder can be used when called for. This increases the likelihood of alumina dissolution in the bath, thus reducing the number of AEs. Ensures that the chisel pushes alumina into the bath, thus increasing the total amount of alumina that dissolves into the bath and consequently decreasing AEs.

Installation of the new system on test cells Four cells in the same potroom were modified for installation of the new chisel-bath contact system. The new breakers were, in fact, standard P155 breakers that were reused and adapted to fit the chisel-bath contact system. Most of the standard breakerfeeder components were therefore recovered. The new breakerfeeders were installed on an operational cell, with tight control of HSE risks. It has been shown that breaker replacements can be completed in less than 45 minutes, as the result of using a quick fastening system. The replacements can be made on several cells at a time, according to potroom lifting equipment availability.

Ultimate objective for the PI55 chisel-bath contact system Determine the technological modifications and control conditions to decrease the anode effect rate and/or duration for PI55 technology. Develop and test the chisel-bath contact detection on P155 cells and adapt PBF's PI55 with the best knowledge Measure the possibility of detecting contact with the bath when using a combined breaker/feeder. Description of the P155 Chisel-Bath Contact System Detection is done by measuring the voltage difference between the chisel in the bath and the cathode, as shown in Figure 2. This requires perfect insulation between the components at cell potential and the breaker-feeder. This new element makes it possible to operate in two separate modes. The "normal operation" mode allows optimising power, air consumption [3], and alumina penetration in the bath. The other mode, "alarm mode", occurs when the voltage difference measured by the system is too low. At this moment, the cylinder pushes down over a longer period, allowing it to reach full power

468

BREAKAGE SIGNAL

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BATH

Chisel-bath detection percentage The percentage of successful contacts is a significant measurement for system performance. It represents the number of times the chisel successfully touched the bath following the crustbreaking signal. In other words, it is the number of times the system was able to measure a voltage difference for each breaker. The higher the contact percentage, the more trustworthy is the data received by the system.

METAL

Figure 2: P155 chisel-bath contact system Details of hot change works: • Cell preparation before hot change: EHS review with team performing risks analysis Adaptation of control system Installation of new electrical connections Prepare all equipments and tool needed • On the hot change day: Cell preparation just before hot change • Put cell in overfeed mode to avoid AE during the modification Stop one crust-breaker and keep in operation the second one Remove the stopped breaker Install adapted head with quick connection Install new crust-breaker with chisel bath contact Do the same with the second breaker feeder EHS observation with feedback to the team

It is theoretically possible to achieve 100% of successful contacts per day. However, operating conditions sometimes make it difficult to measure the contacts with the bath. For example, if a chisel comes into contact with an anode, it is processed as a permanent contact by the system and, consequently, is considered an unsuccessful crust-breaking. The same applies when a hole is blocked. Therefore, when the contact rate is suddenly lower for a breaker, it is a reliable indication that there is an issue with the cell. In future applications this signal can be used to send a message for the operator to take action quickly. The results show an average of 97% for the four cells, from December 2009 to June 2010.

Chisel-bath contact P155

Operational Results On December I s , 2009, four test cells were operating and were representative of the overall plant cell age. A minimum of four control cells were chosen for each test cell.

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Anode effect frequency The anode effect frequency is the number of anode effects occurring per cell per day. The results for the December 2009 to June 2010 seven-month period show a 56 % decrease in AEF with 0.11 AE/cell/day, compared 0.25 AE/cell/day for the control cell group. Figure 3 shows the monthly results from December 2009 to June 2010. The test cell results are always lower than their control groups.

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469

AE reduction due to breaker-feeders A new Pareto analysis was conducted to compare the AEs of cells with chisel-bath contact systems to those of 2006. The new results for anode effect causes shows that the chisel-bath contact system had a major impact on the distribution of AE causes. The breakerfeeders are no longer the major cause.

•

Results for the December 2009 to June 2010 test run show a 56% decrease in AEF, compared to the control group, down to 0.11 AE/cell/day.

•

With regards the duration of the anode effects, the new system made no difference.

•

The average percentage of successful contacts was 97%, from December 2009 to August 2010. This is a measure of system efficiency and proves the trustworthy of the data gathered by the system.

•

Figure 5 shows that the percentage of AEs due to breaker-feeders went from 59% in 2006 to 15% in 2010. This is a near 75% improvement, which leads to the conclusion that the chisel-bath contact system had a positive effect on AE reduction. Cause related to B/F Blocked holes Liquid level Crust on chisel

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With all these positive results, further AEF reductions can be expected with proactive action by the operators in addition to the chisel-bath contact system.

Figure 5: AE reduction percentages relative to breaker-feeders

References

Alarms and Operators

[1]

Based on the information sent back to the control system, two types of alarms are activated. These alarms could allow operators to take immediate action on the cell and be proactive on AEs, particularly multiple AEs. Note that during the testing period, the operators were not asked to respond to alarms since only the impact of the chisel-bath contact was measured. Hypothetically therefore, the AEF could have been decreased even further with operator intervention.

[2]

[3]

The two alarm types are: • Blocked holes Alarm activated when no voltage difference is measured for a complete crust-breaking operation. • Permanent detection Alarm activated when a voltage difference is measured before crust-breaking is initiated. A trouble shooting procedure was established to quickly target the cause of the problem and the corrective action required. This makes it possible to distinguish between process and mechanical issues. Conclusions A Chisel-Bath system was successfully developed for PI55 technology. Our principal conclusions are: •

Voltage differences between the cathode and chisel can be measured efficiently with a PI55 combined breaker-feeder.

•

The new chisel-bath contact breaker feeder can be installed on operating cells while controlling the HSE risks.

•

The breakers are, standard PI55 breakers reused and adapted for the chisel-bath contact system. Most of the standard breaker-feeder parts were hence recovered.

•

Standard cells can be converted to the chisel-bath contact system in less than 45 minutes. Several cells could be modified at the same time.

470

C.Richard, P.desrosiers, L.Lefrancois and B.Gaudreaul, "The Alcan's P155 Smelters now Operating at 195 kA. A Successful Assets Optimization Strategy", Light metal 2008, pp. 267-270 G. Bouchard, J. Kallmeyer, A.T. Tabereaux & J. Marks, "PFC Emissions Measurements from Canadian Primary Aluminium Production", Light Metals, J.L. Anjier Ed., The Minerals Metals & Materials Society, Warrendale, Pennsylvania, U.S.A., 2001, 283-288. Benoit Sulmont, Sylvain Fardeau, Etienne Barrioz and Pierre Marcellin/'The new development of the ALPSYS® system related to the management of anode effect impact", Light metal 2006, pp. 325-330

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ANODE DUSTING FROM A POTROOM PERSPECTIVE AT NORDURAL AND CORRELATION WITH ANODE PROPERTIES. Halldor Gudmundsson Nordural-Century Aluminum Ltd; Grundartangi; Akranes, IS 301, Iceland Keywords: Anodes, Dusting, Anode Spikes Abstract

The removal of dust from a cell is helped by the following [2]:

Anode performance in the cells is ultimately the most important measure of the anode quality and is not always reflected in the quality certificates. Nordural has through the years developed some tools to measure anode performance in the cells to use as feedback to anode suppliers so that they may improve the anode performance. Anode dusting in the cells leading to anode spikes and loss of current efficiency can be the biggest issue in the supplier - customer relationship. This paper shows some examples of anode dusting excursions experienced with three anonymous anode suppliers, how it was measured in the cells and how it was reduced or resolved in cooperation with the suppliers.

• • • •

A slight amount removed with tapped bath Burning in the cell gas flame By manual skimming of the taphole By cavity cleaning during anode change.

Modern prebake cells with point feeder with low anode effect frequency and which have well covered anodes with less open flame have a greater tendency to accumulate carbon dust. Quantifying the amount of carbon dust in cells has been done mainly by weighing the amount of skimmings from cells [1,2,3] but that material is a mixture of bath and carbon. Perruchoud et. al. [3] published formulas for quantifying carbon dust generated by air burn and carboxy burn based on anode properties as measured by R&D Carbon methods.

Introduction Smelters need to maximize production per cell day and minimize energy consumption. Minimizing gross carbon consumption is also necessary for smelters that purchase anodes from an outside supplier. From a potroom perspective the anode quality has a direct effect on current efficiency and also dictates how aggressively the operating parameters can be pushed, for example how thin the anode butts can be while still maintaining current efficiency. Quality of workmanship is also an important factor in pot operations and anode performance. Anode dusting in cells leading to anode spikes and loss of current efficiency is therefore a situation that both anode quality and operating conditions can contribute to [1,2,3]. Any attempt to quantify anode quality excursions from a potroom perspective must also take into account the effects of operating parameters. Carbon dust in the bath can be directly related to increased anode spike frequency and loss of current efficiency [1,2]. The carbon dust increases the electrical resistance of the bath leading to an increase in temperature and a reduction of the anode to cathode distance (ACD) [2].

Measuring the concentration of carbon in the secondary alumina is also a good indicator but not always reflecting the dusting in the cells [3]. Measuring the footprint of anode butts is a good measure of anode dusting as long as airburn due to poor covering practice does not influence the results. According to Stokka [6], carbon dusting in point feed prebaked pots seems to be more dependent on anode calcination level and the quality of the anode covering, than the laboratory measured air and C0 2 reactivity. This paper attempts to quantify dusting using a quick and simple method of observation and data is shown how it correlates with anode quality information. Anode performance in cells at Nordural The Nordural smelter was started in 1998 with an initial production capacity of 60,000 tpy and 120 pots in one potline. Subsequent expansions have been as follows:

The sources of carbon dust in a cell are [1,2]: • Dust from anodes due to selective oxidation of the pitch binder matrix releasing the petroleum coke grains into the bath. • Cathode wear which contributes between 0.5 - 1 kg carbon per tonne of metal produced • Carbon collar paste to protect anode stubs can add between 0.5 and 6 kg carbon per tonne of metal. • The recycled anode cover material can contain up to 5wt% carbon. • Carbon in the secondary alumina between 0.15 and 0.5wt%

• • •

30,000 tpy expansion of existing potline in 2001 130,000 tpy startup of second potline line in 2006 40,000 tpy expansion of second potline in 2007

Through capacity creep the current production capacity is roughly 280,000 tpy from 520 cells. The Nordural smelter has always had an emphasis on minimizing Gross Carbon consumption since the anodes are supplied outside Iceland.

471

Table 1: Physical and chemical properties of three anode types used at Nordural. Physical Properties Apparent Density (2 Stdev) Specific Electr. Resistence (2 stdev) Thermal Conductivity (2 stdev) Air Permeability (2 stdev) C02 Reactivity Residue (2 stdev) C02 Reactivity Dust (2 stdev) Air Reactivity Residue (2 stdev) Air Reactivity Dust (2 stdev) Spent butts fraction Slots in anode Collar paste used on anode pins S Fe Si V Ni Na Ca

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For the past 4 years Nordural has been using anodes from 3-4 suppliers at a given time. Table 1 shows the typical values of anodes from three suppliers which are the subject of this study. The table also indicates the fraction used of spent butts, type of slot made in the anode plus the calculated dust emissions from airburn and carboxy burn according to the formulas of Perruchoud et. al. [3]. The anodes are segregated into sections of cells, in numbers of 80 to 260 cells where they are used exclusively. The anode performance in the cells can thus be monitored and sections compared. The anode performance has been measured by observing: 1. 2. 3. 4.

The level of dusting in the cells as measured by a simple rating system of the dust visible in the tapping hole. The dusting is assigned an index from 1 to 3 (Figure 1) The frequency of anode spikes The footprint and thickness of anode butts. The carbon concentration of the secondary alumina.

Figure 1: The level of dusting evaluated in the tapping hole, in this case a dusting index of 2. This evaluation is done weekly by looking into 40-60 cells from each section using the same type of anode. This takes not more than 2 hours. An average dusting value is calculated from each such campaign. The more cells that exhibit no dusting, the closer this value gets to 1 so this system is designed to measure low dusting. As the average dust value approaches the value 3 the scale gets effectively saturated but then other parameters such as spike frequency, frequency of soft butts and high carbon in the secondary alumina start to increase or be more common. It is very important to do this evaluation at the same time in the work schedule and the most representative time is to do it just before the pot is tapped for metal.

The dusting is evaluated by checking the taphole. If there is dusting it will be visible there due to accumulation in the corners. A rough scale used to evaluate dusting is as follows: • • •

Dusting index = 1 means there is no carbon dust visible in the taphole. Dusting index = 2 means that there is some dust visible but it could easily be skimmed out once. Dusting index = 3 means that there is more dust than can be skimmed out once.

472

Anode dusting excursions Figure 2 shows the number of extra anodes changed per month due to spikes over a period of slightly more than 5 years for the three anode types considered. Anodes from supplier C is in use only part of this period. Figure 3 shows the average monthly dusting index for each anode type.

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A level of up to about 50 extra anodes changed due to spikes per 100 cells per month is considered to be acceptable taking into account the effects of operational parameters and the quality of anode covering. A level of more than 100 to 150 is a serious excursion requiring correction of operating conditions and evaluation of anode properties. Such excursions have lead to drops in current efficiency of about 2-5%.

Figure 5 shows the number of extra anodes change per month due to anode spikes, the dusting index evolution and the ACD evolution. Also shown is the anode stand time and anode butt thickness. The calculations of ACD are checked against measurements made two times during this period. During this period there are some line current fluctuations but the net trend is an increase in line current.

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Looking at the specific electrical resistance (SER) values, the reactive elements, the reactivity dust values (CRD and ARD) for the anodes the possible reasons for increased dusting around month 11 is perhaps a combination of poor baking and an increase of reactive elements. The second dusting and spiking excursion at month 51 looks to be mainly due to an increase in reactive elements since the level of baking had improved. Figure 7 shows that the air and carboxy dusting are typically on the high side which fits well with the elevated level of dusting and spike frequency.

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Supplier B anode were used during a startup and the dusting index remained high for the first 16 months. Spike frequency was initially quite high but dropped as the ACD was increased by dropping the line current and maintaining a high voltage. Around month 16 the dusting and the spikes drop while the ACD is fairly high at 5.1 cm. As the dusting drops during the next 25 months up to month 41, the line current is increased and the ACD decreases. From week 41 to 56 the dusting is fairly constant at 1.5 and the spike frequency remains below 50. Note that the dusting does not increase in the period of month 23 to 33 during a temporary increase in anode stand time with a corresponding reduction in butt thickness. Figure 8 shows that the SER value was quite low during this period.

Figure 5: (a) Dusting index and spike frequency, (b) Line amperage and calculated ACD. (c) Anode butt thickness and stand time.

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475

As for supplier B, the temporary increase in standtime for supplier C between months 15 and 20 does not lead to an increase in dusting.

appears to correlate mostly with the SER value as the reactivity values have consistently been at fairly low levels. A slight increasing trend again in dusting is being addressed in cooperation with the supplier with specific actions on relevant anode parameters.

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According to Table 1 the anodes from Supplier C anodes are the least reactive and this is reflected in the dusting index and the frequency of anode spikes. In this case there appears to be a slight correlation of dusting with SER. Here there is also a slight increasing trend in dusting that is being addressed in cooperation with the supplier.

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Evaluating the dusting level in the tapholes of cells is designed to measure low dusting and to establish a level of dusting that could lead to anode spiking excursions. Nordural uses the dusting index now in communications and meetings with anodes suppliers to spot anode dusting trends and address them with timely specific actions in order to avoid future anode spiking excursions.

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3. Conclusion 4.

In this paper are shown the results of studies over an elongated period of time from real-time data on anode dusting excursions experienced from the use of anodes from three different anode suppliers.

5.

The anodes produced by Supplier A show during this period of study in general more dusting and a higher level of anode spikes when compared to the other two suppliers. The data in Table 1 supports this as the reactivity values are generally higher. The level of baking by Supplier A has greatly improved during this period but a period of increased reactivity due to reactive elements still caused some problems. The moulded slots are a likely source of additional dusting, both due to higher permeability of the anode block around the slot and filler coke that is left behind from baking inside the slots [7]. Furthermore, because the sulphur level of this anode is lower than in the anodes from the other two suppliers the Supplier A anodes are more sensitive to increases in reactive elements [4,5].

6.

7.

The anodes produced by Supplier B anodes improved greatly during this period of study and the level of dusting and spikes

476

References B. Rolofs, N. Wai - Poi, "The Effect Of Anode Spike Formation On Operational Performance", Light Metals 2000. B. Sadler, B. Welch, "Reducing Carbon Dust? - Needs And Possible Directions", 9th Australasian Aluminium Smelting Technology Conference and Workshops, Terrigal, Australia, 2007. R. C. Perruchoud et. al., "Dust Generation And Accumulation For Changing Anode Quality And Cell Parameters", Light Metals 1999. R. C. Perruchoud, M.W. Meier, W.K. Fischer, "Survey On World Wide Prebaked Anode Quality", Light Metals 2004. P. Stokka, "The Effect Of Petroleum Coke Sulfur Content On The Loss Of Sodium During Baking", Light Metals 1994. P. Stokka, T. Berge, "Anode Carbon Dusting in Aluminium Electrolysis Cells", Slovak-Norwegian Symposium on Aluminium Smelting Technology, Stara Lesna, Slovakia 1999 J.J.J. Chen, "Slotted Anodes", University of Auckland 2008 MEng Studies Lecture Notes.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

THE APPLICATION OF CONTINUOUS IMPROVEMENT TO ALUMINIUM POTLINE DESIGN AND EQUIPMENT William Paul Rio Tinto Alcan - Smelter Technology - Centr' Alp - BP 7 - 38341 - Voreppe Cedex - France Keywords: technical package, improvement process, potroom, ancillary equipment visible part of AP Technology™ package improvements. The R&D work focuses on the process side of the technology; such as pot technology as in the case of Reduction. However, there is also a requirement for improving other aspects of the technology. Although these are often referred to as engineering improvements, Rio Tinto Alcan uses the label "industrialisation" to distinguish it from traditional R&D activities. This paper will explore the development and control of industrialisation work. Because of the range of the subject, discussion will be limited to describing the overall process and providing examples drawn from the Reduction.

Abstract The design and supply of technologies to a large number of aluminium smelter projects, over the last two decades, has provided a unique opportunity to benefit from continuous improvement. This improvement process starts with feedback from customers and suppliers, combined with input from R&D and internal reviews, to identify, validate and then incorporate experiences and innovation from many areas. The process can be described as systematic continuous improvement and the results are superior technology packages that incorporate the latest enhancements and better meet clients' needs. This paper describes the methodology used to manage the continuous improvement of technology packages.

Technology Package A smelter technology package is made up of a number of elements. In Rio Tinto Alcan terminology these are referred to as technology bricks. A technology package for an aluminium smelter involves the selection of many technology bricks. (Figure 1) Some bricks comefromprevious projects while others arise out of the needs of a new project. In most cases existing bricks in AP engineering database must be adapted to the specific characteristics of the project. A technology brick is an element of the engineering package detailing one part of a smelter, be it a complete shop or only a single piece of equipment.

It includes examples of improvements such as ancillary equipment redesign to improve safety and reduce forklift truck use. Also discussed are the challenges and opportunities of working with potline equipment suppliers. Effective cooperation enables the development of fit for purpose solutions that minimize potline commissioning issues and maximize pot performance. Introduction Constructing an aluminium smelter requires generating a technology package containing large amounts of documentation including drawings, specifications and procedures. In many cases the documentation is based on previous projects. However with a constant focus on investment cost and cost reduction, new ideas lead to changes and improvements. Expectations, especially health, safety and environmental (HSE), are also increasing from one project to another. Not only is it a challenge to generate the amount of documentation required, but to ensure it contains the latest improvements. Improving a technology package is not an easy task as many factors impact on its development. A big issue is that a smelter project is not the best place to develop new ideas. What a smelter project expects from its technology provider is a demonstrated technology to minimize risk. To meet these expectations the timeframe of a project does not allow for investigation and demonstration of new processes, layout, plant or equipment. In most cases this is because proposed changes need to be validated by trials. Therefore Rio Tinto Alcan, as the supplier of AP Technology™(I), has implemented, alongside project work, a parallel stream of activity to develop new ideas, processes, plant and equipment for inclusion in its packages.

RohUSt PBCk&g®

... provide the best solution for the project

Figure 1. Technology Package Building Block Catalogue The degree of interaction between the bricks is quite high. This adds to the complexity of modern technology packages and makes changes or additions time consuming to implement.

The R&D developments of the AP Technology™ are well known and documented [1,2]. For many decades these have been the AP and AP Technology are trade-marks of Aluminium Pechiney., used under license by Rio Tinto Alcan Inc.

477

In the AP Technology™ delivery process, as the engineering study progresses, batches of documents are delivered to the client in four lots at scheduled points in the project. The document breakdown matches both the need of the client and the EPCM. • • • •

Previous Approaches to Industrialisation Industrialisation work has always been carried out, both formally and informally. Traditionally improvement work was included in a project when the scope and timing of the improvements allowed. Some of the improvements were driven by changes brought about by R&D activities while others from specific cost reduction programs aimed at developing new cost effective solutions. Examples include Design to Cost and Full Economic Cost Reduction (FECRI) programs [3]. In these programs, each technical solution was tested and validated in operation in one of the RTA plants prior to being integrated as a brick in a Master Package.

lot A: Substation, paste plant and civil work design, shop descriptions and process flow diagrams, material specifications lot B: Pot element design: potshell, superstructure, busbars, lining, etc. Anode baking furnace & casthouse equipment, material handling lot C: Annex shop description and operation equipment lot D: Workshop organization and operating procedures

The Master Package concept is a set of reference bricks from which package elements can be drawn for a new project. While a worthwhile concept, in reality it can be difficult to maintain. Also it can be difficult to interface with R&D projects and ideas coming from other sources as these can vary from one project to another and require solutions specific to local conditions.

With the delivery of lot A, experts are seconded to the client's project team to assist in the use of documentation during the detail engineering and the procurement phases, and to forward client requests for alternative solutions. Adaptation of the package to the client needs requires competencies drawn from all sectors of the organisation including electrical, process, control, mechanical and civil design.

The reality is that every project is different and has a different set of needs. Often just as much modification is necessary for a Master Package document as there is in using previous project documentation as the starting point. The Master Package concept, while a useful concept, has been replaced by a new industrialisation system that focuses on continuous improvement.

Industrialisation Difference Between Industrialisation and R&D As shown in Figure 2, industrialisation is a process step in the implementation of R&D outcomes. Its role is to take R&D's new development and create the engineering documentation necessary for inclusion of the idea in a technical package.

- Smelters - Tech. Sales - Engineering

- Technology Feasibility - Economic Feasibility • Research Doc.

• Prototype - Smelter Demonstration - Technology Validation · Technical Support - Development Doc.: - Implementation Doc. - Identified suppliers

- Asset Optimization - High Amperage Pots - Breakthrough Technologies

The Present Approach The new Industrialisation system is built on the theme of continuous improvement. Its aim is to: • • • •

■ FEC final Validation • Tech. Brick Doc. • Work with suppliers

capitalize on information and share it between projects ensure changes are integrated into packages allow changes to be promoted to the client provide an ability to track document changes.

Full Economic Cost •Reduction Process:

New Technology Brick

Construction

Features of this approach include: Technology Brick Validation

Integration I to Package j

• •

Figure 2. Development of Solutions from Ideas to Implementation However the difference between R&D and industrialisation can be difficult to define. In Rio Tinto Alcan, R&D work is generally assumed to include an analysis of the fundamentals of the reduction process. An example is the design of pots which require studies of thermal performance and magnetic stability. But sometimes industrialisation does not follow the path shown in Figure 2 and avoids the R&D step with projects involving proven off the shelf technology that is put together in new or different ways. In practice the distinction between R&D and industrialisation can be somewhat blurred.

A timeframe different to that of smelter projects The type of effort required including: projects, package improvement and fast track

The demand for industrialisation work comes from many sources as shown diagrammatically in Figure 3. Some of the demand is part of normal work flow such as addressing requests from clients and existing plants. Other ideas and suggestions require effort to collect and analyse information, such as feedback from Greenfield commissioning and smelter best practice. However, no matter what the source, a formal process is required to collect, collate and address these ideas. This collection becomes a portfolio of ideas for potential industrialisation projects. The effort required to develop an idea varies greatly as does the timeframe involved. Because of this the ideas are separated into three categories:

Industrialisation is defined as the process for improving a technology by integrating new processes, layout, plant and equipment into a unified package suitable for a smelter project. Its role is to develop an engineering package to provide the commercial delivery of the technology.

1. 2. 3.

478

Fast track issues Package improvement ideas Industrialisation projects

Fast Track Issues These are issues that are identified as being required for the technology packages currently under development for a client where timing is critical. The issue is given a high priority with the industrialisation team and is investigated, resolved and implemented to meet the client schedule.

In the HUB system, an idea is listed and assigned to a person for clarification and evaluation. An on-line ideas form is filled in. Technical Directors monitor the new ideas on a regular basis and, in conjunction with project mangers, assign the Fast Track option on an as needs basis. Fast Track ideas are managed by the project manager.

Because of the time constraints, it is sometimes difficult to develop optimum solutions. Similarly innovation can be difficult because of the risk or lack of time for validation. For these reasons Fast Track can lead to subsequent industrialisation projects.

A review committee meets on a regular basis to review ideas, assign resources and establish priorities. Where possible the industrialisation work is linked to the schedule of an existing smelter project. The HUB database is a multi-project view, tracking changes across all projects and thus providing a traceability function. It provides a record of the evolution of ideas as well as the reasons for adopting the solution. It also documents when a solution is integrated into a particular package and by whom.

Package Improvement Ideas Many ideas involve improvements to the description of the technology in the existing documentation. This can include: • • • •

More detailed description of plant and equipment in the drawings Revision of wording in a specification More details in a specification The need for new specifications and/or drawings

Project team members refer to the HUB to determine which ideas and solutions are ready for including in package documentation currently under development. Status of Progress An industrialisation project system is differentiated from the usual project management systems by inclusion of some specific steps. In an industrialisation system the steps include:

While package improvement can require significant work, the solution is usually already known or defined. The effort is limited to producing the documentation. An example given later is the development of "residual risk" documentation.

1. 2. 3. 4. 5. 6.

Industrialisation Projects Industrialisation projects cover the issues not dealt with by the fast track and package improvement streams. Generally, these are longer duration projects requiring significant engineering input to integrate the solution into a technology package. This can include customising R&D solutions, customising or qualifying suppliers' solutions, or perhaps developing completely new solutions to meet changing requirements.

proposed approved in progress to be validated ready for integration integrated

The methodology for industrialisation projects is similar to other projects including the use of six-sigma methodology. However validation is a key step in industrialisation work. This can take the form of plant or workshop trials, prototype testing or review by subject matter experts. Validation needs to be fully documented to insure the thoroughness of the work and to provide tractability of technology changes. Once validated, it becomes ready for integration into a technology package. Upon integration its status moves to that of fully integrated and it is only here that the project can be considered complete.

The HUB A database called the HUB has been developed to control the flow of industrialisation work. This differs from normal project management software because it must keep track of solutions until such time as they can be integrated into the AP Technology™ package of a project. It also provides a formalised mechanism for this implementation.

Challenges with the Industrialisation Approach Challenges with implementing an industrialisation process can include: • • • •

Resource priorities Integrating with Greenfield project schedules Finding an appropriate method for integration in the technology package Solutions delivered in advance of their requirement.

The issue of resource priorities is traditionally a decision between having the technology package delivery team do the industrialisation work or setting up a separate team which focuses only on industrialisation work. Presently AP industrialisation work is carried out by the technology package delivery team. The challenge is to balance the industrialisation work with the team's

Rejection

Figure 3. Information flow in the HUB

479

primary purpose of delivering a technology package to schedule. This requires strong leadership to balance competing priorities.

These reviews identify hazardous phenomena taking into account HSE related aspects, such as installation and commissioning, normal use of equipment, its reasonably expected misuse, servicing by the user and maintenance by the maintenance personnel, smelter stoppage, demolition and equipment and material recycling. The probability and severity of these risks must then be evaluated. Risk reduction and management measures are then identified.

Integrating industrialisation work into a project provides further challenges. Deciding whether a solution is sufficiently advanced and validated for inclusion in a particular package needs to be part of the formal process of running an industrialisation portfolio. A further challenge arises if there is no immediate technical package for a solution to be incorporated into. The solution needs to be well documented and stored in such a way that it can be readily integrated into a package at a later stage. This leads to issues such as remembering that the work has been done and also determining if the solution is still valid. Again this is where the HUB database is invaluable because it ensures completed work becomes tagged as "ready for integration".

To better help the client understand and deal with the residual risks flowing from the basic engineering stage, new AP Technology™ package documents have been developed to identify residual risks associated with various sections of the smelter such as a substation or potline. An abbreviated example from a recent substation document is shown in Table I. Table I. Example of Residual Risk Documentation

Examples Examples are the best way of explaining the nature of industrialisation work. While the pot is often the focus of attention from a technology perspective, a lot of effort goes into the support plant and equipment and it is from here that the examples are drawn. Residual Risk AP is typically involved in the basic engineering phase of a project. Traditionally it reviews the risks and implements appropriate controls based on its understanding of the technology. However at the basic engineering stage and at subsequent stages not all risks can be controlled. Upon completion of each stage, the remaining risks, called "residual risks", must be handed on to be addressed in the next stage. (Figure 4) "Residual risks" are those risks to health, safety and environment (HSE) that are not eliminated or sufficiently reduced by the protective measures adopted in the previous stage of the project. Consequently residual risks need to be addressed by the client together with the client's project team. It was recognized that the communication of "residual risks" could be better managed and so specific residual risk documentation was developed.

Risk

Hazard explanation

Direct electrical contact

-

Risk of electrocution, contact with live parts: - High current bus, - HVbus& cable ends -

Indirect electrical contact

LV conductors in auxiliary boxes People falling (or objects dropped by people) due to surprise of electric shock Danger of electrocution with metal frames that get accidentally energized People falling due to the effect of surprise caused by an electric shock

Risk Reduction by Engineering & Manufacturer

Residual Risk 1 for client to address

- Fencing off the high current busbar - Signage of electrical rooms and boxes

-

- Fencing off the substation - HV equip, signage and identification

Substation access 1 rules: - Lock-out procedures - Earthing trolley procedure - LV power supply lockout

- As Built electrical drawings. - Equipotential bonding of metal frames + earthing. - Labelling indicating the risk of equipment connected to potline potential.

-

PPE for work on electrical installations

Regular monitoring of earthing and equipotential bonding. - Banning use of coppercorroding herbicide.

1

|

Start-up Tools and Equipment One area that is continually undergoing change is the design of tools and equipment. Smelters are always looking for better ways to carry out tasks. The RTA group in total provides a very rich source of inspiration with a variety of solutions to any given problem to draw from.

Residua) HS£ Risks after the basic design phase

As a result of feedback from a recent project, opportunities were identified for improving the design documentation of some of the start-up tools and equipment. This also provided an opportunity to review equipment safety in light of current standards. Two areas in particular will be discussed, the design of stands and their transportation.

...so making the work of the Project team easier. Residuai HSE Risks after detailed design phase to be managed by procedures...

Process Improvement Analysis of the faulty anode pallet showed that concept of having a single pallet for tools, the removed anode assembly and a bin for the broken butt pieces was not the way operation people worked. The single pallet was originally designed to minimize the transport of equipment to the pot. However at a smelter the butt pieces and anode assembly go to the

Figure 4. Residual Risk As well as considering the residual risks identified during the previous stage, risk analysis reviews are required at each stage involving equipment suppliers and future operators as appropriate.

480

carbon plant while the tools belong to reduction. A more logical approach was to split the pallet accordingly (Figure 5).

Supplier Integration Excellence of smelter operations not only relies on the performance of the pots, but also of the auxiliary equipment. In this context it is of high importance that suppliers' equipment becomes tested and integrated into the technology package. The example given here is that of entry into the cabin of the pot tending assemblies (PTA). In the traditional AP potline design, entry into the PTA cabin is via a set of wooden stairs located in the potroom duct end aisle. Because the PTA travel cannot be impeded by the stairs, the handrails cannot protrude above the top landing. This has meant that as the operator is about to enter the cabin, there is no or little handrail, Figure 8.

New Carbon Pallet

New Reduction Tool Stand

Figure 5. Spitting Up of Pallet for Faulty Anodes Figure 8. Traditional PTA Cabin Entry

Equipment Transportation Many smelters use forklifts for transportation of equipment in the potline. However it is recognized that forklift transportation of large items in the potline is less than ideal. The large size of the object makes visibility difficult for the driver or results in the practice of driving backwards, neither of which are considered best practice. Review of industry best practice has led to expanded use of anode transport vehicles for these tasks. Several equipment stands such as for ladle lifting beams and tapping pipes have been accordingly redesigned. An example is shown in Figure 6.

Original Trailer Design

This small, but significant issue has recently been addressed by a joint effort with ECL, the supplier of the PTAs. A review showed that not only the handrail but also control of access to the stairway was important. As a result there is a new approach to cabin entry that addresses both issues. A small, pivoting handrail section has been added adjacent to the entry point for the cabin. (Figure 9) Because of clearance restrictions during PTA operations, the handrail section needs to fold flush with the cabin when not in use and yet deploy automatically when the cabin is in position with the stair landing.

New Stand

Figure 6. Ladle Lifting Beam Transportation However, not all equipment stands are amenable to this solution. Some are still large enough to require a tow motor and trailer solution. Recent incidences have shown that hitching a trailer can be a hazard because an operator, or more particularly their hands, can be caught between a moving vehicle and a stationary trailer. AP (RTA?)now recommends using self hitching tow motors to address this issue (Figure 7).

Original Design

Figure 9. Pivoting Cabin Handrail The development of the cabin handrail went from concept design, through mock up testing to full scale workshop tests. In parallel the design of the stairway system was altered to include an interlocked door at the foot of the stairs. The interlock works such that the door will not open to gain stair access unless the PTA cabin is in position and the handrail has been deployed. This work highlights several of the previously mentioned issues of industrialisation work. These include the time frame to develop new ideas being different to the project timeframe, the need to integrate a solution into the technology package and the validation of the solution before it is supplied to a client.

New Design

Figure 7. Self Hitching Trailer Solution

481

potroom allows the operator to work safely. This design is a story of continuous improvement and optimization from one project to the next. Recent improvements were driven by the increasing number of pots in modern potlines. This results in increased potline voltage, which, when over 1500V, requires that insulation complies with IEC standards. This along with increased line current which impacts operating temperatures, affects the thickness, the creepage distance and the quality of the material.

Supplier Qualification In this competitive world, suppliers are always improving their products and looking for ways of reducing costs. This can be both an opportunity and a challenge for a technology package. While the aim is always to drive down the costs of smelter construction and operation, any changes must be assessed to make sure technical requirements are still met. An example of supplier qualification is the development of a new generation of ECL anode beam raising frames (ABRF) for AP pots. While the existing generation of ABRFs for AP pots have a well earned reputation, ECL decided to develop a new machine that was not only less costly but also had better functionality. It is not intended to discuss the details of the new design, rather to describe the collaboration with ECL needed to qualify the design for use in a Greenfield project.

The insulation design has also evolved to improve constructability and ongoing maintenance. For example, special neoprene slab bearings are used to avoid cracks in the concrete and in the insulation material. Naturally the material and the construction methods are designed to minimize capital costs. The transitional areas between earthed and insulated slabs have been improved to avoid short-circuits caused by pedestrians or vehicles moving between different potentials. A new busbar and pot support design takes into account their upgraded requirements. In addition, improved levels of engineering detail help ensure correct detail design Finally, the slab design has been changed to reduce cost while at the same time its electrical design is improved to increase operator safety. Conclusion Perhaps industrialisation is not as glamorous sounding as R&D, however, it provides a vital role in improving the content and robustness of the AP Technology™ package. A technology package is an ever changing product. There will always be the desire to improve it and to drive down costs for the benefit of the client.

Figure 10. New Generation of Anode Beam Raising Frame The main technology issues were the positioning of the frame on the superstructure and adequacy of the new clamping system. In ABRFs for AP pots, the frame is lowered onto the superstructure and then clamping arms are lowered against the anode stems. This is a two step operation that requires mechanism to control the raising and lowering of the arms. To simplify the frame, ECL returned to an earlier design which is still in use in many smelters. In earlier designs, the frame is lowered onto the superstructure and at the same time operators must check that the stem clamping system aligns with each stem. In the new ABRF, ECL reverted to the previous design concept but addressed alignment issues by redesign of the guidance system onto the superstructure. In order to be assured that the new design would perform adequately with AP pots, joint ECL and AP reviews were held to consider the new design, visits made and advice sought from smelters operating close variations of the new design.

Industrialisation ensures the proper integration of equipment and processes into the AP Technology™ package through controlled testing and qualification of all improvements. However, tight integration also means that simple answers are not necessarily easy to give. It can take time to establish all the impacts of a proposed change. Industrialisation is an ongoing process used to validate equipment and processes before delivery to the client. Thus, clients can be better assured that proposed equipment and processes are fit for purpose. Acknowledgement

A prototype of the clamping system for a single stem was manufactured and a series of tests conducted at the St Jean de Maurienne smelter to ensure the fitting and clamping of the mechanism on an operating pot in the magnetic field. As well workshop endurance testing was carried out. Issues arising from these tests were resolved with further modifications to the design.

The author wishes to thank all contributors from Rio Tinto Technology and in particular D. Lamant, P. Breme and A. Charbonnier. References [1] O.Martin, L.Fiot, C.Ritter, R.Santerre, H.Vermette "AP30 toward 400 kA" Light Metals 2009, 445-450

This example demonstrates the effort required to qualify changes to critical equipment before its implementation in a technology package can be considered.

[2] B.Benkahla, O.Martin, T.Tomasino "AP50 Performances and New Development" Light Metals 2009, 365-370

Potroom Electrical Insulation

[3] O.Martin, S.Despinasse, C.Ritter, R.Santerre, T.Tomasino "The FECRI approach and the latest developments in the AP3X technology" Light Metals 2008, 255-260

In a potroom operators are in contact with equipment or building elements at a high electrical potential. However, the design of the

482

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Alcoa STARprobe™ Xiangwen Wang, Bob Hosier, Gary Tarcy Alcoa Primary Metals, Alcoa Inc., USA Keywords: Electrolysis, aluminum smelting, aluminum reduction, analysis method, differential thermal analysis, temperature, superheat poor cell control that is often either under or over the target conditions. This under- and over-control of operating target leads to sub-optimal cell performance in current efficiency and energy efficiency. Another issue with the bath sampling and lab analysis procedure is the high probability of information mix-up which will also lead to an inappropriate control action. Bath samples are taken from many electrolysis cells. These samples must be kept in order and numerically tracked through the lengthy process of preparation and analysis. Every time the sample is handled, there is a possibility of sample mix-up and possible sample contamination.

Abstract Alcoa STARprobe™ is a probe device/system used to measure cryolitic bath properties including Superheat, Temperature, Alumina concentration, and cryolite Ratio (acidity), STAR, all together in real time for active pot control. The proprietary measurement principle is based on differential thermal analysis (DTA). This paper shows the fundamentals of operation along with the correlation of all the analysis with the accepted methods of XRD and pyrotitration for acidity, thermocouples for Bath Temperature and LECO and XRF analysis for alumina. The timing of the measurement will be shown to be equal to the traditional methods and the reliability (including reusable use of the probes) will also be described.

In addition to drawbacks of the conventional measurement methods, there is also lack of some key bath physicochemical information, such as liquidus temperature that is critical to efficient operation. Some advanced measurement methods have been a subject of study in the past decades (1, 2). The commercially available measurement tool by Electronite, made it possible to utilize measured bath superheat for active pot control (3, 4). Though it was a step forward from the traditional method, it never reached large scale application across aluminum smelters because: 1. The consumable and expensive probe tip could often not be justified for routine pot control use. 2. Incomplete information from the measurement. The liquidus (or superheat) can be a function of bath chemistry (ratio or %XS A1F3) and/or alumina concentration. Lack of information with respect to the cause for an undesirable liquidus temperature could lead to very different control decisions based on the assumptions made regarding the cause. This in turn could result in inefficient operation if not disastrous consequences. To effectively control an operating cell and to achieve its maximum efficiency, energy state, chemical state, and state of control should be known. These states are represented by core parameters including cryolite ratio (%XS A1F3), temperature, superheat and concentration of alumina. This paper presents TM Alcoa STARprobe , an advanced measurement device developed as a real time analytical tool for measuring the necessary cell information and then instantly supplying the information to the host computer for control decisions.

Introduction The efficiency of aluminum smelting cells relies on sophisticated control in maintaining cell's thermal and material balances by regulating resistance, bath chemistry, and alumina feed. A good control of the cells is dictated by reliable and accurate measurement of key cell's operating parameters, i.e. cell (bath) temperature and electrolyte chemistry (cryolite ratio and alumina concentration). Traditionally, bath temperature measurement, bath sampling, and its subsequent analysis are usually carried out separately by different crews and sometimes at different times. Bath temperature is usually measured using a thermocouple at a frequency of about one per day. This batched process is carried for a whole room or line. Bath compositional analysis is also a batch process that comes from sampling the individual cells, preparing the individual samples (grinding to a suitable particle). Bath analysis is a lengthy process even if some of the preparation and analysis is robotically automated in some modern smelters. Depending on the number of cells, it will take from 6 to 24 hours before results are known and corrective actions can be taken. This separate arrangement of measurements and delay in obtaining results is due to the fact that there are no better measurement options. The drawbacks of the lengthy and tedious temperature and bath sampling/analysis are obvious: When the bath samples are being processed and analyzed, the electrolysis cells are continuously operated and their bath temperature and cell chemistry are dynamically changing due to the variation of power and material input. Control decisions in both material (chemicals) and energy (voltage/resistance) input have to be made primarily relying on old information and empirical guessing at the "current" cell condition. And this "current" cell condition may be significantly different from the "real" cell condition due to a lot of unknown factors during the period. Consequently, the inability to measure cell temperature and real time bath chemistry inherently results in

Alcoa STARprobe™ is a measurement tool for use in the potroom. It measures and gives bath Superheat, bath Temperature, Alumina concentration in bath and cryolite Ratio (excess A1F3) in a single measurement. It also instantly transmits the results to a host computer through wireless communications in the potroom. This procedure unites the conventional processes of temperature measurement and bath sampling-analysis into one online measurement, simplifies and greatly shortens the process and time space from measurement/sampling to pot control decision. The pot control decision can therefore be based on the real time

483

Differential Thermal Analysis (DTA) The principle of DTA in material characterization is illustrated in Figure 2. Any material which goes through phase transformation associated with energy (heat) release, can be analyzed by DTA. The sample is typically placed in a sample cup/container in parallel with a reference material. When the sample together with the reference is heated up or cooled down, the heat adsorption or release due to phase transformation can be reflected through the temperature difference (TC1-TC2).

cell conditions rather than those from few hours ago or from as long as 12 or 24 hours ago. Theoretical Background Crvolitic Melt Cooling Characteristics For simplicity, Na 3 AlF 6 - A1F3 binary phase system shown in Figure 1 is used for discussion.

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Differential thermal analysis (DTA) for materials characterization.

If a cryolitic melt (shown in the phase diagram Na 3 AlF 3 - A1F3 in Figure 1), is placed in DTA (Figure 2) and cooled against an ideal reference in a control environment, cryolite (Na3AlF6) and chiolite (Na5Al3F14) may be easily determined. Shown in Figure 3 are three different compositions cooled from their melting state.

Na3AlF6-AlF3 phase diagram.

When a cryolitic melt is cooled from liquid state to a solid state, it will go through several phase transformation/changes. For example as melt A starts to cool from its liquid state down to its liquidus (melting) temperature (~960°C), solid cryolite starts to form and the remaining liquid will have an increased A1F3 concentration. This will continue changing along phase diagram liquidus line (the red region in the graph): Liq. Bath -> Na3AlF6(s) + Liq.(var. XS A1F3)

(1)

As the temperature is cooled down further to its per-eutectic temperature, solid chiolite forms (the blue region in the graph): Na3AlF6(s) + A1F3(1) -> Na5Al3F14(s)

(2)

As the temperature cools further, the liquid will transform into solid leaving behind liquid at increasing A1F3: Figure 3: Typical DTA patterns of cryolitic melts with three different %XS A1F3 or bath ratio (A, B, and C).

Liq.bath -> Na3AlF6(s) + Na5Al3F14(s) + Liq.(var. XS A1F3) (3)

Three distinct patterns, specifically, the peaks of delta T for cryolite and chiolite appear. The cryolite ratio (NaF/AlF3) or %XS A1F3 is directly related to the heat of release during the phase transformations and therefore the magnitude of delta T:

Finally, when the temperature reaches the eutectic (intersection of the blue green and yellow regions) there is no more remaining liquid: Liq. bath -> Na5Al3F14(s) + AlF3(s)

(4)

%xs AlF3=//[AHA1F3/Na5A13F14/(AHA1F3/Na5A13F14+ AHNa3AiF6)] =/2[S2/(Sl+S2)] =/3[DT2/(DTl+DT2] (5)

As can be seen from the phase diagram, each phase transformation of cryolite components occurs at specific temperature regions, and component phase changes are accompanied by heat release (ΔΗ*). The amount of heat release is directly proportional to the amount of the component.

The Alcoa STARprobe™ was developed based on the principle of DTA so that those unique cooling characteristics can be clearly revealed and the cryolitic ratio accurately determined (5, 6). The ratio result can be known after carrying out a few simplified steps:

484

• • •

temperature was not kept constant, the superheat therefore changed accordingly.

First insert probe tip in molten bath to equilibrate with bath temperature in an electrolysis cell Take probe tip out of molten bath and allow it to cool STARprobe™ analyzes the cooling curve and reports the results. Results and Discussion

The current Alcoa STARprobe™ system is shown in Figure 4. The STARprobe consists of four major components: 1.) reusable probe tip, 2.) portable probe stand/lance to fit various smelters for measurement, 3.) electronics which acquire temperature data, analysis and carryout wireless communications, and 4.) STARprobe™ program which performs all the necessary tasks during measurement.

Figure 5: a.) The probe tip (with molten cryolitic bath) is being cooled, b.) Temperature of bath sample and reference, and c.) differential temperature as a function of sample temperature.

Figure 6: Cryolitic electrolyte temperature and liquidus as revealed by the probe for three ratio conditions: a.) 1.20, b.) 1.10, and 3.) 1.00 from left to right. Table I

Figure 4: Alcoa STARprobe

Test No. 1 2 3 4 5 6 7 8 9 10

measurement system.

Probe Tip Development The probe tip and the material used in constructing/making the probe tip are considered to be critical in effectively revealing cryolite cooling characteristics in a repeatable fashion. Another important consideration is the temperature measurement (7, 8). The type K thermocouples used in the probe tip offer accurate temperature (better than ±0.4%), fast response time to temperature and, most important of all, being able to withstand molten bath for multiple uses. One of the probe tip designs is presented and shown in Figure 5 a.

Avg:

Bath Temperature and Liquidus by STARprobe Temperature, C 961.5 965.3 962.9 960.3 967.3 964.8 965.3 969.3 967.4 963.9 964.6

Liquidus, C 944.8 944.0 944.6 945.1 944.0 944.5 944.0 944.8 944.8 944.0 944.5

Superheat, C 16.7 21.3 18.3 15.2 23.3 20.3 21.3 24.5 22.6 19.9 20.1

1

Bath superheat was compared with a commercially available method. Figure 7 shows comparison of results obtained in a smelter setting. Each point represents an operating cell: The results of Alcoa STARprobe™ agree well with the "reference" method. The difference between the two methods as observed is believed to be mainly due to 1.) different algorithms in determining the "true" liquidus temperature at its "inflection" point, 2) natural variation of smelting conditions especially when there is feeding and 3) the fact that the two were not measured at the "exactly" same time.

A typical cooling curve of a molten cryolitic bath from a molten sate to a solid state in the probe tip is shown in Figure 5 b and c. Bath Temperature and Liquidus: The probe tip is designed so the liquidus temperature is clearly revealed with no supercooling. As an example, Figure 6 shows bath liquidus temperatures for three typical cryolite ratio conditions, i.e. bath ratio of 1.20, 1.10, and 1.00, a range representing aluminum smelting conditions.

Cryolite Ratio: As illustrated in Figure 3, the magnitude of heat released from the cryolite and chiolite phase transformation is a function of their initial constituents. Figure 8 shows how the STARprobe™ measures bath ratio by first utilizing the cooling characteristics bath with different starting ratios. When bath ratio increases, the liquidus temperature increases, the differential temperature for cryolite phase formation (marked as peak 1) increases, and the differential temperature for chiolite phase formation (peak 2) decreases.

Table I shows the typical accuracy for determining liquidus temperature when electrolyte composition was maintained constant. Each measurement was repeated ten times and each time it started with a different bath temperature. The standard deviation of the measured liquidus was 0.43°C. Since bath

485

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as shown in Figure 10, the probe tip is able to detect varied alumina concentration in bath. When alumina concentration increases, the bath liquidus decreases. Most important of all, there is another peak (heat of release) between 800 and 900°C, and this peak increases when alumina concentration in bath gets higher. These two major features plus other features allow the alumina concentration in bath to be measured by the STARprobe™ as shown in Figure 11. Like bath ratio, the STARprobe™ has a non-biased response in measuring %alumina in bath but with some reduced accuracy compared to ratio. The maximum variance was 0.14% and maximum difference between targeted and measured value was 0.66%.

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With a proper analysis of these cooling characteristics the ratio is determined. Shown in Figure 9 is the bath ratio determined using the key cooling curves characteristics. The probe tip produces a non-biased response to the ratio variation in the range of interest. The maximum variance (3 measurements at each ratio condition) was 0.00015 and maximum difference between the target and measured ratio was 0.012. 160

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Figure 11: % Alumina as measured by STARprobe™ using key cooling alumina characteristics. Environmental Impact: The probe tip is cooled in the ambient environment after it equilibrated with bath temperature when removed from the pot. Figures 12 to 14 show the potroom ambient conditions on probe cooling have little impact on the measurement of liquidus temperature, ratio and alumina concentration due to the differential nature of the measurement. Under three ambient temperature conditions, i.e., >150°F (65°C), 78°F (25°C), and <150°F (-101°C), the following results were obtained: • For liquidus: No significant impact was observed. Standard deviations at the two temperatures or electrolyte conditions) were respectively 1.13 and 0.98°C • For ratio: the standard deviations at two ratios were found to be: 0.0123 and 0.043.

W. Ratio (Target)

Figure 9: Bath ratio as determined using key cooling characteristics in STARprobe™. Alumina Concentration: Typical alumina concentration in smelting bath ranges from a low 1.5 to as high as 7.0wt%. The relatively low concentration of alumina represents a small amount of heat release when alumina crystallizes out of a cooling/freezing bath. This small amount of heat release means limited representation on the cooling curve with respect to the other phase transformations. The sensitivity of detecting the heat release is therefore more limited. Nevertheless,

486

For alumina: the standard deviations at two levels were found to be: 0.24% and 0.13%.

Figures 15 and 16 show photos of PDA and tablet PC based STARprobe™ systems. The major difference between the two is that PDA based is a standalone unit while a single tablet PC based unit can control up to two probes (also shown in Figure 4). Figure 17 further illustrates the tablet based STARprobe™ measurement cycle: two probes are inserting in tap hole of each cell in measurement, temperature data transfers to tablet via wireless, the tablet processes the measured data and transfers the results to computer server via the wireless network. The computer server gives control instructions to the pot to complete the control cycle.

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Figure 14: Impact of ambient temperature on alumina. Alcoa STARprobe™ System Alcoa STARprobe™ system was designed and built by incorporating the core probe described in the previous section. It has evolved from early MS-DOS™ based in early 2000 to the current version. It is a portable and yet fully integrated system for making measurement on cells of different technologies: • PDA or tablet PC based electronics which satisfy special potroom conditions (high magnetic field, high ambient temperature, and highly dusty environment). • User friendly and yet robust running program with simple graphical user interface (GUI) which is targeted to all audiences from pot operators to engineers. • Wireless data management which takes advantage of secured wireless technology in both acquiring temperature data and transferring results to the computer server. This makes it possible for real time measurement and pot control. • Self adjustable stand/lance and probe tip assembly which fits all cell technologies (pre-baked or Soderberg, floor or deck plate based) for an easy measurement.

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Figure 17: Schematic showing the STARprobe™ measurement and data management in a potroom setting. Measurement Time: Since the probe tip is reused, a complete measurement requires a few steps from setting probe tip in bath, taking probe tip out of bath to cool, to put the probe tip back to the pot to reheat for dumping bath out of the probe for next new measurement. When

used continuously, average time is just under 4 minutes to complete the measurement cycle. STARprobe™ Repeatability/Resolution

Max Min

Under Controlled Lab Conditions: To determine repeatability or variability of liquidus, three repeated measurements were carried for each bath composition, temperature was intentionally changed at each composition, and 16 bath compositions were studied over a period of time. This was also to determine if there is any possible chance of supercooling in determining liquidus. Figure 18 shows the bath temperature and liquidus of three measurements at each cryolite composition. The bath temperature ranged from a low of 935°C to a high closed to 1000°C while the liquidus ranged from a low of910°Ctoahighof980°C.

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As shown in Table II, for a 16 bath (ratio) compositions, the average standard deviation for the measured bath ratio is 0.0085 (max. 0.02 and min 0.00), while for liquidus, the standard deviation is 1.1°C (max 2.8 and min 0.2°C. The variability of the liquidus at each bath composition is independent of bath temperature variation or the difference between temperature and liquidus, indicating there is no apparent supercooling in measurement.

Figure 20:

Bath ratio as measured by STARprobe™ and analyzed by sampling/XRD. Summary

11

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Alcoa STARprobe™ is designed and developed to measure key cell parameters, i.e. bath superheat, temperature, alumina and ratio for real time active pot control. This paper presents its development, working principles as well as results in comparison with traditional methods. It is being used across Alcoa smelters, replacing the conventional sampling methods.

46

-Bath Temperature

Acknowledgment

Figure 18: Bath temperature and liquidus of 3 repeated measurements at electrolyte composition.

The development of the Alcoa STARprobe™ was only possible with the efforts of many departments. The authors, with deepest appreciation, wish to acknowledge Geff Wood for supporting and supplying talents of his department in coding the STARprobe™ program and choosing the proper electronics, without of which there would not be a STARprobe.

Examples Measured in Industrial Smelting Cells: Three repeated STARprobe™ measurements were carried out for 10 cells, and 3 bath samples were also taken at the same time. Bath samples were then analyzed with conventional analytical method. Figure 19 shows the bath temperature and liquidus of all measurements, and Figure 20 shows the ratio measured by STARprobe and analyzed by XRD method. The variance in ratio is comparable: an average of 0.00012 of sampling/XRD vs. 0.00010 by STARProbe.

References: 1. Y.R Gan et al: "Multifunction Sensor for Use in Aluminum Cells", TMS Light Metals, 1995, p233. 2. P Verstreken and S Benninghoff: "Bath and Liquidus Temperature Sensor for Molten Salts", TMS LM, 1996, p437. 3. M. P. Taylor and J. J. Chen: "Applying Control Principles to Superheat", 1st Superheat user Meeting, 19th - 23rd, March 2005, South Africa. 4. P Verstreken: "Review of 9-box Control - How can we improve", 1st Superheat user Meeting, 19th - 23rd, March 2005, South Africa. 5. C. Bates: US Patent 6,220,748, expires 1/15/2019. 6. B Hosler et al: US Patent 6,942,381, expires9/25/2023. 7. D. J Madsen: "Temperature Measurement and Control in Reduction Cells", TMS Light Metals 1992, p453.

Table II Repeatability (Variance) of Bath Liquidus (Tb-Tliq),C Liquidus, C Bath Ratio

Average Max Min Average Max Min Average

Stddev 10.5 19.3 2.2 1.07 2.82 0.19 0.0085

488

8. X. Wang et al: "Paradox in cell Temperature Measurement Using Type K Thermocouple", TMS Light Metals, 2006, p279.

489

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ACTIVE POT CONTROL USING ALCOA STARprobe™ Xiangwen Wang, Gary Tarcy, Eliezer Batista, Geff Wood Alcoa Primary Metals, Alcoa Inc., USA Keywords: smelting, bath chemistry, pot control, temperature measurement temperature changes from low to high, bath ratio control attempts to maintain the target ratio which will at least lead to sub-optimal operation and at worst a pot failure. It is also true for cases where a pot temperature transitions from a high to a low. Ultimately the bath chemistry control, no matter how sophisticated the algorithm, produces unfavorable conditions due to lack of real time analysis and lack of integration with the rest of the system. Eventually this negatively impacts energy efficiency.

Abstract To run an aluminum smelting cell, routine bath sampling and subsequent chemistry analysis are required along with pot temperature measurement. The sampling and analytical process is lengthy, tedious, and very often results are delayed as long as 24 hours. In addition the results are not coupled to other critical information (e.g. noise, automatic resistance adjustments, etc) at the time of the sample. Alcoa STARprobe™, which was previously described (1), corrects these deficiencies while providing a means to more efficiently and effectively control a smelting pot. This paper presents the background philosophy for an advanced control that has been enabled by the new measurement technique. The control method has been applied in multiple plants and demonstration of improved performance will be shown.

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Pot Temperature

Introduction

Figure 1: Traditional pot ratio and temperature control.

In a previous paper, Alcoa STARprobe™ was described as a real time measurement tool for active pot control (1). The motivation for developing the STARprobe™ is obviously to improve the efficiency of the operation. Since the probe allows a coordinated real time measurement of superheat, temperature, ratio (acidity), and alumina concentration, this information is easily coupled with the other known pot operating conditions such as noise, voltage modifiers, and state of feed control. This paper shows a detailed comparison of STARprobe™ and traditional sampling analytical results on a plant scale in routine production and the principle of an advanced control algorithm which has been enabled by the new integrated technique.

Another deficiency is when a pot suffers problems from anode or cathode issues, raw material changes, or operational upsets (unplanned maintenance). Conflicting decisions between bath chemistry and temperature control can be made that produces an unfavorable condition with excess high superheat. For the worse cases, excess high superheat results in: • Loss of side ledge and potential pot failure • Loss of current efficiency (7) that generates extra heat • Loss of bath cover with loss of alumina control and high emission. Some advanced control algorithms try to overcome the deficiency by using various models to get both bath chemistry and temperature under control in an operating target zone. However, due to lack of key links between material and thermal balances, inefficiency still exists from the inability to control bath chemistry and temperature in a closely coordinated manor.

Deficiency of Bath Chemistry Control Bath chemistry and cell operating temperature are controlled to achieve optimal current efficiency and energy efficiency. Specific chemistry targets and the corresponding operating temperatures are commonly dependent on cell technology. Typically, control of these relies on traditional bath sampling/analysis and separate temperature measurement. The bath chemistry control is done by manipulating chemical additions (A1F3 or soda ash) (2 to 4), while pot temperature control is achieved by maintaining a specific resistance target (5, 6). The two controls are usually carried out independently with very little coordination. As illustrated in Figure 1, chemical additions (A1F3 or soda ash) drive bath ratio to target, and pot resistance (voltage) adjustments drive the pot temperature to its target. Usually this approach is good enough as long as both bath chemistry and temperature are operated within targeted operating ranges. During a transition period when a pot

Dynamic Link Between Material and Thermal Balances Material (chemical) and thermal (energy) properties are continuously in a balanced but dynamic state. Pots under normal operating conditions usually have a fairly consistent heat balance and the heat flow through the pot side and end walls stays fairly constant (8). The relationship of heat flow to and from the sidewall at the bath zone may be approximated by Q = Q (to ledge) = Q (from ledge) = Q (awayfromshell) = h ·Α·(¾ -Τ ) B-L

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where Q = heat flow rate through side-ledge hB L = heat transfer coefficient: bath to ledge heat transfer Tbath = bath temperature TUquidus = liquidus temperature of bath AL = cross-sectional ledge area ΔΤ = bath superheat (Tbath - Thquidus) The heat flow to ledge interface is driven by bath superheat [ΔΤ], if a constant superheat can be maintained during pot operation, the heatflowthrough sidewall would remain constant and: Τ ^ ~ Τ ^ = Constant

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The Alcoa STARprobe™ based control package establishes both thermal (power) and material (chemical) balance at a constant superheat to enable and, more importantly, maintain a high operating efficiency. Figure 4 illustrates the strategy of Alcoa's STARprobe™ pot control: STARprobe™ results are transmitted to a database where proprietary control algorithms will tie the STARprobe™ results together with pot running status (historical and current) and give operation control orders to the potroom floor for chemical alumina and power adjustments. These adjustments will achieve the maximum current efficiency and minimum power requirements.

The liquidus temperature is impacted by bath chemistry (constituents) (9): (3)

Combining (2) and (3), relationship can then be easily derived: Ratio = C + C «Tbi

(4)

Equation (4) is the ratio (XS A1F3) variation purely due to pot thermal change as expressed by bath temperature. Figure 2 shows an example of bath ratio and %XS A1F3 as a function of bath temperature at a constant superheat of 10°C. As the temperature increases due to increasing voltage or resistance, a part of the side ledge will melt whereby the bath ratio increases (%XS A1F3 decreases). It is operationally impossible to run a pot at an exact constant superheat, but it is not impossible to target a range. Figure 3 shows an example of a pot operated at a superheat range of 5 and 15°C. The bath ratio increases as bath temperature increases when the superheat is within this range. In this case, the acidity change is from the freezing and thawing of cryolite in the ledge and not from neutralizing Na20 from alumina during the electrolysis process. Fundamentally, when a pot is in thermal balance its ratio can be known purely based on the temperature and this relationship can be used to judge how well and efficiently a smelter is operated.

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Figure 4: Alcoa Proprietary STARprobe™ pot control. Results and Discussion The STARprobe™ is deployed by carrying out plant wide "side by side" comparison to the traditional analytical method to assure a smooth transition to the STARprobe™ method of operation. Then, after the measurement technique has been proven, STARprobe™ control is phased in as a second step. To date, Alcoa STARprobe™ has been successfully deployed in eight Alcoa smelters. The longest time the first smelter starting to use STARprobe™ for pot control has been over 3 years (by the time this paper was written).

Pot operation is all about maintaining a thermal and material balance. A dynamically well-kept pot (both material and energy are maintained well balanced) is at peak operating efficiency when the ratio, temperature, and superheat are at the optimal condition. 1

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y^.\.... Figure 2: Ratio (XS AIF3) as a function of temperature for a thermal and material balanced cell at 10°C superheat.

492

Smelter A: Figure 5 shows the comparison of %XSA1F3 results from two of typical daily measurements. Bath samples were taken and

analyzed by pyro titration while STARprobe™ was used to make measurement for the same pots. The %XS A1F3 typically varied from a low of 6% to a high of 15% and varied randomly from pot to pot without specific order/pattern. The %XS A1F3 by the two methods overlapped and agreed well. One D a y M e a s u r e m e n t s

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Figure 8 shows the typical bath temperature and superheat distribution by STARprobe™ during the period of "side by side" comparison before the new control algorithm was used.

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Plant B: A similar measurement comparison campaign was carried out in smelter B for one month to make sure the measurements were capable prior to turning on the new STARprobe™ control. The campaign was also used to determine the superheat distribution during the normal pot control conditions

Figure 6 shows the %XS AlF3 by pyro titration vs. STARprobe™. The distribution is exactly as expected indicating STARprobe™ gives a non-bias analysis against the pyro titration. There was no statistical difference between the two methods. The average %Xs AlF3 by STARprobe™ was 10.55% with a standard deviation of 1.98 compared to an average of 10.66% with standard deviation of 2.02 for the pyro titration method.

Figure 9 shows a comparison of typical daily measurements while Figure 10 summarizes the results for the month long campaign. The bath ratio ranged from a low of 1.04 to a high of 1.22. Generally, the STARprobe™ agreed well with XRD analysis. Minor differences were due to die fact that bath sampling and STARprobe™ were not exactly carried out at the "same" time. Bath sampling was done in batch mode taking about one-half hour to one hour while STARprobe™ measurements were done over a period of four hours. During this four hour period some pots experienced some operational activities that caused a bath ratio change. During the month long measurement campaign, STARprobe™ measured an average ratio of 1.1239+0.036 vs. an average ratio of 1.12435±0.031 by bath sampling/XRD method. Thus, there was virtually no difference between the two methods.

Fewer comparisons were made for alumina due to limited availability of the LECO analyzer. Figure 7 shows the %alumina results as measured by STARprobe™ and LECO analysis. The STARprobe™ trended the %A1203 well compared to the LECO but the accuracy is not as good as the ratio measurement. The typical correlation (R2) between the two alumina concentrations was 75% with a typical standard deviation of 0.5%. All R e s u l t s ( o n e m o n t h )

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and material states are likely to change. The energy change will result in a change of pot temperature while the material change will result in a change of bath chemistry (ratio). Either change individually or combined will either increase or reduce the superheat. If the perturbation is large, a new energy and material balance state is represented by a different bath chemistry and pot temperature. Pot control is aimed at maintaining a targeted energy and material balance to minimize operational impact, which means that the superheat needs be controlled and maintained to achieve the maximum efficiency.

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%XS A1F3 from a shift of sampling campaign.

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Plant C: A third large scale "side by side" capability study was carried out in smelter C before STARprobe™ was deployed. Figure 12 shows the results from a group of measurement (66 pots). The acidity results by STARprobe™ and sampling/XRD method matched well except for two pots as indicated in Figure 12. After further examination, it was found that bath samples were switched during preparation for XRD analysis. After the correction, the two results all matched well as shown in Figure 13.

%XS A1F3 after correction of the mixed bath samples.

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The results of the overall comparison are shown in Figure 14. Again, with the smelter's typical operating range from a low of 5% to a high of 13.5% XS A1F3, the STARprobe1 agreed well with the bath sampling/XRD analysis method. Alumina concentration in bath was compared with XRF and shown in Figure 15. Again, the STARprobe™ agreed with sampling XRF analysis method fairly well.

13 17 21 25 29 33 37 41 45 49 53 57 61 65 • STARprobe

Figure 15:

Pot Behavior Profiling During an operational disturbance such as tap, anode change, cover application, feeding change or instability, the pot energy

%A1203 by STARprobe™ and sampling XRF method.

Figure 16 shows profiles of temperature, liquidus, and superheat over a short period of 5 hours during which there were no major operational activities. During this period of normal operation the

494

pot temperature varied from a low of 938°C to a high of 956°C. The liquidus generally followed the temperature and varied from a low of 930°C to a high of 942°C. The superheat varied from a low of 6°C to a high 13°C, a normal operating range - not high enough to melt excessive ledge while not low enough to cause any operational problems.

can be achieved). In other words, by actively taking the correct measures, a correction can be made to bring the pot back into the optimal multi-dimensional performance condition. This is achieved by factoring in STARprobe™ information along with all the information available about the condition of the pot when it was sampled and its most recent historical performance.

Bath Temperature, C

Figure 19: Bath superheat vs. 1) bath ratio, and 2) pot temperature. Figure 16: Variation of bath temperature and liquidus.

STARprobe™ Based Active Pot Control - Performance An example is given below for one of Alcoa smelter locations where STARprobe™ based control is deployed.

Figures 17 and 18, respectively, show variations of temperature and bath ratio, and temperature and superheat in a pot for a 42 hour period. During this period, the pot underwent chemical additions, metal tapping, anode setting, and other operations. The pot temperature varied from a low of 935°C to a high of 958°C while the bath ratio varied from a low of 1.03 to a high of 1.11. Superheat changed from a low of 5°C to a high of 20°C. 1.14

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Figure 20 presents voltage behavior for a period of 5 months in a 66-pot scale trial at this smelter. 33 pots were controlled using the new control algorithms while the other 33 pots were used as standard control group. STARprobe™ measurements were applied to all pots. The average pot voltage of the test group was reduced by at least 15 mV while the voltage of the control group increased by almost 10 mV. Paired t-test of the periods before and during test has a p-value much lower than 5%, which means that the pot voltage reduction is statistically significant. This improvement was consistent in all phases of the trial. One important observation was that the reduction on voltage did not create pot instability.

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Figure 20: Voltage reduction in a 5-month period trial. Figure 21 presents the overall pot voltage trend before and after the implementation of the control algorithm using STARprobe™. The pot voltage first experienced a gradual reduction and then stabilized after a few months of deployment. During this period, a voltage reduction of about 50 mV was achieved. It should be mentioned that the settings in the control algorithm were still being fine-tuned after the deployment while at the same time pot operation was being adjusted to maintain thermal balance.

Figure 18: Temperature and superheat variation within 42 hours. Figure 19 shows another pot survey across a smelter including pot temperature, bath ratio, and superheat in a specific time period of operation. Optimal operating ranges for both bath ratio and bath temperature are also shown. As can be seen, a considerable amount of cases (pots) are out of the targeted ranges, indicating that there is a high potential for improvements if we are able to move both bath ratios, temperatures, as well as bath superheat to optimal operating zones (tighter ratio and temperature operating range with optimal superheat so that maximum pot performance

The new control algorithm also minimizes over-control in bath chemistry (undershoot or overshoot of chemical (A1F3) additions). The end result is that the overall fluoride (AIF3) consumption is

495

those pots identified with excess high superheat (>15°C), and 2) Automatic parameter adjustment based on the integration of the STARprobe™ measurement with the pot status information for all those pots that were operated under sub-optimal conditions. After deployment the current efficiency was maintained at an average of 94.6% for over a 6 month period during which at times there was a significant raw material degradation. The increased current efficiency was primarily obtained with reduced pot voltage and increased line amperage during this time.

reduced. This can be seen in Figure 22 that shows the A1F3 addition change before and after the implementation of STARprobe™ based control. The estimated net reduction on the AIF3 consumption is estimated about 5%. i i ! j1

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Figure 22:

A1F3 consumption before and after the active control.

Alcoa STARprobe™ has been successfully deployed in Alcoa's plants to replace the traditional bath sampling/analysis methods and pot temperature measurement practice. The detail comparisons of measurement results on plant scale for several smelters further verified that STARprobe™ measures bath ratio (acidity) as accurately (if not better) as the analytical methods. A constant superheat based control algorithm was developed and deployed for pot control to take advantage of the real time STARprobe™ measurement results. Higher pot performance, meaning a reduced pot power consumption, a reduced chemical consumption, and an increased current efficiency, has been achieved by using STARprobe™ and the integrated STARprobe™ based pot control.

Probably the most important pot performance indicator is current efficiency. Figure 23 shows the current efficiency (plant monthly average) prior to and after deployment of both STARprobe™ and the STARprobe™ based control algorithm. The detail values are also listed in Table I.

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Acknowledgement Authors wish to thank Alcoa Global Primary Products for allowing publication of Alcoa STARprobe™ and its related applications in active pot control. We also wish to thank those contributors (technicians, operators, engineers and management across Alcoa R&D and smelter operations - too many to be mentioned here) for their dedication, support, and efforts during the technology deployment.

a

References

—> STARprobe deployed

1/1

4/11 7/20 10/28 2/5

5/15 8/23 12/1 3/11 6/19 9/27

1/5

X. Wang et al.,"Alcoa STARprobe™", TMS Light Metals, to be published in 2011. P. Entner, "Control of A1F3 Concentration", TMS Light Metals, 369, 1992. M. Wilson, "Practical Consideration Used in the Development of a method for A1F3 Addition", TMS Light Metals, 375, 1992. D. J. Madsen, "Temperature Measurement and Control in Reduction Cells", TMS Light Metals, 453,1992. P. Desclaux, "A1F3 Additions Based on Bath Temperature Measurement", TMS Light Metals, 309,1987. P. Entner, "Control of Bath Temperature", TMS Light Metals, 229, 1995. W. Haupin, "The Liquidus Enigma", TMS Light Metals, 477, 1992. J. Thonstad and S. Rolseth, "Equilibrium Between Bath and Side Ledge", TMS Light Metals, 415, 1983. R. D. Peterson and A. T. Tabereaux, "Liquidus Curves for Cryolite - A1F3 - CaF 2 - A1 2 0 3 System in Aluminum Cell Electrolysis", TMS Light Metals, 383, 1987.

4/15 7/24

Figure 23: Average current efficiency prior to and after STARprobe™ deployment and control. Table I: Pot Performance Comparison

Baseline (13 months) STARprobe™ only and partial control (12 months) STARprobe based control (6 months)

CE

Amperage (A) 240185 240414

Pot Volts (V) 4.551 4.586

94.2 94.3

240887

4.575

94.6

(%)

The current efficiency (13 month average) under previous pot control prior to STARprobe™ deployment was 94.1%. The current efficiency was steadily improved after STARprobe™ deployment with an average over 94.5% (12 month period). The increase was mainly due to: 1) Daily manual interventions on

496

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

TECHNOLOGY & EQUIPMENT FOR STARTING UP & SHUTTING DOWN ALUMINIUM POTS UNDER FULL AMPERAGE Yang Tao, Cao Bin, Li Meng, Yi Xiaobing (Guiyang Aluminum Magnesium Design and Research institute, Guiyang 550083, China ) Key words: status breakup; voltage grade; idle power consumption; aluminum production decrease; emission decrease of equivalent weight carbon dioxide

Abstract This paper presents the basic principle, technical solution and key equipment of the technology of pot startup/stoppage under full amperage, based on the design and production properties of aluminum smelting potline. In addition, it provides for a benefits analysis brought by this technology on energy saving, production increase and emission decrease. Also presented is the current application status of this technology.

produced, which may cause personal harm and damage to equipment. It's different from AC that a DC arc has high energy and is difficult to be extinguished.

Theoretical Basis

Based on electric contact theory111, the following basic concepts can be obtained:

Therefore, it's a must to close and open the short-circuit mouth, in order to startup and/or stop a pot under full potline current. However, it is the key problem for theoretic study that how to reduce the DC arcing energy between the short-circuit piece and busbar riser during short circuit mouth operation.

The modern large aluminum electrolysis potline is basically composed of a few hundreds of pots connected in series, this kind of series potline is powered by high amperage DC power source, potline voltage is above 1000V, but pot voltage is normally around 4.2V. In order to carry out overhaul for certain single pot without interrupting other pots' normal production, parallel bypass—short circuit busbar at pot bottom is designed for each pot, each busbar at pot bottom is connected to the short-circuit piece via flexibles, shown as the following circuit schematic diagram, See fiure 1.

The less the electric potential and its variation gradient of the short circuit mouth contact surface before and after opening and closing is, the less the total arcing energy is. The less the current and its variation gradient of the short circuit mouth contact surface before and after opening and closing is, the less the total arcing energy is. The shorter the opening or closing duration of the short circuit mouth contact surface is, the less the total arcing energy is.

When pot is under normal production status, a electric insulation plate is inserted between short-circuit piece and busbar riser, and no current goes through, but current only goes through pot for electrochemical reaction; However, when pot is stopped, the short-circuit piece and busbar riser are pressed together via bolt for current going through, then the potline current goes via the busbar at pot bottom to the downstream pot. The gap between the short-circuit piece and busbar riser is called short-circuit mouth.

Based on the theoretic analysis conclusion, the following technical targets can be determined: To reduce the electric potential of the short circuit contact surface before and after opening and closing121. To reduce the current variation gradient of the short circuit contact surface before and after opening and closing. To open and close the short circuit mouth rapidly, non-manually, and with remote control.

If the short-circuit mouth is directly opened and/or closed under high amperage of DC, a DC arc with high energy will be

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Technical Solution the electric potential at the short circuit mouth, then open the short circuit piece by mechanical driving, and finally shut off the current diversion switch to connect the pot to potline.

The following technical solution can be determined as figure 3, based on the above mentioned technical targets: To fit one parallel bypass between the busbar riser and short circuit piece, to reduce the electric potential and current variation gradient of the short circuit mouth contact surface before and after opening and closing, so as to reduce the DC arc energy produced when opening and closing the short circuit mouth.

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To design mechanical device to quickly and simultaneously open/close the short circuit mouth, to realize the quick opening and closing of short circuit mouth, reduce the duration, and reduce the total DC arcing energy; to realize the opening and closing of short circuit mouth non-manually, with remote control, to increase security; in addition, to realize the gang control with current diversion switch.

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Procedure Analysis Procedure of Pot Stoppage

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Pot stoppage is to short out the on-line running pots for overhaul, and the short circuit mouth is closed, instead of opened.

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Schematic Diagram of Technical Solution Procedure Breakup

To design a control cabinet to manipulate the action of switch group and device group remotely.

The short circuit mouth of on-line running pot is under switch is closed, close and tightly press the short circuit piece with the mechanical open/close device, and then fix it by screwing down the bolt; finally, disconnect the current diversion switch, and the pot stoppage procedure is finished. The procedure of pot stoppage is divided into 6 current through-flow statuses, and these statuses along various paths are shown as follows:

This kind of device is directly installed at the short circuit mouth when starting and stopping a pot. For stopping a pot, first, close the current diversion switch to reduce the electric potential at the short circuit mouth, and then quickly close the short circuit piece by mechanical driving to short out the pot from potline; For starting a pot, first, close the current diversion switch to reduce

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various paths under various statuses during pot stoppage are shown in following table 1.

Status Analysis Take one 240 kA pot for example, the current distribution along

Table 1 Status Status 1 Status 1_1 Status 1_2 Status 1_3 Status 2 Status 2_1 Status 2_2 Status 2_3 Status 2_4 Status 2_5 Status 2_6 Status 2_7 Status 3

Pot 240000 64889 37781 26649 20504 17651 17027 15880 14463 13729 12687 11613 11276

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Current Distribution List during Pot Stoppage (unit: A)

Busbar at Pot Bottom Switch Group Short Circuit Piece Group 0 0 175111 0 0 202219 213351 0 219496 0 33393 188956 40696 182276 54121 169999 154824 70713 146965 79306 135813 91500 104072 124315 120706 108018

Voltage (mV)

Operation

4200 1136 661 466 359 309 298 278 253 240 222 203 197

Normal running Close switch 1 Close switch 2 Close switch 3 Close switch 4 Close short circuit piece 1 Close short circuit piece 2 Close short circuit piece 3 Close short circuit piece 4 Close short circuit piece 5 Close short circuit piece 6 Close short circuit piece 7 Close short circuit piece 8 Screw down the bolt at short circuit mouth Disconnect switch 1 Disconnect switch 2 Disconnect switch 3 Disconnect switch 4 Take anodes out

Status 4

9963

106650

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Status 4_1 Status 4_2 Status 4_3 Status 5 1 Status 6

11220 12809 14922 17931 0

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138954 158634 184808 222069 240000

196 224 261 314 339

The short circuit mouth is pressed in two closed statuses by two ways, one is by mechanical open/close device and another is by bolt, the former is just the transition status during work process, and the latter is the target status of pot short circuit, the pressing

Duration

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force of the former is less than that of the latter, so, there is more current passing the short circuit mouth when the bolt is screwed down. Short-circuit pieces in status 3 are closed by mechanical open/close device, and in status 4 they are closed status by bolts.

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It's indicated by table 1 and figure 5: normal pot voltage 4200mv (status 1), pot current 100%; when current diversion switch group is all connected (status 2), voltage reduces to 359mv, i.e. reduces 12 times, and 91.5% current is bypassed by switch group; subsequently, when short circuit piece group is closed in turn (status 3), the current bypassed by short circuit piece group increases from 0% to 45.0%, and the current bypassed by switch group reduces to 50.3%; when switch group is disconnected in turn, the current bypassed by switch group reduces gradually to 0%, however, the current bypassed by short circuit piece group again increases gradually to 92.5%; when anodes are taken out, the current at short circuit group increases to 100%, and current diversion from pot to short circuit mouth group is finished.

Procedure of Pot Startup Pot startup is to power on the overhauled pot and connect to the potline, and the short circuit mouth is disconnected, instead of closed. Procedure Breakup The short circuit mouth of the pot which is to be powered on for baking is under closed status, current flows via the short circuited busbar at pot bottom to downstream pots, and pot voltage is within 0.2v~0.3V; during pot startup, first install shunts between the side B big busbar of the target pot and the side A busbar riser of the downstream pot; close the current diversion switch; press the short circuit mouth tightly with the mechanical open/close device; loosen the bolt; close the current diversion switch; open the short circuit mouth with the mechanical open/close device, and insert insulation plate; finally, disconnect the current diversion switch, and pot startup procedure is finished. Similarly, the procedure of pot startup is divided into 6 current through-flow statuses, and they are shown as follows:

From status 2 to status 5, i.e. switch group status changes from closed to disconnected, voltage only changes within lower and small range of 197~309mv, therefore, the arcing energy at short circuit mouth is effectively controlled, and the operation security at short circuit mouth is also ensured.

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Current Through-flow Status during Pot Startup various statuses during pot startup are shown in following table:

Status Analysis The current distribution percentages along various paths under

Current Distribution List during Pot Startup (unit: A)

Table 2

Busbar at Pot Bottom Switch Group Short Circuit Piece Group 227161 0 188319 41037 161214 69674 140929 91105 124943 107995

Status

Pot

Status 1 Status 1_1 Status 1_2 Status 1_3 Status 2

6420 5322 4556 3983 3531

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4003

122431

4116 4517 4714 5219 5451 6099 6317 7365 9648 13879 24716 120000 240000

125894 138168 144177 159622 166714 186544 193194 225269 220703 212242 190568 0 0

Status 3_1 Status 3_2 Status 3_3 Status 3_4 Status 3_5 Status 3_6 Status 3_7 Status 4 Status 4_1 Status 4_2 Status 4_3 Status 5 1 Status 6

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Voltage (mV)

Operation

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321 266 228 199 177

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200

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206 226 236 261 273 305 316 368 482 694 1236 6000 4200

Install shunts Close switch 1 Close switch 2 Close switch 3 Close switch 4 Loosen the bolt at short circuit mouth Open short circuit piece 1 Open short circuit piece 2 Open short circuit piece 3 Open short circuit piece 4 Open short circuit piece 5 Open short circuit piece 6 Open short circuit piece 7 Open short circuit piece 8 Disconnect switch 1 Disconnect switch 2 Disconnect switch 3 Disconnect switch 4 Dismantle shunts

Short-circuit pieces in status 2 are closed by bolt, and in status 3 they are closed by the mechanical open/close device.

Duration

1

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Current Change Curves of All Paths during Pot Startup

It's indicated by table 2 and figure 9: voltage of stopped pot is 305mv (status 1), 89.8% current is bypassed by short circuit piece group; when current diversion switch is all connected (status 2), voltage reduces to 172mv, 43.7% current is bypassed by switch group; subsequently, loosen the bolt at short circuit mouth, and open short circuit piece group (status 4), the current bypassed by short circuit mouth group reduces from 50.6% to 0%, and the current bypassed by switch group increases to 88.4%; finally, when switch group is disconnected in turn, the current bypassed by switch group reduces to 0%, and the current bypassed by pot increases to 50.0%, voltage increases to 6000mv (impact voltage); after running for certain time, as shunt temperature keeps increasing and the current bypassed by it keeps reducing, dismantle the shunts, pot is powered on 100%, and current diversion from short circuit piece to pot is finished.

References 1. Cheng Lichun, "Electric Contact Theory and Application" [M], "Electric Engineering Theory", Oct, 1985, Mechanical Industry Press; 2. Zhou Wei, Zhang Chunyan, He Junjia, "Influence on Relay DC Arc by Electric and Mechanical Parameters", [J], 2006, PI9 of China Electrical Engineering Periodical; 3. Fu Wanan, Song Baoyun, "Study on High-Voltage Circuit Breaker Permanent-magnet Operation Device", [R], 2000, P8 of China Electrical Engineering Periodical; 4. Qiu Zhuxian, "Prebaked Pot Smelting" [M], Aluminum Smelting and Mechanical Industry Press, Oct, 1985: 247-249;

From status 3 to status 4, i.e. the short circuit piece group status changes from closed to opened, voltage only changes within lower and small range of 194~347mv, therefore, the arcing energy at short circuit mouth is effectively controlled, and the operation security at short circuit mouth is also ensured.

5. Feng Naixiang, "Aluminum Reduction" [M], Aluminum Smelting and Chemical Industry Press, Jul, 2006: 83-86.

Conclusions By adding current switch bypass, the electric potential, current and their change gradients at short circuit mouth before and after the short circuit mouth is opened and closed can be reduced, so as to reduce effectively the potential energy at the short circuit mouth. The short circuit mouth mechanical open/close device can open/close the short circuit mouth quickly, simultaneously, non-manually and with remote control, shorten the arcing duration of the short circuit mouth during open and close period, and also increase security greatly. Pot startup/stoppage technology with full potline current amperage can bring for enterprise social and economic benefits on energy saving, production increase and emission decrease.

502

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Study on solution of A1203 in low temperature aluminum electrolyte Hongmin Kan1, Ning Zhang1, Xiaoyang Wang1 !

Key Laboratory of Advanced Materials Technology of Liaoning province, Shenyang University, Shenyang 110044, China Keywords: Low temperature aluminum electrolyte, Alumina, Image analysis expression was given. Welch [10] et al. studied crust and alumina powder dissolution in aluminum smelting electrolytes. Yushu [7] measured the solubility of A1203 in basic cryolite melts experimentally for 3
Abstract Current efficiency can be increased and energy consumption can be lowered by low temperature aluminum electrolysis. However, many problems will occur, such as low electrical conductivity, cathode shell, low alumina solubility and alumina-solution rates if the temperature is too low. Of these problems, low alumina solubility and alumina-solution rate are difficult problems. In this paper, a novel method that can measure the solubility and dissolution rate of A1203 is introduced based on early research. The double rooms' transparent quartz electrobath is used for low temperature aluminum electrolysis. The image sequence is obtained by taking from the double rooms' transparent quartz electrobath. Image analysis techniques are used to compute the change of solubility and dissolution rate of A1203. The method can analyze the influence factors of solubility and dissolution rate of A1203 intuitively and roundly by transparent quartz electrobath and image analysis techniques.

In this paper, a novel method that can measure the solubility and dissolution rate of A1203 is introduced based on early research. The double rooms' transparent quartz electrobath is used for low temperature aluminum electrolysis. The image sequence is obtained by taking from the double rooms' transparent quartz electrobath. Image analysis techniques are used to compute the change of solubility and dissolution rate of A1203. The method can analyze the influence factors of solubility and dissolution rate of A1203 intuitively and roundly by transparent quartz electrobath and image analysis techniques.

Introduction Low temperature aluminium electrolysis has been one of active research fields in recent years I15]. Traditional Hall- Heroult electrolysis process for aluminium production usually operates at very high temperature (nowadays about 950 °C) and unavoidably shows high energy consumption, complicated operation and emissions. By the introduction of low-melting baths one might expect an increase in current efficiency and lower energy consumption. It is important to reduce liquidus temperature for study on low temperature aluminium electrolysis. But many problems will occur, such as low electrical conductivity, cathode shell, and low alumina solubility and alumina-solution rates when the temperature is too low. These problems have seriously affected the application of low-temperature aluminum electrolysis in industry. In many problems, the single largest problem is the solubility of alumina.

Computing dissolution rate of A1203 based on image sequence analysis Factors and constraints Here the double rooms' transparent quartz electrobath is used for a container. Alumina powder is added into the transparent quartz electrobath. Factors of dissolution rate of A1203 are considered. Then constraints which the following algorithm will use are set. These factors are temperature of the transparent quartz electrobath, quality of A1203, rate in which A1203 is added, stirring, additives used, composition ratio of additives. Although capacity of the transparent quartz electrobath does not affect shape of dissolution rate varying curve of A1203, it does affect the critical value of dissolution rate varying curve of A1203. Suppose that Vtranseiectmiyzer denotes capacity of the transparent quartz electrobath, Ttranselectroly7jer denotes temperature of the transparent quartz electrobath, Qaiumim denotes quality of A1203, Vaiumina denotes rate of which A1203 is added, Tstir denotes times of stirring, Vstir denotes speed of stirring, Vadditive denotes additives used, Padditive denotes composition ratio of additives, then the factor set affecting dissolution rate of A1203 5rfliJ0/Maii0nraie can be denoted as

An important indicator one determines the performance of alumina in electrolytic aluminum production is the solubility of alumina. When the temperature of aluminium electrolysis is lowered, low alumina solubility and alumina-solution rates cause alumina can not be completely dissolved or quickly dissolved in molten cryolite. Alumina will deposite on bottom, thus the physical and chemical processes and heat balance of the electrolysis cell are severely affected. Many previous measurements of alumina solubility have been published and significant progress has been made t610]. The solubility of alumina in molten Na3AlF6 containing various amounts of A1F3, CaF2, and LiF was determined by measuring the weight loss of a rotating sintered corundum disc. The empirical

(i)

?={Vt transelectrolyzer» ·* transelectrolyzen i&alumina> "alumina'-

503

Here image sequence of A1203 dissolution is taken from the transparent quartz electrobath at a fixed time interval. Then information region is computed in each image. It can be found in Fig.2 that A1203 dissolution image contain three regions: information region (A1203 dissolution region, Fig.2(a)), interference region (Transparent quartz electrobath region, Fig.2(b)) and background region (Black region, Fig.2(c)). As far as Computation of A1203 dissolution rate is concerned, the information region is most important. Here rectangle window, such as A and B, is used to compute the information region.

* stin *stin *additive* * additive)

In the paper, the method that computes the dissolution rate of A1203 is proposed based on image sequence analysis. Each element is set in the factor set Sdissoiutionrate. Image analysis and understanding is used for image sequence of A1203 which is taken from the transparent quartz electrobath. The dissolution rate varying curve of A1203 is developed. Then the dissolution rate varying equation of A1203 is devloped. Here the constraints for Sdissoiutionrate are set as followed. Suppose ' transelectrolyzef~ **> * transelectrolyzer~®i \oalumina=(~"> * alumina~Qi * stir^&i

*stir^f

* additive~§>>

Padditive=h then dissolution rate of A1203 DVaiumina can be denoted as: DVnl

= U V (\ Vtranselectrolyzer~&)\*

(2)

transelectrolyzer~V)'\

Qalumim=c)iVaiumina=ay(Tstir=e)iVstir=f)' ( Vadditive=g)' (P

additive-^))

Here, DV denotes dissolution rate function, operator "·" denotes dependence relation among constraints.

Fig.2 A1203 dissolution image (a)information region (b)interference region (c)background region

It should be noticed that the factor set Sdissoiutionrate can be extended dynamically and parameter value can be changed dynamically. This makes it possible that constraints can be changed by application.

When A1203 dissolution images are taken, the position of camera and distance from camera to the transparent quartz electrobath is relatively fixed. This means that the position and scale of a rectangle window can be decided in advance. Here the point at the upper left corner is selected as a reference point. The scale of rectangle window A is Al x A2 and the position is (PI, P2). The scale of rectangle window B is Bl x B2 and the position is (P3, P4).

Algorithm computing dissolution rate of AMD? Here the image sequence of A1203 dissolution is taken from the transparent quartz electrobath. It can be found that each image contains a large amount of texture information. Our idea is as follows: First, the image sequence of A1203 dissolution is taken at a specific time interval. The image sequence is regarded as a point set where each point corresponds to an image at a specific time. Texture features of each image are used to characterize a point. Then, the similarity between neighboring points is used to describe dissolution rate variation of A1203, so that the dissolution rate varying curve of A1203 can be drawn. Then, approximation and interpolation techniques are used to establish the dissolution rate varying equation of A1203. Finally, dissolution rate of A1203 can be computed by the equation. The flowchart by which dissolution rate of A1203 can be computed based on image sequence analysis can refer to Fig.l. The error of the method is caused mainly by sampling time of one time and time interval of sampling.

Suppose thatf(x, y) denotes an A1203 dissolution image of size M x N. The equation (3) can be used to decide whether a point (x, y) is in information region. PI <= x <= PI + Al and P2 <= y <= P2+A2

Here x = PI and x = Pl+Al are left boundary and right boundary of the information region respectively, y = P2 and y = P2+A2 are upper boundary and bottom boundary respectively. Extracting texture features from information region of Al203 dissolution image It can be found in Fig.2(a) that the information region of the A1203 dissolution image contain a large amount of texture. Here texture features are used to characterize an image, and the cooccurrence matrix method is used to compute the texture features.

Sample at a Transparent

fixed time

quartz

Image

sequence

represent

electrobath Similarity computation between neighboring points

Dissolution

rate

varying equation ofAl 2 0 3

Point set and Characterization 1

of A1 2 0 3 solution

Computation

(3)

First, the image sequence is taken through the transparent quartz electrobath. Then, color images are transformed into gray images by gray-scale transformation. Suppose that / (JC, y) denotes gray image of A1203 dissolution. If Ng denotes gray scale of the image, then co-occurrence matrix can be denoted as:

Dissolution rate of A1 2 0 3

Fig. 1 Flowchart for computing the dissolution rate of A1203 based on image sequence analysis

P (i,j, d,6) = count {((xh yx), (x2, y2)) I (xu y{) e (L r xL c ),(x 2 ,y 2 )e (L r xL c ),

Computing information region in Al203 dissolution image

504

(4)

(*2» yi) - (χι> yi) + (4 cosφ, d

Texture feature vector

sm0),f(xu yi) = i,f(x2, y2) =;, 0 < i,y < Ng}

Here count (') is used to denote number of pixel pairs which is to satisfy the condition in the set. The parameter d denotes the distance between pixels. The parameter È denotes the angle between line denoted by pixel pair and horizontal direction.

Fig.3 Image sequence of A1203 dissolution

Five parameters, which are angular second moment, contrast, correlation, inverse difference moment, entropy, are computed according to the co-occurrence matrix P (i, j , d, È). The five feature can be denote as: Ng-lNg-l

Angular Second Moment £ £ p 2 ( / , y )

(5)

i=0 7=0 ",-1

Ng-Wg-l

Contrast = £ n 2 ( £ £ P(i, j)), n=0

J=0

1

I i - j 1= n

y'=0

J~

J~

Correlation = —— £ £ ijP(i, j) - ìχìγ °x°y

i=0 ;=0

(6)

Our approach is to use the similarity between texture feature vectors of two image points to describe dissolution rate variation of A1203. If t' is sufficiently small, then the dissolution rate of A1203 can be denoted as quotient between similarity and t\

(7)

1

PQJ) Inverse Difference Moment = £ £ ^—— ,=0 7=0 1 + 0 ' - j)

(8)

Suppose that Euclidean distance is used as similarity measure, then the similarity between texture feature vectors of two image points can be denoted as:

Ng-lNg-l

Entropy = - £ £ P(i, j) log(P(i, j))

(9)

j=0 7=0

Similarity(Tn,Tl2) = ( £ [ ! ; , (s)-7; 2 (i)] 2 )*

Here the parameters μχ, μγ, σχ, ay denote mean and standard deviation of probability density function Px and Py, which can be denoted as N,

Px (0 = £ P& J)

i = 0,1, · · ·, Ng -1

Here f which is sampling time of one time denotes time interval between points in the ordered point set. If f is sufficiently small, then it can be assumed that the dissolution rate of A1203 can be changed at a fixed rate at t' time interval. Time interval between two samples is called as time interval of sampling. If time interval of sampling is too small then this will result in a large number of calculations. If time interval of sampling is too large then there is a risk that the rate of alumina dissolution changes within the selected interval of sampling.

(12)

Here Tn and Tt2 denote the texture feature vectors of two image points at t' time interval in the ordered point set respectively. The dissolution rate of A1203 DV^fne can be denoted as:

(10)

DV alumina

7=0

_ Similarity(TtVTt2)

(13)

Ng-l

Suppose that the image point set of A1203 dissolution consists of num elements. The k resampling is carried out for image sequence of A1203 dissolution in Fig.3. The information can be regarded as a sample in the dashed box. Then equations (12) and (13) are used to compute dissolution rate of A1203 so that the dissolution rate varying point set of A1203 can be developed. Then a polynomial approximation is used for the point set to compute dissolution rate varying equation of A1203. Finally, the dissolution rate of any time can be computed by initial value of dissolution rate and dissolution rate varying equation of A1203.

1=0

Here the d is set as 1, 3 and 5 respectively in equation (4). The values for Θ is set as 0°, 45°, 90° and 135°, respectively. Then equation (4) is used to compute 12 co-occurrence matrices. Then the five features are computed from 12 co-occurrence matrices, and texture feature vector consists of these features. Computation of dissolution rate equation ofAl203 Here the dissolution rate equation of A1203 will be computed. After the image sequence of A1203 dissolution is obtained from the transparent quartz electrobath using a fixed time interval, the images can be regarded as an ordered point set. Furthermore, texture feature vectors can be used to characterize a point. Refer to Fig.3.

Experiments and analyses In this paper, dissolution rate of alumina is studied. The electrolyte system is Na3AlF6-AlF3-Al203-CaF2-LiF-MgF2. The liquidus temperature is about 900 °C. To compute the varying curves of dissolution rate about alumina in cryolite molten salt, a prototype system is designed by the above approach, and compiled by Java.

505

temperature, agitation and other factors on the dissolution of alumina by the novel method.

The configuration of computer for test is Intel(R) Celeron(R) mainboard, 1.60GHZ CPU main frequency, and 256 MB memory. Test result for alumina is shown in Fig.4.

Acknowledgements

The value in the first textbox is sampling time of one time, and described by the number of image points in sampling of one time. The result computed by the value minus 1 multiply sampling time of image sequence is f in Fig.3. The value in the second textbox is time interval of sampling which refers to time interval between two sampling, and described by the number of image points.

The authors wound like to thank reviewers for their helpful suggests. The authors wound like to thank Dr. Gang Zhang of Shenyang University of Technology for participating in simulation of algorithm. The authors also wish to thank Professor Zhaowen Wang, Dr. Xianwei Hu and students from the molten salt electrolysis research group of Northeastern University for their help. The authors wound like to acknowledge the financial supported by "National Natural Science Foundation of China (51072121)" and "Key Laboratory Foundation of Liaoning Province (LS2010109)".

Here the dissolution of alumina is sampled at a fixed time interval of one second. When the value is 2 in the first textbox and the value is 3 in the second textbox, the dissolution rate varying curve is shown in Fig.4 for alumina. pissotutkm rate cnage curve

References 1. Pomiakov (ΠΟΛΗΚΟÂ) et al., "Method of aluminum production: Russia", 1708933-A1 [P], 1989-06-12 2. Beck T. R., "A non-consumable metal anode for production of aluminum with low-temperature fluoride melts", Light Metals 1995, The Minerals, Metals & Materials Soc, Warrendale, USA, 1995: 355-360 3. Rolseth S., Gudbrandsen H., et al., "Low temperature aluminum electrolysis in a high density electrolyte", Aluminum, 2005, 81(5): 448-450 4. Qiu Zhuxian, Gao Bingliang, "Aluminum electrolysis in a 200A bench-scale cell at 845 °C with 96%~98% current efficiency", Aluminum, 2001,77(12): 974-976 5. Lu H., Fang K., et al., "A new electrolytic aluminum production process", Aluminum, 1999,75(12): 1113-1118 6. Gerlach J., Hennig U., Kern K., "The dissolution of aluminum oxide in cryolite melts", Metallurgical transactions B, 1975, 6B(3): 83-86 7. Yunshu Zhang, Xiaoxia Wu, Robert A. R., "Solubility of alumina in cryolite melts: measurements and modeling at 1300K", Metallurgical and materials transactions B, 2003, 34B(4): 235-242 8. Solheim A., Rolseth S., Skybakmoen E. et al., "Liquidus temperature and alumina solubility in the system Na3AlF6AlF3-LiF-CaF2-MgF2", Light Metals 1995, The Minerals, Metals & Materials Soc, Warrendale, USA, 1995: 451-456 9. Egil Skybakmoen, Asbjorn Solheim, Asmund Sterten, "Alumina solubility in molten salt systems of interest for aluminum electrolysis and related phase diagram data", Metallurgical and materials transactions B, 1997, 28B(2): 8186 10. Welch B. J., Gerda I. K., "Crust and alumina powder dissolution in aluminum smelting electrolytes", JOM, 2007, 59(5): 50-54 11. Yang Zhenhai, Gao Bingliang, Xu Ning, Qiu Zhuxain, Liu Yaokuan, "Dissolution of Alumina in Molten Cryolite: A Video Recording Study", Journal of Northeastern University (Natural Science), 1999, 20(4): 398-400 12. XU Junli , SHI Zhongning , GAO Bingliang, QIU Zhuxian, "Dissolution of Alumina in Molten Cryolite", Journal of Northeastern University (Natural Science), 2003, 24(9): 832834 13. QIU Zhuxian, WANG Zhaowen, GAO Bingliang, YU Xuguang. Physical and chemical processes of low temperature aluminum electrolysis [J]. Mining Research and Development, 2003, 8: 9-12.

^

jtoput sampling t»ne of one tiff»

}2

jlnptir taue interval for sampling

[3

Dissolution Rate of Alumina

^

1 15 L

j — 0j[

,

,

0

5

10

^

„ 15

20

^ ,

v

25

30

j

Intervals) Computing regression equation of dtssototion rate |y=-0 0017Qc(3?) + O.G984Qs(2)? -1.2661 (x(1)) * 15.3509 scatter point graph

ji

Computing equation

' i

Saw« to Hie

i 1 drawing curve

Fig. 4 Test results for alumina The equation about the dissolution rate of alumina is y2 = -0.0017X3+0.0984JC2-1.2661JC+15.3509

(15)

Here x denotes sample, Vi is the dissolution rate of alumina at sample JC which is computed by the method in the paper. Conclusions A novel method which can measure the solubility and dissolution rate of A1203 is introduced. The method can analyze the influence factors of solubility and dissolution rate of A1203 intuitively and roundly by transparent quartz electrobath and image analysis techniques. Image analysis techniques are used to compute the change of dissolution rate of A1203. The dissolution rate equation of alumina is established base on this method: y2 = -0.0017x3+0.0984x2-1.2661x+15.3509 We will further study on the solution of A1203 in low temperature aluminum electrolyte and discover impact of the additives,

506

14. FAN liman, QIU Zhuxian, Grjotheim, "A direct observation of the process of aluminium in cryolite-alumina electrolyte through a transparent quartz cell", Journal of Northeastern University, 1986, 46(1): 97-106 15. XU Junli , SHI Zhongning , GAO Bingliang , QIU Zhuxian, "Bubble behavior on metal anode of aluminum electrolysis", The Chinese Journal of Nonferrous Metals, 2004, 14(2): 298301

507

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

APPLICATIONS OF NEW STRUCTURE REDUCTION CELL TECHNOLOGY IN CHALCO'S SMELTERS Liu Fengqin1 Gu Songqing2 Wang Jiangmin3 Yang Hongjie1 1. Zhengzhou Research Institute of Chalco Zhengzhou Henan Province China 450041; 2. China Aluminum Corporation Beijing China 100082; 3. Lanzhou Branch of Chalco, Lanzhou Gansu Province China 730060 Keywords: Chalco, aluminum reduction, new structure cell, wetted cathode, ACD, low cell voltage, heat preservation lining, DC consumption, energy saving Northeastern University in China developed the technology to change cross section shape of the cathodes for low cell voltage in recent years. This technology is based on blocking walls in the molten metal in the cells by changing cross section shape of the cathodes to retard the metal flow so that ACD can be reduced. But the cracking and breaking are possibly occurred for the abnormal shape cathodes due to the impact and erosion by molten metal flow and high stress inside the cathodes.

Abstract A new generation of key energy saving technology for aluminum reduction has been implemented by Chalco. A new structure aluminum reduction cell technology has been successfully developed by Chalco and is now widely applied in several of Chalco's smelters. This new structure includes key technologies such as the new specially designed cathode installation, a heat preservation lining and a low voltage operation control system. The molten metal fluctuation that is driven by the great magnetic field is greatly restrained in the cells, based on reducing the anode to cathode distance (ACD) without any loss of current efficiency. This realizes a lower energy consumption and less GHG emissions. The industrial tests and applications of this technology have been carried out successfully in the various scale aluminum reduction cells in Chalco's smelters and about 12000 kWh of DC consumption per ton of aluminum and 3.70-3.75 V of cell voltage have been achieved.

Shenyang Aluminum and Magnesium Engineering Institute developed the technology to put the barrier blocks on the cell cathodes to block the molten metal electro-magnetic flow for lower ACD. But selection of the block materials are still very difficult for they have to undertake the serious smelting operation conditions without great corrosion and erosion and to be stable enough with suitable density and shape. Chalco has put a great importance on the new structure reduction cell technology and set up a series of R&D projects around the technology development. About 10 years have been spent on laboratory tests, 4kA cell simulation experiments, 160kA new structure cell tests and on large scale industrial applications. A great achievement has been made by the new structure reduction cells that the DC consumption is reduced by about 1000 kWh per ton aluminum while the cell operation is kept very stable for sustainable results.

Introduction The large amperage prebaked reduction cell technology has been wide applied and optimized since 280kA prebaked cell technology was successfully developed in China in 1996. The design technology with magnetic field simulation, modern computer control systems, high graphite cathodes, new lining materials and high quality anodes are used in Chinese aluminum industry so that the DC consumption is reduced to 13-13.5 kWh per kg of aluminum.

Development of the new cell structure The theoretical considerations for energy saving in reduction cells were made during R&D of the new structure cells.

Nevertheless, a great energy saving pressure is put on Chinese smelters due to the higher and higher energy prices. In addition, there are enhanced requirements for energy saving and emissions reduction from Chinese government and environmental protection agencies. The energy cost in the most Chinese smelters occupies more than 40% of total production cost. To develop the new generation reduction technology aimed mainly on great energy saving and sustainable development becomes urgent and more important.

The energy efficiency mainly depends on DC in the aluminum reduction process, while the direct current consumption can be calculated by the following formula: DC = 1000ν/(ηχ0.3356) With DC-direct current consumption, kWh; V~cell voltage, V; η—current efficiency, %.

The study on energy consumption reduction in the smelting process has been carried out by many aluminum companies.

It can be seen from formula (1) that cell voltage reduction should be major approach to reduce energy consumption and improve energy efficiency. And the most efficient way to reduce DC is ACD reduction. The cell voltage will be reduced by more than 300mV by only 10mm reduction of ACD and DC will reduced by about 7% if no current efficiency loss happens. It is concluded

Comalco Research Centre and CSIRO, Australia make a great effort on the drained cell technology and the great progresses in such aspects as cathode TiB2 coating technology, cathode structure with slope surface and low ACD operation etc.

509

The horizontally installed cathode and side wall lining structures of the cells are changed so that the fluctuation of molten metal caused by the magnetic field is greatly reduced and the CO2 is removed quickly from anodes areas.

that ACD reduction can be the great potential and important approach for energy saving in aluminum reduction. Nevertheless, the reduction of ACD and cell voltage is restricted by the molten metal electro-magnetic flow in the common cells. About 4.5-5.5 cm of ACD has to be kept for reducing aluminum secondary reaction loss caused by fluctuation of bath—metal interface and metal diffusion to anode area. The general cell voltage distribution is shown in Figure 1.

This is the technical basis for the development and design of the new structures in the cells. In order to get the better operation results it is necessary to develop suitable operation technology to fit to the new structure cell.

External VR

Cathode VR __^ Anode VR

The cell voltage reduction should correspond to the heat loss reduction to keep heat balance in the cells. As Figure 2 shows the reasonable way is reduction of heat emission from the cells. Cathode CPV Anode Overvoltage

o

>

The most feasible approach to reduce heat losses is to use heat preservation materials for the areas of the side walls and enhance the coverage over bath, which is the heat preservation lining structure design concept of the new structure cell technology.

Anode CPV Reaction Voltage

It will become difficult to keep uniformity and homogeneousness of ACD for the different anodes and aluminum content in the different areas in the cells under low ACD and less fluidity. The technical challenge can be solved by optimizing both cell structure and operation conditions.

Fig 1 Voltage Distribution in Reduction Cell With: ACV—Voltage Between Anode and Cathode; VR—Voltage Reduction; CPV—Content Difference Polarization Voltage

The application of TiB2-C composite wetted cathodes developed by Chalco is a great benefit to the operation of the new structure cells due to high wettability, high resistance to sodium corrosion and high mechanical strength.

It can be found in Figure 1 that the voltage reduction in the bath is up to 1.5-1.6V and occupies almost 35-40% of total cell voltage under 4.5-5.5 cm of ACD.

The wettability of TiB2 composite cathode and common cathode is compared in Figure 3.

Most of this part of energy input will change into useless heat and the heat emitted. The reduction potential for the bath voltage might be 400-500mV at least according to the study results. Therefore the focus is on finding some way to reduce ACD and bath voltage reduction without any efficiency loss.

Common cathode

The current efficiency mainly depends on the aluminum secondary reaction happened in the anode areas and caused by the reaction between the aluminum dissolved in bath and C0 2 or even anodes. So reduction of aluminum content in the bath and its diffusion under a lower ACD is by us regarded the best solution for low cell voltage without efficiency loss. For the present Hall-Heroult aluminum reduction technology it is only through changing the design concept and the key technological route that the aluminum content in the bath and its diffusion can be reduced. Energy input Bus Bars

Fig.3 Wettability Comparison of Cathodes under Vacuum

Chemical Reaction

Heat Loss External Heat Loss

Cathodes Anodes Bath ACD Anode Preheat and Alumina Dissolution Fig.2 Energy Balance in Reduction Cell

510

It is concluded from Fig. 4 and Fig. 5 that the cell voltages of all the new structure cells, whether 280kA or 200kA cells, will be reduced quickly after their start up and to the target cell voltage of about 3.7-3.75 volts in two months. All the new structure cells have been operated stably and controlled constantly at the target cell voltage.

The operation technologies of new structure cells The baking and starting up methodologies for the new structure cells have been developed. The cathode structure permits to apply the simple, reliable and low cost coke resistance heating practice for the cell baking with the advantages of homogeneous baking, and therefore the benefit to longer cell life. During the development procedure lots of technical problems were solved for reducing the surface oxidation and thermal stress impacts.

It is confirmed by the tests and applications that the new structure cell technology developed by Chalco has the better applicability, operation stability.

A quick voltage reduction methodology after cell start up has been developed as well. The cell temperatures and bath chemistry should be adjusted synchronously with the cell voltage reduction so as to keep the cell operation stable and regular ledge formation on the side walls. The energy saving can be achieved during the first period after start up by the quick reduction of cell voltage.

The working cell voltages of the smelters using the new structure cell technology are usually reduced to 3.7-3.75 V without obvious efficiency loss. The electricity consumption for the new structure cells in a long term operations is accounted and compared with the average consumption of the Chinese smelters in the past years from 2006 to 2009 as shown in Fig. 6.

A new cell control system has been developed for the new structure cells that is adapted to the new requirements for low cell voltage and low ACD. One of the advantages of the new structure cells is to permit higher alumina concentrations in the bath without too much sludge formation on the cathodes, which will provides better conditions for reducing AE and PFC's emission. This is proven by the industrial application.

15000 14000 Q 13000

Main results of the industrial applications in Chalco's smelters

I* 12000

The new structure cell technology is applied in almost all the Chalco's smelters. The industrial tests were carried out in the various types of cells, such as 160kA, 200kA, 240kA, 280kA and 300kA cells. Cell Volts

|

11000

0

10000 2006

2007

»AC Consumption '

2008 Years

2009

N. S. C

»DC Consumption

Fig. 6 AC and DC Electricity Consumption Comparison N.S.C— for New structure cells in 2010 Others-Average for Chinese smelters in the past years It can be found in Fig.6 that the AC and DC energy consumption is reduced to about 13000 and 12,000 kWh per ton of aluminum respectively for the new structure cells, which is at least 1000 kWh/ton less than the common technologies. 1

4

7

1

1

1

1

2

2

2

3

3

3

4

4

4

4

Davs after Start UD Fig. 4 Cell Voltage Change of 280kA Cells after Start up Cell Volts

The input energy into the cells is reduced by about 7% and the heat balance is kept by heat emission loss reduction. The overall energy utilization efficiency is improved by 6.34%.

53

Furthermore the anode effect coefficient of new structure cells can be reduced to less than 0.05 owing to their higher alumina content in the bath without sludge on the cell bottom.

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New structure cell technology has been expended over the Chalco's smelters since its great technological and economical advantages. More than 300 cells of 160kA, 200kA, 240kA, 280kA and 300kA have been set up and put into successful operation in Chalco's Lanzhou Branch, Liancheng Branch, Qinghai Branch and Jiaozuo Wanfang Smelter etc. in 2009-2010.

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Nevertheless, the cell temperature and alumina content fluctuations easily take place due to the smaller heat capacity of the new structure cells. This results in that all the operation parameters should be more accurately controlled and the various

Days after Start up Fig.5 Cell Voltage Change of 200kA Cells after Start up

511

operations, especially the aluminum tapping and anode changing, have to be more careful and prepared. The cell structure, operation parameters and control system of the new structure cell technology has to be optimized continuously in the future applications to further reduce energy consumption, improve efficiency and aim to realize zero anode effects. Conclusion [1]

The new structure aluminum reduction cell technology has been developed and widely applied by Chalco through the study and tests for nearly 10 years.

[2]

The key technical solutions for the low cell voltage without efficiency loss are the following: a. horizontal installed wetted cathodes; b. new structure of cathodes; c. heat preservation lining; d. special baking and start up technologies; e. new control system and operation technology.

[3]

The great technical and economical achievements have been made for the new structure cell technology. The DC power consumption is reduced by more than 1000 kWh per ton to 12000 kWh/ton of aluminum, and the energy utilization efficiency is increased by about 7%. In addition, PFC's and CO2 emissions are alsogreatly reduced.

[4]

Coninuous improvement of the new structure reduction cell technology will further optimize energy consumption, improve efficiency and minimization of anode effects.

References [1]

Liu Yexiang, Li Jie, "Modern Aluminum Reduction". Metallurgy Industry PublishingHouse, Beijing, 2008 : 9092

[2]

Liu Fengqin., Gu Songqing et al., "Study on the New Reduction Technology for Energy Saving"// John A. Johnson. Light Metals 2010 : 387-390

[3]

Qiu Zhuxian, Principles and Application of Aluminum Reduction, China Mining University Publishing House, Beijing, 1998

[4]

Gu Songqing, Wu Lichun, Liu Fengqin et al. "Progresses of Nonferrous Metals in China", Central South University Publishing House, Changsha, China, 2007

[5]

Yao Shihuan, "Discussion on the Aluminum Reduction Technology Development Roadmap in China", Chinese Aluminum Industry, 2009(2): p. 2-14

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Transport Numbers in the Molten System NaF-KF-AlF3-Al203 Pavel Fellner1, Jan Hives1, Jomar Thonstad2 Slovak University of Technology in Bratislava, Radlinskeho 9, Bratislava, 812 37, Slovakia 2 Norwegian University of Science and Technology, Trondheim, 7491, Norway Keywords: Transport Number, Cryolite, Aluminium Chemicals Aluminium fluoride was purified by sublimation in a platinum crucible (inner diameter 75 mm, height 270 mm). The temperature at the bottom of the crucible was 1250 °C; the Pt lid covering the crucible was cooled in an argon atmosphere. In each run, 100 g of sublimated A1F3 was prepared. Potassium fluoride of grade "pro analysis" was dried in a vacuum furnace for 10 days. The drying procedure starts at ambient temperature. At the bottom of the vacuum furnace a dish with P 2 0 5 is placed. After 4 days the temperature is gradually increased up to 200 °C. This procedure ensures a very low moisture content in the system. Sodium fluoride and alumina were of analytical grade, and prior to use they were heated to remove any moisture.

Abstract Transport numbers in the molten system NaF - KF - A1F3 (A1203, CaF2) were investigated by the Hittorf method. It was confirmed that in molten cryolite, Na3AlF6, at 1010°C, the current is transported almost exclusively by the Na+ cations (t(Na+) = 0.99). When A1F3 is added to a Na3AlF6 melt, the transport number of sodium cations decreases to 0.74 at the composition corresponding to NaAlF4. In molten K3A1F6 the transport number of K+ cations equals 0.836 at 1005°C. In melts containing both Na+ and K+, the cations contribute to the charge transport approximately in the ratio of the squares of their ionic radii. Introduction

Apparatus and Experimental Procedure The scheme of the apparatus used to measure the transport numbers of Na+ and K+ in the molten system NaF - KF - A1F3 Al203(sat.) is shown in Figs. 1-2. The cell was made of hot-pressed boron nitride (BN). Open channels between the anode and the neutral compartments and between the cathode and the neutral compartments were made, the diameter of the hole being 2.1 mm. The anode was made of platinum, and a graphite cathode was used. The temperature was measured by means of a Ptl0Rh-Pt thermocouple.

The transport number tj of the ion j , is defined as the fraction of the electric current carried by that ion in a solution of uniform composition, i.e. without a concentration gradient. Other names for tj are "transference number" or "electrical transport number". This parameter is always positive, and it does not reflect the direction of transport. Transport numbers are called "internal" when another ion of the system or a neutral solvent is theframeof reference. "External" transport numbers have a porous plug or a wall as reference. The sum of the transport numbers of all ions equals one. A detailed discussion of this topic together with examples can be found in the paper by Ratkje et al. [1]. Principal data on the transport phenomena in cryolite melts was discussed in the monograph "Aluminium Electrolysis" [2]. Transference (transport) numbers are discussed also in the 3rd edition of "Aluminium Electrolysis" [3]. The treatment is based on results published by Frank and Foster [4], Tual and Rolin [5, 6] and Dewing [7]. Frank and Foster investigated transport phenomena in cryolite-alumina melts by means of a radioactive tracer method. It was found that t+Na = 0.99. Tual and Rolin applied the classical Hittorf method. These authors also came to the conclusion that in neutral or basic electrolytes the transference number of the sodium cation is close to unity. With increasing acidity of the bath, the transport number of Na+ decreases. This is often explained by participation of the F" ions in the conduction [2, 3]. Even in electrolytes with an excess of 7 mass % A1F3, the transference number of the sodium cation did not drop below t+Na = 0.9 [5, 6]. For Li3AlF6 melt at 1030 K, Dewing [7] found that the transport number t+Li = 0.957 + 0.08. This strongly supports the notion that electrical charge in cryolite-based melts is transported mainly by cations. Sterten et al. [8] reported transference numbers for Na+ based on emf measurements for CR = 2 - 3 (molar ratio n(NaF)/n(AlF3)) to be 0.96 - 0.99. It was assumed that fluoride anions carry the remainder of the current. In the present work we used the Hittorf method in a similar way as was done by Tual and Rolin [5, 6].

Figure 1. BN electrolysis cell. A - top view, B - side view. Outer diameter 51 mm, inner diameter of compartments 17 mm, crucible height 59 mm. Diameter of channels between compartments (dotted lines) 2.1 mm. The principle of the measurement is as follows: Electrolyte of desired composition was pre-melted in a platinum apparatus, (ca 55 g of electrolyte was prepared.) The sample was pulverized. The electrolyte (8 - 11 g) was placed into each compartment of the cell. The cell was placed in a vertical resistor-heated furnace with a controlled argon atmosphere. When the desired temperature was reached, it was kept constant ± 1 K. Electrolysis was carried out with 100 - 200 mA current for 30 60 minutes. (Duration and current should be as low as possible. However, the total charge passed through the cell must be large enough to provide measurable changes in the composition in the cathodic and anodic compartments.) Initially 60 min electrolysis

Experimental

513

time was used. However, we found that 30 min electrolysis is optimal for obtaining reliable data on transport numbers. It was found that a charge of about 550C was optimal. After the experiment, the cell was withdrawn from the furnace and cooled in a stream of argon. Samples of the electrolyte were drilled out of each compartment for analysis.

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Prior to the analysis the sample was ground in an agate mortar and homogenized. 0.2 g of the sample was transferred to a PTFE dish (Teflon) and 5 ml of concentrated H2S04 was added. The sample was heated with stirring for 30 min in a sand bath until white fumes appeared. After cooling, 5.0 ml of concentrated HC1 was carefully added. The sample was again heated in the sand bath for 30 min. After cooling, the sample was transferred to a 100 ml flask and diluted to 100 ml of solution (exactly) with distilled water. This procedure was repeated 4 times for each sample.

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Determination of Na and K Na and K were determined by flame AAS (Perkin Elmer 1100, USA) in a C2H2-air flame. In order to depress ionisation of the anolyte, addition of Cs (1 g/1) was used. Calibration was made using one-element standard solutions (Merck) with Na concentrations (0.5; 1.0; 1.5) mgl"1 and K concentrations (0.2; 0.5; 1.0; 2.0) mgl 1 .

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Determination of Al Aluminium was determined by flame AAS (Perkin Elmer 5000, USA) in a C2H2-N20 flame. For calibration, solutions having the concentrations 10.0; 20.0; 50.0 mg Γ1 were used. The conditions of the measurement are given in Table I. The accuracy of the determination and the uncertainty characterizing the analytical procedure are given in the Table Π.

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Table I. Conditions of AAS analysis.

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Results and Discussion All results are summarized in Table II.

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Table II. Transport numbers in the system NaF-KF-AlF3 A1203 CaF2 n(NaF) + CR /wt.% t(K+) /wt.% n(KF) t / ° C t(Na ) only 0.990 3.0 0 0 1018 ±0.003 NaF 0.676 only KF 920 2.0 sat. 0 ± 0.006 0.504 0.264 1.3 sat. 0 1 767 ± 0.002 ± 0.005 0.394 0.520 1 1.3 sat. 5 800 ± 0.009 ±0.031 only 0.742 1.22 795 ± 0.003 NaF only 0.816 1.5 825 ± 0.007 NaF only 0.883 2.2 1004 NaF ±0.005 0.863 3.0 - onlyKF 1005 ± 0.009

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Figure 2. Scheme of the apparatus. Analysis The main difference in comparison with previous studies consists in the analysis of the samples. While in the past radioactive tracers were used for the determination of the composition of the samples, in this work samples of the solidified electrolyte were dissolved in aqueous solutions and analysed by atomic absorption spectroscopy. Dissolution (decomposition) of cryolite-based electrolytes was carried out as follows (procedure ISO 2366:1974).

514

A1203.

t(F) 1 0.010 ±0.003 0.323 ± 0.006 0.232 ± 0.007 0.086 ± 0.040 0.258 ± 0.003 0.184 ± 0.007 0.117 ±0.005 0.137 ± 0.009 |

It can be seen that the reproducibility of the measurements is good. We always found aluminium deposited on the bottom of the graphite cathode (see Fig. 3).

Conclusions An apparatus for measurement of transference numbers in cryolite-based electrolytes was constructed and tested. The heart of the apparatus is a BN cell having three compartments connected with narrow capillaries. The cathode is made of graphite; the anode is made of platinum. Electrolysis is carried out in argon atmosphere in a vertical furnace. After electrolysis the cell is cooled in a stream of argon. The current passing through the cell was (100 - 200) mA. (60 30) min of electrolysis was sufficient for obtaining reproducible data on the transference numbers. After cooling, the electrolyte from each compartment is drilled out, homogenized, dissolved (treated with sulphuric and hydrochloric acid) and analysed by AAS for the contents of Na, K, and Al (and Ca if present). From the difference in the composition of the electrolyte in the cathodic and middle (neutral) compartments, the transference numbers of Na+ and K+ can be determined. The transference number of anions is obtained by difference: 1 - t(Na+) - t(K+). The obtained transference numbers at different composition of the cryolite-based electrolyte are summarized in Table II. We assume that the temperature does not have a pronounced influence on the transference number. It follows that the transference number of Na+ is about twice as high as that of K+. The more acidic the electrolyte, the higher is the role of the fluoride anions in charge transfer. It was found out that addition of CaF2 to the electrolyte diminishes the contribution of the fluoride ions to the charge transfer.

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Figure 3. A view of a used graphite cathode with Al deposit. Analysis showed high current efficiency for Al (94-97 %); while in the case of potassium cryolite it was lower, i.e. 91.8 %. In the case of a melt prepared by melting pure cryolite, the transport number of sodium was found to be 0.99, in agreement with previous studies [1,5, 6]. When A1F3 was added to the melt, the transport number of sodium cations decreased, as shown in Fig. 4.

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Acknowledgement The authors are grateful for financial assistance from the Grant Agency of the Slovak Republic under the project VEGA 1/0535/08. Financial support from ALCOA INC. is gratefully acknowledged.

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Figure 4. The transport number of sodium (squares) and potassium (circles) cations as a function of the cryolite ratio (CR). Temperature range 795 - 1018 °C.

References 1. S.K. Ratkje, H. Rajabu, and T. F0rland, "Transference coefficients and transference numbers in salt mixtures relevant for the aluminium electrolysis," Electrochim. Acta, 38 (1993), 415423.

When extrapolating the data for the NaF - A1F3 system to the NaAlF4 composition (CR = 1), we obtain t(Na+) = 0.73. From models of cryolite-based melts [2, 3] it follows that this melt contains almost exclusively A1F4" anions. These anions apparently contribute to the transport of charge. On the other hand, in Na3AlF6 melts the fluoride ions, as predicted according to the cited model, do not contribute to charge transfer. The ionic radius of the sodium cation is 100 pm and that of the A1F4" anion is 165pm [9, 10], so that r(AlF4)/(r(Na+) + r(AlF4)) = 0.626. This estimate gives the order of the transference number of Na+ cation observed experimentally. This fact should be taken into account in ionic models of molten aluminates with complex anions. When the electrolyte was composed of an equimolar mixture of NaF and KF (+ AIF3), the sodium cation transferred about twice as much electrical charge as the potassium cation. The ionic radius of the sodium cation, r(Na+) = 100 pm and that of the potassium cation r(K+) = 133 pm. The ratio r(K+) / r(Na+) = 1.33. The ratio of the squares of the ionic radii r^K*) / r^Na*) = 1.77. It appears that the ratio of the transport numbers of t(Na+) / t(K+) is approximately proportional to the square of their ionic radii.

2. K. Grjotheim et al., Aluminium Electrolysis - Fundamentals of the Hall-Heroult Process (Düsseldorf, Germany: AluminiumVerlag, 1982), 175-178. 3. J. Thonstad et al., Aluminium Electrolysis - Fundamentals of the Hall-Heroult Process (Düsseldorf, Germany: AluminiumVerlag, 2001), 127-128. 4. W.B. Frank, and L.M. Foster, "Investigation of Transport Phenomena in the Cryolite-Alumina System by Means of Radioactive Tracers," J. Phys. Chem., 61 (1957), 1531-1536. 5. A.Tual, and M. Rolin, "Etude des nombres de transport ioniques dans les melanges cryolithe-alumine selon le principe de la methode de Hittorf—I. mise en oeuvre de la methode," Electrochim. Acta, 17 (1972), 1945-1954. 6. A.Tual, and M. Rolin, "Etude des nombres de transport ioniques dans les melanges cryolithe—alumine fondus selon le principe de la methode de hittorf—II. Resultats," Electrochim. Acta, 17 (1972), 2277-2291.

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7. E.W. Dewing, "Thermodynamics of the System L1F-AIF3," /. Electrochem. Soc, 123 (1976), 1289-1294. 8. A. Sterten, K. Hamberg, and I. Maeland, "Activities and Phase Diagram Data of NaF-AlF3-Al203 Mixtures Derived from Electromotive Force and Cryoscopic Measurements. Standard Thermodynamic Data of ί -Al203(s), Na3AlF6(s), and NaAlF4(l)," Acta Chem. Scand., A36 (1982), 329-344. 9. R.D. Shannon, "Revised Effective Ionic Radii and Systematic Studies of Interatomie Distances in Halides and Chaleogenides," Acta Cryst. A32 (1976), 751-767. 10. Z. Akdeniz et al., "Ionic Interactions in Alkali - Aluminium Tetrafluoride Clusters," Z. Naturforsch., 54a (1999), 570-574.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

Cells Process Modeling SESSION CHAIR

Marc Dupuis GeniSim Inc. Jonquiere, Canada

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DEVELOPMENT AND APPLICATION OF AN ANSYS BASED THERMO-ELECTRO-MECHANICAL COLLECTOR BAR SLOT DESIGN TOOL Marc Dupuis GeniSim Inc. 3111 Alger St., Jonquiere, Quebec, Canada, G7S 2M9 marc.dupuis@ genisim.com

Of course, the development of the TEM eliminates the need to make this kind of arbitrary assumption by calculating the contact pressure and then the corresponding contact resistance value based on some temperature and pressure dependant relationship [5].

Abstract After the successful development and application of an ANSYS based thermo-electro-mechanical anode stub hole design tool [1], an ANSYS based thermo-electro-mechanical collector bar slot design tool has been developed. Since the average contact resistance at the cast iron/cathode block interface is higher than the contact resistance at the cast iron/anode carbon interface, the potential for mV savings is even greater.

ANSYS® version 12.0 based TEM cathode collector bar slot model development As for the TEM anode stub hole design tool developed and presented last year [1], the TEM cathode collector bar slot model is based on the usage of ANSYS® SOLID226 3D thermo-electromechanical second order element together with CONTA174 and TARGE 170 thermo-electro-mechanical contact pair elements. CONTA174 element supports the setup of a pressure and temperature TCC (thermal contact conductance) and ECC (electrical contact conductance) values through the %table% option.

A demonstration model has been developed and used to study different collector bar slot configurations. The results obtained are presented. Introduction Contrary to the anode stub hole cast iron/carbon contact resistance problem, issues related to the cathode collector bar slot cast iron/carbon contact resistance have not been the subject of numerous publications in recent years.

Essentially, the only difference between the TEM anode stub model and the TEM cathode collector bar slot model is the topology which is quite easy to build and when needed to modify using ANSYS® parametric design language (APDL).

It is a bit strange in a way because since the introduction of 100% graphitized cathode blocks, the voltage drop due to the contact resistance represents more in percentage of the total cathode lining drop than the voltage drop due to the contact resistance represents in the total anode voltage drop. There should be room for further reduction of that lining voltage drop like it is the case for the anode voltage drop by using the thermo-electro-mechanical (TEM) collector bar design tool to optimize the cathode slot design.

One particularity of both TEM models that was not described in last year paper [1] is the selection of the thermal expansion reference temperatures. Contrary to Richard [5] who is creating a model geometry that corresponds to the room temperature geometry and hence is incorporating an air gap between the cast iron and the carbon (anode carbon in his case), the model geometry in the present work was constructed without incorporating an air gap between the cast iron and the carbon corresponding to the geometry when the cast iron has solidified.

Unfortunately, again contrary to the anode case [2], it is not so easy to instrument the cathode lining in order to measure the contact resistance between the cast iron and the cathode carbon in the collector bar slot. Boivin [3] did instrument a collector bar (see Figure 1 and 2 of his 1985 TMS paper) and indirectly measured 6.6 μΩηι2 assuming a uniform contact resistance value on all three contact interfaces.

In order to do that and still be able to accurately calculate the contact pressure of the unit (either anode or cathode) in operation, the material reference temperatures to calculate the thermal expansion must be set differently than in Richard's model. In that model incorporating a room temperature air gap, the reference temperature of all materials is the room temperature while in the present model that does not incorporate a room temperature air gap, the reference temperature of the cast iron is its solidification temperature (Ts in Equation 2 of [6]). The reference temperature of the other materials (stub and anode carbon or collector bar and cathode carbon block) is the average temperature of those materials when the cast iron solidified (like Ta in Equation 1 of [6]; notice that Equations 1 and 2 in [6] assume that the effective anode carbon temperature at cast iron solidification is T0 the ambient temperature which is a simplification not made in the present work).

Yet as argued by Sorlie [4], since contact resistance is very dependent upon applied pressure, one have to assume that most of the current passes through the vertical cast iron-to-carbon contact interfaces but there are no references on experimental measurements that will confirm that. This is the reason why over the last 20 years, per lack of measurements to confirm what is the true situation, when developing a thermo-electric cathode lining model, the author kind of arbitrarily assumed that the contact resistance on the top horizontal interface was twice the value of the contact resistance of the two vertical interfaces.

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Base case model Figure 1 is showing the geometry of the base case model. It is a quarter cathode block model of a "single slot per block" design type. Actually, there are two collector bars per block because the block is 3.67 m long and the two bars are 2.175 m long each leaving a section without bar in the middle of the block. Those two collector bars have a square cross-section of 160 mm x 160 mm. The cathode block has also a square cross-section of 48 cm x 48 cm. The size of the collector bar slot is 176 mm of height leaving room for 16 mm of cast iron above the bar and on average 200 mm of width leaving 20 mm of cast iron on each side of the bar. Yet, because of the typical "V" shape of the vertical faces of the slot, the cast iron thickness actually varies from a minimum of 15 mm to a maximum of 25 mm. It is assumed that there is 28 such cathode blocks in a cell running at 300 kA, so the current in each bar is 300/28/2 = 5.36 kA for a maximum current density in each collector bar of 5360/16/16 = 20.92 A/cm2. Figure 3: Temperature solution of the base case model As afirststep, the cathode voltage drop is calculated using constant user defined contact resistance values as in the TE model. Typical values of 4 μΩπι2 for the vertical interface and 8 μΩπι2 for the horizontal interface were selected (still using that arbitrarily factor of 2 between vertical and horizontal contact resistances). As presented in Figure 4, for setup, the model predicts a cathode lining drop of 212 mV.

Figure 1: Mesh of the base case model In a typical TE cathode side slice model [2], the collector bar and the slot are not represented in that much details but the full lining and potshell are also represented (see Figure 2). This is required in order to be able to accurately calculate the cathode heat loss. That calculation is not a requirement of the TEM cathode model, yet computation of the temperature is still required. Fortunately, it is possible to compute that temperature without having to represent the full lining by using appropriate boundary conditions (see Figure Figure 4: Voltage solution with constant contact resistance values Figure 5 is presenting the resulting current density at the edge of the cathode block. Some current is travelling vertically straight down from the top of the slot into the top section of the cast iron. This may or may not be real, no measurement being available to confirm or disprove that. The only thing that is known is that this would be the current density, if the value of the horizontal contact resistance would be twice the value of the vertical contact resistance. Assuming that the 4 and 8 μΩπι2 were selected to match measured cathode lining drop, the next step is to activate the temperatureand pressure-dependent contact resistance property in the model and calibrate the model so that it can predict close to 212 mV of cathode lining drop. Many parameters could be used to do that calibration. The one selected in the present work is Ta the effective collector bar temperature at cast iron solidification: a value of 750 °C was required to get the results presented in Figure 6.

Figure 2: Mesh of a standard TE cathode side slice model

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It is clear that far less current enters from the top horizontal interface section. Now at least that tool can be used to investigate how to improve the situation. Base case model,finermesh At this point, the model can be considered as validated since after calibration it is reproducing "measured" data. Yet, before starting to use the model as a design tool, it is also a good idea to test the model mesh sensitivity. The initial model is using a mesh that is much finer that the one of a standard TE cathode side slice model. But is the meshfineenough to well represent the contact behavior?

Figure 6: Voltage solution with variable contact resistance values

To answer that question, a second mesh was developed. The initial mesh has 2592 3D solid elements and 1065 2D contact elements. It took only 566 seconds to solve on a 64 bits dual core Intel Centrino T 9300 Cell Precision M6300 portable computer running ANSYS® 12.0 version. The refined mesh has 10924 3D solid elements and 2760 2D contact elements. Solving the same problem with that refined mesh took 5225 seconds, so about ten times more than solving for the initial mesh.

So after calibration, the total cathode voltage drop prediction is close of being equal. But is the current density in the cathode block edge very different now? In order to answer that question, one only has to compare Figure 5 with Figure 7 presenting the current density solution obtained while using the temperature and pressure contact resistance property in the model.

The predicted cathode lining voltage drop is identical; so as far as the accuracy of the solution is concerned the initial mesh is clearly good enough. But the current density vectors presented in Figure 8 indicate that the finer mesh is helping a lot in the interpretation of the results. In Figure 8, the current is concentrating itself in three points where the contact pressure is concentrated.

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The model predicts 206 mV using the constant contact resistance setup and 197 mV while using the temperature- and pressuredependent contact resistance setup. So a saving of about 6 mV came from the fact that there is less voltage drop in the collector bar section outside the cathode block. Then, according to the TEM model, an additional reduction of about 9 mV can be expected due to the improved contact in the top horizontal interface section that resulted from the decrease of the cast iron thickness and hence the increase of the contact pressure. Since it is not expected that a direct steel collector bar/cathode carbon block interface contact would behave any differently than a cast iron/cathode carbon block contact, this run is really testing the option of not putting any cast iron above the bar. Figure 10 shows the corresponding current density in the cathode block edge.

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Figure 8: Current density with variable contact resistance values finer mesh version Same slot, higher collector bar As it is not clear at this point if having any cast iron on top of the bar is useful, the aim of the first design change run is to test that. In this run, the 160 mm wide x 160 mm high collector bar is replaced by a 160 mm wide x 174 mm high collector bar leaving only 2 mm of cast iron above the bar (completely eliminating the cast iron above the bar would require a new model topology). The new model geometry is presented in Figure 9.

Figure 10: Current density with variable contact resistance values for the same slot, higher collector bar case Same slot, higher and wider collector bar It is clear that the maximum vertical interface contact pressure will be achieved using the minimum cast iron thickness possible. This reduction must be done by increasing the collector bar section, not by decreasing the collector bar slot width because, inside the block, the effective collector bar section is the slot section as the current travels in the cast iron too. So this second design change run is testing a 174 mm wide x 174 mm high collector bar using the same collector bar slot leaving on average only 13 mm of cast iron. Figure 11 is presenting the corresponding model geometry. Figure 9: Mesh of the same slot, higher collector bar case

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due to the improved contact, in addition of the 13 mV reduction due to the increase of the collector bar section: hence a grand total of 20 mV reduction over the base case value for a reduction of 9.4% while still keeping the same collector bar slot aspect ratio and cross-section. New collector bar aspect ratio This of course is only the beginning of a multitude of new collector bar slot configurations that can now be tested using this new TEM collector bar slot design tool, like testing if a "W" profile would provide a better contact than the standard "V" shape profile. Yet, testing a "W" profile would require a little change in the model topology while there are still many new cases that can be analyzed using the current model topology.

Figure 11: Mesh of the same slot, higher and wider collector bar case The model predicts 199 mV using the constant contact resistance setup and 195 mV while using the temperature- and pressuredependent contact resistance setup. So a saving of about 13 mV came from the fact that there is less voltage drop in the collector bar section outside the cathode block. Then, according to the TEM model, an additional reduction of about 4 mV can be expected due to the improved contact, which is less than the previous case.

Per example, it is well known that a rectangular collector bar crosssection is more efficient than a square collector bar cross-section. But it would be interesting to see if the TEM collector bar slot model confirms this. In this forth design change, the 174 mm x 174 mm collector bar is replaced by a 144 mm wide x 210 mm high collector bar keeping about the same cross-section by significantly changing the aspect ratio. Figure 13 is presenting the resulting model geometry still keeping 13 mm of average cast iron thickness on the sides and the minimum thickness area at the upper quarter point.

New slot design, higher and wider collector bar Next, the most difficult thing is to come up with a collector bar slot design change that improves the contact and hence decreases the cathode lining drop. As a simple example, it is possible to study the impact of changing the position of the minimum thickness area of the slot. In this third design change run, that position is moved upfromthe mid point position to the top quarter point position still keeping the bigger 174 mm x 174 mm collector bar and still keeping the same average 13 mm cast iron thickness on the two side sections (see Figure 12).

Figure 13: Mesh of the new collector bar aspect ratio case The model predicts 192 mV using the constant contact resistance setup and 187 mV while using the temperature- and pressuredependent contact resistance setup. It is fair to compare those results with the ones of the previous case as only the collector bar aspect ratio has been changed. It is also fair to compare the two constant contact resistance results and the two variable contact resistance results between themselves.

Figure 12: Mesh of the new slot design, higher and wider collector bar case The model obviously still predicts 199 mV using the constant contact resistance setup but now predicts 192 mV while using the temperature- and pressure-dependent contact resistance setup. So this is an additional decrease of 3 mV for a total of 7 mV decrease

According to the constant contact resistance model setup with an arbitrarily ratio of 2 between the horizontal and the vertical contact resistance, that change of aspect ratio should reduce the cathode voltage drop by 7 mV. According to the variable contact resistance model setup, that change of aspect ratio should reduce the cathode voltage drop by 5 mV. So there is no strong disagreement between the two versions of the model which is a good thing for the user of the standard TE cathode side slice model. Of course, changing the collector bar aspect ratio will also affect the lining life so maybe in that context this design change is not an improvement!

A very quick design optimization study has revealed that it is possible to reduce the cathode lining drop of a typical single collector bar slot per block design having a square collector bar section of 160 mm x 160 mm by 40 mV or about 19%. This is done by keeping the same amount of carbon above the collector bar by shifting to a double collector bar slots per block design. This design is obtained by removing the cast iron above the bars with an increase of the bar height while keeping the same collector bar slot height and also reducing the cast iron thickness on the bar sides by increasing the bar width while keeping the same slot width.

Two collector bar slots per block It is also well known that it is better to use two collector bar slots per block instead of one. So for this fifth design change, the single 174 mm x 174 mm square collector bar has been replaced by two 87 mm wide x 174 mm high rectangular collector bars. The average cast iron thickness on the two sides of the two collector bars has been decreased to 11.5 mm which adds up to 46 mm of cast iron as opposed to a total of 26 mm in the single collector bar slot per block design. Figure 13 is presenting the resulting model geometry. That it is still the same model topology but this time the model represents 1/8 of a full cathode block instead of 1/4 as in all the previous cases. Typically, the two collector bar slots are located a bit closer to the block centerline than the two block quarter points in order to have thicker carbon wings, but it is not possible to test this case using the current model topology of course.

It was also demonstrated that changing the collector bar slot profile design had some influence on the cathode lining drop. Performing a true collector bar slot profile optimization study would have required the development of a multitude of alternative model topologies which was not done in the present study. References 1. M. Dupuis, "Development and Application of an ANSYS® Based Thermo-Electro-Mechanical Anode Stub Hole Design Tool", in Proceedings ofTMS Light Metals, (2010), 433-438. 2. M. Dupuis and C. Fradet, "Using ANSYS® Based Aluminum Reduction Cell Energy Balance Models to Assist Efforts to Increase Lauralco's Smelter Productivity", Proceeding of the ANSYtf® 8th International Conference, volume 2, 2.233-2.240, (1998). 3. R. F. Boivin, P. Desclaux and J. P. Huni, "Cathode Collector Bar Temperature and Current Pickup:", in Proceedings ofTMS Light Metals, (1985), 625-635. 4. M. Sorlie and H. Gran, "Cathode Collector Bar-to-Carbon Contact Resistance", in Proceedings of TMS Light Metals, (1992), 779-787. 5. D. Richard, "Conception des tourillons d'anode en usage dans une cuve de Hall-Heroult ä l'aide de la methode des elements finis", M.Sc. Thesis, Universite Laval, Quebec, Canada, (2000).

Figure 14: Mesh of the two collector bar slots per block case The model predicts 178 mV using the constant contact resistance setup and 172 mV while using the temperature- and pressuredependent contact resistance setup.

6. D. Richard, P. Goulet, O. Trempe, M. Dupuis and M. Fafard, "Challenges in Stub Hole Optimization of cast iron rodded anodes", in Proceedings of TMS Light Metals, (2009), 10671072.

So as far as the constant contact resistance version of the model is concerned, replacing a single 174 mm x 174 mm collector bar by two 87 mm x 174 mm should result in a reduction of 21 mV while the variable contact resistance version of the model is predicting a reduction of 20 mV. So the two versions of the models are in fairly good agreement. Conclusions An ANSYS® version 12.0 based fully coupled TEM collector bar slot design tool has been successfully developed and is now available to the whole aluminium industry through GeniSim Inc. The ANSYS® based APDL model is parametric, which means that for a given model topology, it is possible almost instantaneously to edit the APDL model input file to change the model geometry and submit another run. The finer mesh quarter block model presented here solves in only around 5200 CPU seconds on a 64 bits dual core Intel Centrino T 9300 Cell Precision M6300 portable computer running ANSYS® 12.0 version. So this parametric ANSYS® based TEM collector bar slot model is a very efficient tool to study alternative collector bar and collector bar slot design.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Impact of Amperage Creep on Potroom Busbars and Electrical Insulation: Thermal-Electrical Aspects Andre Felipe Schneider, Daniel Richard and Olivier Charette Hatch, 5 Place Ville-Marie, Suite 200, Montreal, Quebec, Canada, H3B 2G2 Keywords: Amperage Creep, Potroom Electrical Safety, DC Busbars, Finite Element Analysis Abstract

Pot-to-pot Busbars Heating & Cooling Mechanisms

Busbars electrical insulation is a critical aspect of aluminium smelters potrooms electrical safety. With amperage creep, busbars are typically running hotter than they were at start-up. The longterm reliability of busbars and the integrity of insulating materials is therefore of concern.

In a continuous busbar, heat is internally generated by the passage of electrical current through it, as per Equation ( 1 ) .

To assist smelters evaluate the performance of the busbars systems under realistic operating conditions, a methodology was developed using ANSYS™-based numerical simulation, laboratory testing and in situ measurements. This approach has been validated on different pot technologies and smelters.

where: q gen,busbar is the volumetric heat generation in a continuous busbar, [W/m3]; p=f(T) is the temperature-dependant material's electrical resistivity, [Ω.ηι]; and J is the conductor's current density, [A/m2].

a'

fc fc

τΓ gen,busbar

=o-J2 H

(Π

*

On the other hand, when surfaces from different bodies are in contact, a localized heat generation occurs at the interface due to the imperfect contact, represented by a contact resistance, shown in Equation ( 2 ).

A realistic test case based on a demonstration busbar system is presented and the typical impact of line current, ambient temperature and selected operational procedures on thermalelectrical performance and reliability is discussed.

a

Introduction

f gen, contact

=R AX

ctc

J2

*

(2)

where: q gen,contact is the interfacial heat generation at the contact between different conductors, [W/m2]; Rctc is the contact resistance between neighbor conductors, [Ω.ιη2].

Busbars are an integral part of the aluminium reduction technology and their design has a profound impact on the stability and performance of cells, notably through magneto-hydrodynamics (MHD) effects. However, on the most basic level, their purpose is to collect current from the cathodic part of a cell and feed it to the anodic part of the next. Busbars are also needed to connect groups of cells, for example at passageways and between potrooms, and to carry the electrical current to and from the rectifiers. The reliable operation of these conductors and their insulating materials is therefore of capital importance to sustaining smelter operations and workers safety.

By inspecting these equations it is obvious that increasing current density / (through amperage creeps for instance) or increasing contact resistance Rctc (for example on risers bolted connections, bypass wedges and start-up shunts) leads to increased Joule Heat generation and, consequently, leads to increased busbars and insulators temperatures. Heat can also be conducted from the hotter cathode block/collector bar assemblies to the colder pot-to-pot busbars through the flexibles linking them. Furthermore, conductor surfaces facing the potshell will be exposed to radiant heat loads.

Typically, busbars design will be dictated by a target electrical current distribution mostly based on MHD considerations and acceptable collector bar current unbalance (bar-to-bar, upstream/downstream, tapping end/duct end). To reduce cost, sizing of the bars is often based on the minimum bar cross-section at the maximum allowable temperature. Insulators are also selected based on that allowable temperature.

Busbars are mainly cooled by means of convection and radiation to the ambient. Potroom building ventilation will generally determine the air flow around the conductors. Heat is also redistributed in the busbars through conduction. For example, a large section busbar with a low current density can act as a heat sink for a conductor with higher current density. Heat is also transferred between neighboring busbars by means of bus-tobus radiation.

Past experience has shown that the window for a trouble-free operation of these components tends to reduce with increasing ambient temperature, pot line current and contact resistance between non-welded assemblies, such as in bolted connections and short-circuiting stations. In particular, excessive heat generation due to increased current or poor contact resistance has led to busbars systems damage and, in extreme cases, to catastrophic failure of both conductors and insulators. Poor cooling conditions, for example due to excessive covering material in the potroom basement, can also contribute to busbars overheating. Amperage creep and operational procedures therefore play an important role in the performance of busbars.

Note that all these heat generation and transfer mechanisms (depicted in Figure 1) are dependent on the potline current, busbar system, pot room ventilation and cell designs, and operational procedures.

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resistivity of aluminum are considered to be those of the pure element. During bypass analyses, short-circuiting wedges and equalizer components (e.g., wedges or flexibles) can be included while the pot shell and collector bars flexibles can be removed. Note that the worst condition is when the hot pot shell is still in place (thus transferring heat through radiation to the busbars), the collector bar flexibles are cut and the extra equalizer components (if any) are not included (which forces the whole potline current to flow through a selected group of busbars, leading to increased current densities). Contact resistances at the risers bolted connections and between the wedges and shunting-clamping stations are also considered. Reference voltage is applied at the bath/metal pad interface of the cell of interest and uniform anode current is prescribed at the next considered downstream interface, which is modeled by equivalent resistors representing the anode, rod, yoke and bath. Pot-to-pot, cathode assembly, busbars external and equivalent anode resistances are calibrated based on measured voltage break downs.

Legend Conduction from cathode Convection to ambient Radiation from shell Radiation between busbars Radiation to ambient

Temperatures are typically prescribed at the following locations:

\

•

\

• • •

Volumetric heat generation Interfacial heat generation

*

Busbars supports bottom (basement's ambient temperature); Cell of interest's bath/metal pad interface (~960°C); Cell of interest's cathode assembly's bottom (~900°C); Next downstream riser flexibles/anode bridge interface (measured or estimated anode bridge temperature).

The following heat generation and transfer mechanisms are considered (as per Figure 1):

Figure 1 - Heating and cooling mechanisms acting on pot-topot busbars.

•

Cell-to-Cell Busbars Thermo-Electrical Problem Modeling

• •

The calculation of the temperature and voltage distributions in the busbars is a non-linear coupled problem through Joule heating and the temperature-dependent nature of aluminium electrical resistivity. To solve this thermo-electrical (TE) problem and assess the operational performance of busbars systems, a methodology was developed using ANSYS™-based numerical simulation, laboratory testing and in situ measurements.

•

Volumetric and interfacial heat generation - see Equations ( 1 ) and ( 2 ) , respectively; Busbars-to-busbars and cathode-to-busbars conduction; Shell-to-busbars, busbars-to-busbars, busbars-toambient and collector bars-to-ambient radiation - note that the sum of all view factors for any given surface necessarily equals to 1; Busbars-to-ambient and collector bars-to-ambient convection.

In Situ Measurements

The proposed approach, which has been validated on different pot technologies and smelters, can be used to assess the performance of busbar systems and their insulators regarding the impacts of potline current intensities, contact resistances and ambient conditions.

For existing potlines, an extensive in situ measurements campaign is carried out in order to determine suitable boundary conditions for the pot-to-pot busbars TE model. A non-exhaustive list of the measurements typically performed includes:

Finite Element TE Model Assumptions and Characteristics

•

The fully parametric pot-to-pot busbars TE Finite Element (FE) model spans two consecutive metal pads (i.e., two consecutive equipotentials) allowing the electrical current to redistribute itself according to the equivalent resistance of each circuit. The pot shell, busbars supports and insulators are included in the thermal calculation. Normal operation (Figure 2), single (Figure 3) and multiple bypass conditions can be analyzed. The system is assumed to be in thermal equilibrium (i.e., steady-state) and the temperature-dependent thermal conductivity and electrical

•

526

Air velocities and free-stream temperatures on the vicinity of cathode and anode busbars. Used on the determination of the temperature/velocity-dependent convection film coefficient hc=f(T,V) to be applied on busbars surfaces, which are calculated using wellknown semi-empirical correlations [1]; Thermographic survey of the pot shell's temperature distribution. The measured data is mapped onto the shell surface by using in-house Visual Basic and APDL macros, thus establishing the temperature profile to be considered in the shell-to-busbars heat exchange;

• • • •

After continuous potline current increases, a major busbar retrofit transformed the downstream (DS) end risers into quarter side risers in order to reduce the magnitudes of both transversal and vertical magnetic flux density components. Note, however, the shunting-clamping stations, located close to the head walls upstream (US) corners, were kept untouched.

Pot-to-pot, cathode, busbars external and equivalent anode voltage drop (for model calibration purposes); Risers bolted connection and wedge-to-shunting clamping stations equivalent contact resistances Rctc (model calibration); Collector bars, risers and main cathode busbars current distributions (model validation); Cathode and anode busbars temperatures at several locations (model validation).

This modification brought the busbar design a step closer to typical 1980s concepts [2] and enabled, along with other important modifications, to operate the cells at 200 kA. Figure 2 and Figure 3 show the modified busbars and supports arrangement in normal operation and single bypass modes.

Laboratory Testing Ideally, the data sheet of an electrical insulator should specify the maximum allowable temperature (temperature rating) that this component can be submitted to without degrading its insulating and mechanical properties. However, different suppliers may provide different temperature ratings for the same class of material or this data may even not be provided at all. In order to assess the operational limit of electrical insulators when available information is either contradictory or incomplete, laboratory testing under realistic operation conditions (regarding service temperatures and pressures) can be performed following wellestablished testing standards (e.g., ASTM D3755).

Anode Bridge Riser #1

Coupling with Cell Heat Balance and Potroom Ventilation Models

Concrete Supports & Insulators

The pot-to-pot busbars TE model can be further coupled with external cell heat balance and potroom ventilation analyses. Two kinds of coupling can be considered: •

•

Riser Contact Resistance

Figure 2 - Fictitious side-by-side busbars system with quarter and end risers in normal operation

One-Way Coupling: both the cell heat balance and the potroom ventilation analyses are performed considering loads and boundary conditions obtained independently from the pot-to-pot busbars model. Their outputs (pot shell and cathode assembly temperatures; air velocities and temperatures, and convective heat transfer coefficient distributions at the vicinity of cathode and anode busbars) are then used as boundary conditions for the busbars TE analysis. This is generally what has been done; Two-Way Coupling: while the cell heat balance and the potroom ventilation analyses provide the required boundary conditions for the busbars TE model, they would in turn have some of their loads and boundary conditions defined by pot-to-pot busbars model's outputs. Potroom ventilation models can be fed with proper busbars heat generation rates. Cell heat balance models can use the predicted collector bars current distribution and busbars surface temperatures to fine tune the cathode heat generation rates and the pot shell's radiant heat losses. This is an iterative procedure that repeats itself until convergence is achieved. This would be an improvement on the approach used so far.

Equalizer Wedges Bypass Wedges

Wedges-to-ShuntingClamping Stations Contact Resistance Figure 3 - Single cell bypass condition Long-term operation at 200 kA led, however, to systematic failure of busbars and insulators in our fictitious smelter. A project was then carried out to investigate the damages probable causes and propose a viable, cost-effective mitigation strategy to avoid further damage. Upon the completion of the numerical tools, an in situ measurements campaign was performed during summer time to allow for model calibration and validation under worst-case ambient conditions1. Laboratory testing showed that electrical insulators start to degrade at 215°C and catastrophic failure occurs at 250°C.

Test Case Model In order to illustrate the proposed approach's capabilities, a fictitious pot-to-pot busbars circuit will have its performance assessed when running under different amperage, ambient and operational conditions. The considered system was designed to operate at 150 kA and originally consisted of side-by-side cells with a symmetric design and 4 identical end-risers - a typical 1970s topology, according to [2].

1 Minimum and maximum historical ambient temperatures are, respectively, 0°C and 30°C.

527

Normal Operation

163°C

Figure 4 shows the predicted busbars and supports temperatures for normal operation at 150 kA, 30°C. Note that even though the current is evenly distributed (-25%), the quarter risers (#2 & #3) are considerably hotter than the end ones (#1 & #4) due to increased radiant heat received from the pot shell's side walls.

IJ.U'tiJ

C^.i^φ

23.^1

iZi.-WO

11/.til

Figure 4 - Normal operation at 150 kA, 30°C: temperature distribution [°C].

Figure 5 - Single bypass at 150 kA, 30°C: temperature distribution [°C] when following standard bypass procedures.

Table 1 summarizes the predicted temperatures for normal operation under design and actual potline currents. It can be seen that the components temperature increases with both potline current and ambient temperature. Note that electrical insulators maximum numerical results are well within the maximum acceptable temperature (215°C) for all cases.

Practice in our fictitious smelter has shown, however, that the integrity of the electrical insulators could be maintained even when the pots were short-circuited without the systematic usage of the equalizer wedges at 150 kA, as shown in Figure 6. This resulted in a decrease of the labor involved on the bypass/preheating procedures and became the smelter's modus operandi. Furthermore, note that when the equalizer wedges are not in place, the quarterrisers#2 & #3 are disconnected.

Table 1 - Normal operation: busbars and insulators maximum predicted temperatures [°C]. End Risers #1, #4 Quarter Risers #2, #3 Shunting-Clamping Stations Insulators Supports Insulators

150 kA, 30°C 117 148 72

0°C 142 179 75

200 kA 30°C 161 203 106

115

134

161

!

Single Bypass Condition at 150 kA Figure 5 shows the predicted temperatures when bypassing one cell at 150 kA, 30°C, according to the standard operational procedures (use of equalizer wedges and typical shuntingclamping stations-to-wedges contact resistance Rctcwdg=0.0S0 μΩ.πι2). Maximum global temperatures occur at the end risers (#1 & #4) due to increased loading (32% of potline current each while each of the quarter risers #2 & #3 carry 18%). The bypassed cell's upstream head busbars evidently show increased temperatures for the same reason (50% of potline current each). Maximum electrical insulators temperatures at the shunting-clamping stations and supports are, respectively, 114°C and 125°C. Figure 6 - Single bypass at 150 kA, 30°C: temperature distribution [°C] without equalizer wedges.

528

Coming back to our fictitious potline, one also notes that the insulating materials of the shunting-clamping stations were also damaged. If enough electric potential difference between the cathode rings of two adjacent pots is experienced (e.g., a severe anode effect), the current may flow through the steel spacers and/or tie rods. Again, Figure 9 shows a real-life severely damaged shunting-clamping station that conducted part of the potline current through its steel spacers under similar conditions. Part of the aluminium busbar melted while the steel spacer and the bottom tie-rod and insulating tube assembly fell in the basement.

Single Bypass Condition at 200 kA The fictitious plant's established practice of not installing the equalizer wedges led to severe degradation of the insulating materials when operating at higher amperages, as can be seen in Figure 7. End risers #1 & #4 temperatures can be as high as 490°C when bypassing the cells at 200 kA, 30°C without the equalizing wedges. The temperatures of the electrical insulators supporting the bypassed US head busbars can be higher than 250°C, as shown in the detail. The maximum predicted temperature of the shunting-clamping stations insulators is 290°C. As previously mentioned, the system's insulting materials experience catastrophic failure at 250°C.

Figure 9 - Example of damaged shunting-clamping station. It's obvious that this shunting-clamping station is not able to provide an adequate electrical contact with the wedges anymore. Data gathered during a real-life in situ measurements campaign showed that the contact resistance at damaged shunting-clamping stations can be an order of magnitude larger then the typical value. Figure 10 shows the predicted temperatures for a cell with poor electrical contact at the right-side shunting-clamping station being bypassed at 200 kA, 30°C without equalizer wedges. Note that end riser #l's temperature increased to 513°C due to current rerouting. Figure 7 - Single bypass at 200 kA, 30°C: temperature distribution [°C] without equalizer wedges. To illustrate the real-life implications of this, Figure 8 shows a busbar-to-support electrical insulator that has experienced severe damage under similar conditions.

Figure 10 - Single bypass at 200 kA, 30°C: temperature distribution [°C] without equalizer wedges and with a damaged shunting-clamping station.

Figure 8 - Example of damaged electrical insulator.

529

There is a critical riser contact resistance that would lead to a riser failure. Figure 11 shows the predicted temperatures for the single bypass at 200 kA, 30°C for the same cell if end riser #1 has a contact resistance -6 times larger than the average value. Note that the endriser#1 reached the aluminum's melting point.

bypass procedure is followed, even at 200 kA, 30°C - see Figure 13. 244°C

Critical riser contact resistance 666°C

n

-3

30

7a

1 "& b d 201 B 225 ™ 250

Damaged shuntingclamping station (leads to increased contact resistance)

Figure 13 - Single bypass at 200 kA, 30°C: temperature distribution [°C] when following standard bypass procedures. Finally, the fictitious smelter's bypass/pre-heating procedures were reviewed and a rigorous voltage drop and temperature follow-up routine was established to control the shuntingclamping stations and risers contact resistances.

Figure 11 - Single bypass at 200 kA, 30°C: temperature distribution [°C] without equalizer wedges, with a damaged shunting-clamping station and critical riser contact resistance.

Conclusions

A real-life damaged end riser under similar circumstances can be seen in Figure 12.

A methodology based on numerical simulation, laboratory testing and in situ measurements was presented for the assessment of the thermal-electrical performance of busbars systems. It was shown that busbar design, insulating materials selection, amperage creep, operational procedures and potline conditions have important effects on the reliability of these systems.

" \

\ ^Ρρ, WL^ä

AH

%

Evidence was shown that busbars systems should not be ignored when planning amperage creep or when setting operational control procedures. Although not presented here, experience has also shown that liaison busbars between pot groups, between substation and potline and between potrooms, need to be included in the planning of amperage creep.

Mw..^''

Figure 12 - Example of damaged end riser.

Future work includes the evaluation of the thermal-mechanical performance of the busbars systems, including expansion joints.

Proposed Mitigation Strategy

References

Initially, in our fictitious smelter the damaged busbars have been repaired either by reconstructing the original conductor's crosssection with aluminum weld plates or by replacing the damaged part. The damaged insulators were replaced by materials with higher temperature rating and new tie rods and spacers were installed wherever required. An additional bypass analysis showed that all electrical insulators should operate within the allowable range (maximum predicted temperature=193°C) if the standard

1. 2.

530

F. P. Incropera & D. P. Dewit, Fundamentals of Heat and Mass Transfer - 5th Ed. John-Wiley & Sons (2002), 981 pages. A. R. Kjar, J. T. Keniry & D. S. Severo, Evolution of Busbar Design for Aluminium Reduction Cells, 8th Australiasian Aluminium Reduction Technology Conference, 3Γί1-8Λ October 2004.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Modern Design of Potroom Ventilation Anastasiya Vershenya1, Umesh Shah1, Stephan Broek1, Tom Plikas1, Jennifer Woloshyn1 and Andre Felipe Schneider2 1

Hatch, Sheridan Science & Technology Park, 2800 Speakman Drive, Mississauga, Ontario, Canada, L5K 2R7 2 Hatch, 5 Place Ville-Marie, Suite 200, Montreal, Quebec, Canada, H3B 2G2 Keywords: Potroom Ventilation, Cell Emissions, Heat Release Rates, Computational Fluid Dynamics (CFD) ventilation flow rate predictions. Since neither ambient temperature nor wind characteristics can be assumed constant over long periods of time, it is necessary to identify specific outdoor design conditions for the potroom ventilation. Based on weather data, a set of prevailing wind conditions has to be selected, thus affecting the flow inside the potroom in the most adverse way. It is up to the CFD modeling expert to identify those scenarios and use known meteorological data to define the ambient conditions for those simulation runs.

Abstract A typical natural ventilation design of a potroom incorporates incoming ambient air flow dividing into two main streams. The 'under the pot' stream passes through basement infrastructure and sweeps the heat/emissions from the pots upward to roof. The 'over the pot' stream flows through the Claustra wall and pushes the emissions/heat away from the working zone. This paper describes the use of CFD modeling to accurately predict the ventilation air flow split between these two main streams and improve the ventilation design, thus meeting the defined workplace hygiene standards.

The second CFD model represents the inside of the potroom, including detailed geometry of the pot infrastructure, to accurately capture flow resistance. In addition to applying boundary conditions based on external model predictions, the internal potroom model needs to include pot cell emissions and heat release rates that reflect the total calculated heat load per pot. Since the internal model is used to determine whether working zone conditions are meeting the workplace hygiene standards, Hydrogen Fluoride (HF) emissions need to be captured as well. Both heat releases and HF emission are specified as boundary conditions for surfaces within pot geometry. Separate pot heat balance and busbar thermoelectric analyses can be used to determine the correct heat generation rates.

The model's success greatly depends on reliable input conditions such as cell emissions and heat release rates. An innovative approach for the calculation of cell emission rates is thus introduced. Furthermore, the ventilation model can be coupled with pot heat balance and busbars thermoelectric analyses in order to obtain proper Joule heat generation and release rates. Introduction During development of a smelter, potroom ventilation design needs to address both heat and fugitive emission issues in order to comply with workplace standards. Computational fluid dynamics (CFD) is a commonly used technique which aids in design and further evaluation of the ventilation system performance as discussed in [1] and [2]. In addition, to correctly incorporate the cell emissions and heat release rates inside the potroom, CFD modeling needs to account for ambient conditions outside of the potroom that can effect ventilation such as local terrain, wind speed and direction. Often two separate CFD models are built in order to properly predict wind velocity, temperature and pressure outside of the building with one way coupling to the building's detailed interior model. Due to complex interactions between multiple boundary conditions, it is critical to use correct inputs for both models. To ensure success, different disciplines have to come together to outline the various boundary conditions and use CFD analysis to comprehensively design the ventilation system.

Cell Emissions Cell emissions are an important input parameter for any ventilation model, which traditionally are taken care of by prescribing arbitrary cell emission rates at specific locations. In order to obtain a set of boundary conditions that holds a better connection to actual potroom practice, an in-house computer model for the prediction of emissions profiles was developed. Such a model is based on: • • • •

External and Internal CFD Models

• • •

Purpose of the external model is to establish wind flow patterns in close proximity to the building based on ambient temperature, wind velocity and direction. The geometry of an external model should include the main buildings on plant site and the surrounding terrain. Results of the external CFD model will provide pressure and velocity profiles for the two streams identified in the potroom slice model: 'under the pot' and 'over the pot'. An improper external model set up could result in an incorrect representation of the pressure differentials between the air inlets and the roof vent, which would lead to erroneous

The main production parameters; The gases produced from the electrolysis process; The configuration of the cell including covers and possible gaps; The duration of each step in the production process (tapping, anode change, sampling, etc); Normal draft and/or high draft ventilation; Anode effects; Secondary emissions from spent anodes and bath material removed from the cell.

All of these lead to a set of emission values that are applied at specific points on a cell where it is most likely that the emissions occur (see also Figure 9).

531

by the industry. The plant discussed in this case study is Aluminerie Alouette located in Sept-Iles, Quebec.

Cell and Pot-to-pot Busbars Heat Releases The incoming fresh air that enters the building cools down the cells and busbars and, in turn, has its own temperature risen by the released thermal energy. One possible way to take these heat sources into account is by estimating the heat generation through measurements and well-known analytical and empirical relationships: •

•

Firstly, the external model of the plant and surrounding terrain was used to simulate the flow patterns around the buildings for a series of particular weather conditions reported on site. External model geometry with applied boundary conditions is shown in Figure 1. The wind is treated as a turbulent, isothermal and incompressible flow with a velocity profile at the inlet specified to match the expected shape for an atmospheric boundary layer. Properties are assigned for air at atmospheric pressure, measured local ambient temperature and humidity. Upon convergence of the external model, the wind patterns around the potroom are determined for each scenario. The information extracted from the results of the external model simulation includes the velocity inlet and pressure outlet profiles that are used as boundary conditions for the internal model.

The energy dissipated by the cell can be estimated by a voltage breakdown. The amount of heat released by each cell component (collector bars, pot shell's side walls and end walls, pot covers and gas collecting ducts) can be obtained through a thermal blitz campaign; The internal heat generation on each busbar can be estimated through the Joule Heat Equation. Current distribution and busbars equivalent resistances can be obtained by means of voltage drop and temperature measurements.

3 de and top faces: P= 0 Model Inlet: Specified velocity profile

Coupling with Cell Heat Balance and Pot-to-pot Busbars ThermoElectrical Models On the other hand, the CFD ventilation model can be further coupled with external cell heat balance and pot-to-pot busbars thermo-electrical (TE) analyses. Two kinds of coupling can be considered: •

•

One-Wav Coupling: both the cell heat balance and the busbars TE analyses are performed considering loads and boundary conditions obtained independently from the CFD model. Their output heat release rates are then used as fixed heat sources for the ventilation analysis. This is generally what has been done; Two-Way Coupling: while the cell heat balance and the busbars TE analyses provide the heat sources for the ventilation model, they would in turn have their convective boundary conditions defined by the air velocities and temperatures obtained from the CFD code. This is an iterative procedure that repeats itself until convergence is achieved. This would be an improvement on the approach so far.

Figure 1. Boundary conditions for the external model geometry with the domain configured for a parallel wind. The actual weather condition instances used to represent prevailing wind directions of interest were selected based on an extensive analysis correlating meteorological data from the onsite weather station (Pointe-Noire) to reported anemometer measurements. It was found that the maximum and minimum anemometer readings for a given period correlated well with certain wind directions. Based on predictions obtained from the external model, an internal model of a slice of the potroom was used to simulate ventilation conditions corresponding to the identified scenarios. Since the velocity and pressure from the external model vary along the length of the potroom, the location of the profiles for application to the slice model is selected to simulate the worst case conditions. Generally, adverse wind conditions correspond to pressure at the roof vent being greater than pressure at the air inlet or pressure on one side of the potroom being significantly greater than on the other side, as can be seen in Figure 2.

Case Studies Over the years, CFD modeling has become an accepted practice which is used to improve control of contaminants, such as HF and SO2 gases, and provide a method of controlling cooling air flow in potrooms. Through extensive experience in the aluminum industry, Hatch has developed a methodology that ensures proper modeling procedures to be used for analyzing potroom ventilation systems. To share the expertise on how to use CFD analysis properly, a list of case studies supporting the approach is presented. Environment Canada - CFD Modeling of Potroom Ventilation (2010) The objective of the Environment Canada study was to investigate the impact of wind speed and direction on the flow exiting the roof vents in an aluminum smelter. Specifically, the study was focused on wind effects on the velocity reported at the roof ventilator, which is used in the calculation of declared emissions

532

Front and back faces Symmetry boundaries

| Pressure outlet based on ♦ results from external model 175 225 X Position (m)

325

Figure 2. Predicted pressure profile along the north and south basement panels and the roof monitor of the potroom due to wind. A more rigorous approach to the internal model setup, than described above, is implemented in this case in order to investigate the influence of wind conditions on the ventilation system performance. The internal model includes detailed geometry of the potroom as well as an additional domain outside of the building. Profiles obtained from the external model were applied to the domain of the internal model, thus capturing air velocity and pressure directly outside of the air ingresses and outlets. This approach more accurately predicts the division of 'over the pot' and 'under the pot' air flows. It also allows the model to capture re-entrainment of flow exiting the roof at the basement inlets when it occurs.

Figure 3. Boundary conditions for the internal model geometry with the domain configured for a wind approaching normal to the potroom. In case of the wind direction being parallel to the potrooms, internal model has been setup as shown in Figure 4. A specified wind velocity is applied to the front face with a pressure profile applied to the back face, while the top and both side faces external to the potroom are set to the atmospheric pressure to allow air to exit or enter the domain, as required. For the front and back faces inside the potroom, periodic boundary conditions are used for the normal wind direction. Both of these configurations assume uniform internal conditions along the length of the potroom and neglect end effects, thus being representative of the potroom's central portion.

Based on whether wind direction of interest was parallel or perpendicular to the potrooms, the internal model setup had two different configurations. Figure 3 displays the model configuration when wind direction is perpendicular to the potroom. A specified wind velocity is applied on one side boundary with a pressure profile from the external model on the opposite side face, while the top face is set to atmospheric pressure to allow air to exit or enter the domain, as required. Due to the repetitive structure of the long potroom buildings, a slice model encompassing the two halves of the neighboring reduction cells with symmetry boundaries on the front and back faces of the potroom has been used. Plant anemometer data measuring the velocity at the roof vent has been used to validate the CFD models.

- S de and top faces: P= 0

Back Face: Söt pressure based On **~~~jp

external jyiQ^jgj

^^"^^^^|

jii

Front face: Velocity

1 inlet based on reailt

of external model

Front arid back faces through potroom: Periodic boundaries

Figure 4. Boundary conditions for the internal model geometry with the domain configured for a wind approaching parallel to the potroom. The flow inside the potroom is driven by the heat released from the pots. Therefore, the model must account for the energy transport. The flow is treated as a turbulent, non-isothermal, incompressible ideal gas. Hydrogen fluoride (HF) emissions are specified at the pot as a separate species with properties of HF. The air and HF species are free to mix and disperse together in the model.

533

Alcoa - CFD Modeling of the Proposed Potline Ventilation Design (2010)

As shown in Figure 5, emissions of HF were specified from the pot cover and anode butt surfaces. Heat release surfaces were the bottom and sidewalls of the pots as well as the superstructure with the total heat losses calculated using thermo-electric analysis. HF emissions ^Decified heat flux on specified on * ^ ^ the superstructure pot cover surface.,

Removable Basement Panels

The purpose of the study commissioned by Alcoa was to determine if a new potline, given its close proximity to the neighbouring casthouse and maintenance buildings, would cause a fatal flaw with respect to meeting the pot and bus cooling, and industrial hygiene criteria. Alcoa smelter discussed in this study is located in Baie Comeau, Quebec. An external CFD model of the smelter was first run to determine the ambient air pressures and temperatures adjacent to the new potline. It provided pressure and temperature values at the potroom air intakes and outlet boundaries. An internal CFD model of a section of the potroom was run using the results from the external model simulation. Other boundary conditions were set based on known process conditions (pot heat release, emission rates, etc). The potroom slice model predicted the ventilation flow rate through the potroom, the heat stress and emissions level in the worker area, pot and busbar temperatures.

HF emissions specified on ode butt surface, if applicable.

Based on the climatological data for Baie Comeau from 2003 to 2009 and corresponding wind rose for each season, several wind conditions were selected to use as inputs to the external model. The simulation results were analyzed to identify the pressure/temperature conditions at the potroom air intakes (basement panels) and outlet (roof monitor) that could produce the worst ventilation performance. The selected pressure/temperature conditions were applied at the two air ingresses and a roof ventilation outlet as shown in Figure 7. The internal model included a slice of the potroom from the basement panels up to the roof gravity ventilator.

Figure 5. Boundary conditions on internal surfaces of the potroom slice model. Additional sub-models have been run to calculate the loss coefficient for the basement grating, opening in the gas side and tapping side Claustra wall. They were treated as porous jumps in the potroom slice model, as theflowpasses through these surfaces with specified loss coefficients. Model predictions for each scenario have been validated using the anemometer data collected at the roof vents. Summary of the results for all cases is presented in Figure 6. Although the CFD model generally under-predicted the roof velocity, as compared to the anemometer data, it was able to successfully capture the trend in changes of the roof velocity due to variations in wind speed and direction.

Figure 7. Boundary conditions for the internal model geometry. Dec. 27 (B)

May 24 (A)

May 24 (B)

Jun. 10 (A)

Jun. 10 (B)

Due to the small scale of the gratings, the internal CFD model did not include their actual geometry; however, in order to maintain the accuracy of the results, detailed pot geometry reflected the exact open grating area. The pressure drop caused by the gratings was accounted through the use of an appropriate loss coefficient value. A separate CFD sub-model, which included complete geometry of the gratings around the pot, was created to determine the loss coefficient. Similarly, a loss coefficient was calculated for the Claustra walls and implemented in the potroom slice model.

Figure 6. Comparison between predicted roof vent throat velocity and the recorded plant anemometer data. The simulation results showed that, using the outlined methodology with reliable boundary conditions, CFD models are capable of capturing the effects of wind direction of the ventilation flow and gas velocity at the roof vents.

534

whether potroom conditions for each case met the criteria for the heat stress and contaminant threshold limits. Predicted values of wet bulb globe temperature (WBGT) in the working zone have been compared against set heat stress levels to ensure that both temperature and humidity levels were acceptable for the smelter personnel. The working zone was identified as the area up to 2 m above the operating floor between the pots and in the traffic aisle in front of the pot for tapping and pedestrian movement. Additionally, HF, S0 2 and dust concentrations throughout the entire potroom have been evaluated to guarantee that the predicted range does not exceed specified threshold levels.

The air flow within the potroom was treated as an ideal, incompressible, non-isothermal, turbulent gas with natural ventilation occurring due to thermal buoyancy. For the winter operation of the potroom, all 6 basement panels, shown in Figure 8, were kept closed. For the summer operating scenario, all 6 basement panels were open, thus the corresponding walls were removed from the model setup. It is important to establish the exact area of air inlets as the reduction in ventilation air flow in winter has a significant effect on the amount of heat removed from the pot surfaces and results in an increase of the concentration of HF, SO2 and dust in the work zone.

Completed CFD modeling work provided a comprehensive analysis of the proposed ventilation system for multiple operating scenarios. Various combinations of wind direction, speed and ambient temperature have been used to assess performance of the ventilation system and ensure that the efficient pot and bus cooling is observed while meeting industrial hygiene criteria. Conclusions CFD isfrequentlyused to assess current operational conditions at aluminum smelters and furthermore predicts the effects of the proposed geometrical or process changes. Due to the complex interaction between known boundary conditions and performance of the ventilation system inside the potroom, it is critical to establish the correct methodology for CFD analysis. The presented case studies outline the factors that need to be taken into consideration in order to ensure proper modeling predictions are obtained. The approach has been validated using measured site data, which allows for the potroom models to be used with confidence to predict air flow in the potroom, while taking into account outdoor ambient conditions and plant terrain unique for each smelter.

Figure 8. Internal potroom geometry for summer and winter configurations.

Acknowledgements

Sources of heat loss present in the model are the busbar, superstructure, side and bottom of the pot. The mass flow rate of HF, S0 2 and dust particulates emitted into the potroom were specified over surfaces within the internal CFD model. Emission points, as displayed in Figure 9, include: (1) the end gap between the cover and side wall; (2) collar around the anode stems; (3) the gap above the doors; (4) the openings around the anode stems in the pen butt container. The distribution of HF and S0 2 were calculated as a result of ventilation air flows within the potroom.

The authors would like to acknowledge the support and input from the staff of Environment Canada, Aluminerie Alouette Inc. and Alcoa. Daniel Richard is gratefully acknowledged for the valuable inputs and comments. References [1] A. Van Maarschalkerwaard, "The Use of CFD Simulations to Optimize Ventilation of Potrooms", Light Metals 2010. [2] J. Berkoe, et al, "CFD Modeling of the Fjardaäl Smelter Potroom Ventilation", Light Metals 2005. 373-378. [3] J. Woloshyn, "CFD Modeling of Potroom Roof Ventilation - Impact of Wind", Hatch report 334517-RPT-CA0110002 for Environment Canada, October 2010.

Figure 9. Location of modeled emission points on the pot structure. A total of 6 cases were modeled to encompass a variety of wind conditions in both summer and winter seasons. Each case predicted air flow throughout the potroom as well as concentration of all emissions at various heights above the operating floor. CFD modeling results were analyzed to determine

535

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

A PRELIMINARY FINITE ELEMENT ELECTROCHEMICAL MODEL FOR MODELLING IONIC SPECIES TRANSPORT IN THE CATHODE BLOCK OF A HALL-HEROULT CELL Frederick Gagnon1, Donald Ziegler2 and Mario Fafard1 ^SERC/Alcoa Industrial Research Chair MACE3 and Aluminium Research Centre-REGAL, Laval University, Quebec City, Canada 2 Alcoa Primary Metals , Laval University, Quebec City, Canada Keywords: cryolite, aluminium, cathode, transport, ionic equilibrium, electroneutrality, finite element modelling current efficiency, the need for modelling metal species dissolution could lead to improvement in the cell operating procedure. In the same view: a mushy layer, sludge or muck behaviour model will improve cell operation. All three of these physico-chemical phenomena, in some physico-chemical conditions, are kinetically controlled[l,2,3,4].Consequently, there is a need for a model that could include a kinetic reaction formulation.

Abstract A first 2D transient isothermal model was developed for the modelling of bath penetration in a Hall-Heroult cell cathode block. The model simulates the 'non convective' ionic species transport. The molten cryolite solution system is defined with an ionic model. The migration and diffusion of the ionic species is followed in the cathode block pores and in an overlying bath layer. The evolution of the potential is simultaneously obtained using charge conservation equations. The model includes a linear kinetic model for the electrochemical formation of metallic aluminium. The porous cathode behaviour is modelled using volume averaging based methods. The aluminofluoride ionic equilibrium is implemented by the penalty method and the electroneutrality criterion as a constraint on the FEM weak form. This is the first application of porous electrode theory, adapted for a molten salt, to Hall-Heroult cell cathodes. The results show the early stage evolution of voltage and ion distributions.

Most models [1,5,6] related to bath penetration are not very complex, being steady state, non-ionic and mostly empirical. This kind of model has a very limited range of application and is not useful for detailed studies of bath component behaviour in the cathode block. The main problem of non-ionic, or diffusion based, models is the impossibility of modeling some reactions related to a particular ionic species and thus it is difficult to insert kinetic formulations for chemical or electrochemical reaction. To avoid this limitation, then the model is developed within an ionic frame. In our case, there is also a need to avoid removing one of the species conservation equations by using the electroneutrality criteria or thermodynamic equilibrium equations. The electroneutrality criterion was implanted as a constraint and the ionic equilibrium by the FEM penalty method.

Introduction Presently, the most important research areas in aluminium industry are improvement of energy efficiency and minimisation of waste. An important way to minimize waste is to extend electrolysis cell life. Most cells breakdown due to cathode failure or related problems. Because of the high temperature, electromagnetic environment and highly corrosive properties of the cryolite melt, it is extremely difficult to assess the cathode block state using experimental methods during cell operation. Consequently, it is very difficult to correct operating parameters to avoid too fast degradation of the cathode bloc. Numerical simulation used with more indirect measurements can help to assess cathode block problems. Developing a cathode block model can also help to optimise cells design or materials properties. In this study, a preliminary model is developed for modelling early life of cathode block for start up operations of new cells. The model is a base for developing more complex models for following bath penetration in cathode block and related degradation reactions.

Developing this kind of model is difficult due to the lack of data for the bath properties over CR 3 and the scarcity of studies of carbon/cryolite interfaces under cathodic polarisation[l,2]. Reaction system The reaction system modeled is NaF-AlF3-Al0. The model is mainly based on ionic species. The species modeled are Na+, F", A1F4", A1F6"3 and dissolved Al°. The metallic aluminium behaviour is extremely simplified. In the real system, the metallic sodium is always present as a second product from the following equilibrium: 3NaF + A 1 ^ 3 N a ° + A l F 3

(1)

There are disagreements related to the last reaction[2]. It is not clear if metallic sodium comes from a homogenous reaction or from a parallel reaction at the electroactive cathode. Other difficulties are related to the unclear state of the dissolved metallic

After electric or fuel preheating, the cell start up procedure differs slightly among companies[l]. The bath is poured onto the cathode blocks with the current on. The bath slowly begins to penetrate and after some time, the carbon surface can become sufficiently kinetically activated for metal formation, and eventually the first metal droplets will appear. However, in most practices, some metal is poured to form a stable metal layer before it can form by droplet agglomeration.

species (Al , Na ) that make the development of a conductivity model coupled to species conservation equations a very difficult task. Moreover, layers of the bath containing dissolved reduced metals seem to evolve in two zones [7]. Explicitly including metallic sodium is a possible follow up to the model presented in this paper.

This model was not only developed in view of modelling bath penetration but also for assessing some other related physicochemical process, in particular, modelling phenomena related to aluminium carbides formation, current efficiencies and stability of bath mushy layers. The modelling of aluminium carbide formation could lead to a cathode erosion model. In the case of

537

Electrochemical reaction The electrochemical reaction for Al° formation is a two steps process[2,8,9]. In this model, for simplification, we only wrote a global interface reaction: A1F +3e"

AI + 4 F

4

ionized as Na+, F , A1F4 and A1F6 3 ; this allows calculation of ionic molar fractions and subsequently the molar concentration and density[11]. The ionic molar concentration was used to calibrate the ionic species diffusion and mobility coefficients with experimental conductivity data. Following that, the activity data for different CR were defined as a new function of Na+ concentration. Those functions and equation 8 were used in the penalty formulation of the modeled ionic equilibrium modeled. The system is considered to be isothermal.

(2)

At this step, it was judged too complex to include interface modelling with an extra equation to account for any intermediate species that could be either diffusing outside or adsorbed onto the interface. The dissolution of Al° was considered instantaneous and unaffected by its activity as a dissolved species in the bath. Other than the evaluation of density by molar concentration and molar mass, we did not include any effect of Al° dissolution in the bath, implying:

Electrochemical reactor The model is built with a large stagnant zone overlaying the cathode block because it preserves the possibility of modelling an experimental system of metal dissolution and migration. This could subsequently allow comparison of more complex model results with the two-layer dissolved metal zones mentioned earlier[7]. This configuration will also able us to add different effects of turbulent fluid movement.The last point is important because in the industrial conditions, there is a turbulent zone over the cathode block. The cathode below the bath saturation front is not modeled, nor is the advance of this front into the block.

a) Transport properties of other species are unaffected b) The conductivity is not corrected c) Real molar volume can be different; d) Effect on electrochemical kinetic parameters is neglected; e) Modelling of metal nucleation is excluded. Ionic equilibrium The ionic equilibrium modeled is the following: A1F

:A1F

+2F

4

Mathematical models The model is mainly based on species and charges conservation equations. As mentioned earlier, the domain is split into a fluid only zone and a zone representing the bath saturated porous cathode block. The equations describing the transport in the block are based on a volume averaged model, in which the properties of the carbon and the bath are described in terms of a small number of structure related parameters, such as the porosity, å, and the intrinsic average species concentrations c,. The carbon phase is considered non reactive and the porosity independent of time. The general species conservation equations for the species in pore fluid are define by the following equation 8c S — + VJ-sR-r =0 (9)

(3) +1

The only cation included in the system isNa . The ionic equilibrium in molten salt was defined by using the result from the model of Zhang et al[10]. They used the hypothesis that the activity for NaAlF and Na3AlF are equal to their molar fraction and the following equilibrium were formulated: A1F' + NaF ^ NaAlF 3

(4)

4

A1F + 3 NaF ^ Na A1F

(5) 3 3 6 The associated equilibrium constants are then formulated as following: K

A&4ffr

=

NaAIFA

,ΑÏ?

Ä**

^ ^ F ,

NojAB^

dt

(6)

3

6

4

KN

K =-

"

(7)

The electroneutrality constraint is supposed to be valid over the entire domain. S^z.c. = 0 (10)

(8)

„Na3AlF6 Na3AlF6

'

velocity. The species sources R( and r. account respectively for homogenous and heterogeneous reactions. All convective effects are neglected. The flux is averaged on the cross section of the homogenisation volume. The effect of the porous media on transport is modeled with effective transport coefficients. This is discussed later in the transport model section. The species concentration is averaged over the pore fluid and enables us to have continuous species concentration variables between the overlaying fluid domain and the cathode pore fluid.

Using complementary equations and the measured activity for NaF, A1F and other data, Zhang et al.[10] were able to evaluate the equilibrium constants. Using the same hypothesis for the activity of NaAlF and Na3AlF6we define the following equilibrium for the ionic equilibrium of interest in our model: Na A1F ^ NaAlF + 2 NaF

'

The superficial flux of species Ji is referred to the barycentric

"Na3AlF6

As mentioned earlier, the model has to be defined for ionic species concentration. Then we had to transform the system in molar concentration. The molar fractions were calculated from data on activity and equilibrium constants obtained by Zhang et al.[10] The effect of neutral species (A1F , NaF) and of alumina on the equilibrium constant were neglected. The components' molarfractionswere calculated from the equilibrium constant and activities. The components were assumed to be completely

As the electroneutrality criterion is considered valid and we don't take account of convective transport, there is non convective current. The electroneutrality is not implemented in this model in the conventional way by removing one specie's conservation equation; this is discussed more thoroughly in a later section. The charge conservation equations for the porous media are the following: V. / = r

538

V · / = -r

(Π)

The potential Φ 5 and Φ, are respectively those of the carbon solid phase and liquid bath phase. The open-circuit voltage U is defined as if using a reference electrode of the same kind [12] as our cathode electrode for assessing the voltage. The reference electrode bath for U is CR 1 and saturated with alumina. Using the overvoltage definition, we define the interphase electrochemical reaction charge term with a simple linear form:

The superficial current densities is and it are respectively for the graphitised carbon phase and the bath in the pores. As the capacitive effect is neglected, the charge source term r only includes interphase electrochemical reactions. These general equations are the same as the ones generally used in simple porous model[12,13]. The equations for species and charge conservation for the overlaying bath are almost the same as the ones for pore bath. In fact, the sole differences are that we remove the interphase electrochemical reaction source r and instead of using effective diffusion and mobility coefficients in the transport equation, we use standard coefficients.

r = inA

The transport parameters D. and ui are respectively the diffusion coefficient and mobility. The Nersnt-Equation is considered valid [2,7]: Di = RTüi (13) The species flux for the overlying bath is defined identically, except that transport coefficients are uncorrected for the porous media effect. In the block, a simple porosity effect model is used [12]:

( = FXZJ

(19)

By substitution of flux definition in the last equation and using equation 11, we obtain the full charge conservation equation for bath in the cathode pores:

Transport coefficient calibration The ionic species diffusion coefficients for the bath were calibrated from experimental conductivity measurements. The classic ohm law model for ionic solution was used:

V-f F j J - ^ V c , -zfqFuS/Ö,] 1=**(Φ, -Φ,) (20) The same equation for the bath overlying the cathode is easily obtained by removing interphase electrochemical reaction terms and using standard transport coefficients. The carbon phase is considered to only transport electronic current by ohm's law.

(15)

æ = "*, ν φ ,

Migration coefficients were obtained using the Nernst-Einstein relation. Bath conductivity êι data necessary for the calibration over a wide range of bath molar ratio (CR) without alumina were obtained from an equation calibrated with experimental conductivity measurements collected from literature [14]. The transport properties of dissolved aluminium come from Haarberg etal.[4].

(21)

The full charge conservation equation for the carbon is easily obtained with left equation 11: V · ( - F 5 ν Φ 5 ) = -kelc (Φ 5 -Ö,-U)

(22)

The parameter Jcs is the effective conductivity of the carbon phase.

Electrochemical reaction implementation The interphase carbon/bath surface is considered already kinetically activated for A1F~ ion discharge as if it were a carbon surface soaked in bath for some significant time [3] or a molten aluminium surface. For now, this hypothesis is used because almost no kinetic electrochemical studies have been done for the system modeled. Most experimental data are for an aluminium surface. The activation overvoltage η was defined as follows:

ö-öι-õ

(17)

r =± — r (18) nF The stoichiometric coefficient sf is for the species "i" in the electrochemical reaction. As discussed earlier, the aluminium dissolves instantaneously and most effects related to metal dissolution are neglected. Current density The superficial current density in cathode pore bath is defined from the species flux definition as follow:

Dt = D/5 (14) The porosity is defined as constant excepted in a thin layer near the interface between the porous media and free bath. Across this thin layer, the porosity is smoothed to 1 with a polynomial near step function. Normally, for the thin layer, a non zero gradient of porosity should appear in equation 9 but we choose to neglect it for this preliminary model.

η=

-Ö-U)

The parameters / 0 , A and n are respectively the exchange current density, specific surface and the number of electrons exchanged in the related electrochemical reaction. The choice of a simple linear model is justified by the fact that the electrochemical reaction is considered to have a small activation overvoltage when occurring on the liquid aluminium surface or on a carbon phase soaked in bath [2,3,8,9]. The electrochemical reaction term for the species conservation equation is defined as follow:

Transport properties flux model and porosity effect The species flux in the bath in the cathode pores is driven by migration and diffusion, and is defined by the following equation: J. = -DNc. - ziciFuiVOl (12)

2^ ztF ct TRT

(Φ

Ionic equilibrium-reaction implantation The ionic equilibrium is implanted with the FEM penalty method: R = ±S.kw

K —

(23)

The stoichiometric coefficient St is for the species "i" in the ionic equilibria. This approach is analogous to a linear reactiontransport model or a kinetic term based on a non-linear

(16)

539

thermodynamic force. As the ionic equilibria is supposed to be fast relative to transport properties and theAlF~ion discharge

is 1283 K. All other calibrated parameters or literature sources for them can be found in Gagnon [14].

reaction, the penalty parameter kloc must be high enough. The

FEM. Initial and Boundary Conditions

activity data for NaF, NaAlF and Na3AlF6 were redefined as function of Na+ molar concentration obtained as discussed earlier. Each activity function used was defined as a series of radial functions[15,16] of the following form:

■^t^t-cNJ^)\f

a8

Ka+y

The radial function parameters( b% , d^ , e , / . .

^.

Λ

f

NaF

NaAlF4

Na3AlF6.

...

(24)

) for each AΛ

.

activity function (a ,a \a ) were calibrated using Na+molar concentration data and a in-house program in Matlab[14]. The three activity functions were calibrated together to minimize the error on the equilibrium constant K. Simultaneous calibration helped to avoid oscillation between calibration data and give a betterfitfor K. Electroneutrality constraint As mentioned earlier, the electroneutrality equation was used to simplify some equations terms, such as the convective current, but not for removing one species conservation equation. The studied system is not very well suited for removing a species because of the many reactions that can be related to those species. Also, the way that ionic equilibrium is dealt with does not really make the species concentration available for such an equation elimination procedure. The electroneutrality criterion is instead applied as a constraint in the FEM formulation. A constraint can be applied to a PDE system solved by FEM using a variational formulation [14]. The applied constraint is the following:

ÌÓc,.z, ) dV = 0

V

\ '

The constraint is inserted following weak form:

J

Λ

\

The cathode pores are supposed to be completely filled with bath of the same uniform composition of the overlaying bath. The system is initially perfectly electroneutral. The model is started with the bath already in ionic equilibrium and of CR 1.72. The initial bath voltage is of-0.31 V relative to aluminium reference electrode (CR 1 alumina saturated bath). The bath top boundary Dirichlet condition is the initial ions concentration for CR 1.72. and 1 for the Lagrangian variable. The bath top voltage Dirichlet boundary is set to -0.31V. The internal boundary between the cathode block and the overlaying bath is the continuity of all variables except for the graphite current which is isolated. A current of 0.4 A cm"2 is put as a bottom boundary condition on the graphite. Any other boundary is the zero flux of Neuman kind. For solving the model by FEM we used COMSOL. A boundary layer mesh is generated near the cathode block and overlaying bath interface. The PDE General Form applications mode was used for the model implementation. The weak form terms related to the electroneutrality constraint were inserted by using the weak form contribution features of COMSOL. Result and discussion General result Figure 1 shows the metal concentration as a function of depth. In a future application, the metal produced inside the pore could be linked to a kinetic relation describing aluminium carbide behaviour. The calculated concentration of aluminium metal is too high. At such a concentration, the nucleation of droplets or formation of aluminum carbide should already be started. The aluminium production reaction would then probably shift onto the metal droplet surface. The real aluminium metal concentration should also be lower because of the formation of metallic sodium as a co-product. The metallic sodium will dissolve and react following other kinetics. For a model more in agreement with reality, metallic sodium modelling will have to be included. Also, the solubility of aluminium metal at high CR is not well known. The study of metallic aluminium in molten cryolite almost always implies side reactions related to metallic sodium. Because of this metallic side reaction, the modelling and the experimental measurements of this kind of system have the same related difficulties.

(25)

the FEM formulation by the

SÄdV + Ó pLz,£c. dV = 0 (26)

J However, using this formulation unfortunately introduces a new equation and its associated Lagrangian variable λ that is well defined mathematically but not necessarily on the chemicalphysical point of view. We make the hypothesis that the formulation should be good for the averaged concentration inside the cathode pores. Solving the equations The system is constituted of the homogenised cathode block and the overlaying bath. The 2D domain for the system is of 0.1 m by a depth of 0.045 m. It is split in two relative to depth to give a cathode block of 0.015 m thick and 0.030 m of overlaying bath. The system is modeled for a temperature of 1300 K. All parameters or data used are for 1300 K excepted for dissolved aluminium transport properties and the exchange current density. The temperature for the dissolved aluminium transport coefficient^] is 1273 K and for the exchange current density[8,9]

As observed in figure 2, the CR is rising in the electrochemical reaction zone. The metallic sodium rate or fraction produced will increase in the vicinity of a high CR zone [1,2,6-9]. This fact justifies by itself the necessity to add the metallic sodium modelling for an accurate simulation of aluminium current efficiency. As the CR crosses 3, the bath melting point will rise; this is consistant with a mushy layer formation. The figure 3 shows that density changes substantially near the interface. The formation of a stagnant bath layer with a higher density could help stabilizing a mushy layer. It is also important to observe that in the real system, the diffusion-migration boundary layer thickness will be smaller because of turbulent mixing by thermal and magnetohydrodynamic forces. It is clear that the model needs to be improved by adding fluid dynamic modelling.

540

cathode block surface defects could be relevant in the reaction distribution and the aluminium carbide related erosion process. The overvoltage and bath voltage are shown in figure 5. The voltage across the bath phase and the overvoltage follows the normal trend that should be observed [1,2,8,9].

An interesting result will be to assess the mixing inside the porous block by density flow. It is not evident to predict if the density flow will develop or not inside the block and consequently a density layer could be relatively stable inside the cathode bloc. The last remark is true only if the initial bath is of CR lower than 3 because for higher CR, the density begins to lower again. Also, the effect of metal dissolution should be taken in account. The metallic sodium and aluminium dissolution behaviour vary with CR, The behaviour of the cathode and overlaying bath interface zone could also be significatively influenced by thermal Peltier effect.

Table I. Calculated activities at 200 s for two depths. Depth (m)

NaF

C

Na+

(mmol cm"3) 20.99 21.50

0.035 0.016

20*0

Porous Cathode 2060

a

a

0.081 0.086

Λ

1

2Ö2Ö

a

0.61 0.59

1 j

2040

NaAlF4 4

I960

y!

1920 1900

Figure 1. Calculated weight % of Al metal dissolved in the bath at 200 s.

0

j, «.005

h

i O.OI

j

..;.....

Bath

1

\

Ί

\ \ H .......: .... .J

-H;-"-^-f!~™ • r j

1940

(%)

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The ion distribution curves shown in figure 4 mostly follow what is predicted by a more simple model [6,9]. Very near the electrochemical reaction zone, the pattern can be explained by the simultaneous effect of the enforced ionic equilibrium, the electrochemical reaction used and ion transport. The table I shows the calculated actvities for two points. Activities are calculated from Na+ concentration using the three activity functions defined in the form of equation 24. Table I also shows the relative error for the non respect of the equilibrium constant. The free bath is mostly following Ohm's law. The layer where the charge transfer reaction occurs is very thin, about 2 mm. Thus, the height of

Penalitv method for ionic equilibrium performance The maximum relative error for not respecting the ionic equilibrium is 0.47%. The largest error is near the strong concentration and voltage gradients but is still within a very good range. The quality of the solution is also function of the size of the penalty constant used in the model. The penalty constant cannot be raised too high because it induces too much numerical instability. In a future model, adding ionic equilibrium related to alumina dissolution should be straight forward but the

541

optimisation of the penalties constants could be difficult because the ionisation kinetic differences between the different equilibria is not well known.

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The authors gratefully acknowledge the financial support provided by Alcoa Inc. and the Natural Sciences and Engineering Research Council of Canada. A part of the research presented in this paper was financed by the Fonds quebecois de la recherche sur la nature et les technologies by the intermediary of the Aluminium Research Centre - REGAL

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1. Morten Sorlie and Harald A. 0ye, Cathodes in Aluminium Electrolysis Tdedition (Düsseldorf: Aluminium-Verlag, 1994). 2. J. Thonstad et al., Aluminium Electrolysis: Fundamentals of the Hall-Haroult Process 3rd edition (Düsseldorf: Aluminium-Verlag, 2001). 3. E. Y. L. Sum and M. Skyllas-Kazacos, "Dissolution Study of Electrolytically Deposited Aluminium at a Graphite Electrode," Electrochimica Acta, 36 (5/6) (1991), 811-817 4. G.M. Haarberg and J. Thonstad, "Electrochemical properties of metal-molten salt mixtures," Journal ofApplied Electrochemistry, 19 (1989), 789-801. 5. Y. Sun et al. "3-D Modelling of Thermal and Sodium Expansion in Soderberg Aluminium Reduction Cells," TMS Light Metals 2004, 5S7-592. 6. A. Solheim, "Crystallization of Cryolite and Alumina at the Metal-Bath Interface in Aluminium Reduction Cells," TMS Light Metals 2002, 225-230. 7. J. Thonstad and R. Oblakoowski, "On the Migration of Dissolved Aluminium in NaF-AlF3-Al203 Melts," Electrochimica Acta, 25 (1980), 223-227. 8. J. Thonstad and S. Rolseth, "On the Cathodic Overvoltage on Aluminium in Cryolite-Alumina Melts-I," Electrochimica Acta, 23 (1978), 223-231. 9. J. Thonstad and S. Rolseth, "On the Cathodic Overvoltage on Aluminium in NaF- A1F3-A1203 Melts-II," Electrochimica Acta, 23 (1978), 233-241. 10. Y. Zhang and R. A. Rapp, "Modeling the Dependence of Alumina Solubility on Temperature and Melt Composition in Cryolite-Based Melts," Metallurgical and Materials Transaction B, 35(B) (2004), 509-515. 11. A. Solheim, "The Density of Molten NaF-LiF-AlF3-CaF2A1203 in Aluminium Electrolysis," Aluminium Transactions, 2 (1) (2000), 161-168. 12. John Newman and Karen E. Thomas-Alyea, Electrochemical Systems 3rd edition (Hoboken, N J: John Wiley & Sons, Inc., 2004). 13. J. A. Ochoa-Tapia and S. Whitaker, "Heat Transfer at the Boundary Between a Porous Medium and a Homogeneous Fluid," International Journal of Heat and Mass Transfer, 40 (11) (1997), 2691-2707. 14. F. Gagnon, "Modelisation par elements finis du transport electrochimique des especes ioniques dans une cuve Hall-Heroult "(Ph.D. Thesis, Laval University, 2011). 15. R. Franke, "Scattered Data Interpolation: Tests of Some Methods," Mathematics of Computation, 38 (157) (1982), 181200. 16. F. Gagnon, "Methode numerique de resolution de Γequation de diffusion par collocation de functions radiales"(M.Sc, Thesis, Laval University, 2006). 17. T. Bryk and I. Mryglod, "Nonhydrodynamic Collective Processes in Molten Salts: Theory and Ab Initio Simulations," Internationaljournal of Quantum Chemistry, 110 (2009), 38-45.

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0 0.005 0.01 0.015 0.02 0.025 0.03 0.035 0.04 0.045 Depth (m) Figure 5. Simulated bath voltage relative to aluminium electrode CR = 1 and charge transfer overvoltage between cathode graphite and pore bath. Electroneutralitv constraint performance The system remains mainly neutral over the entire domain. It appears that the error follows the element mesh size. This means that we could probably have a smaller charge excess by using smaller element and by the way improving the model solutions. The charge excess seems randomly distributed over the domain. The related residual electric field is calculated with a relative permittivity of 10000 by solving the following equation: V.(-e 0 e r V<J>)=Fl>,.z,. I

(27)

The maximum voltage error is 0.2 mV. Formally, the relative permittivity of an ionic conductor for non alternative current is almost infinite [17]. Conclusion The model developed was judged to give good preliminary results relative to the scarce availability of parameters for CR over 3. We were able to observe ionic distributions conforming to what most of the literature predicts from a qualitative point of view. The rising of the CR near the cathode block top was the most interesting result because experimental observations[8,9] suggest indirectly such behaviour. The use of the penalty method for implanting ionic equilibrium was successful. The implantation of electroneutrality as a constraint was also judged numerically efficient for our model. For better modelling of a real system, the model needs to be improved by adding metallic sodium species conservation equation and fluid modelling. The base model developed opens the way for modelling complex phenomena in cathode block pores of a Hall-Heroult cell.

542

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CFD MODELLING OF ALUMINA MIXING IN ALUMINIUM REDUCTION CELLS Yuqing Feng, Mark A. Cooksey, M. Philip Schwarz CSIRO, Box 312, Clayton South, VIC 3169, AUSTRALIA Keywords: computational fluid dynamics, aluminium reduction cell, alumina mixing effectively predict the spatial and temporal variation in alumina concentration in an aluminium reduction cell.

Abstract Aluminium reduction cells have been continuously improved to reduce energy consumption and increase metal production rates. One method of reducing energy consumption is to operate the cell at low anode-cathode distances (ACDs). Line currents have been increased to increase production rates, which often requires larger anodes to maintain an acceptable current density. These changes may have a significant impact on aspects of cell performance such as bath flow and alumina mixing.

In previous work by these authors [4 and references therein], the bath flow and alumina mixing in aluminium reduction cells has been studied using physical modelling and computational fluid dynamics (CFD) modelling of part cell and whole cell geometries. Afindingof the previous work was that, for the simple geometry considered, the lowest alumina concentrations are often found in the centre channel. Experimental work by Moxnes et al. [1] using large A1F3 additions in an industrial cell also suggests spatial variation in the cell. As a result, the alumina feed rate was increased at the two end feeders and decreased in the three central feeders to produce a more even alumina concentration, which suggests that the original situation was lower alumina concentrations at the ends of the cell. These modelling and experimental results suggested that it was worthwhile to investigate the spatial distribution of alumina concentration in the cell in more detail, including modelling the effect of modifying the relative feed rates from different feeders. This is complex, as Moxnes et al. were unable to reproduce the measured flow patterns.

In previous work, a computational fluid dynamics (CFD) model of bath hydrodynamics and alumina mixing in an aluminium reduction cell has been developed. The bath hydrodynamics has been validated using water model data. In the present work, the CFD model has been used to study the effect of alumina feeder location and feeding strategy in a typical full-scale industrial cell. Introduction Bath flow in aluminium reduction cells is important for several reasons. In particular, it dictates the distribution of chemistry and temperature within the cell. Control of alumina dissolution and distribution is important for conventional cells. It has been calculated that a 1 wt% increase in the alumina concentration can increase the metal height (and decrease the bath height) by ~5 mm, i.e. the ACD is reduced by ~5 mm [1]. Significant amperage increases often require a reduction in ACD and/or an increase in anode size, which reduces the amount of bath available to dissolve and distribute the alumina. Furthermore, better alumina control is possibly critical for the implementation of advanced cell designs. For example, it is likely that a cell using inert anodes will require tighter control of alumina concentration to prevent corrosion of the anode [2].

Model Description Bath flow in an aluminium reduction cell is a typical multiphase flow process, involving strong interaction between gas bubbles and liquid bath due to buoyancy effects, and between liquid bath and liquid metal due to strong electro-magnetic forces. Two types of CFD models have been developed at CSIRO to study bubble driven flow in the bath phase: a small-scale resolved bubble model and a macro-scale time-averaged cell model. The former approach tracks the interfaces around each of the bubbles using, for example, the Volume-of-Fluid (VOF) method, and detailed transient bubbling behaviour can be obtained. However, this model requires a very fine mesh that presents a major hurdle for current computing power. The time-averaged model represents the flowfieldaveraged over time and hence steady state equations are solved. The model also averages over small-scale phase structure (i.e. bubbles) using the so-called two-fluid or Eulerian-Eulerian approach, where gas and liquid are described as interpenetrating continua and equations for conservation of mass and momentum are solved separately for each phase. The model requires less computing power, but the detailed bubbling hydrodynamics cannot be obtained. The former model is suitable for fundamental studies, the latter for process simulation, and has been widely used in various multiphase flow systems, e.g. gas stirred baths [5, 6].

Bath flow can be driven by several mechanisms [3]: • • • •

Release of gasfrombeneath the anodes; Interaction of the magnetic and electric fields in the bath; Dragfrommagnetically induced movement of the metal pad; Thermal convection.

It is difficult to directly measure the effect of cell and anode design on bubble-driven bath flow, because of the corrosive nature of cryolite at -960 °C. Thus when considering an amperage increase, there is often no alternative to trialing the new operational parameters in a number of cells, which is timeconsuming and costly. Also, the opposite problem can occur: potential gains from a reduction in ACD and/or an increase in anode size are missed because a multi-cell trial is deemed too time-consuming and costly. Thus there is value in being able to

The time-averaged two-fluid modelling approach has been adopted for this study. The governing equations are Navier-Stokes equations in the form for multi-phase flow systems, facilitated with extra source terms to account for the inter-phase actions (e.g. the bubble drag force, bubble induced turbulence). The details of the modelling approach and validation using PIV measurement can be found elsewhere [7, 8].

543

surface. The standard dump weight for each feeder is 1.65 kg. The model assumes that the alumina dissolves instantaneously.

Model Parameters There have been very few published CFD studies of alumina mixing in an entire aluminium reduction cell because of the large amount of computer time required. However, to gain a realistic understanding of phenomena such as alumina mixing, it is important to simulate the whole cell. A full-scale cell with 18 anodes was chosen for this study: the geometry represents a typical pre-bake smelter, but is not related to any specific cell design. Figure 1 shows a plan view of the cell geometry. There are four alumina feeding positions, which are modeled in two different arrangements:

In this simulation the lowest alumina concentrations are found in the centre of the cell; therefore modified feed rates to produce a more even alumina concentration need to be higher in the centre of the cell and lower at the ends. In the experimental work of Moxnes et al. [1], the feed rates were adjusted significantly; up to 50% at one feeder. For this reason, two additional feed rates are modelled; ±10% and ±30%. A summary of the three feed rates is shown in Table 1. Since there are two feeder locations modeled, this gives a total of six simulations. Table 1: Alumina feed rates

1. Roughly equally spaced in the centre channel and in line with the mid-point of anodes ('mid-anode feeders') 2. Located at inter-anode gaps ('inter-anode feeders')

Feed Strategy Standard ±10% ±30%

The ACD is 40 mm, the bath depth is 200 mm and the gas flow rate is set to be equivalent to a current density of 0.9 A/cm2. A commercial CFD code (ANSYS-CFX12) [9] was used to obtain a numerical solution, where user-defined subroutines have been applied to account for the bubble induced dispersion force and turbulence.

It is important to note the following simplifications in comparison to an industrial cell: •

The alumina concentration is set to be 3.5 wt% throughout the cell at the start of the simulation. The alumina feeding strategy is: 1. 2. 3.

Dump weight (kg) Two centre feeders Two end feeders 1.65 1.65 1.82 1.48 2.14 1.16

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The anode base is flat and the anode edges/corners are square. This represents a new anode (without chamfers) but most anodes in a cell at any one time have a more rounded bottom profile; The side channel and end channel profiles are square and there is no ledge; The metal layer is not simulated in the model.

These effects can be important if considering the absolute liquid flows. As this study is comparative (i.e. examining the effect of feeder location and feed rate), the importance of the above effects is reduced.

Alumina is added as a source term in a region of 200 mm x 200 mm x 20 mm high, with the top of the region being at the bath

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544

the alumina in the centre channel can be depleted faster than in the side channel. The present results show that increasing the feed rate of the central feeders addresses the depletion of alumina in the central channel, but has little effect on the side channel.

Results The average alumina concentration in the ACD under each pair of anodes through a feeding cycle, for each case, is shown in Figure 3. Only five pairs of anodes are shown because the cell is symmetrical. The data is at one minute intervals.

This can be explained by the fact that the ACD is only 40 mm and the inter-anode gap is only 20 mm, so bath tends to preferentially flow through the centre channel (240 mm) and end channels (200 mm) to the side channel.

In all cases, the range of alumina concentrations is low during underfeeding, increases rapidly during overfeeding, and gradually converges during normal feeding. The pattern over multiple cycles was investigated in the CFD model, and it was found that the concentrations tended to converge at the end of overfeeding in each cycle, so the results presented here can be considered representative of repeating alumina feeding cycles.

The contour maps highlight that feeder location is extremely important. In both cases considered here (mid-anode feeders and inter-anode feeders), it appears that the feeders are too close to the ends of the cell. Significant changes to the relative feeding rates are unable to overcome the effect of feeder location.

The range of alumina feeding concentrations (the difference between the highest pair and lowest pair of anodes) at each one minute interval for each case is shown in Figure 4. Overall the range is greater for the inter-anode feeders than for the mid-anode feeders. For the inter-anode feeders, the maximum range is 1.0 wt% alumina for the standard feeding. This range decreases for the ±10% and ±30% feeding; down to 0.7 wt%. For the midanode feeders, a different trend is that the range is greatest for ±30% feeding.

The results suggest that there are two broad types of spatial variation in alumina concentration in the cell: 1. 2.

From Figure 3, it appears that the minimum and maximum alumina concentrations are generally under the centre and end anodes respectively. A notable exception is the mid-anode ±30% feeding case, where the highest concentrations are under the 6/15 pair of anodes.

Variation along the length of the cell, which can be influenced by feeder location and the relative feeding rates of different feeders; Variation from the centre channel to the side channel, which can be influenced by the ACD and inter-anode gap, and presumably anode slots.

The alumina concentration changes quite quickly in the centre channel and end channel following a feed, but it is slower to change in the ACD and side channels. This gives a guide to the relative speed of alumina processes in the cell, as shown in Figure 2.

As stated, the largest range of alumina concentrations tends to occur at the end of overfeeding; at approximately 30 min. Contour maps of the alumina concentration in the horizontal mid-plane of the ACD at this time are shown in Figure 5. These contour maps show that the spatial variation in the cell is even greater than is apparent from the numerical results discussed to date. There is not just variation from the central anodes to the end anodes, but also from the centre channel to the side channels.

Alumina Mixing via Centre Alumina Mixing via ACD Feeding and End Channels Consumption and Side Channels A Fastest Slowest Figure 2: Relative speed of alumina processes in cell.

It is striking that the lowest alumina concentrations are found towards the side channels for the central anodes. It appears that the alumina concentration can be 1 wt% higher under the inner half of the anode (towards the centre channel) compared to the outer half.

The results shown that the alumina concentration can vary by 1 wt% due to both types of spatial variation. Given that, as stated previously, a 1 wt% change in alumina concentration can affect the ACD by ~5 mm, the effect of this variation on cell performance may be significant.

Increasing the relative feeding rate in the two central feeders does little to increase the alumina concentration in this region. All it seems to do is increase the concentration in the centre channel.

Limitations The model assumes that alumina dissolves instantaneously, which is almost certainly a significant simplification, since sludging of cells is a known problem. Water modelling work by Walker et al. [11] found that the use of crushed ice was superior to a dye tracer in modelling the behavior of alumina in an industrial cell, probably because the crushed ice forms 'rafts' on the liquid surface, as does alumina on the bath surface.

Discussion The results illustrate the importance of the bath flow in the cell. Note that the bath flow is same for the six cases considered here; it is only the alumina feeding locations and feeding rates that are being modified. The bath flow has been reported previously [10], and the bath velocity vectors for this case are shown in Figure 6.

The possible risk of sludging would have to be considered when considering increasing feeding rates at some feeders.

In previous work [4] it was described that, between feeds, the centre, side and end channels serve as reservoirs of alumina as it is consumed in the ACD, and also that because the centre channel is effectively feeding more anodes than the side or end channels,

545

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Figure 6: Bath velocity vectors in horizontal mid-plane of ACD (40 mm ACD, 240 mm side channel width) [10]. Conclusions A time-averaged bubble driven CFD model, validated using water model data, has been extended to study bath flow and alumina mixing in a full scale aluminium reduction cell. The main findings of this study are: • •

[4] Y. Feng, M.A. Cooksey and M.P. Schwarz, CFD Modelling of Alumina Mixing in Aluminium Reduction Cells, Light Metals 2010, Seattle, WA, 14-18 Feb, 2010,455-460. [5] M.P. Schwarz, & W.J. Turner, Applicability of the standard k-ε turbulence model to gas-stirred baths, Applied Mathematical Modelling, 72(3), 1988, 273-279.

There can be ~ 1 wt% spatial variation in the alumina concentration, both along the length of the cell and from the centre channel to the side channels. Varying the relative feeding rates between different feeders can influence the variation in alumina concentration along the length of the cell. However, this has little influence on the variation between the centre channel and side channels, because this is primarily controlled by the ACD and inter-anode gaps.

[6] G.L. Lane, M.P. Schwarz, & G.M. Evans, Numerical modelling of gas-liquid flow in stirred tanks, Chemical Engineering Science, 60(8-9), 2005, 2203-2214. [7] Y.Q. Feng, W. Yang, M. Cooksey, & M.P. Schwarz, CFD model of bubble driven flow in aluminium reduction cells and validation using PIV measurement, Fifth International Conference on Computational Fluid Dynamics in the Process Industries, Melbourne, Australia, 2006.

These findings demonstrate the effectiveness of CFD modelling for studying the likely effect of various design parameters (e.g. alumina feeder location and feeding rate) on cell performance (e.g. bath flow and alumina mixing). However, the specific results of this modelling should not be over-interpreted, as many of the results are specific to the geometry being considered.

[8] Y.Q. Feng, M. Cooksey, & M.P. Schwarz, CFD modelling of electrolyte flow in aluminium reduction cells, Light Metals 2007, Orlando, FL, 2007,339-344. [9] ANSYS INC CFX12: 2009

References

[10] Y. Feng, Y., M.A. Cooksey & P. Schwarz, Development of Whole Cell Model of Bath Flow and Alumina Mixing, 9th Australasian Aluminium Smelting Technology Conference, 2007,14pp.

[1] B. Moxnes et al., Improved Cell Operation by Redistribution of the Alumina Feeding, Light Metals 2009, San Francisco, CA, 15-19 Feb, 2009,461-466. [2] J. Thonstad, & E. Olsen, Cell operation and metal purity challenges for the use of inert anodes, JOM, 53(5), 2001, 3638.

[11] M.L. Walker et al., Design Considerations for Selecting the Number of Point Feeders in Modern Reduction Cells, Light Metals 1995, Las Vegas, NV, 12-16 Feb, 1995,363-370.

[3] K. Grjotheim, & B. Welch, Aluminium smelter technology, 2ndedn (Düsseldorf, Germany: Aluminium-Verlag, 1988).

548

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

BUBBLE TRANSPORT BY ELECTRO-MAGNETOPHORETIC FORCES AT ANODE BOTTOM OF ALUMINIUM CELLS

Valdis Bojarevics and Alan Roy University of Greenwich, School of Computing and Mathematical Sciences, 30 Park Row, London, SE10 9LS, UK [email protected]

However a less recognized but significant contribution to the force balance on a moving bubble arises from the so called electromagnetophoretic forces which were first theoretically analyzed by Leenov and Kolin in 1954 [6]. They derived expressions for electrically conducting and nonconducting particles (or bubbles) in the presence of an electric current and magnetic field in the surrounding liquid. They suggested possible applications like separation of particles and opposing the gravity effects. Since then metallurgists were keenly exploiting the electro-magnetophoretic force for a variety of applications, such as removal of inclusions from steel melt [7], concentrating the insulating bubbles and inclusions [8], and for many other purposes [9]. It is important to note that not only the electrically conducting particles experience the effect of the electromagnetic force due to the current passing through them and interacting with the overall magnetic field. The electrically nonconducting particles in an electric current carrying fluid will experience significant force due to the pressure distribution gradient arising with the electromagnetic force action in the surrounding fluid. The physical effect is very similar to the buoyancy force created by the vertical pressure stratification in the fluid in the presence of gravity. With the electromagnetic force the pressure gradient can be created in any direction in the fluid carrying electric current.

Abstract Electrically conducting and nonconducting particles and bubbles experience additional forcing in a liquid which carries electric current. These so called electro-magnetophoretic forces are well known in metallurgical applications, like metal purification in vacuum-arc remelting, electro-slag processes, impurity removal or concentration change in special castings. However, the effect of electro-magnetophoretic forces has never been considered for aluminium cells where the gas bubbles evolving in the liquid electrolyte are surrounded by an electric current and significant magnetic fields. We present models to estimate the effect of electric current flow in the vicinity of the bubbles and the additional pressure distribution resulting from the magnetic forces in the surrounding liquid electrolyte. According to the estimates, this force becomes important for bubbles exceeding 2 mm in size, and could be sufficient to overcome the typical drag force associated with electrolyte flow thereby opposing motion of the bubble along the base of the anode when it is inclined at a slight angle. The effect could explain certain features of the anode effect onset. Mathematical models and numerical results are presented and a further implementation in the general MHD code for the aluminium cell design is discussed.

In this paper we attempt to derive a very simple mathematical model for a bubble of hemispherical shape attached at the bottom of the anode in an aluminium electrolysis cell. The mathematical expression for the electro-magnetophoretic force is used to estimate a possible effect on the bubble transport and compared to the viscous drag force from the integral large scale bubble driven and electromagnetically driven flow . By comparing the buoyancy force at the bottom of a sloped anode with the electro-magnetoforetic force it is possible to find conditions when a bubble of larger size could be trapped in a stationary position for some time. This observation can yield further insight to the anode effect mechanisms and the conditions triggering the onset of the anode effect [10].

Introduction The presence of gas bubbles is an inherent feature of the electrolytic aluminium production cells. Typically C0 2 gas bubbles are produced at a rate proportional to the electric current magnitude yielding approximately 2.5 m3 of gas per kg of aluminium produced. A detailed description of the bubble creation, detachment and transport is given in many publications ([1-5] are recent examples). The detached bubbles during the aluminium reduction are typically of 3-5 mm in size [1] and can grow due to collision and coalescence. On extreme occasions their volume can reach 100 ml. The presence of bubbles contributes significantly to the overall voltage loss of an individual cell (about 0.25 Vfromthe total of 4 - 4.5 V) [2]. The usual assumption is that the bubbles are transported due to the buoyancy driven flow originating from the bubble escape into the side channels at the edges of the anodes [35]. The shape of the anode bottom is recognized as an important factor to facilitate the initial buoyancy force moving the bubbles [3]. In addition to this the electromagnetically driven flow exerts an additional drag force which contributes to the bubble transport along the flow streamlines, and could change the drag coefficient due to local nonuniformity of the electrical current around the bubble [9].

The inclusion of the electro-magnetophoretic effect with the bubble distribution models is a possible development for the wave model and the dynamic interaction with the electromagnetic field as implemented in the MHD numerical code [11].

549

The setup for a mathematical model Before considering the local bubble model it is important to have a view to an integral picture of the whole aluminium cell. For this purpose we will use the results obtained from our previous coupled MHD models for the whole cell [11]. The interface between the liquid metal and electrolyte forms a nearly stationary dome like shape, as shown in Figure 1 for the case of a 500 kA cell. The interface shape can be variable if the cell is close to the stability limit when the waves start appearing. For a stable cell the time average deformation of the aluminium-bath interface causes redistribution of the electric current, which accelerates consumption of the anode bottom in areas where the anode-cathode distance (ACD) is reduced. The time average shape of the anode block above the metal interface shown in Figure 1 will eventually assume a shape similar to that computed in Figure 2. Note that the bottom shape is shown with an exaggerated scaling in the z-axis direction (about 15 times).

Figure 3. The computed magnetic field in the electrolyte layer for the 500 kA cell. A simple model for a hemi-spherical bubble In order to obtain a quantitative description of the electromagnetophoretic force acting on a gas bubble attached or slowly moving along the anode bottom, let us consider an idealized situation shown in the Figure 4. A hemispherical shape bubble is positioned somewhere in the middle of a mildly sloped carbon anode,fromwhich the electric current flows uniformly downwards. The actual current distribution is obtained using the commercial 3d electromagnetics module in the package COMSOL.

Figure 4. The hemi-spherical bubble at the bottom of carbon anode with the computed electric current lines.

Figure 1. The interface shape of a 500 kA cell.

For the analytical derivation we can use the problem setup shown in the Figure 5. The electric current J = - J0zez has only the vertical component of constant value 70z· The magnetic field is assumed to have only the Bx component given as B = B0x y/Ly e x , where B0x is the magnitude of the field at the external edge of the anode whose width is Ly. The Lorentz force acting in the fluid (electrolyte) has only a single y-directed component given by

F y =-*Ä.

(1)

Figure 2. The view to the bottom of all anode blocks (without divisions between individual anodes) as computed for the electric current adjustment to the constant ACD condition.

11L 111., 11 i

The electric current distribution in the electrolyte layer below the anodes is approximately uniform at the density 0.7 A/cm2. The magnetic field computed for the 500 kA cell is shown in Figure 3. From the direction of the arrows and the colour flooding of the contours it is evident that the Bx component (along the long side of the cell) is the dominant contribution to the total magnetic field. This could be generalized for various other cells we have considered modelling. The Bx field is unavoidably present, it is the dominant component, and it can not be affected significantly by rearranging the cell bus bar network supplying the electric current. The Bxfieldis mainly produced by the vertical current in the whole cell and the horizontal currents in the liquid metal and the cathode collector bars. Typically the Bxfieldhas a nearly linear gradient in the y-axis direction, increasing in magnitude away from the cell central axis.

< -

<

ψ

y

y

,111

z

/////Sf,?.^ y

Z

i

Figure 5. Setup for the analytical model with the uniform electric current Jz and linearly growing magneticfieldBx.

550

Π I I I I1 I 1i i 1I1 l l l l l l l

Then by replacing y' = rb sinÈsirup in local spherical coordinates after the transformation, we can express the y-component of the force as:

•»Oz;

dy

F p -e y =-JJ/idS R -e y = = - Ι Ú prb2 sin2 0sin öάθάö.

The expression (6) for p{yr) contains even and odd terms in powers of y' = rsinOsinip, of which only the odd powers (linear in this case) will contribute to the total integral over the symmetrical bubble surface. This means that the total force in the y-direction is

Figure 6. The model for obtaining the pressure p(y) - created by the uniform electric current Jz and linearly growing magnetic field Bx. The electromagnetic force (1) acting in the fluid will create a pressure p(y) distribution according to the boundary conditions chosen to represent the situation under the anodes. Figure 6 is representative of one half of the cell with the symmetry axis at y = 0, where the appropriate symmetry condition of zero normal derivative is imposed. At the left edge of the electrode we assume a constant hydrostatic pressure under the given depth of the electrolyte. Since mathematically the pressure is defined to the accuracy of an arbitrary constant, we can choose p(y=Ly) =0. To find the pressure distribution we need to solve the hydrostatic equation

V-(Vp) = V - ( / x H ) ,

B

0xJ0z

_Vφz

à

a« dy

dy

(2)

2πη

n

0xJ0z

yb j

Fp=-(JxB)

y

0;Λæ

^ΟΑΛÆ

2L

y

2 , * p

2

0x

j

(3)

j

0 z r

(4)

b'>

(9)

v

v

f

Let us estimate the relative values of the electro-magnetophoretic force and the drag force acting on a bubble in a large scale flow driven by either the total effect of bubbles escaping at anode edges (bubble driven flow) or the electromagnetic force-driven large scale horizontal flow. Suppose that, for simplicity, the bubble is of spherical shape. The drag force on a small spherical bubble is given by Stokes formula, modified for the slip boundary condition:

(5)

where the minus sign appears because the normal vector to the fluid facing the bubble is opposite to the eR direction in a spherical coordinate system with the origin at the bubble centre. To compute the integral (5) we first shift the origin to the bubble centre by transformation y =y-yb'

p=-^(y+yb)2+\ίoJozLy

V

Returning to the electrolysis cell situation, it is instructive to analyze the expression (8). We can see that the force increases linearly with the bubble distance ybfromthe cell centre. The force grows as the third power of the size of the bubble. It will increase proportionally to the magnitudes of the current density Jz and the maximum magneticfieldB0x.

Suppose that the centre of the bubble is at y = yb . The total force acting on the bubble from the fluid is the surface integral over the hemi-spherical surface:

F P =-JJ/*S R ,

(8)

The expression (9) shows that there will be a force acting on a nonconducting bubble which is equal to that computed over the fluid volume in the absence of the bubble, but with the minus sign. Some authors [8] introduce form factors to represent different shapes of particles or bubbles compared to the spherical volume expression. The expression in (9), however, is more general.

The solution of (3) with the given boundary conditions is

P

π/2 0 3 b _

where Vb is the volume of the hemi-spherical bubble. It is easy to make more general conclusions, remarking that the expression (8) coincides with that derived by Leenov-Kolin in the case of uniform current and uniform field, in fact,

which implies that

_

yb-

π 2π

yb ί ί rb3 sin3 0sin 2 öάθάö =

fluid

2

(7)

Vd=-4Mjrbv9

(10)

where η is fluid viscosity. When the electro-magnetophoretic force is equal in magnitude to the drag force, the bubble is trapped in the flow or deviates from thefluidflow line direction (if the forces are not aligned). The condition for a spherical bubble can be expressed by multiplying expression (8) by a factor of 2:

(6)

551

A

*-d

distribution. The original derivation by Leenov-Kolin [6] assumed an infinite volume of fluid, and the authors proved analytically that the arising viscous flow due to the pressure distribution does not affect the resultant force. We have set up an idealized fluid dynamic model using COMSOL to investigate possible flow effects on the resulting pressure distribution. The fluid flow field shown in Fig. 8 was computed using the boundary conditions for pressure in Figure 6 and the previous distribution of electric current and magnetic field which yielded the force distribution (1) in the liquid electrolyte. This flow has no similarity to a real aluminium cell where multiple anodes are present and the flow results from the integral effect of all electromagnetic interactions. The purpose of this simple model is to compute possible variation in the pressure field with velocity present.

/?>

i i B.J,

Ami

(ID

Assuming typical values for the electrolyte circulation velocity Ivl equal to 0.1 m/s, fluid viscosity 0.002 kg/(s.m), the maximum magnetic field B0x = 0.02 T, the electric current density / = 7000 A/m2, and the position of the bubble close to the edge yb = Ly, the minimum possible size for a bubble, stopped by the action of the electro-magnetophoretic force opposing the dragfromthe flow, is

rb=2.01l0~3

(m) « 2 mm.

(12)

This estimate is well within the range of typical bubble sizes ( 3 - 5 mm) after detachment in the electrolysis process [1]. For a 4 mm bubble the magnetic pressure force will be 4 times larger than the drag force under the assumed conditions, and the bubble will start moving against the incoming flow.

Figure 8. The computed pressure and the flow driven from the side channel along the anode bottom to the central channel of the cell. Figure 7. Schematic representation of a possible force balance for a bubble positioned at the bottom of a mildly sloped anode.

0.165

Figure 7 indicates another situation when the bubble at the bottom of a mildly sloped anode is trapped in a stationary position due to the balance of the buoyancy force Fg and the magnetic pressure force Fp:

F =F P

0.164

0,163 ft'

(13)

0.162

where peg sin a is the effective buoyancy force driving the bubble upwards along a gently inclined slope of angle a to the horizontal. A simple calculation, using similar data as in the previous example (assuming an electrolyte density of 2100 kg/m3) would suggest that when the anode bottom is sloped at about 0.45 degrees the bubble motion could be stopped. This is a very mild slope, but nevertheless quite realistic in operating conditions.

0.161

0.16

Figure 9. The computed pressure distribution on the surface of the hemi-spherical bubble whenfluidflowis present..

Numerical results for the local hydrodynamics In this section we investigate the effect of the flow which could arise due to the pressure distribution obtained from the expression (4), which can affect the validity of the pressure

Figure 8 demonstrates the pressure distribution in the whole bottom space, showing a similar pressure gradient to that obtained

552

analytically (4). Locally at the hemi-spherical bubble surface the computed pressure distribution is shown in the Figure 9. Figure 10 presents the computed pressure as a function of position along a line in the electrolyte just below the bubble (solid blue line); the dotted black line represents a fitted quadratic function. The fitted quadratic function has the equation

p = -35.597y 2 +14.752y+ 0.1694,

0.155, < >

(14) Figure 11. Schematic 3D geometry for the anode and the electrolyte.

Encouragingly the pressure is well-described by a quadratic function. We may compare the coefficient of the quadratic term with the analytical expression derived earlier (4), substituting B0x = 0.02 T:

(

p=-

o.ouOz

Λ y2+0.0V0zLy.

_ , <*=S.99xl0' S/m ''copper anode

(15)

The 3D model has end effects associated with the flow which would explain why the fitted curve has a linear term in y which is not present in the analytical expression. Additionally, since only the gradient of pressure drives the flow there is a constant offset associated with the gauge. If we assume that the current density in the 3D model is j = 7. JQ3 A/m2 » t n e n m e analytical coefficient of the quadratic term approximately matches the value in (14), which would suggest that the 3D simulation is producing numbers of the correct order of magnitude.

Figure 12. Streamlines of electric current around a large bubble (r=20 cm) attached at the bottom of the anode.

End effects associated

Initially the current was computed in the absence of any bubbles at the bottom. Fig. 13 (blue curve) shows the percentage of the total current exiting the anode on the side surfaces as a function of the ACD distance. Also shown in Fig. 13 are, for comparison, results for the case when a large bubble of radius 20 cm is present. The local distribution of electric streamlines around the bubble is shown in Figure 12. The results indicate that the presence of the bubble adjusts the effective ACD distance and makes the current path through the side faces of the anode more favorable. The nonlinear resistance due to the electrochemical voltage drop was not accounted for as this is not the subject of the present study.

with 3D mode!

Figure 10. The computed pressure distribution at the mid-plane of the electrolyte passing through the position of the bubble as computed for the flow in Figure 8.

Numerical results for the electric current distribution It is of certain interest to compare the electric current distribution on the anode when a bubble or cluster of bubbles are attached to the bottom of the anode. For this purpose we used the simple 3D model shown schematically in Figure 11.

0.01

0.02

0.03

0.04

z-depth electrolyte (cm)

Figure 13. The computed ACD dependence of the electric current percentage exiting the side wall when the large bubble is present compared to the situation without the bubble.

553

Conclusions Analysis of the electro-magnetophoretic forces acting on bubbles in the aluminium reduction cells suggests that their presence could significantly affect bubble transport, concentration and detachment. The models presented give numerical estimates of the effect of electric current flow in the vicinity of the bubbles including the additional pressure distribution resulting from the magnetic forces in the surrounding liquid electrolyte. According to these estimates, this force becomes important for bubbles exceeding 2 mm in size. The force is sufficient to overcome the typical drag force due to electrolyte flow, and could potentially prevent translational displacement of the bubble along the base of the anode when it is inclined at a gentle gradient. The effect could explain certain features of the anode effect onset. A further implementation in the general MHD code for the aluminium cell design is considered for future work.

References 1. K. Vekony and L. I. Kiss, "Morphology of Two-Phase Layers with Large Bubbles". Metallurgical and Materials Transactions, 41B, 1006-1017 (2010). 2. M.A. Cooksey, M.P. Taylor and J.J. Chen, "Resistance due to gas bubbles in aluminium reduction cells". Journal of Metals, Feb. (2008), 51-57. 3. Z. Zoric, T. Thonstad and T. Haarberg, "The influence of the initial shape and position of an anode and the curvature of the aluminium on the current distribution in prebaked aluminium cells". Metallurgical and Materials Transactions, 30B, 341-348 (1999). 4. Y. Feng, M.A. Cooksey and M.P. Schwarz, "CFD modelling of alumina mixing in aluminium reduction cells". TMS Light Metals (2010), 455-460. 5. D.S. Severo, V. Gusberti, "A modelling approach to estimate bath and metal heat transfer coefficients". TMS Light Metals (2009), 557-562. 6. D. Leenov and A. Kolin, "Theory of electromagnetophoresis". Journ. Chem. Phys., 22, N 4, 683-688 (1954). 7. T. Toh, H. Yamamura, H. Kondo, M. Wakoh, Sh. Shimasaki and S. Taniguchi, "Kinetics Evaluation of Inclusions Removal During Levitation Melting of Steel in Cold Crucible", ISIJ International, 47, No 11, 1625-1632 (2007). 8. J.W. Haverkort and T.W.J. Peeters, "Magnetohydrodynamic effects on insulating bubbles and inclusions in the continuous casting of steel", Metallurgical and Materials Transactions, Online First (2010). 9. V. Bojarevics, J. Freibergs, E. Shilova and E. Shcherbinin, Electrically Induced Vortical Flows, Kluwer Academic Publishers, Dordrecht, Boston, London, p 248 (1989). 10. J. Thonstad and H. Vogt, 'Terminating anode effects by lowering and raising the anodes". TMS Light Metals (2010), 461466. 11. V. Bojarevics and K. Pericleous, "Shallow Water Model for Aluminium Electrolysis Cells with Variable Top and Bottom". In Proceedings of TMS Light Metals (2008), 403-408.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ANODIC VOLTAGE OSCILLATIONS IN HALL-HEROULT CELLS Kristian Etienne Einarsrud 1 and Espen Sandnes 2 Department of Energy and Process Engineering, Norwegian University of Science and Technology, N-7491 Trondheim, Norway. 2 Hydro Aluminium AS, Primary Metal Technology, N-6882 0 v r e Δrdal, Norway Keywords: Anodic gas bubbles, lab scale measurements, industrial scale measurements, optical methods Abstract Experiments on lab- and industrial scale cells have been conducted in order to study the behaviour of anodic gas bubbles under various operating conditions. Traditional voltage measurements have been supplied with high-speed video recordings of the bath surface showing a good correspondence between voltage fluctuations and escaping gas bubbles. On average, 0.5 and 2 bubbles were observed per second in each respective case. Average frequencies obtained by a F F T of the voltage signal however show significantly lower values, approximately half of that observed. It is shown that this discrepancy can be due to large variations in the bubble release times. Observed bubble events can be related to F F T frequencies by means of a frequency based on statistically significant periods. For industrial anodes, the possibility of overlapping bubbles is investigated as alternative effect resulting in the mismatch between observed and calculated frequencies. Introduction As stated in the introduction of the recent review by Cooksey et al. [1], "the contribution of gas bubbles to electrical resistance in aluminium reduction cells is becoming increasingly important as smelters attempt to reduce energy consumption". The importance of anodic gas bubbles arises from their negligible conductivity, effectively screening out sections of an operating anode. In practical models the influence of the gaseous bubbles is often treated as a gas coverage factor È and a gas layer thickness d& and a review of the various models used for determining cfo and θ is found in Hyde and Welch [2]. For typical operational values, the bubbles can be responsible for as much as 10% of the voltage drop, which is a pure loss of energy through Ohmic heating. The Hall-Heroult cell has a strong dynamic nature due to buoyant bubbles and strong MHD coupling. Hence, it should come as no surprise that the anodic voltage varies with time. There are several sources to these variations, ranging from high frequency ripples in the DC current applied to the cell, to low frequency MHD-instabilities

creating large scale wave motion of the metal pad. A special frequency band ranging from 0.5 to 5 Hz has been credited to bubbles and the produced signal denoted as bubble noise (cf. for instance Wang and Tabereaux [3] and Kaigraf et al. [4] and references therein.). The relation between gas coverage fraction and bubble noise has recently been treated by Kiss [5] in a numerical model, showing that the nature of the bubble noise is highly dependent on the number of bubbles present beneath the anode. A regular pattern in È is observed for single bubbles, corresponding to its regular growth and departure. When many bubbles are present, the combined effect of growing, coalescing and detaching bubbles results in random like fluctuations. However, under certain conditions, bubbles can manifest a concerted movement due to very big gas pockets sweeping longside the anode, engulfing lesser bubbles in its motion, yielding a so called quasi periodic self organized motion of bubbles and hence variation in the cell voltage. Although periodic behaviour has been observed by several authors (cf. Wang and Tabereaux [3], Wang et al. [6] and Xue and 0 y e [7]), results appear to depend significantly upon electrolysis parameters as well as the geometry of the anode. The last point is somewhat alarming as the results in the references above are all obtained on small lab-scale anodes which not necessarily produce the same physics as that encountered on industrial scale. Actually, Wang and Tabereaux [3] even point out that no direct studies have been reported regarding the bubbling phenomena occurring in commercial cells. Hence, even though it is crucial for validation purposes, the relation between lab- and industrial scale anodes appears to have been given little focus. A complete treatise on the relation between lab- and industrial scales is, necessarily, beyond the scope of the present work. Instead, the main focus will be to study the most significant properties of anodic gas bubbles and determine whether or not it is meaningful to transfer observations made on lab scale to an industrial setting.

555

Experimental setup In order to see the relation between industrial- and lab scale measurements, both settings have been investigated. Industrial measurements were conducted at Hydro Reference Center in Δrdal, October 2009, while lab scale measurements on a moderate sized anode were conducted at SINTEF Materials and Chemistry in Trondheim, March 2010. On both scales, two parallel sets of observations were conducted; high frequency measurements of anodic voltage and high frequency observations of escaping gas bubbles. In essence, the high frequency observations are video recordings of the (upper) bath interface taken at such high sampling rates that the interface movement due to individual escaping bubbles could be observed. A sketch of the experimental setup is shown in figure 1.

1.6 μ ν and an accuracy of 25ì\1 for lab scale experiments, while the frequency was 10 Hz on industrial scale. The motion of the bath interface due to escaping bubbles was recorded using a PHOTRON 1024PCI FASTCAM digital camera, controlled by a laptop PC by means of a PCI bus, at 250 fps. The camera was supplied with a Nikon 28-85 mm zoom-lens and was mounted on a tripod. The industrial cell operated at approximately 315kA under standard electrolysis condtitions, while the lab scale experiments were set up as follows: A 100 by 100 mm anode was made from industrial carbon and placed in a cylindrical graphite lined crucible lined with S13N4SiC with inner diameter 220 mm. The anode was fixed to a steel rod so that the anode-cathode distance could be varied. The anode was connected to a power supply (type LAMBDA ESS) allowing for currents up to 500 A. The experiments described in the following where conducted at up to 110 A yielding current densities of 1.22 A/cm 2 . Initially, 6.7 kg of bath (aiming at l l w t % excess AIF3, 5wt% CaF2 and 4wt% AI2O3) was melted and standard industrial alumina was added at regular intervals. Eight lab scale and three industrial scale experiments were conducted. Lab scale experiments are summarized in table I. Table I: Summary of lab scale experiments

Metal level Cathode

Figure 1: Simplified sketch of experimental setup.

For practical reasons, the camera could not be placed directly above the bath surface in the industrial setting. Instead, the bath surface was recorded through a tap hole at the end of the cell. This necessarily restricts observations to the bath interface between the two end anodes. This is however not believed to significantly influence the bubble noise, as confirmed from measurements on anodes not situated at the end. The cell voltage was logged at 50 Hz using a CR23X multilogger from Campbell Scientific with resolution

Experiment

Amperage

ACD

Temperature

(#) 1 2 3 4 5 6 7 8

(A) 80 80 110 80 110 110 80 80

(mm) 40 40 15 40 40 15 40 15

°C 975 975 969 969 970 966 966 966

Data from the voltage measurements was filtered before a statistical and spectral (FFT) analysis was performed using MATLAB. Due to bad contrast, video recordings were analyzed manually in order to obtain data for comparison. Results Identification of bubbles As noted in the previous section, industrial scale recordings were performed through an (enlarged) tap hole at the

556

end of the cell. As a consequence, all bubbles escaping from either of the two anodes adjacent to the hole in question. Fortunately, large gas bubbles typically accumulate a large portion of momentum during their travel beneath the anode, a motion which is continued even after release at the anode edge (cf. Einarsrud [8] for details), resulting in right or left-bound bubbles on the bath-surface, depending upon the anode of origin. Bubbles originating from different (industrial) anodes are shown in figure 2.

"——Voltage signal ® Observed event

Leftbound 15 time (s)

Figure 3: Measured voltage signal (solid line) and observed bubble events (circles) for lab scale experiment # 2 . Rightbound

200 mm

As video recordings and measurements were started simultaneously, the time of an observed event can be compared directly to the measured time series. Such a comparison is shown in figure 3. As seen from figure 3, the sharp decrease in the measured voltage corresponds very well to events related to detaching and escaping bubbles, as for micro-anodes such as the one used by Xue and 0 y e [7]. Lab-scale observed events account for approximately 95% of the fluctuations with amplitude larger than 3% of the average voltage, suggesting a close relationship between small and moderate size anodes.

where NOB is the number of observed bubbles over the observation time To, which is equal to 25.6 seconds. Given a periodic sequence of events, the observed bubble frequency should be close to the most dominant frequency, / i obtained from a F F T of the corresponding signal. Most and second most dominant F F T frequencies, f\ and /2, observed frequencies, foB> as well as (average) relative magnitude of voltage oscillations, Ä% are given in table II. As seen from table II, high frequencies appear to be related to small amplitudes in voltage oscillations. As noted by Kiss [5], high frequency regimes are related to small amplitudes in the bubble coverage factor 0, resulting in smaller voltage fluctuations. The same tendency is observed by Wang and Tabereaux [3], postulating inverse proportionality between the magnitude of the fluctuation and bubble frequency. The relative magnitude of the oscillations as well as typical frequencies fit well within the values expected for bubble noise. Table II shows that observed frequencies are significantly higher that those obtained from the F F T of the corresponding signal; almost four times higher in the extreme case of experiment I2a.

Significant properties of bubble noise

Interpretation of frequency

Figure 2: Distinction between bubbles originating from different anodes (denoted a and b) from escape direction. Left-bound bubble originates from anode b while rightbound bubble originates from anode a. View is approximately at 45° with horizontal.

Relation between signal and observation

The average observed bubble frequency JOB is simply defined as foB

= -777—5

(!)

The apparent discrepancy between observed and FFTfrequencies can be due to the lack of periodicity in the bubble noise signal from moderate and industrial size anodes. The average escape time for each set of experiment

557

is 1.99 and 0.51 s for each respective case, suggesting frequencies of approximately 0.5 and 2 Hz, which is close to the values expected. The corresponding standard deviation is (on average) found to be 0.99 s on lab- and 0.35 s on industrial scale. This is a significant variation suggesting that the assumption of a periodic signal does not hold, making a meaningful spectral analysis of the signal challenging. Table II: Comparison of frequencies in bubble noise for lab- ( # ) and industrial (I#) measurements. Exp. (#) 1 2 3 4 5 6 7 8 11 I2a I2b

A% (%) 5 6 4 3 5 2 2 2 3 4 2

ίθ (Hz) 0.36 0.31 0.38 0.36 0.47 0.85 0.66 0.61 1.88 1.95 1.95

h

(Hz) 0.20 0.15 0.39 0.22 0.24 0.71 0.37 0.55 0.76 0.55 0.88

h

(Hz) 0.29 0.34 0.34 0.17 0.56 0.56 0.24 0.45 0.86 0.32 0.55

f1

J ss (Hz) 0.25 0.17 0.31 0.25 0.35 0.71 0.37 0.50 1.32 1.45 1.33

Large scale effects As noted by Kiss and Poncsak [9], the frequencies of individual bubbles are very difficult to observe. Besides the influence from their release at the anode edge, interactions between moving bubbles and their coalescence dominate the spectrum of the voltage fluctuations, i.e. it is the collective bubble behaviour under the anode which is measured as a voltage signal. Based on this a hypothetical collective bubble signal is reconstructed based on the observed release frequencies. The procedure used is as follows: • The time from which a bubble appears on the surface to the time it escapes is registered. • The bubble escape time, combined with the extent of the splashing gives an indication of the bubble size, which is divided into three classes (0.25, 0.5 and 1). • The residence time under the anode is approximated from the bubble escape time.

The interpretation of frequency used to compute JOB should strictly only be considered as an average property and a one to one correspondence to the F F T frequency is expected only if the signal is purely periodic. Evidently, this is not the case for the signals obtained in the present work. In order to capture the variations in bubble release time, the frequency of statistically significant events is computed as

fi = ^ ,

frequencies from industrial scale measurements still compare poorly.

(2)

where To is the observation time and Nss is the number of statistically significant events. An event is denoted as statistically significant if its value is in the range ì ± σ, where ì is the mean value and ó is the standard deviation of the dataset in which the event in question occurs. The range is chosen so that at least 50% of the events in the dataset are included (given that the dataset follows a normal distribution). Frequencies obtained from the number of significant bubble periods (fjs) are shown in table II. As seen from table II, frequencies based on statistically significant events reproduce the calculated F F T frequencies with an average absolute error of less than 20% for lab-scale measurements. Although somewhat improved,

• Based on the approximated residence time and the size class, a resistance-curve is obtained for each individual bubble. • The total signal is computed from the sum of each individual bubble. Figure 4 shows the time of escape for bubbles originating from experiment I2a. The magnitude of each peak is determined from the size of the corresponding bubble. As expected, the time of escape is irregular, bubbles appearing as bursts rather than at a distinct frequency. Furthermore, several of the peaks are separated by very small time intervals, close to the sampling time used in the voltage measurements. This overlap becomes even more visible when a linear relation between the first appearance of the bubble and its time of escape is plotted (assuming the appearance of the bubble has zero magnitude), as in figure 5. As noted by Kiss and Poncsak [9] and observed in water models and simulations (cf. Einarsrud [8]), the nature of the bubble changes dramatically when detaching at the anode edge; from an elongated flat bubble under the anode to a more spherical shape in the center channel. As a consequence, the residence time of the bubble is greater under the anode, compared to its residence time in the center channel (i.e. the time from its appearance to its escape).

558

1.5

1.5 —» Bubble event

10 15 Time of escape (s)

20

25

3

Figure 4: Escape time and estimated bubbles size from experiment I2a.

The bubble residence time for an equivalent volume bubble under the anode can be approximated by the simple relation tanode

—

utchannel

^anode

U anode

^channel

^channel'

(3)

where C'channel and άanode represent average bubble velocities in the channel and under the anode and Lano^e and Lchannei represent typical lengths travelled by a bubble. The velocity in the channel is greater than under the anode, though the magnitude is the same (cf. Einarsrud [8]). The length scales are however significantly different; the typical anode length being of order 100 cm, while the typical bath height is 20 cm. Hence, the approximated bubble residence time under the anode is tanode

~

^channel·)

observed bubble residence time

time of appearance

Figure 5: Linear bubble growth from time of bubble appearance to escape for experiment I2a.

of 1.0 yields a frequency of 1.1 Hz. It is thus clear that the bubble residence time has a large impact on the resulting spectrum; lower residence times yielding higher frequencies. Recalling how the bubble residence time is defined (equation 3), a reduced residence time is equivalent to a shortening of the anode in the bubble flow direction.

(4)

yielding the bubble growth curve shown in figure 6. Comparing figures 5 and 6, it is clear individual bubbles in principle can overlap if the anode is of large size. A hypothetical bubble signal is constructed by adding the contribution of each individual bubbles shown in the above figure. Performing a F F T on this signal results in dominating frequencies in the range of 0.3 to 0.70 Hz, which is remarkably close to the measured values (cf. table II). The most critical parameter in the above analysis is the approximation of the bubble residence time. Changing the value of the pre factor from 5.0 to 2.5 increases the maximum frequency from 0.49 to 0.64 Hz, while a value

Figure 6: Linear bubble growth from approximated bubble residence time under anode for experiment I2a.

A similar increase in dominating frequencies is predicted by Kiss [5], showing an increase in frequency of a factor close to two when the aspect ratio of the anode is changed correspondingly, for a given flow regime.

559

The final point is important as the flow regime necessarily depends upon the anode geometry, as the present lab scale (and hence smaller length) experiments show a tendency towards lower frequencies than their (larger) industrial counterparts. A dependency upon anode size is also found by Wang et al. [6], suggesting the possibility of an optimum anode size for a given current density.

gratefully acknowledged. The authors would like to thank SINTEF Materials and Chemistry and Hydro Aluminium for inclusion in their ongoing experiments and for all valuable help with equipment and experimental setup. Also, all suggestions and recommandations given by Stein Tore Johansen, Asbj0rn Solheim, Egil Skybakmoen, Iver H. Brevik and Eirik Manger are greatly acknowledged.

Concluding remarks

References

Voltage measurements combined with visual observations show that anodic bubbles are responsible for a distinct class of voltage fluctuations on both lab- and industrial scale. In both cases, anodic bubbles yield voltage fluctuations with average frequencies in the range of 0.5 to 2 Hz and amplitudes as large as 10% of the mean voltage. In the cases studied higher frequencies were in general found on industrial scale. Although the present results appear to, on average, have properties similar to those obtained in small scale lab investigations, a detailed analysis reveals that large variations in bubble release times are present; bubbles releasing in seemingly random bursts rather than at regular periodic times. These phenomena are believed to be coupled to anode size and geometry, which in turn influence the bubble flow regime. For a given flow regime, the bubble residence time under the anode has a large influence on the resulting spectrum, indicating that the lowfrequency oscillations observed in industrial Hall-Heroult cells are due to the collective behaviour of several large anodic bubbles. This finding suggests that measurements on small lab scale anodes not necessarily represents the reality encountered on large scale industrial anodes. Small scale measurements could however be used to validate the local behaviour on an industrial anode. Throughout this work (and in the overall literature) the term "frequency" has been used to describe anodic bubble noise. Results presented herein however indicate that typical signals are not strictly periodic, hence violating the basic assumption of the traditional F F T analysis. This suggests that a frequency based description not necessarily is the best way of treating bubble noise and a more general framework is needed to describe bubble noise in industrial processes. This will however be the topic of future research. Acknowledgements The present work was financed by Hydro Aluminium, Primary Metal Technology with support from the Research Council of Norway. Permission to publish the results is

560

[1] Cooksey M. A., Taylor M. P. and Chen J. J. J., Resistance Due to Gas Bubbles in Aluminium Reduction Cells. Journal of Metals February 2008, 51-57. [2] Hyde T. M. and Welch B. J., The Gas under Anodes in Aluminium Smelting Cells Part I: Measuring and Modeling Bubble Resistance under Horizontally Oriented Electrodes. Light Metals 1997, 333-340. [3] Wang X. and Tabereaux A. T., Anodic Phenomena observations of anode overvoltage and gas bubbling during aluminium electrolysis. Light Metals 2000, 19. [4] Kalgraf K., Jensen M. and Pedersen T. B., Theory of bubble noise, bath height and anode quality. Light Metals 2007, 357-361. [5] Kiss L. L, Transport processes and bubble driven flow in the Hall-Heroult cell. 5th Int. Conf. on CFD in Proc. Ind., CSIRO, Melbourne, Australia, 2006, 1-7. [6] Wang Z., Gao B., Li H., Shi Z., Lu X. and Qiu Z., Study on Bubble Behavior on Anode in Aluminium Electrolysis - Part I. Light Metals 2006, 463-466. [7] Xue J. and 0 y e H. A., Bubble behaviour - Cell voltage oscillation during aluminium electrolysis and the effects of sound and ultrasound. Light Metals 1995, 265-271. [8] Einarsrud K. E., The Effect of Detaching Bubbles on Aluminum-Cryolite Interfaces: An Experimental and Numerical Investigation. Metallurgical and Materials Transactions B, Vol. 41:3, 2010, 560-573. [9] Kiss L. I. and Poncsak S., Effect of the bubble growth mechanism on the spectrum of voltage fluctuations in the reduction cell. Light Metals 2002, 217-223.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY

Energy Savings by Cell Design Improvements SESSION CHAIR

Bijorn Moxen Hydro Aluminum

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ELECTMCAL CONDUCTIVITY OF THE KF - NaF - A1F3 MOLTEN SYSTEM AT LOW CRYOLITE RATIO WITH CaF2 ADDITIONS A. Dedyukhin, A. Apisarov, P. Tin'ghaev, A. Redkin, Yu. Zaikov Institute of High Temperature Electrochemistry, S. Kovalevskaya st. 22, Yekaterinburg, 620219, Russia Keywords: Cryolite melts, electrical conductivity, calcium fluoride Two rigidly fixed parallel molybdenum electrodes were immersed in the melt. A unit for the load of the additives to be charged to the molten electrolyte in the flow of inert gas was attached to the small alundum gas-feeder tube. After each addition, the submersion depth was corrected.

Abstract The effect of calciumfluorideon the electrical conductivity of low temperature electrolytes for aluminum electrolysis has been investigated. Calcium fluoride addition was found to decrease the electrical conductivity of the molten mixtures of potassium and sodium cryolite melts.

Chemicals

Introduction

The electrolytes under investigation were prepared from individual salts (chemically pure) A1F3, NaF and KF (taken as KFHF). First, the electrolytes KF-A1F3 and NaF-AlF3 with the molar ratio required were prepared. A1F3 was initially heated in a mixture with NH4F and kept for 6 hours at 450-500°C for performing oxidefluorination.In order to obtain KF-A1F3 electrolyte the mixture of KFHF and A1F3 was heated up to 700°C in the glassy carbon container and kept at this temperature over four hours. HF was removed from the melt as a result of the thermal KFHF decomposition (T = 238.7°C). The NaF-AlF3 system was prepared by melting the mixture of A1F3 and NaF in presence of NH4F. The ternary system KF-NaF-AlF3 was obtained by mixing binaries NaF-AlF3 and KF-A1F3. Before conducting the experiments the electrolytes obtained were analyzed on potassium, sodium and aluminum content by an ICP method.

The electrical conductivity is a crucial property for the electrolytic process because it influences the charge and heat transfer in the electrolytic cell. The electrical conductivity decreases in the sequence of cryolites Li3AlF6>Na3AlF6>K3AlF6at the temperature of 1000°C (1). Calcium fluoride usually decreases conductivity but the precise influence has not yet been investigated well. According to different authors the electrical conductivity drops with CaF2 additions to sodium cryolite [2-4]. The increase of electrical conductivity with calcium fluoride additions was found for KF-LiF-AlF3 mixtures [5]. Experimental The investigation of the electrical conductivity of fluorideoxide melts is very difficult due to their high corrosion activity. In order to obtain reliable experimental data special requirements to the construction materials of the conductivity cell must be kept.

Measurement procedure Zahner electric IM6E was used for impedance measurements. Impedance diagrams were recorded in a frequency range from 100 Hz to 100 KHz using a signal with amplitude of 5 mV. The electrolyte resistance determined from the impedance diagrams was used for calculation of the electrical conductivity. The constant of the cell was determined from the electrical conductivity values obtained for potassium cryolite with CR 1.3 or 1.5 in the capillary type cell. The electrical conductivity was calculated taking into consideration the cell constant temperature dependence.

Experimental cell The investigations were carried in an experimental cell with two parallel electrodes [6]. A glassy carbon crucible filled with a weighed amount of the electrolyte to be investigated was put at the bottom of the quartz test tube, which was tightly closed by a vacuum rubber plug with holes for the electrodes, thermocouple, and the inert gas inlet and outlet. Small tubes for the gas inlet and outlet and the protection shields for the thermocouple and electrodes were made of alundum. The platinum - platinum-rhodium thermocouple without any shield was immersed directly into the molten salt.

563

Results and Discussion The experimental values of the electrical conductivity obtained in the KF-NaF-AlF3-CaF2 system with CR 1.3 and 1.5 at different temperatures are presented in Figures 1-3. The results were treated in the form of temperature dependence equations: 1ηχ=Α-Â/Τ

[1]

here χ - specific conductivity, Sm-cm"1; T - temperature, K; A and B - empirical coefficients. The results obtained for different electrolyte compositions in a broad temperature range (> 50°) are presented in Table 1.

Figure 2. Electrical conductivity of the (KF-NaF-AlF3)CaF2 mixtures (CR=1.5) 1 - 0 mol.% of CaF2?, 2 - 1 . 6 mol.% CaF2, 3- 3.2 mol. % of CaF2, 4- 5.7 mol. % CaF2. The NaF concentration in the initial salt was 43.86 mol.%.

Table 1. The empirical coefficients of the electrical conductivity temperature dependences. CR 1.5 1.5 1.5 1.5 1.5 1.5 1.5 1.5 1.5 1.5 1.5 1.3 1.3 1.3 1.3

NaF mol.% 60.00 57.25 56.31 55.37 54.41 53.45 52.48 43.86 43.17 42.47 41.41 44.44 43.73 43.01 41.92

KF mol.^

16.14 15.89 15.63 15.24 12.09 11.89 11.70 11.40

A1F3 mol.% 40.00 38.16 37.54 36.91 36.27 35.63 34.98 40.00 39.37 38.73 37.77 43.48 42.78 42.08 41.02

CaF2 mol.% 0.00 4.59 6.14 7.72 9.31 10.92 12.54 0.00 1.58 3.17 5.59 0.00 1.60 3.21 5.66

-B 2.86 4.42 3.74 4.15 5.45 6.65 7.27 2.05 2.13 2.73 2.23 2.05 2.09 2.23 2.47

2507 4292 3607 4091 5538 6885 10784 1861 1971 2634 3223 1963 1930 2173 2471

Figure 3. Electrical conductivity of the (NaF-AlF3)-CaF2 mixtures (CR=1.5) depending on CaF2 (mol.%) concentration: 1 - 2.3; 2 - 4.6; 3 - 7.8; 4 - 9.4; 5-11. The CaF2 additive reduced the electrical conductivity of the melts under investigation. The specific conductivity change with calcium fluoride content in the NaF-AlF3 system (CR=1.5) at T=800°C is shown in Fig. 4. The conductivity polytherms presented in Figures 1-3 for different electrolyte compositions show the temperature slope increase with increasing calcium fluoride concentration. The specific conductivity temperature coefficient change with CaF2 in the NaF-AlF3-CaF2 system (CR=1.5) is shown in Fig. 5.

770

T, °C

The electrical conductivity activation energy for the binary NaF-CaF2 system was investigated too. It slightly changes with CaF2 additions. Aluminum fluoride additions significantly change the state of the melt, complicating its structure due to the joint presence of the two strongly complexing ions Al3+ and Ca2+. Both of these do not exist as single cations, but are bound tofluorideions in the melt.

790

Figure 1. Electrical conductivity of the (KF-NaF-AlF3)CaF2 mixtures (CR=1.3) depending on CaF2 (mol.%) concentration: 1 - 0; 2 - 1.6; 3 - 3.2; 4 - 5.7. The NaF concentration in the initial salt was 44.44 mol.%.

564

References

Figure 4. Electrical conductivity of the NaF-AlF3 system (CR=1.5)atT=800°C.

§

8000

0.04

0.06

0.08

0.10

Molar fraction of CaF2

Figure 5. Electrical conductivity temperature coefficient, B. 1 - NaF-AlF3-CaF2 (CR=1.5); 2 - (KF-NaF-AlF3)-CaF2 (CR=1.5), the NaF concentration in the initial salt was 43.86 mol.%; 3 - (KF-NaF-AlF3)-CaF2 (CR=1.3), the NaF concentration in the initial salt was 44.44 mol.%. It leads to the great increase of activation energy with calcium fluoride additions in the NaF-KF-AlF3-CaF2 system. XRD analysis of the solid samples showed the presence of elpasolite K2NaAlF6, NaCaAl2F9 and KCaAl2F9. Most likely, such huge complexes can be present in the liquid state leading to ion mobility decrease. Conclusions The CaF2 additives decrease the specific conductivity of the KF-NaF-AlF3 mixtures with CR=1.3andl.5. The temperature coefficients of the specific conductivity of the melts under study increase with calciumfluorideadditions. 3. The electrical conductivity decrease with CaF2 in the mixed sodium-potassium cryolite systems can be explained by the formation of large complex ions containing calcium ions.

1. E.W. Yim and M. Feinleib. J. Electrochem. Soc. V.104, 1957, pp. 626-630. 2. J.D. Edwards, C.S. Taylor, L. Cosgrove, and A.S. Rüssel. J. Electrochem. Soc. V.100, 1953, pp. 507-512. 3. A.V. Vakhobov and A.I Belyaev. In Physical Chemistry of Molten Salts. Metallurgizdat, Moscow, 1965, pp. 99-104. 4. J. Hives, J. Thonstad, A. Sterten, and P. Fellner. Light Metals 1994, pp. 187-194. 5. A.E. Dedyukhin, A.P. Apisarov, A.A. Redkin, O.Yu. Tkacheva, and Yu.P. Zaikov. Light Metals 2008, pp. 509-511. 6. V. Kryukovsky, A. Frolov, O. Tkatcheva, A. Redkin, Yu. Zaikov, V. Khokhlov, and A. Apisarov. Light Metals 2006, pp. 409-413.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

STUDY OF ACD MODEL AND ENERGY CONSUMPTION IN ALUMINUM REDUCTION CELLS Tian Yinfu, Wang Hang Chongqing Tiantai Aluminum Industry CO., Ltd. Chongqing, 401328, China Keywords: aluminum electrolysis, ACD model, critical ACD, DC energy consumption However, according to a recent study [3], a 180 kA potline, where the current has been increased by 25 kA, has achieved 93.48% CE at 3.83 to 3.88 V, with 0.83 A/cm2 anodic current density. The bath composition is listed in table 2.

Abstract In this paper a model of the anode-cathode distance (ACD) is built according to actual production in traditional aluminum reduction cells. The critical ACD is also studied here. Based on the ACD model and the critical ACD it can be described why some cells can produce at low cell voltage in the present aluminum industry. It is possible that only 9500 kWh/t of DC energy will be required at 2.0 cm ACD and 0.8 A/cm2 anode current density.

Table 2 The bath composition used for the amperage increase study Molar ratio A1203 concentration CaF2 content LiF content 3% 2.5 2.07% 5%

For one cell in the 180 kA potline operated at 935°C, the voltage drops are listed in table 3.

Introduction Energy consumption for one ton of aluminum produced in HallHeroult cells can be calculated from the following formula: W=2980*i//;/ Here, W is the direct current (DC) power consumption in kWh per ton, and U is cell operating voltage drop in volts, and η represents current efficiency (CE).

Table 3 The voltage drops Cathode voltage drop Anode voltage drop Drops in all bars Voltage drop between anode and cathode Cell voltage drop

Cell operating voltage drop is composed of several parts shown in table 1.

The voltage drop between anode and cathode listed the table 3 includes decomposition voltage, polarizing voltage and bath voltage. If the decomposition voltage and polarizing voltage is 1.65 V, the bath voltage is 1.253 V. According to the literature 1, the resistivity is 0.44Q.cm, and thus: ACD=1.2526/(0.44x0.83)=3.43 cm

Table 1 Typical example of cell operating voltage compositions ( for 0.75 A/cm2 anodic current density ) Anodes Cathodes Busbars Reaction Polarization Bath Total

0.376 V 0.349 V 0.259 V 2.902 V 3.887 V

Voltage Drop/V 1 -0.33 -0.32 -0.2 -1.20 -0.45 -1.55 -4.05

Therefore, there are two questions: how much is the ACD with high CE, and how much is the lower of energy consumption. The three layers ACD model

Note: semigraphitic blocks are used as cathodes here. Based on the bath voltage drop, Ub, and resistivity, Rb, the anodecathode distance (ACD) can be calculated: ACD=cV(Qxfl») Here, Cd means current density.

In order to answer the questions above, a model of ACD, named three layers ACD, is introduced. It is assumed that the ACD is made up of b, c and a, present the height of three layers: the top layer of bath mixed with anode bubbles, the middle transition bath layer, and the bottom metal pad fluctuation layer, respectively, shown in figure 1. So A, the value of ACD, should be equal to sum of a, b and c. h =a+b+ c

The value of Rb can be found in the literature [1]. For example, it is suggested that Rb is 0.47Ω·àΐη when the cell is operated at 950°C with bath composition of 2.3 molar ratio, with 2%~3% alumina and 5% CaF2. If Ubis 1.55 V: ACD = 1.55/ ( 0.75x0.47 ) = 4.4 cm Tian Yingfu studied the optimum ACD for a cell operated at more than 4.2 V, and he introduced a model for ACD [2]. According to that, it is suggested that the optimum ACD is 4.2 to 4.5 cm for traditional cells operated at a voltage of 4.0 to 4.1 V.

Fig. 1 Schematic diagram of the three layers ACD model

567

The value of a is dependent on the stability of the magnetofluid and the height of metal pad. For example, a is lower with steady magnetofluid and reasonable bars distribution. In addition, it is also lower with a larger height of the metal pad, which causes it to flow slowly and weakens the waves. The value of a is ordinarily lower than 1.5 cm. For example, with a 1.42 cm height of the aluminum waving layer, the maximum value for the Chongqing Tiantai cells designed by Shenyang Aluminum & Magnesium Engineering & Research Institute, was taken into account [4].

DC consumption for aluminum production Based on the three layers ACD model, in order to reduce the DC consumption for one ton of aluminum produced, the three heights should be decreased as much as possible. For the NSC cells the metal pad fluctuation height is less than 0.7 cm, as mentioned above, so c may be reduced to 0.5 cm in order to decrease the bath layer voltage drop without any CE loss, and the ACD can be controlled at about 3.2 cm. The cell voltage drop would reach 3.709 V with 0.8 A/cm2 Cd and 0.45 Ω-cm Rb. So the DC consumption of aluminum producing in the NSC cell would then be 11760 kWh/t-Al with 94% CE.

The value of b is dependent on the width of the anode carbon and the immerged depth in the bath, the density and viscosity of the bath, and the surface tension of the bath to anode bubbles. The b layer would become thin due to the anode gas discharging in time under the following cases, reducing the width of the anode carbon, or decreasing the bath level, or using the low-density bath, or using the low-viscosity bath, or using the bath with high surface tension. For a commercial cell with 660 mm wide anode, the b layer is about 2 cm thick [5, 6].

Based on the ACD model, if the anode is optimized to reduce the height of the b layer, for example, by diminishing the gas bubbles by reducing the width of the anode, or by other methods releasing the anode gas in time. The voltage drop of the b layer can be decreased by 0.3 V due to the height decreasing to 1 cm from 2 cm. It means that the DC consumption can be decreased by about 1000 kWh/t-Al. In addition, the anodic overvoltage can be decreased also with decreasing of the b layer.

In order to avoid the reaction occurring between anode gas and aluminum produced, the c layer should exist to separate the a layer and the b layer. However, the bath voltage drop would be higher with an increase of the c. For a commercial cell with a total voltage of more than 4.0 V, if Cd is less than 0.75 A/cm2, the value of c is about 1 cm. For example, a cell operated with 0.72 A/cm2 Cd and 1.55 V ACD voltage drop, if Rb is 0.47 Ω·αη, the ACD, h, is calculated by the following equation: H= 1.55/(0.47x0.72) = 4.58 cm

Therefore with the method decreasing b and some optimal technology, for example, using graphitized blocks and decreasing contact voltage drop of Fe-C in anodes, the NSF cell can operate at less than 3.0 V cell voltage, and the DC consumption for aluminum producing would not be more than 9500 kWh/t-Al. It means that the energy efficiency would reach 66%.

So, the value of c,

References

C = h-a-b = 4.58-1.5-2 = 1.08 cm Therefore, if c is small, the cell voltage drop would be decreased.

1. Zhang Mingjie, Qiu Zhuxian. Advanced Technology of Aluminum Electrolysis (in Chinese). 2001.

The Critical ACD

2. Tian Yinfu. Discussion on Best polar distance of industrial aluminum cell (in Chinese). Non-ferrous mining and metallurgy, Vol. 25(6), 2009: 23-25.

Based on the three layers ACD model as mentioned above, h would become smaller when c decreases. When it decreases to zero, h would reach its critical value, named the critical ACD. So for a commercial cell the critical ACD, 3.5 cm, is equal to sum of a and b as: he = a + b = 1.5 + 2 = 3.5 cm

3. Li Jie. A report on Meeting of New Cathode Aluminum Reduction Cell Technology. Shenyang, 2010. 4. Shenyang Aluminum & Magnesium Engineering & Research Institute. Design scheme of Chongqing Tiantai Aluminum smelter, 2009.

If the cell is designed badly for the magnetic field or it is operated poorly, the height of the metal pad fluctuation layer would be more than 1.5 cm, and he would be higher than 3.5 cm.

5. Zoric J, Solheim A. On gas bubbles in industrial aluminium cells with prebaked anodes and their influence on the current distribution [J], Journal of Applied Electrochemistry, 2000, 30: 787-794.

For a commercial cell with an optimal design of the magnetic field, a can be diminished to 1 cm, so he would be close to 3 cm. In order to keep a high CE, it is suggested that the c layer should be 0.5 cm high, so the cell voltage drop would be less than 3.85 V. This has already come true in China.

6. Zhou Naijun, Xia Xiaoxia. The research development of electrolyte flow in aluminum reduction cell and its direction to the research on flow field in drained cell (in Chinese). Light metals, 2004, (12): 26-32.

For the novel structured cathodes (the NSC cells), invented by Professor Feng Naixiang [7], the height of the metal pad fluctuation layer could be controlled to less than 0.7 cm, so he should be less than 2.7 cm. Therefore, for the NSC cell operated at 2.7 cm he and with 0.72 A/cm2 Cd and 0.45 Q»cm Rb, the cell voltage drop could be 3.42 V, which is almost the same as the critical cell voltage drop, 3.428 V, measured by Yang Xiaodong, professor at SYAMI, China [8]. In that experiment, the cell voltage drop is decreased till the critical value, at which the cell become very unstable, or noisy..

7. Feng Naixiang. Aluminum reduction cathodes[P]. CN 20072001223l.X

cell

with

novel

8. Yang Xiaodong. A report on Meting of New Cathode Aluminum Reduction Cell Technology. Shenyang, 2010.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Modeling of Energy Saving by using Cathode Design and Inserts Rene von Kaenel, Jacques Antille KAN-NAK SA - Sierre - Switzerland Keywords: Cathode design, Inserts, Energy saving, Magneto-hydrodynamic modeling, Cell stability

The cell voltage I J ^ can b e decomposed into the following

Abstract:

contributions ( Κ . Θ Γ & Η.Κ. [ 1 ] ):

The use of modified cathode shape and inserts in the cathode allow drastically redistributing the current density in the aluminium liquid pool. As a result both the cathode voltage drop and the anode to cathode voltage drop can be significantly decreased. Electrical modeling can easily explain the energy savings in the cathode. Magnetohydrodynamic modeling shows why and how specific energy can be saved in the anode to cathode distance. The energy saving inside the cell depends on the cell initial design but may exceed 1 kWh/kg AI.

I.

Ucell Uext U^ Ubath UCVD

Introduction

U

bath

U . + U n + U h h + UAF,with:

Udec U^ Uaco Ucco Uί Ubub UAE Uelec

Decomposition voltage A n o d e reaction overvoltage A n o d e concentration overvoltage Cathode concentration overvoltage Ohmic voltage drop in A C D Bubbles voltage drop Average anode effect voltage Electrochemical voltage

bath

Prod P^ EcelI E

Cell production Cell power Cell energy for the period At Cell specific energy cons.

=U„U*I

Been = U w i l * I * A t Prod = F * η * I * At

\ Curr. eff. 1 Voltage

[V] [kA] [1] [hours] [kg/kAh]

AE'

energy

96%

Specific Energy [kWh/kg] 12.807 BL. 0.708 E_ |u_ 1.080 E„„„ |u 10.215 Ewk 1.840 5.711 1 u.,„ E.,„ u 3.846 1.239 B„ 0.162 0.503 uk„t ^hnh 0.050 0.155 EAB uAB 0.804 0.259 I-~C.VI1_— In fact the specific energy consumption has decreased from about 6 0 kWh/kg at the end of the 19th century to about 12.5 kWh/kg AI in 2010 [10].

u„„

[kg] [kW] [kWh] [kWh/kg AI]

1

Total power [ k W ] Energy [kWh] Production [kg Al]

[kWh/kg]

bub

Table 1: Voltage and related specific

[V]

4.126 0.228 0.348 3.291

-

Are there more solutions for saving voltage?

And from there the specific energy is found equal to: E = UceU / ( F * T I )

Ω

Table 1 presents a typical situation for a modern cell:

Basic physics tells us that (K.Grjotheim, H.Kvande m ) : PceU

elec

Each contribution can b e determined rather accurately if the bath temperature, t h e chemical composition and t h e local current density are known.

Specific energy consumption

Cell voltage Line current Cell current efficiency Time Cst. (Faraday's law) 0.3356

U_ + u_+ u + u + u„+ ii.,. + u„

U biΔ =

In order to discuss the different components of the specific energy consumption let us first introduce some notations: Ucell I η t F

Total cell voltage Ext. voltage from collector bars to the anode stubs From anode stubs to the bottom of the anodes F r o m bottom of anodes to the liquid metal pool Cathode voltage drop (metal pool to coll. bars)

And the bath voltage is:

The aluminium industry is continuously working on reducing the specific energy consumption as it is one of the most important contributions to the production cost. To reduce the specific energy consumption one must act on two parameters; the current efficiency and the cell voltage. It is trivial to say that the current efficiency must be increased and the cell voltage decreased. The current efficiency is mainly a result of the cell design, the cell operation and the quality of the raw materials. Today, the best cells achieve 96% current efficiency. Further improving the current efficiency is mostly linked to the materials properties and the operation of the cell. This study is focusing on the cell design and is assuming a constant current efficiency. Π.

with:

u =u +u

The external voltage can only b e reduced b y changing the busbars electrical resistance. This means for example

(Eq. 1)

569

increasing the mean busbars cross-sectional area and/or changing the busbars topology and/or using copper. Today modern side by side cells are already rather well optimized in term of external voltage and additional voltage reduction is very often not economical.

consumption in the range of 12.2-12.5 kWh/kg AI [10]. New cathode shapes allow decreasing the ACD without reaching the magneto-hydrodynamic instabilities [11,12]. The electrical conductivity of molten cryolite is very low, typically 220 Ω^ηϊ1 and the ACD cannot be decreased too much due to the formation of magneto-hydrodynamic instabilities leading to waves at the metal-bath interface. The existence of waves leads to a loss of the current efficiency of the process and does not allow decreasing the energy consumption under a critical value. In average in the aluminium industry, the current density is such that the voltage drop in the ACD is minimum 0.3 V/cm. As the ACD is 3 to 5 cm, the voltage drop in the ACD is typically 1.0 V to 1.5 V. A thorough study of cell magneto-hydrodynamic instabilities led to the conclusion that there is still a large possibility of improving the interaction of the magnetic field with the local current density in the liquid metal by modifying the local current density. The magnetic field inside the liquid metal is the result of the currents flowing in the external busbars and the internal currents. The internal local current density inside the liquid metal is mostly defined by the cathode geometry and its local electrical conductivity. In other words, changing the cathode geometry and conductivity will lead to new current distribution inside the liquid metal. The magnetic field and current density produce the Lorentz force field,

The anode voltage drop has a number of components [3] but depends mostly on the anode electrical resistivity and the anode height. Since long, the anodes have been analyzed and optimized in term of baked density and carbon recipes (Sheralyn Marie Hume[2]). Only small improvements to the anode voltage drop are possible today, unless using different materials such as metallic anodes which are out of the scope of discussion of this paper. The electrochemical voltage components can be slightly modified by the current density in the cell and by the bath composition and temperature. These components cover the minimum theoretical energy needed to convert alumina into aluminium (1.43/η +4.91 kWh/kg AI) and there is not much to do at that level. The bubbles voltage drop can be reduced by minimizing its thickness layer. The use of slots in the anodes for helping the gas to escape is one solution [4]. However, today the trend is to increase the anode current density and this means most of the time increasing the bubble voltage drop. The average anode effect voltage drop is low and all efforts to reduce the number of anode effects further reduce its value. There is no big challenge left when considering only the energy.

f = jXb[N/m3]

j Current density vector field. b Induction magnetic vector field generated by internal + external currents. f Lorentz force vector field.

The only two remaining voltage contributions that were not discussed so far, are the ohmic voltage drop in the anode to cathode distance (Un) and the cathode voltage drop (UCVD). Both together they still represent about 1.5 V or 4.7 kWh/kg Al. III.

That itself generates the metal surface contour, the metal velocity field and defines the basic environment for the magneto-hydrodynamic cell stability. The cell stability can be expressed as the ability of lowering the ACD without generating unstable waves at the surface of the metal pad. The level of stability depends on the current density and induction magnetic fields but also of the shape of the liquid metal pool. The shape of the pool depends on the surface of the cathode and the ledge shape. In order to eliminate potential sludge problems this study was restricted to smooth cathode surfaces.

Cathode voltage drop minimization

Many efforts have been taken to minimize the cathode voltage drop (CVD) [5,6]. On one hand the carbon properties of the carbon blocks are strongly modified by changing from amorphous blocks to graphitized, graphitic, impregnated or even using variable resistivity properties blocks. On the other hand the current collector bars design has been upgraded both from a geometrical and material point of view. A number of patents cover different aspects [7,8,9]. The use of copper inserts has allowed an interesting step in the CVD minimization. IV.

(Eq.2)

V.

Cathode design and inserts

Different geometries have been described [12], using both a change in the surface shape and the use of conductive inserts. Figure 1 shows a possibility of using a deeper pool ("A") machined out of the standard cathode ("B").

ACD ohmic voltage drop minimization

At first glance the minimization of the ohmic voltage drop in the anode to cathode distance (ACD) looks very simple as it is only necessary to reduce the ACD. This is without considering the thermal balance and the magnetohydrodynamic effects. In fact, for some cell designs, one should consider increasing the CVD for minimizing the total cell voltage. This can be understood when analyzing the whole thermal and magneto-hydrodynamic effects. More recently, new cathode designs have shown real improvement in the cell voltage, leading to a specific energy

570

is increased. The decrease is such that the total internal heat is kept constant. Figures 2 and 3 show the reference cell.

Figure 2: Case - Reference cell design (end to end)

Figure 1: Modified metal pool The impact of the busbars on the cell magnetohydrodynamic stability is well known. As an example, reference [13] analyses an end-to-end cell technology using three different busbars topologies. In this study, the cell stability was determined by using sophisticated models that were published earlier [14,15,...,25]. The impact of the cathode shape and the use of inserts are analyzed. Instead of changing the magnetic field generated by the external currents of the cell, the internal current density in the liquid metal is modified by changing the cathode geometry and conductivity. Obviously due to the Lorentz force equation (2) the change of current density in the cell will have a strong impact on the cell magneto-hydrodynamic state.

Figure 3: Case 1- Cathode block of the reference cell Figure 4 shows the modified cathode block surface.

Three cases are compared: Case 1: Reference cell with a standard cathode block (Figure 3). Case 2: Modified cathode surface but using the same collector bars (Figure 4).

Figure 4: Case 2 - Modified cathode block surface

Case 3: Addition of a conductive inserts above the collector bars while using the same collector bars (Figure 5).

In order to further improve the cell stability (magnetohydrodynamic cell state) one can use conductive metallic inserts in the cathode to drive the current through a different path. Figure 5 shows the inserts on top of the current collector bars. The cathode is not visible but the metal pool is shown.

Stationary quantities are analyzed for all three cases such as metal upheaval, metal velocity field and electrical field. The stationary state is further analyzed to determine the cell magneto-hydrodynamic stability.

LIQUID METAL

The followings parameters are considered for the cell:

I JA

K K

Qint

170 kA 0.9 A/cm2 19.0 cm 22.0 cm 310 kW

Line current Mean anode current density Mean metal height Mean bath height Internal heat

The ACD has been kept constant under each anode, i.e., the anode surface is deformed in order to follow the metal surface. The total volume of metal has been kept constant when modifying the cathode surface.

Figure 5: Case 3- Conductive inserts on top of the current collector bars Figures 6, 7 and 8 show the metal upheaval for the three cathode designs.

For the cell magneto-hydrodynamic stability calculations, the anode to cathode distance is decreased while the current

571

Figure9: Velocityfieldof the reference cell (Maximum velocity = 0.13 m/s) Figure 10 shows the impact of modifying the cathode surface. The global pattern flow is kept similar but the maximum velocity is slightly higher at one position. This is most likely due to the lower level of metal above the cathode on the long sides of the cell.

Figure 6: Metal upheaval of the reference cell (Min = -7.3 cm, Max = 2.5 cm) The use of a modified cathode surface leads to a much lower metal upheaval which is favorable for the anode setting.

(Maximum velocity = 0.15 m/s) Figure 11 shows the impact of the inserts. The change is drastic, the flow has changed its sign at one end of the cell, and the maximum value of the velocity field is less than half of the previous cases.

Figure 7: Metal upheaval of the modified cathode surface (Min = -2.6 cm, Max = 1.8 cm) The additional inserts do not further help to reduce the metal upheaval, which remains almost unchanged when compared to the modified surface cathode.

Figure 11: Velocityfieldof the cathode using inserts above the current collector bars (Maximum velocity = 0.06 m/s)

Figure 8: Metal upheaval of the cathode using inserts above collector bars (Min = -3.1 cm, Max = 2.0 cm) Figure 9 shows the horizontal components of the velocityfield8 cm above the surface of the cathode for the reference cell. The maximum value is 0.13 m/s.

572

Figure 12 shows the electrical potential in the liquid metal for the reference cell. The current is flowing perpendicular to the isopotential lines (see vectors). It can clearly be seen and it is well known that the current density has a horizontal component leading current from the center of the cell towards the sides inside the liquid metal.

^Í,×

-

Ledge

Center of the cell

Figure 12: Electrical potential in the liquid metal of the reference cell (CVD = 0.34 V) Figure 13 shows the electrical potential in the liquid metal for the modified cathode surface. The impact on the CVD is not very important but the current density is more vertical in the liquid metal.

\ \ vv

3

170

175

180

185

190

195

Current (KA)

200

Figure 15: Refative change of cell stability for the modified cell compared to the reference cell

Ledge

Center of the cell

Figure 13: Electrical potential in the liquid metal of the modified cathode surface. (CVD = 0.32 V) Figure 14 shows the impact of the inserts. The CVD shows a small change but the current density is strongly modified. This is of prime importance for the cell stability and also explains the change of metal velocity field.

Finally the most important results are summarized in table 2. By increasing the current, the specific energy consumption is decreased and this is possible due to the new cell stability. There are other constraints to current increase that were not discussed in this paper such as the minimum ACD and the maximum current density in the anodes and busbars. Each technology must be analyzed in order to determine the global specific energy saving which is in the range 0.4 -1 kWh/kg AI while the productivity increases are 10% to 20%.

fz

Table 2: Summary of resi dts Ledge

Center of the cell

Figure 14: Electrical potential in the liquid metal of the cathode using inserts above the current collector bars (CVD = 0.32 V) For any given design, the cell stability decreases when the current is increased. In order to keep the same thermal state and save energy using the new cell designs, the current is increased while keeping the same internal heat production. In fact the ACD is lowered while increasing the current. Figure 15 shows the relative change of cell stability for the reference cell and for the modified cathode surface cell to the cell stability of the reference cell. As expected, the instability increases with the current for both cases. However, the current can be increased by more than 15% in the modified cathode cell to reach the same level of stability as for the reference cell.

[Current 1 Internal heat 1 Anode Cur. Dens. CVD Specific energy saving in ACD & CVD | Total energy saving | Productivity increase

Unit

Reference

Cathode surface

Insert

kA kW Afcm2 mV

170.00 307.5 0.90 340

187.00 307.1 1.00 353

200.00 291.4 1.06 374

kWh/kg

0.00

-0.63

-1.29

kWh/kg

0.00 0

-0.38 10

-0.85 18

%

In this study the current efficiency was assumed to be constant. In fact, an improved cell magneto-hydrodynamic state may contribute to increase the current efficiency and therefore further decrease the specific energy consumption. Many more cathode shapes and positions of inserts can be considered to improve the cell stability and the solutions can be applied for end-to-end or side-by-side cell topologies. VI.

Conclusions

New designs of cathode blocks allow for an important reduction of the specific energy consumption and offer great opportunities for production increase. The energy saving is marginally due to the direct energy saving in the cathode but is mostly realized through the possible current increase due to a better cell magneto-hydrodynamic state. The current efficiency might also be increased by the improved cell stability but this must be quantified by operating the cells.

573

VII.

On the Analysis by Perturbation Methods of the Anodic Current Fluctuations in an Electrolytic Cell of Aluminium, Light Metals 1989, edited by Paul G. Campbell, pp. 237243. [15] J. Descloux , P. Maillard An electromagnetic free-boundary problem. Equadiff 7, Teubner-Texte zur Mathematik 118, 1990, pp. 240-242 [16] J. Descloux, M. V. Romerio, M.Fliick, Linear stability of electrolysis cells Parts 1,11, EPFL, DMA, November 1990 [17] J. Descloux , M. Flueck, M.V. Romerio. Modeling of Stability of the electrolysis cells for the production of aluminium. Numerical Methods in Engineering and Applied Sciences. Alder and al. CIMME, Barcelona 1992, pp. 30-38. [18] J.Descloux, M.Fliick, M.V.Romerio Modeling for instabilities in Hall-Hιroult cells: mathematical and numerical aspects Magnetohydrodynamics in process metallurgy Light Metals 1992, Ed. by E.R.Cutshall, ppl 195-1198 [19] J.Descloux, YJaccard, M.V.Romerio, Stability in aluminium reduction cells: a spectral problem solved by an iterative procedure, Light Metals 1994, pp 275-281, Ed. U.Mannweiler

References

[I] Κ. Grjotheim and H. Kvande Understanding the Hall-Hιroult Process for Production of Aluminium Aluminium-Verlag, ISBN 3-37017-181-2, 1986 [2] Sheralyn Marie Hume Influence of Raw Material Properties on the Reactivity of Carbon Anodes Used in the Electrolytic Production of Aluminium Aluminium-Verlag, ISBN 3-87017- 237-1,1993 [3] Wangxing Li, Jieming Zhou, Numerical Analysis of the Anode Voltage Drop of a Reduction Cell, Light Metals 2009, ppl 169-1171, TMS, Edited by Geoff Bearne [4] Markus W. Meier, Raymond Perruchoud, Production and performance of slotted anodes Light Metals 2007, pp 277-281, TMS, Edited by Morten Sorlie [5] F. Hiltmann, P.M. Patel, Influence of Internal Cathode Structure on Behavior during Electrolysis Parti : Properties of Graphitic and Graphitized Cathode Materials, Light Metals 2005, pp 757-762, TMS, Edited by Halvor Kvande [6] Zhongning Shi, Junli Xu, Test of Various Graphitic Cathode Blocks materials for 300 kA Aluminium Reduction Cell, Light Metals 2007 pp 849-852, TMS, Edited by Morten Sorlie [7] Graham E. Homley, Donald P. Ziegler ALCOA Cathode collector bars Patent US 6,231,745 Bl, May 2001 [8] Jacques Antille ALCAN Carbon bottom of an electrolysis cell for production of aluminium, WO 02/064860 Al, August 2002 [9] Frank Hiltmann SGL Cathodes for aluminium electrolysis cell with expanded graphite lining WO 2007/071392 A2, June 2007 [10] Li Jie, Lu Xiao-jun, ... Industrial Test of Low-voltage Energy saving Aluminium Reduction Cell, Light Metals 2010 pp 399,404 TMS, Edited by John A. Johnson [II] Wang Ziqian, Feng Naixiang,... Study of Surface Oscillation of Liquid Aluminium in 168 kA Aluminium Reduction Cells with a New Type of Cathode Design Light Metals 2010, pp 485,488 TMS, Edited by John A. Johnson [12] Renι von Kaenel, Jacques Antille Hιroult Cell Cathode Design, PCT/IB2010/052394, May 2010 [13] J. P. Antille, R. von Kaenel Busbar optimisation using cell stability criteria and its impact on cell performance Light Metals 1999, pp 165-170

[20] J. Descloux , M. Flueck, M.V. Romerio. Spectral aspects of an industrial problem. In Spectre Analysis Of Complex Structures. Collection Travaux en cours 49, Hermann, Paris, 1995, pp. 17-34. [21] R. von Kaenel, J. P. Antille On the stability of alumina reduction cells", Fifth Australasian aluminum smelter conference 1995, Sydney, Australia, Ed. B.Welch & M.Skyllas Kazacos, pp 530-544 [22] J.Descloux, M.Fliick, M.V.Romerio, Spectral aspects of an industrial problem, Spectral analysis of complex structure, Ed Hermann Paris, coordinator E.Sanchez Palencia 1995,ppl7-33 [23] J. P. Antille, P. Snaelund, J. M. Stefansson, R.von Kaenel, "Determination of metal surface contour and improved anode consumption", Light Metals 1997, pp 469-476 [24] J.Descloux, M.Fliick, M.V.Romerio, Modelling of the stability of aluminium electrolysis cell, Non-linear partial differential equations and their applications, College de France, Sιminaire Volume XIII, Ed Longman 1998, pp 117133 [25] J. P. Antille, J. P. Descloux, J. Flueck, M. Romerio Eigen Modes and interface description in Hall Hιroult cell Light Metals 1999, pp 333-338

[14] J. Descloux , M. Romerio,

574

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EXPERMENTAL INVESTIGATION OF SINGLE BUBBLE CHARACTERISTICS IN A COLD MODEL OF A HALL-HÉROULT ELECTROLYTIC CELL Subrat Das1, Yos Morsi, Geoffrey Brooks, William Yang2, John J.J. Chen3 Subrat Das, Yos Morsi, Geoffrey Brooks; Swinburne University of Technology, FEIS, PB 218, Hawthorn, VIC 3122, Australia 2 William Yang; CSIRO, Minerals, Box 312, Clayton, VIC 3169, Australia 3 John J.J. Chen; 20 Symonds Street, Dept. of Chem. and Materials Engineering, Auckland 1142, New Zealand Keywords: Bubble dynamics, Wall induced forces, Hall-Hιroult process, High speed photography highly discontinuous properties of the interface. Thus most of these analyses have been limited to single drops and bubbles with small density ratios in inviscid flow.

Abstract Understanding the characteristics of the bubbles generated within a Hall-Hιroult electrolytic cells, can assist greatly in the optimization and the operation of the process. One of the significant factors that greatly influence the bubbles formation is the vertical walls formed by the anodes. In this paper we used a high speed camera to investigate the effect of vertical walls on the shape of a single bubble rising in two liquids of high and low viscosity namely glycerol and water respectively under various offsets of the vertical wall and gas injection rates. The images of the bubble rising were recorded at the speed of 5000 frames per second and subsequently processed using the classical imageprocessing algorithms incorporated with MATLAB. The data related to various parameters such as aspect ratio, equivalent diameter of the bubble and bubble distance from the vertical wall are presented and discussed for various flow rate regimes. The findings showed that the presence of a vertical wall on one side of the bubble has a significant effect on the bubble shape, orientations and the trajectory path. In addition, it was found that as the bubble moved away from the wall, the velocity of the fluid between the bubble and the wall increased relative to the surrounding fluid, which created an asymmetric flow field around the bubble. Still the aspect ratios of the bubbles were found to be a function of the rate of gas injection as well as wall offset.

A number of experimental studies have addressed aspects regarding bubble shape and the velocity field around individual bubbles using particle image velocimetry (PIV) and high speed cameras [7-9]. Several techniques such as laser induced fluorescence (LIF) and infrared shadow technique (1ST) have been used to enhance the camera and PIV output [10, 11]. Previous mathematical modelling studies have shown that when the distance of the bubble interface to the wall is more than twice as large as the bubble diameter, there is practically little effect of the wall boundary conditions on the simulation results[12, 13]. But if the walls are not symmetrical, i.e. if the bubbles are nearer to the wall of one side of the container then the shape is greatly influenced by the surrounding hydrodynamic field. A spherical bubble rising near a vertical wall exhibits a creeping flow due to migration forces. A larger bubble, closer to a wall at one side, will be affected more due to the wall induced asymmetric flow and may undergo significant deformations. Details of these migration forces that stem from nonlinearities in advection momentum transport and interfacial deformability to break the symmetry are discussed by Magnaudet et al. [14]. Moreover, bubbles can fluctuate in response to oscillations in the pressure field [15] in the liquid. This pressure gradient will have a significant effect while the flow field behind the bubble is asymmetric.

Introduction The rising of bubbles in a viscous liquid due to buoyancy is a common phenomenon in chemical and metallurgical processes. The influence on volume, shape and terminal velocity of bubbles upon the rheological properties of fluids (density, viscosity, surface tension etc.) are thought to be of key importance in designing many such multi-fluid systems. The interaction between bubble and the wall is complex due to the wall induced drag forces, bubble shape and related oscillations. Thus, an understanding of the interactions between bubbles and solid walls is important for many practical applications.

In many situations, a bubble encounters a boundary wall during its transport process, and a hydrodynamic interaction occurs between the bubble and the wall. One example of such application is an aluminium electrolytic cell, where the bubbles nucleating from anode surface move along the vertical wall of anode inducing a strong circulatory motion. There have been several studies on bubble induced flow in Hall-Hιroult cell [6,7], but most of these works have mainly addressed the bulk behavior of the flow rather than the bubble movement itself. In this paper, an attempt is made to analyse bubble behavior as it moves along the vertical side of the anode. The main objective of this study is to determine the influence of a solid wall on the shape and the trajectory of the bubbles travelling parallel to the vertical wall.

The rise of a bubble in a viscous liquid is accompanied by deformation of the bubble, resulting in spherical, ellipsoidal and in some cases toroidal shapes. Numerical simulations render possible the prediction of overall shape development of several techniques such as front tracking method [1], the volume of fluid method [2], the Lattice-Boltzman method [3] and the level set method [4] are available for solving problems with moving boundaries and are discussed in an excellent review article by Hua et. al. [5] and Chen et. al. [6]. Complications associated with the advection of the interfaces between two dissimilar fluids have largely limited numerical studies to simplified cases. The primary obstacle is the numerical diffusion which results in the sharpness of the advancing phase front being blurred due to the presence of

Experimental Apparatus Experiments were conducted in a 0.2 x 0.2 m Perspex square tank of height 0.15 m. The bubble train was generated using a 6 mm diameter orifice placed near the sidewall. A schematic representation of the experimental setup is shown in Fig. 1. A continuous air supply with a rotameter is attached to the orifice to

575

ellipse onto the circle. Their contours are extracted and measured by calibrating the image pixel length to millimeters.

generate a chain of bubbles (bubble train) using different air injection rates ranging from 0.1 to 0.5 1pm. These flow rates (of air) are generally very high for bubble formation in water and very low in viscous media like glycerol. Initially, the sidewall distance was chosen to be twice the size of orifice diameter where the wall effects are negligible. Provisions were also made to decrease the sidewall distance from the bubble train. Diffuse light source

4.

-

Results and Discussions As has been reported by many authors, the shape of the bubbles is very different with fluid property and the rate of injection. But in all non-Newtonian fluids, bubbles evolve from spherical shape to the teardrop shape when it detaches from the bottom surface [18, 19]. During this transition (from spherical to teardrop shape), bubbles exhibit different shapes as well as growth rate[20] depending upon the properties of the surrounding fluid. In the present work, five different flow rates of air (0.1 1pm to 0.5 1pm) are considered to create a chain of bubbles (bubble train) near a solid wall. At such injection rates of air, the deformation of the bubble and the flow pattern of the surrounding fluid would be completely different due to the viscosity of the fluid. In the case of water, the bubble deformations are accompanied with wobbling and oscillation exhibiting typical shape characteristics of the high Reynolds numbers regime. In the highly viscous situation like glycerol, the deformations are negligible at the initial growth period as the flow is in the low Reynolds number regime. In the following table the properties of glycerol and water are given:

High Speed Camera and IDA measurmenft

ΗΠΙΠ

%

SIOEV1ΙW x & y axes control

Image Processing

Figure 1. Experimental setup

The shape of bubble was studied using a high speed camera (FASTCAM-APX RS 250KC) at different wall offsets (WO). The camera was put on an automatic leveling shock-proof platform with precise x and y axes control. This enabled us to minimise the parallax error to a large extent. The continuous flow of bubbles was photographed at 5000 frames per second (fps) and images were processed using classical image-processing algorithms. The major source of uncertainty in the bubble diameter is the definition of the threshold in the setting of the digital image system. However, this error is a decreasing function of increasing bubble diameter. Bubble diameters in this study are fairly large (> 7 mm dia.) and thus the measurement of bubble dimensions are within reasonable accuracy.

(kg.m-3) Glycerol (99%) Water Electrolyte for H-H cell

b

=

Jfb(x,y)dxdy

where this integration is over the whole image 7. The position of the centre of bubble in each frame can be obtained as follows:

Ab U

Ab

J

y = — JJyb(x, y)ixdy = — Jxh(x, y)dx

(m2/s)

1.2 le-3 2.5e-3

9.5e-4 le-6 1.2e-6

62 72 136.0

V

1260 1000 2100

Figure 2 shows the teardrop shape of a detached bubble at sidewall distance of 12.5 and 6.5 mm for the experiment using glycerol. The symmetry-axis remains along the direction of bubble motion at WO of 12.5 mm showing negligible interaction. Bubble tends to incline towards the solid wall, as side channel width (wall offset) reduces showing the departure of symmetrical shape, which may be attributed to the pressure gradient causing fluid to move towards the base of the bubble from the wall. In glycerol, the onset of departure from the symmetric line is observed even before the detachment of bubble (Figure 2b). At a wall offset of 6.5 mm, the mass flow between the wall and bubble prevents the bubble from contacting the wall. The bubble elongates due to the shear force that it generated in the gap which is an important observation of this study.

(1)

x = — f fxb(x, y)dxdy = — fxw(x, y)dx

(Pas)

Surface Tension σ (mN.ni·1)

Analysis of shape of the bubble

The images from high speed camera are analysed binary and/or edge bubble images. Based on binary image-processing techniques, in which a binary image is defined by a characteristics function b(x,y) that takes on the values 0 or 1 and the area of the bubble or the binary image is: A

Table 1. Fluid Properties Density Viscosity

Fluids

(2)

ä^ütttefe H|ttt#i* 535 TÜ ■ IXltiiär ,-C\ · InS 11Mirti 1 [ i ^ Ί Ä '■■ fSf M! Tvtf H 1 7 n M ] T ttiifcfcfc Hfl I x i i n x t : 47 \\ Ð ft i l l 1 l\y : ' t } m j | Wem I | }1 FM 1 { | 1 LJJL : ffljflj 11 ftI 1 1 \ 1 iXt j ^\Li~ T! H I IÄJ 1 ðæñøß÷ : .4^.vi„4^i™Lv.i.J.~, ' lll+Nrh '"["'["'{fll H 4444444r llilJ4^ It PNi I I I 4M44 ...1...

(3)

| M |

Where w(x,y) and h(x,y) are the horizontal and vertical projections of the bubble on the x-axis and y-axis.

il il

1Wall offset = 12.5 mm

Matlab Image Processing toolbox is used to calculate area, centroid position, perimeter of the bubble in each time frame [16, 17]. The eccentricity and orientation are calculated by fitting an

1

1

r

I lil-UIWall offset = 6.5

Figure 2. Bubble shape in glycerol at the time of detachment

576

When the bubble begins to rise owing to buoyancy, the pressure gradient at the lower surface of the bubble is higher than that at the top surface, which deforms the lower surface significantly. The contour plots in figure 3 clearly show that the distance between successively images decreases with time indicating an initial deceleration. The initial, higher velocity created by air injection is affected by the viscosity of the fluid.

The reference frame is one when the bubble is just detached from the base of the tank (tref = 0) which is shown in all the figures given below. In figure 3, the bubble trajectory is plotted sequentially at different sidewall offset for water and glycerol media respectively. Bubble contours are plotted at 40 frames apart, i.e. at 8 ms time interval. In glycerol, the bubbles move upward steadily as shown in Fig. 3a without any wobbling giving rise to a trajectory of rectilinear path[21].

Figure 4 and 5 shows the contours of bubbles at a lesser time interval for the case of water to bring out the finer changes in shape and orientation due to the presence of the solid wall. Each row of the figure 4 and 5 shows the first 15 frames at 2 ms interval right after the detachment for WO of 12.5 and 6.5 mm respectively.

The velocity of the fluid in the gap (between wall and bubble interface) increases with decreasing side channel width. The imbalance between the velocity of the fluid in the gap and the surrounding fluid gives rise to an asymmetric flow pattern at the trailing edge of the bubble. At a wall offset of 6.5 mm (Fig 3a-III) the bubble not only changes it shape and orientation but also elongates significantly due to the shear force arising from the high velocity within the gap. The initial deformation of bubble in both water and glycerol is observed due to the induced wall migration forces indicating dominating effect of the wall in general. The bubble is subjected to both rotational and translational displacement due to the presences of asymmetric forces caused due to the presence of a solid wall in the near vicinity. In the case of the less viscous fluid, water, shape deformation and oscillation becomes a dominating mode due to the pressure gradient in the surrounding fluid. In this case, bubbles rise upward in a jig-jag trajectory pattern. The general features of the rise of a gas bubble in a liquid, and its distortion at a high Bond number and Reynolds number are discussed elsewhere [6].

I. WO = 12.5 mm

II. WO = 9.6 mm

Figure 4. Firstfifteenconsecutiveframesat 2 ms apart for WO of 12.5 mm and 0.5 1pm gas injection rate in water.

III. WO = 6.5 mm Figure 5. First fifteen consecutive frames at 2 ms apart for WO of 6.5 mm and 0.5 1pm gas injection rate in water. The initial contours (first 5 frames) are quite symmetrical when the wall offset is twice the size of the bubble diameter in all sizes of bubbles. The bubble loses this symmetric nature as it rises further along the plume. It is to be noted that the change in shape of bubble, no matter how small it is, may initiate a lift force along with other migration forces. In the case of larger bubble the asymmetric shape is more apparent; this may be due to the higher bubble induced flow. When the bubble is closer to a wall the mass flow in the gap is much higher than in the surrounding. The tip tends to elongate as shown in the figure. This downward flow (in the gap) does not continue further going downward it rather tends to follow the bubble as a negative pressure field is created behind the bubble by the bubble deformation itself.

Figure 3. Bubble contours at 8ms interval of time in (a) Glycerol and (b) water respectively at 0.21pm

577

The aspect ratio (AR) and other parameters of the image are calculated using following equations.

published data [13, 24]. The aspect ratio are calculated (equ. 4 and Fig. 6) for the bubble in each time frame starting from the reference frame. A high aspect ratio at initial position (tref = 0) in figure (Fig. 8) indicates a teardrop shape of bubble in glycerol [18]. The rate of change of aspect ratio with time is very slow due to the viscosity of the glycerol. The influence of presence of the sidewall in the vicinity can be realised as the bubble evolve from bottom of the tank (tref=0, Fig.2b and Fig. 5). And also, infigure8 the percentage change in aspect ratio is nearly 10% due to the sidewall. The aspect ratio (AR) of the bubble is always higher at lower channel width indicating significant elongation bubble tip (Fig. 5). The bubble tip elongates due to the shear force that generated between the wall and bubble.

(4)

w

The rise velocity (v) is calculated from the mass centre (x and y coordinates) of the bubble over two consecutive frames (i, i+1) as:

V =

V(xi+i-xi)2+(yi+i-yi)2 At

(5)

100

It is to be noted that in 5000 frames per second data acquisitions, At is equivalent to 0.0002 s.

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Figure 6. The bubble (at t = 14ms) and the best fitted ellipse

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d e =(hw 2 ) 1 / 3

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The bubbles have a visible indentation or a dimple at the base right after the detachment causing a shape transformation from spherical to asymmetric toroidal shapes within a small span of time. The shape change is obviously a manifestation of toroidal wake accompanying the bubble. This is the characteristic of bubble shapes at high Reynolds number flow. Thus, it becomes difficult to predict the volume from the 2-D projected images using equ. 7. However, the projected area of the image may be a better representation of the bubble shape during initial growth. Figure 7 represents the projected area vs. time in glycerol showing significant changes in shape with the decrease in wall offset.

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Figure 8. Aspect ratio vs. Time for bubble in glycerol Bubbles in a water column with similar flow rates behave differently (Fig. 9). The bubble changes its teardrop shape (prolate) to spherical shape, within 3 ms of time after the detachment, and then to ellipsoidal and toroidal capped bubbles (oblate) as it rises further [19]. The width of the bubble changes

Figure 8 plots the variation of aspect ratio with time for the bubble in glycerol and found to be in good agreement with previously

578

produced due to the bubble departure from symmetric shape to asymmetric one. Furthermore, the predicted data of various operating conditions will be validated using Particle Image Velocimetry (PIV).

significantly when compared with glycerol. This can be attributed to the lower viscosity of water. If one compares the results in terms of aspect ratio between glycerol (Fig. 8) and water (Fig. 9), it can be seen that the influence of sidewall on the geometric properties and velocity of the bubble can be measured at higher viscosity fluid but very much unpredictable in a lower viscosity fluid like water. The unpredictable influence of the sidewall in water is mainly attributed to the oscillation in the bubble-fluid interface. But the influence of sidewall cannot be ignored. The wall migration force changes the shape of bubble which is the main cause of wall induced lift force. This, in turn gives rise to a bubble trajectory path and bubble induced flow in the surrounding fluid. This phenomenon has a direct implication in Hall-Hιroult cell where the bubbles near the vertical side of carbon anode would experience similar forces and can contribute significantly to the circulation flow patterns that occur in the side channel of the electrolytic cell. With reference to Table 1, the kinematic viscosities of water and the electrolyte of the Hall-Hιroult cell are of similar magnitude. Thus if Reynolds Number is the major influencing factor, the results found for water may be assumed to be applicable. Further experimental and modeling work is required to fully quantify these effects. 1.3

1.2

References 1. S. 0. Unverdi, "A front-tracking method for viscous, incompressible, multifluid flows," Journal of Computational Physics, 100(1) (1992), 25-37. 2. W.J. Rider, and D.B. Kothe, "Reconstructing Volume Tracking," Journal oj Computational Physics, 141(2) (1998), 112152. 3. N. Takada, et al., "Numerical simulation of two- and threedimensional two-phase fluid motion by lattice Boltzmann method", Computer Physics Communications, 129(1) (2000), 233246. 4. S. Osher, and R.P. Fedkiw, "Level Set Methods: An Overview and Some Recent Results," Journal of Computational Physics, 169(2) (2001), 463-502. 5. J. Hua, J.F. Stene, and P. Lin, 'Numerical simulation of 3D bubbles rising in viscous liquids using a front tracking metho," Journal of Computational Physics, 227(6) (2008), 3358-3382. 6. L. Chen, et al., "The development of a bubble rising in a viscous liquid,"Journal ofFluid Mechanics, 387 (1999), 61-96. 7. W. Yang, and M.A. Cooksey, "Effect of slot height and width on liquid flow in physical models of aluminium reduction cells," TMS Light Metals, (2007), 451-456. 8. K. Sakakibara, et al., "Measurement of the surrounding liquid motion of a single rising bubble using a Dual-Camera PIV system," Flow Measurement and Instrumentation, 18(5) (2007), 211-215. 9. Z, Liu, and Y. Zheng, "PIV study of bubble rising behavior," Powder Technology, 168(1) (2006), 10-20. 10. A. Tokuhiro, et al., "Turbulent flow past a bubble and an ellipsoid using shadow-image and PIV techniques," International Journal ofMultiphase Flow, 24(8) (1998), 1383-1406. 11. A. Fujiwara, D. Minato, and K. Hishida, "Effect of bubble diameter on modification of turbulence in an upward pipe flow," International Journal of Heat and Fluid Flow, 25(3) (2004), 481488. 12. I. Chakraborty, I., et al., "Computational investigation on bubble detachment from submerged orifice in quiescent liquid under normal and reduced gravity," Physics of Fluids, 21(6) (2009) 13. C. Chen, and L.S. Fan, "Discrete Simulation of Gas-Liquid Bubble Columns and Gas-Liquid-Solid Fluidized Bed," AIChE Journal, 50(2) 2004, 288-301. 14. J. Magnaudet, S. Takagi, and D. Legendre, "Drag, deformation and lateral migration of a buoyant drop moving near a wall," Journal ofFluid Mechanics, 416 (2003), 115-157.

• 12.5 mm I

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. 6.5 mm

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This work was supported by the Breakthrough Technologies for Primary Aluminium Research Cluster - a collaborative research effort between CSIRO Light Metals Flagship, Swinburne University of Technology, University of Auckland, University of New South Wales, University of Wollongong and the University of Queensland, focused on developing technology to lower the overall energy consumption associated with primary aluminium production.

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Acknowledgements

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Figure 9. Aspect ratio vs. Time for bubble in water Conclusion The effect of sidewall on the shape of the bubble was studied experimentally in both water and glycerol. Different flow rates of air were used from the bottom of the tank to generate chains of bubbles. In general the results indicated that the bubbles lost their symmetric shape right after the detachment when sidewall was at a close proximity. The bubble also showed tendency to elongate as the side channel width decreased. The data presented in this paper by our visualization technique and image processing was sufficiently accurate to allow us to predict not only the velocity but also the shape as well the orientation of the bubble. In future we are planning to analyze the complete hydrodynamic field including wall induced lift force that

579

15. K, Lunde, and R. J. Perkins, "Shape oscillations of rising bubbles," Applied Scientific Research, 58(1-4) (1997), 387-408. 16. C. Beggan, and C.W. Hamilton, "New image processing software for analyzing object size-frequency distributions, geometry, orientation, and spatial distribution," Computers & Geosciences, 36(4) (2010), 539-549. 17. C. Igathinathane, et al., "Shape identification and particles size distribution from basic shape parameters using Image J," Computers and Electronics in Agriculture, 63(2) (2008), 168-182. 18. D. Funfschilling, and H.Z. Li, "Effects of the Injection Period on the Rise Velocity and Shape of a Bubble in a Non-Newtonian Fluid," Chemical Engineering Research and Design, 84(10) (2006), 875-883. 19. W. Fan, et al., "Study on bubble formation in non-Newtonian fluids by laser image technique", Optics & Laser Technology-,40(2) (2008), 389-393. 20. A. Vazquez, et al., "A look at three measurement techniques for bubble size determination," Experimental Thermal and Fluid Science, 30(1) 2005, 49-57. 21. G. Brenn, V. Kolobaric, and F. Durst, "Shape oscillations and path transition of bubbles rising in a model bubble column," Chemical Engineering Science, 61(12), 2006, 3795-3805. 22. W. Kracht, and J.A. Finch, "Effect of frother on initial bubble shape and velocity," International Journal of Mineral Processing, 94(3-4) (2010), 115-120. 23. Y. Saito, et al., "Shape measurement of bubble in a liquid metal," Nuclear Instruments and Methods in Physics Research Section A: Accelerators, Spectrometers, Detectors and Associated Equipment, 605(1-2) (2009), 192-196. 24. R. Clift, J. R. Grace, and M. E. Weber, "Bubbles, drops and Particles,". 1978.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

LARGE GAS BUBBLES UNDER THE ANODES OF ALUMINUM ELECTROLYSIS CELLS Alexandre Caboussat1'2'4, Laszlσ Kiss3, Jacques Rappaz4, Klβra Vιkony3, Alexandre Perron5, Steeve Renaudier6, Olivier Martin6 ^coor Systems SA, Technopτle 10, 3960 Sierre, Switzerland university of Houston, Department of Mathematics, 4800 Calhoun Rd, Houston, TX 77204-3008, USA Departement des sciences appliquιes, Universitι du Quιbec à Chicoutimi, 555 Boulevard de l'Universitι, Chicoutimi, Canada G7H 2B1 4 Chaire d'Analyse et Simulation Numιriques, Ιcole Polytechnique Fιdιrale de Lausanne, 1015 Lausanne, Switzerland 5 Rio Tinto Alcan, CRDA, 1955 Boulevard Mellon CP 1250 Jonquiθre, Canada 6 Rio Tinto Alcan, LRF, BP 114, 73303 Saint-Jean-de-Maurienne Cedex, France Keywords: aluminum electrolysis, PIV, numerical simulations, two-phase flow, large bubbles, Fortin bubble. notion of bubbles. The present paper deals with these so-called Fortin bubbles, studying their morphology and behavior both by experiments and by mathematical modeling and numerical simulations.

Abstract The gas bubble laden layer under the anodes during the electrolysis of alumina plays an important role in the hydrodynamics and the voltage balance of the reduction cells. Under certain geometrical and operational conditions, very large gas pockets, in the order of hundred cubic centimeters can be formed. The particular shape of these large gas bubbles was first described by Fortin et al in 1984 [1]. In the present paper the results of a combined experimental and numerical approach are described. In the experiments, the shape and the kinematics of the Fortin bubbles were analyzed by videography and Particle Image Velocimetry (PIV). A finite element method (FEM) combined with a Volume of Fluid (VOF) method was used to reproduce the experimentally observed phenomena, with particular attention to the reduction of the numerical diffusion of the liquid-gas interfaces. The morphology of the large bubbles and their movement including the velocity field around them are described.

Experiments Experimental setup Large bubbles under the anodes were studied using lowtemperature hydrodynamic models. Two different setups were constructed. In the first, the large bubbles were produced by coalescence under a porous plate that represents the anode bottom. Such an arrangement produces a bubble layer and bubble driven flow similar to that in real reduction pots, but it does not permit studying the hydrodynamics of the large bubbles separately from other effects as the continuous interactions with smaller bubbles, the fluctuating velocity of the bubble layer, etc. For this reason, a second experimental rig was developed that makes possible to study of the morphology and kinematics of the Fortin bubbles in its pure form. In the experiments presented here, water and air were used as fluids. The bubbles were produced one-by-one under the lower end of the inclined solid surface (Figure 1), by using the so-called inverted cup technique that permits to produce bubbles precisely with the desired volume of the gas.

Introduction The gas bubbles generated during the electrochemical reduction of aluminum occupy a part of the horizontal inter-electrode space. Being electrical insulators, these bubbles induce a parasite voltage drop and cause an increase of the energy consumption of the cell. The electrical resistance of the bubble-laden layer under the anode is influenced by the quantity of the gas (volume fraction, covering factor) and by the morphology of the bubble layer. The electrical resistance of the molten electrolyte follows the fluctuations provoked by the random events of bubble nucleation, coalescence and escape of the bubbles [2]. In the past decade, significant efforts were invested into the clarification of the mechanisms of bubble-induced overvoltage by several researchers (see a review in [3]). The electrical resistance of the bubble layer is considered by most of the researchers as an equivalent resistance of a heterogeneous medium with the electrolyte as the continuous conducting phase. The classical approach is that of Maxwell, who gave the first expression to calculate the equivalent (sometimes called "apparent" or "bulk") conductivity for a conducting medium with sparsely distributed spherical, identical isolating voids [4,5]. Later, Bruggemann developed a formula for spheres of nonuniform diameters [6], but still for voids that are sufficiently far from each other compared to their diameters. Physical modeling in laboratory, as well as numerical simulations show [1,7] that the real situation is far from the case of sparsely distributed spherical bubbles. Under the anodes with dimensions common in today's industry, the growth by coalescence produces very big bubbles with shapes that are not similar to the usual

600 mm

Figure 1. Schema of the experimental setup In order to follow the movement and shape evolution of the bubbles, they were recorded by high-speed video cameras. Images were taken from the side, using a rail mounted, moving camera, as well as from the bottom by the help of an inclined mirror under the water pool. Image processing was extensively used for extracting quantitative information about the shape and velocity of the bubbles. The necessary algorithms and codes were developed by using MATLAB®. Different edge-detecting algorithms were tested to

581

follow the contour of the Fortin bubbles. An illustrative example is shown below in Figure 2.

the image, the magnitude of the velocity distribution is shown by superposing a colored contour plot of the magnitude of the velocities on top of the vector plot.

Figure 2. Image before and after using the threshold filter Particle image velocimetry Figure 5. Mean bath velocity close to the side channel at 2° of angle of inclination

Besides the analysis of the morphology and movement of the large bubbles, the velocity field in the liquid phase was also studied quantitatively by using particle image velocimetry (PIV).

Morphology of bubbles Since the first description of the morphology of large bubbles by Fortin et al. [1], its shape is characterized mostly by the side view. It is particular that around the leading edge of the moving bubble a so-called "head" is formed, which has a penetration depth at least twice that of the static or slowly moving bubbles. This is a typically dynamic phenomenon - stationary bubbles never have this head, independently of their volumes. Another interesting feature of the shape of Fortin bubbles is that in the plane of the solid surface (see right view of Figure 6) the leading edge takes a nearly perfect circular form. In three dimensions, the surface of the Fortin bubble is complex, the typical contour, as it is shown in Figure 4, is a longitudinal, central section. When filming laterally, this central section is mostly hidden, depending on the method of illumination (see the left image in Figure 6).

Figure 3. Using PIV in a two-anode hydrodynamic model. The measurement is illustrated in Figure 3 for a two-anode arrangement, but the principle is the same for the single bubble setup shown in Figure 1. The flow is seeded with small tracer particles whose settling velocity is negligible. A pulsed laser sheet illuminates the selected planar section of the fluid volume and a camera records consecutive images of the tracer particles. Two of these images are used to determine the velocity vectors in the illuminated plane.

Figure 6. Side and bottom views of a 150 ml Fortin bubble, 4° inclination Figure 4. Instantaneous velocity field around a Fortin bubble

The head of the Fortin bubble follows the curvature of the leading edge in the plane of the solid surface; its shape reminds of a croissant cut in half. Analyzing many Fortin bubbles, the shape was approximated by a schematic geometry as shown in Figure 7. The cross section of the head is described by a parabola that follows the circular arc of the leading edge.

Both instantaneous and time-averaged velocity fields were determined. The example in Figure 4 shows an instantaneous velocity pattern around a large Fortin-bubble whose contour is marked by the dashed line in the image. Figure 5 was obtained by averaging instantaneous velocity distributions obtained during the passage of several bubbles. In

582

volume-of-fluid approach is used to track the liquid domain and the free surface. Let φ be the characteristic function of the liquid domain (φ equals one if liquid is present, zero if it is not). In order to describe the kinematics of the free surface, φ satisfies the following transport equation in the cell A:

mean height of tail Ht

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where v denotes the velocity of the two-phase fluid in A. The incompressible Navier-Stokes equations are assumed to hold in the liquid domain Qt (varying in time), which implies that the velocity v and the pressure p of the liquid satisfy, in Qt, the equations of conservation of mass and momentum:

height of head

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Figure 7. Analytical approximation of the geometry of a Fortin bubble Using the notation of the schema, the volume of the gas in the bubble can be calculated according the formula below v y

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A two-phase flow model and an Eulerian algorithm are proposed for the numerical simulation of the hydrodynamics of large bubbles under an inclined plane. This numerical model is inspired from [8,9]. Unlike in [10,11], where the complete magnetohydrodynamic problem is considered, this model focuses on the hydrodynamics of a single large bubble only and is used to describe the morphology of these bubbles. Physical model Let A be the cavity (aluminum cell, more precisely the electrolytic bath) in which the fluid is confined, and let T > 0 be thefinaltime of simulation. For any given time t, let Qt be the domain occupied by the liquid and let Tt be the free surface defined by 5QtV5A. The notations are shown in Figure 8 for a large bubble initially located under an inclined plane, under the effect of gravity forces.

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where σ is the (constant) surface tension coefficient between the liquid and gas, ê is the curvature of the interface, and n is the normal vector to the interface, oriented outside the liquid domain. Appropriate essential boundary conditions on the boundary of the cell dA and initial conditions on the bubble position and velocity are added to close the mathematical model. Note that the numerical method does not include any representation for the mechanism of the nucleation of bubbles.

We advocate a numerical method based on a splitting algorithm for the time discretization, and a two-grid method for the space discretization. The splitting algorithm allows decoupling the physical phenomena and solving each of them sequentially at each time step. It is illustrated in Figure 9. At each time step, two advection problems are solved first, leading to a prediction of the new velocity, together with the new approximation of the characteristic function of the liquid domain. They consist in solving

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This allows determining the new liquid domain, the new gas domain and the new liquid-gas interface. Then the bubbles of gas are tracked with an original numbering algorithm (see [8,9]) and a

Compressible gas bubbles in the incompressible fluid in modeling represent the bubbles nucleating during aluminum reduction. A

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constant pressure is computed in each bubble by applying the ideal gas law. The surface tension effects are taken into account by computing an approximation of the curvature of the interface together with the external normal vector, by using the so-called continuum surface force model. Finally, a Stokes problem is solved, in the liquid domain only, to compute the final velocity in the liquid and its pressure. The corresponding equations are:

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Numerical setup and morphology indicators for Fortin bubbles Numerical simulations are undertaken in order to compare the experimental results in [13,14,15,16] with those obtained by mathematical and numerical modeling. Such comparisons are achieved thanks to a set of indicators describing the shape and morphology of the Fortin bubble under an inclined anode. These indicators are namely the length of the Fortin bubble L, the maximal height of the Fortin head Hh, the maximal length of the Fortin head Lh, the mean height of the bubble Ht and the width W of the bubble in a plane parallel to the surface of the anode. The notation is detailed in Figure 7. The mean height is measured at the middle of the Fortin bubble.

Surface tension Diffusion

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Figure 9. The splitting algorithm decouples the physical phenomena: From left to right: transport equations, tracking of bubbles and computation of gas pressure, computation of surface tension effects, and diffusion equations (Stokes problem).

Several configurations for the Fortin bubble and the inclined anode have been considered in [16]. We consider here inclination angles of 1, 2, 4 and 8 degrees, and bubbles of different sizes. Experiments in [16] have been undertaken with bubbles of volumes 50, 150 and 250 ml. The geometry and dimensions of the computational domain of the corresponding numerical simulations is a subset of the one illustrated in Figure 1. The anode-cathode distance (ACD) is set to approximately 4 cm. The density, viscosity and surface tension coefficients are those of the water and water-air interface and respectively given by p=1000 kg/m3, μ=0.001 kg/(ms) and σ=0.0728 N/m. The gravity forces are g = -9.81 ez m/s2 (oriented vertically).

A two-grid method is used for the spatial discretization, as illustrated in Figure 10: a regular grid of small cells is used to solve the transport equations (left), while the solution of the Stokes problem is performed on an unstructured finite element mesh composed of tetrahedrons (right). The solutions are interpolated at each time step between the two grids with projections operators. In order to reduce numerical diffusion effects of the free interface motion and to balance accuracy and computational cost, the structured grid is typically five to ten times finer than the finite element mesh.

Numerical Results

A method of forward characteristics allows to obtain piecewise constant approximations of a prediction of the velocity and of the volume fraction of liquid on the structured grid of small cells. It is coupled with geometric reconstruction algorithms of the liquidgas interface to reduce numerical diffusion. Adaptive subdivision techniques for the finite element mesh have been proposed in [12] to further reduce the numerical diffusion.

An illustration of the Fortin bubble obtained by numerical simulation is given in Figure 11. A single bubble of initial volume 250 ml evolves under an inclined plane with an angle of 8 degrees. It illustrates a three-dimensional contour of the bubble (A), a cut in the middle vertical section, showing the typical Fortin head (B), a cut along an inclined plane parallel to the anode surface (C), and a representation of the mesh that is purposely finer in the layer under the anode where the bubble evolves. One can observe that small residual parts of the large bubble are left behind during the process, similarly to the behavior found experimentally.

Stabilized (Galerkin least-squares-style) piecewise linear finite elements, based on piecewise affine finite element approximations of the velocity and pressure, are used for the approximation of the solution of the generalized Stokes problem. The tracking of the bubbles and the computation of their pressure has been described in [9]; it is achieved by solely solving Poisson problems with piecewise linear finite element approximations. The approximation of the surface tension effects has been detailed in [8]; it relies on variational arguments for the computation of the mean curvature of the interface. The complete algorithm is an order one algorithm in space and time.

584

(A)

(B)

Figure 13 illustrates the other morphological indicators, namely Hh, Lh and Ht for no-slip boundary conditions only. Conclusions are similar, as the match between simulations and experimental data is consistent. Sources of errors include measurement issues, but also discrepancies in the exact volumes of bubbles (as the volume of the simulated bubble is slightly larger than 250 ml), differences of ACD distances and water height, and questions related to the influence of the meshes. Maximum height of the head

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Figure 11. Illustration of a Fortin bubble of volume 250 ml under an inclined plane of 8 degrees. (A) contour of the bubble; (B) cut in the middle vertical plane; (C) cut in the plane parallel to the inclined anode; (D) illustration of the 3D mesh used for the numerical simulation. Comparisons and discussion A comparison of the morphology of the Fortin bubble between the numerical results and the experimental ones is based on the indicator variables L, Hh, Lh, Ht and W. We consider first a bubble with initial volume 250 ml and various inclination angles for the anode. We compare the length L and width W of the bubble. Results are illustrated in Figure 12. Experimental results are taken from [16]. Numerical results are obtained by applying either slip or no-slip boundary conditions on the surface of the anode. When enforcing slip boundary conditions on the anode, the bubble can slide on its surface; when enforcing no-slip boundary conditions, the bubble sticks to the surface.

angle of inclination [degree]

Figure 13. Comparison of the maximal height of the Fortin head Hh, the maximal length of the Fortin head L^, and the mean height of the bubble Ht for a bubble's volume of 250 ml and various inclination angles. In a second step, we consider an inclined angle of 8 degrees and various volumes for the bubble, ranging from 140 ml to 300 ml. Figure 14 shows comparison of the results for the quantities L, Hh, Lh and W. Simulations and experiments exhibit the same trends for the morphology of the Fortin bubble.

Bubble length

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These numerical results also express the physical instability of the Fortin bubble, which is similar to the one of a rising buoyancy bubble (see for instance [8,17] and references therein).

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Although not quantified here, the measurement errors are relatively large for both the experiments and the simulations. We can observe that the length of the simulated bubble is slightly overestimated, while its width is underestimated, indicating that the numerical solution may not be completely stationary. Note also that the influence of the type of boundary conditions is of the same order as the error on the measurements.

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Masserey, Dr G. Steiner (Ycoor Systems SA) for fruitful discussions. The first author, A. Caboussat, gratefully acknowledges the financial support of the Mathematics Institute of Computational Science and Engineering (EPFL) during his sabbatical leave. The fourth author, K. Vιkony, expresses her thanks for the financial support of the government of Quebec (FQRNT) during her PhD studies.

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References

Volume of bubble (mfj

Height of the head Hh Length of the head Lh Figure 14. Comparison of morphology indicators between experiments and simulations for an inclined plane with an angle of 8 degrees and various bubble volumes. Finally a comparison between the numerical results and the velocity data obtained with particle image velocimetry (PIV) techniques is briefly sketched. The quantitative interest is in the terminal velocity of the bubble (in the quasi-stationary state of the Fortin bubble). For a Fortin bubble with volume 250 ml and an inclination angle of 8 degrees, the terminal velocity obtained by PIV and reported in [16] is given by 32 cm/s. In the numerical simulations, the bubble velocity is computed by considering the displacement of the bubble in a given time interval. Simulations lead to a terminal velocity in the range 28-35 cm/s for volumes between 250 and 290 ml. Instantaneous velocities in a cut in the middle of the bubble are illustrated in Figure 15. Note that the mathematical model does not compute a velocity inside the bubble, and therefore the velocity in the gas phase is set to zero. These preliminary results allow us to conclude to the consistency of the velocity values between experimental and numerical results, and will be detailed in future works.

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Figure 15. Instantaneous velocity field around the Fortin bubble obtained by numerical simulation. Conclusions Large bubbles under an inclined anode have been studied both experimentally and numerically. These bubbles are formed during the aluminum electrolysis. From the experimental point of view, videography and particle image velocimetry (PIV) allow to describe the morphology of Fortin bubbles and terminal velocities. From the numerical viewpoint, a volume-of-fluid method and finite element approximations have been used to reproduce experiments. Comparisons of the results have exhibited similar morphologies of bubbles, comparable quantified results and similar trends in the morphology of bubbles, when varying the inclination angle of the anode or the volume of the bubble. Acknowledgments This research project is supported by Rio Tinto Alcan. The authors particularly thank Prof. M. Picasso (EPFL), Dr A.

586

[I] S. Fortin, M. Gerhardt, and J. A. Gesing. Physical modeling of bubble behavior and gas release from aluminum reduction cell anodes. TMS Light Metals, pp. 721-741, 1984. [2] M. V. Romerio, A. Lozinski and J. Rappaz. A new modeling for simulating bubble motions in a smelter. TMS Light Metals, pp. 547-555, 2005. [3] M.A. Cooksey, M. P. Taylor and J.J.J. Chen, Resistance Due to Gas Bubbles in Aluminum Reduction Cells, Journal of Metals, pp. 51-57, 2008. [4] A. Perron, L. I. Kiss, and S. Poncsàk. Mathematical model to evaluate the ohmic resistance caused by the presence of a large number of bubbles in Hall-Hιroult cells. J. Applied Electrochemistry, 37, pp. 303-310, 2007. [5] J.C. Maxwell, A Treatise on Electricity and Magnetism, 3rd ed. New York, Dover Pub., 1954. [6] D. Bruggeman, Annalen der Physik, 24, pp. 636-664, 1935. [7] S. Poncsβk, L. Kiss, D. Toulouse, A. Perron and S. Perron, Size distribution of the bubbles in the Hall-Hιroult cells, TMS Light Metals, pp. 457-462, 2006. [8] A. Caboussat. A numerical method for the simulation of free surface flows with surface tension. Computers and Fluids, 35(10), pp. 1205-1216, 2006. [9] A. Caboussat, M. Picasso and J. Rappaz. Numerical simulation of free surface incompressible liquid flows surrounded by compressible gas. J. Comput. Phys., 203(2), pp. 626-649, 2005. [10] M. Flueck, T. Hofer, M. Picasso, J. Rappaz, and G. Steiner. Scientific computing for aluminum production. Int. J. Numer. Anal, and Modeling, 6(3), pp. 489-504, 2009. [II] M. Flueck, A. Janka, C. Laurent, M. Picasso, J. Rappaz, and G. Steiner. Some mathematical and numerical aspects in aluminum production. J. Sci. Comp., 43(3), pp. 313-325, 2008. [12] A. Caboussat, P. Clausen, and J. Rappaz. Numerical Simulation of Two-Phase Flow with Interface Tracking by Adaptive Eulerian Grid Subdivision, submitted to Int. J. Num. Meth. Fluids, 2010. [13] A. Perron, L. I. Kiss, and S. Poncsβk. An experimental investigation of the motion of single bubbles under a slightly inclined surface. Int. J. Multiphase Flow, 32, pp. 606-622, 2006. [14] K. Vιkony and L. I. Kiss. Morphology of two-phase layers with large bubbles. Metallurgical and Material Transactions B, pp. 1-12, 2010. [15] K. Vιkony and L. I. Kiss. Velocity measurements in a real size model of an aluminum electrolysis cell model using the PIV techniques, submitted to J. Applied Electrochemistry, 2010. [16] K. Vιkony. Large bubble moving under a solid surface. Ph.D. thesis, Universitι du Quιbec à Chicoutimi, 2009. [17] S. Hysing, S. Turek, D. Kuzmin, N. Parolini, E. Burman, S. Ganesan, and L. Tobiska. Quantitative benchmark computations of two-dimensional bubble dynamics, Int. J. Num. Meth. Fluids, 60(11), pp. 1259—1288, 2009.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INITIATIVES TO REDUCTION OF ALUMINUM POTLINE ENERGY CONSUMPTION ALCOA POΗOS DE CALDAS/BRAZIL Andrι L.T. Abreu1, Mauro H. D. Salles1, Ciro R. Kato1 Alcoa; Km 10; Poηos de Caldas, MG, 37701-468, Brazil Keywords: Energy Consumption, Voltage, Process Management, Alcoa This means that energy consumption reduction can be achieved in either two ways:

Abstract Energy is one of the most important inputs for aluminum production and is responsible for approximately 40% of the cost of aluminum production (CAP) in Soderberg pots. Facing the 2008/09 global economic downturn, Alcoa Poηos de Caldas Plant, Brazil, has focused its efforts on a planned project, counting on its personnel's potential, to reduce the energy consumption. Main initiatives taken along this process were: workshops on energy (thermal balance and energy consumption), STAR Probe measurements (pot control focused on thermal balance), new cathode design, financial model development and changes in automatic pot control. Through this project a reduction of 77mV/pot and 0.30 kWh/kg AI were achieved, the best ever result reached at the plant at the present load level. In financial terms, in 2008/09, US$ 1 million was saved without any extra investment.

1) 2)

Reducing total energy use (Voltage/Pot); or, Increasing current efficiency (%).

The "Business Case" of this project was to focus on the first item. Fundamental effort was placed on CVD (Cathode Voltage Drop), AVD (Anode Voltage Drop) and Bath Voltage Drop, according to the voltage drops given in Figure 1.

Introduction

CVD 0.350V

Alcoa is a world leader in the production and management of primary aluminum, fabricated aluminum, and alumina combined, through its active and growing participation in all major aspects of the industry: technology, mining, refining, smelting, fabricating, and recycling (Alcoa Inc. Website, 2010). The increasing global competition gives a steady demand to reduce production costs. The electrolytic production of primary aluminum metal in Hall-Heroult cells is highly energy intensive and accounts for nearly 40% of the production cost (Tandon and Prasad, 2005). This means that energy saving should be the driving force to reduce production costs. To ensure Alcoa Poηos de Caldas' future as a competitive aluminum producer, the facility has constantly researched for solutions and development that guarantee the minimum energy consumption and maximum

Bemf* 1.650V

Total: 5.000V

* Bemf: When the pot line amperage is to as the battery effect or Bemf.

Bath Voltage 2.180V

i^k

led off the pots still ha\

Fig. 1 - Break up of total pot voltage (typical). The largest part of the pot voltage is the voltage drop of about 2.180 V in the molten bath between the carbon anode and the metal pool, which is essentially the cathode. The anode to cathode distance (ACD) is around 5 cm. Reducing ACD, to reduce pot voltage for decreasing energy consumption usually is not adopted in practice because of the danger of losing stability and workability of the cells and consequently the current efficiency. The voltage drop across the ACD depends upon bath resistivity, current density and bubble resistance (Tandon and Prasad, 2005).

benefit.

As the aluminum industry has faced the 2008/09 global economic downturn and any investments were postponed, Alcoa Poηos de Caldas has focused its efforts on a planned workforce, involving its personnel's potential to reach the best financial results in a short-term through energy saving, without any extra investment. The objective of this paper is to present the main initiatives to reduce energy consumption at Alcoa Poηos de Caldas, Brazil.

The main benefits of reduced voltage per pot are: Financial: energy saving; Market: different alternatives for use of the surplus energy, including availability to the electricity spot market; increased metal output by means of amperage increase without substantial increase in pot voltage. Sustainability: long-term energetic sustainability to the Company; Process: reduced bath temperature (since pot stability is maintained); reduced superheat (increase in potlife).

Background The energy consumption is related to the total energy employed (kWh) to produce a certain amount of molten metal (kg Al), i.e., kWh/kg AI. In other terms, as is derived from Faraday's law, the energy consumption for aluminum electrolysis is: 298.06 * Pot Voltage Current Efficiency (%)

587

Cathode Voltage Drop

Project Approach

0,500 -

In 2007, pot energy consumption at the Poηos de Caldas plant presented numbers that positioned the plant in disadvantage within Alcoa Soderberg Technology System. Thus, in 2008/09, a systematized approach based on Alcoa's personnel potential focused on energy reduction was initiated. What most supported this work was the synchronism between operational and process teams.

0,450 0,400 >" 0,350 -

3 0,300 -

v

?

\

1

I

V-

/

X

J

0,250

The first step was an "Energy Consumption Workshop". This initiative counted on representative of the Global Alcoa Technology Team (LTT). Through the action plan, several additional initiatives catalyzed reduction of energy consumption and can be divided into 4 major topics: 1) Anode Voltage Drop (AVD); 2) Cathode Voltage Drop (CVD); 3) Bath Voltage Drop; 4) Other Initiatives.

Main projects developed on this voltage drop were:

1) Anode Voltage Drop Changes in stub pulling reference and tolerance for stub set in range were implemented to reduce zero height towards the target, since the real results were consistently higher (Fig. 2). Actions taken: Stub in range limits squeezed from 300mm to 200mm; Tolerance of stub setting reference was changed from 0300mm than target to ± 100mm near the target.

• Superheat Workshop: aiming at better knowledge concerning the variables that compose the cell heat balance: voltage, bath chemistry and liquid levels control. From this point on, pot doctors started pot control taking into account superheat numbers. Calculated Superheat and STAR Probe (equipment developed by Alcoa Technical Center employed to measure superheat, temperature, alumina and bath ratio) were widely used and trends daily charted.

0,200

0,150

—Old Project

— N e w Project

Fig. 3 - Cathode Voltage Drop. 3) Bath Voltage Drop

• Reduction of Bridge Moves: fine-tuning of the anode bridge move time and step for lowers and raises resulted in a 40%decrease of anode bridge moves. Thus, pot stability was significantly improved making it easier to make voltage management. Moreover, such a change reduced bridge move travels by over 50% in pots out of control, bringing them into control faster (50% less time). • Limit for Noisy Pot: for computer control purpose, limits for noisy and quiet pots (extra resistance addition), were increased by 2 units. This adjustment diminished in 90% the total extra resistance for temporary instability control. Trial pots ran for over three months under these new parameters and it wasn't observed any significant current efficiency loss.

Fig. 2 - Illustrative sketch of changes in stub setting reference. Theses changes resulted anode voltage drop of 5mV. 2) Cathode Voltage Drop In this area, the main development was a new cathode design. This project, by itself, reduced cathode voltage drop by 47mV/pot (Fig. 3).

4) Other Initiatives • Financial Model: Calibration and application of a financial model developed by LTT. Basic premises of the model are energy price and LME (London Metal Exchange). Based on these principles the model was utilized as a guideline for the most profitable direction: Current efficiency vs. Voltage per pot (Fig. 4).

588

p^i RΔnnnnonn

CE x Voltage x Profit

Change in bath correction table, improving bath level in range from 88% in 2007 to 94% in 2009.

^ ^ ^ ^ ^ B

8.600.000,00

«

Implantation of fortnightly report of external losses.

H

1

8.200.WJ

,.«-*** ^

8.000.000X10

D

7.200£«,.00

• Process Training: three trainings applied for tapping and stubbing operators and pot tenders:

9X)00X»0 00

^ g ^ m

8.800.000,00

7.800.000,00 WEB - " £ 7.600.000JD0 V g H

H U ^ V MW

S P

It 8.400.000,00

8.600.000,00

% 8.200.000,00 8.400.000,00 m 8.000.000,00

Resistance Control Fundamentals;

8J000J0O0.O0

1

»C^

U

89.4

■

m 7.600.000,00 7300000,00

^Mf

^ H j ^

^ ^ ^ J ^ B

^

.s

S/.4UU.UÜU,UU

| ! 9 S.0O0

, 7.200.000,00 7400 £00,00

^

«7.000.000,00

438

Resistance Control and Graphs;

u.

(:E(%)

Basic Electrolysis Fundamentals.

/' 4,92

91,4 91.6

9 1

%

3

Voltage (V)

These actions are part of a large Process Management System and are considered truly important to keep a long-term improved performance in energy efficiency.

Fig. 4 - Surface graph from the financial model. • Energy Management: since this was a project with no technological investment, its structuring depended a lot on the staff creativity and innovation. Hence, all the ideas were supported by the ABS (Alcoa Business System) Methodology, mainly some stability tools, as follows:

Results The entire project was based on Process Management Tools and, therefore, it was not invested any money in it. As a result, in 2009 the best result ever was registered for the Alcoa Poηos de Caldas Plant since 2001 (from 2001 on, the load increase exceeded 120kA).

• Daily Management System (DMS) focused on resistance target: - Visual boards; - Resistance targets (periodically reduced); - Daily meetings focused on reduction of energy consumption. •

Voltage per pot was gradually reduced 77mV in 2009 (Reference: 2007). Figure 5 shows the voltage per pot evolution over 3 years. In 2009 the energy efficiency achieved was 0.30kWh/kg Al below the result recorded in 2007. The energy reduction provided a total profit of US$ 1 million.

Practical Problem Solving: Voltage breakdown study (comparison with the Alcoa Benchmark: Avilιs); Establishment of an action plan to reduce the voltage drop.

Voltage (V/Pot)

É

• Kaizen/Standardized Work: it was carried out a Kaizen for Spike Pulling Activity, which minimized deviations in this activity. A new procedure was developed for clamp voltage drop correction, where the operators adjust clamp drop as soon as they pull a set of spikes.

1° Energy Workshop Beginning of the Project

S§ §

• Process Management: focus on continuous improvement through the integrated actuation of teams of Resistance Control, Liquid Levels Control and Bath Chemistry with the mission of put under control and capable variables that affect directly energy consumption. Main developments were:

Fig. 5 - Voltage per pot evolution. Energy Efficiency (kWh/kg AI)

Quarterly, Cp and Cpk (process capability indexes which measure how close a process is running to its specification limits, relative to the natural variability of the process) are calculated for the most critical variables for energy (voltage, stability, AVD, CVD, External Voltage Drop and Clamp Drop), including control charts. Implantation of a new noise metric (MHD Noise), that has shown better correlation with magnetic stability, given that SPPN (peak-to-peak noise) has strong influence of bubble noise. Statistical analysis showed an optimum point for current efficiency.

Fig. 6 - Energy efficiency evolution.

589

Conclusion Given the world economic crisis, this project implementation has shown the many opportunities smelters can pursue in their own capabilities and personnel's potential. Through a well-organized and integrated Process Management System, by using the correct tools, the energy efficiency improved by 0.30kWh/kg Al, reaching the best results in recent times. Such initiative provided the company with a saving of US$ 1 million and the benefits of this energy reduction are many (sustainability, market, financial and process) References 1. Alcoa Inc. Website. Available in: Accessed in: 09/25/2010. 2. S.C. Tandon and R.N. Prasad, Energy Saving in Hindalco's Aluminum Smelter. Light Metal 2005, pp 303-309.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Overview of High-Efficiency Energy Saving for Aluminium Reduction Cell Xi Canming (Guiyang Aluminum Magnesium Design & Research Institute, Guiyang, 550081) differential prices of power in 8 industries, and a price markup separately to the confined enterprises and rejected enterprises by 0.10 Yuan and 0.30 Yuan. Further, Kyoto Protocal stipulates the validated age of C0 2 emission reduction by year of 2012 to 39 industrial developed countries. And, this the protocal may also be called for C0 2 emission imposition all-round to developing countries.However, the aluminium smelter produces C0 2 by 6750—7500kg per ton. In case of postulation that the sale price in Europe of certified volume of C0 2 emission reduction is €8.35 per ton and Chinese aluminium industry yields 20.62 million ton aluminium accepted, the C0 2 emission reduction fee will charge €1.162-1.291 billion.

Abstract : The paper indicates the potentiality of energy saving for aluminium reduction cell and comprehensive factors involved in execution of energy saving. Based on the application of recent developed higji efficiency technologies, the major features and differences are exposed. Furthermore, the paper also indicates the world advanced level of Chinese energy saving techonology in aluminium reduction. Key words : aluminium reduction cell, power cost, energy saving at all levels, material balance, electroheating balance, stability, current effciency

Aluminium industry, called "Electricity Guzzler" became the major object of state micro-contro since 2008 because its power consumption took up 5.51% of China. Although the average DC consumption of aluminium industry decreased from 14085kWh in 2001 to 13258kWh in 2008, the increasing power price made the average cost of power in aluminium industry reaching 44.7%, which exceeded the defined 25% warming line of power cost by 20% in developed countries(See table 1-1)

1 Preface Chinese aluminium industry suffers 3 restrictions of energy shortage, power hike, and environmental protection & emission reduction currently. In 2008, reserve of bauxite in China was about 3.223 billion ton, of which provides 0.529 billion ton mining, taking upl.96% of the world. On May 2010, NDRC , SERC and State Bureau of Energy jointly issued "Notification about favourable power price and other issues governing on high energy consumption enterprises", which prescribed

Calculated average production cost in 2008 of China

Table 1-1 Project Cost price.Yuan/t Indication of nsumption,

Manpower and depreciation

Anode Carbon

Cryolite

Al- fluoride

Comprehensive power consumption

2650

2750

5300

5600

0.43 kWh

—

1950

500

5

27

14323kWh

—

5168

1375

27

151

6159

800

9.98%

0.02%

1.10%

44.70%

5.81%

kg/t Cost of primary aluminium, Yuan/t Percentage

Table 1-2

Total cost

Alumi na

37.50

%

power price paid by aluminium enterprises of China in recent years

1378 I 0 100%

Unit : Yuan/kWh

Year

2002

2003

2004

2005

2006

2007

2008

2009

World I

Average price

0.306

0.316

0.350

0.351

0.392

0.422

0.433

0.461

0.172

Followed with state policy of coal-power linkage at the end of 2004, compounded by the increasing collision of "market-oriented coal" and "planning power", the regulation on structure of power vs reform on power price could be inevitable, in despite of being the largest customer of power generation plant, aluminium industry was involved in the

difficulty of surviving So, for purpose of self-saving, aluminium industry in China should greatly conserve the energy at first, especially develop the technology of "energy saving at all levels" to be the only option.

591

2 Track of techonolgy "energy saving development" on aluminium industry in China

592

Table 2-1 The following chart shows the track of techonolgy "energy saving development" in China: Energy saving pre-baked pot

Heat radiative pre-baked pot

"Tri-high" soderberg

Environnent protection pre-baked pot

DC consumption >14500kWh

DC consumption>12000j<[/yiT__ DC consumptions 13000kWh ">» ! Inertia anode Slotted anode \ / Graphitized carbon MHD, Low tern pera tu re N 3-variable Control Technology^ I self-control .eJec_trolYsis__ < ALCI3 \ Low anode e f f e c t / electrolysisy

r

Irregular cathode Low votage

High-efficiency fume treatment technology

Drained cell

3

Principle of energy saving in aluminium industry

V = 2.98- = 2 . 9 8 —

To have a better development and application of technology" energy saving at all levels" at alumiium cell, the necessity of analyzing the principle of energy saving is conducted.

(1)

Apparently, the way to save the energy is either decrease of resistivity R or increase of current efficiency η. To have a clean analysis of ptentiality of energy saving on aluminium cell,the voltage balance and energy balance of cell AP30 is adopted for illustration.

3.1 Cellentiality of energy saving in aluminium industry Principlly speaking, energy consumption in aluminium industry could be calculated by the following Ohm theorem and formula:

Busbar V-drop —

0.15V.«

Cathode V-drop

0.35V^

Anode V-drop

0.35V*'

AC distance

0.25V^

Bubble V-drop

Saving energy area AV=0.842Vv

1.334V*<

Bubble V-drop /

«*«, LüMCFv

0.032V

Anode current density

0.036V.

1.026V,

0.466V^

>

1.222V* W

CXPLH

2.248V.«

T

7 Ì : CE = 96%, V = 4, 19V, T = 960°Η, a - A 1 : 0 , £ 10%*< Table 3-1 : Pechiney voltage balance and energy balance of cell AP30 1: cathode concentration overvoltage 2. anode concentration overvoltage 3: anode overvoltage 4: anti-equilibrium potential 5: excessive A1F3 6. heat loss 2.013V

From table3-l, the basic voltage to keep the cell in production is: : VI + V2 + V3 + V4 + V5 = 3.348V VI: equivalent voltage

593

V2: bubble voltage V3: anode voltage V4: cathode voltage V5: busbar voltage However, the voltage of cathode-anode to keep the aluminium reduction production is:

each cathode-anode by 1cm accepated, the minimum ACD is equally to the theoretical value of 1.6cm. From table3-l, the voltage drop of electrolyte 1.334V (about4.45cm cathode-anode gap) requires 1.334 - 0.492 = 0.842V (equal to 2.8cm ACD ) for compensation of heat loss, formation of profile, and compensation of energy consumption at cathode-anode gap for fluctuation of metal. The movement of electrolyte and ACD could be indicated by the chart:

2.248- 1.756 = 0.492V In case of postulation that voltage drop 0.300V of electrolyte at

Concentration of turbulent layer 0.0'

Air film gap

Concenration of al reaction layer 0.05%

Heating ratiative gap

Concentration of laminar layer 0.1%

Fluctuating gap

Table3-2 : movement of electrolyte and ACD 1 ) Better the composition of electrolyte, and improve the conductivity;

Table 3-2 shows that , a minimum ACD , namely "distance of heat radiation" (could guarantee the normal production of aluminium reduction, and keep heat balance of cell), as well as"air film gap" required by bubble turbulent layer and "fluctating gap" caused by distorted metal (affected by magnetic field) and fluctuation. The analysis mentioned-above indicates the 4 ways to save the energy by alumium cell.

2 ) Take measures to releas the anode bubble and reduce the resistivity of air film; 3 ) Optimize the configuration of busbar, and improve the distortion and speed of circulatory flow of metal interface, as well as increase the statical stability of metal interface; 4 ) The drained cathode with flow-resist structure of lowers the interference of metal and improves the statical stability of metal interface. 3.3 Factors affect the current efficiency

1 : Decrease of busbar voltage drop, cathode voltage drop and anode voltage drop ; 2 : Decrease of electrolyte voltage drop, including ACD reduced

According to the integrative mechanism model formula of CE ^1 :

or increase of conductivity of electrolyte ;

CE(%) =100—219CT1en Dme°-67-M°-5-ue °-83d-°17puCAl

3 : Lessen the fluctuation of metal ;

cathode . kA/m

3.2 Relation between voltage and current efficiency A query is usually raised if the"energy saving at all levels" would affect thecurrent efficiency of aluminium reduction cell. Refers to the formula of heat balance by Mr. Shen Shiying I1J : ( a + b - η ) = V h e a t radiation

(3)

whereas : CE—current efficiency , % ; den—current density of

4 : Increase current efficiency.

I'Rsystem +

(1—f)

Dm

effective diffusion coefficient of

interface tension correction ; ì—density of electrolyte ; ue—average velocity of bath flow compared with metel , m/s ; d—ACD , m ; p—desity of bath , kg/m3 ; CAi —saturated concentration of Al in bath ;/—proportionality coefficient of metal strength.

( 2 )

That is to say, decrease of ACD means the decrease of R^tem, meanwhile, the current efficiency will be lower in case of Vheat radiation remained. So the limit of decrease of ACD will be happened to the case when the current efficiency remained to each cells. This limit is the "heat radiative gap" foregoing. The following 4 factors are necessary when decrease the ACD for having both "energy saving at all levels" and high current efficiency :

Table 3-3 Dependence of factors aftect CE : CE%

594

4.1 Grooved anode technology The slotted anode technology began with experiment of water model at inertia anode with groove at bottom by American Shekhar in 1990. It showed that anode bottom with groove was helpful to decrease the coverage of bubble at anode bottom and facilitate the mass transfer of alumina between anode and cathode. There are horizontal and lenthwise types of grooves at bottom, of which the lenthwise is more popular. The experimental result of 300kA prebaked cell with grooved anode in China indicated that the grooved anode brought about 4 lmV i31 voltage drop than common anode.

1 . ExcessvieÄlF3 2. Concentration of AI 2 0 3 3. ACD 4. Current density of cathode 5. Velosity of surface metal flow 6. Distance of non-anode projection 7. Divergence of anode current distribution 8. Growth of factors invloved

4 New energy saving technology of aluminium reduction

4.2 Graphitized cathode technology

The innovation on developing a number of technologies about the energy saving in aluminium reduction cellis is constantly taken. Mainly concludes:

Typical index of carbon blocks for aluminium

Table 4-1 Material Index

Amorphous base

Graphite (10%)

Graphite (30%)

Graphite (50%)

Graphite (100%)

Graphitized carbon

Resistivity pQm

55-60

42-45

33

30

20

8-14

10

14

18

31

73

Heat conductivity

1

The cathode block of aluminium reduction cell comprises amorphous base, graphite, and graphitized carbon.

W/m-k Expansion of sodium , %

0.6-1.5

Price , Yuan/t

2400

0.3-0.5 7000

3500

About 10-15% graphite of cathode has lower voltage drop than amorphous by 15~20mV , and 30% graphite of cathode is lower than amorphous by 45mV, from Pechiney API8 cell by 2 years experiment. The graphite content 40% of cathode applied in Guizhou aluminium smelter 230kA cell brought about voltage drop by 53 - 83mV lower than 9% graphite of cathode. The graphite content 50% of cathode applied in 22 cells in Yunnan aluminium smelter showed the voltage drop of cathode block was lower by 50mV than common cathode block, as the voltage drop of cathode block around282 - 338mV, with average value 325mV. The graphitized carbon cathode applied in 16x 300kA cells of Wanji Aluminium smelter saved 80mV voltage compared with the graphite content 10% of cathode. Furthermore, the life up to 1216 days of cell could realize the profit and loss break-even point through calculation of application of graphitized cathode.

0.05-0.15 10000

12000

0.82A/cm2 with current density up to 330kA

17000

0

4.3 Flow resist technology of drained cathode The flow resist techonology of drained cathode was early presented as "a series parallel channel or grooves struacture" of drained cathode in US patent (application number 6093304) by Vittorio de Noro and Nassau Bahamas in 2000., consisting of following 4 structures:

Due to the low thermal dilatation ( < 0.25% ) of graphitized cathode, the non-heating paste with low shrinage value ( < 0.15% ) is adopted to get rid of lining breakage at early stage.. After the partial 400kA pilot cells of Lanzhou aluminium smelter adopted graphitized cathode, the temperature at bottom of cell was inclinable to fall off somewhat as the current density of anode less than 0.8 - 0.82A/cm2. However, after Wanji aluminium smelter adopted graphitized cathode on 300kA potlines, the current density of anode was increased to about

595

side (<200kA), is inclinable to decrease in temperature and current efficiency when voltage drop > 0.2V. 5) Subject to to the heat balance formula (2) of cell, the voltage drop will be over 0.3V, and current efficiency will be lower by 0.5-2%. Although the flow resist of cathode techonology is required to be improved somewhat, the great achievement of power saving become a technical attraction point in energy saving of aluminium industry in China.

Table4-1 : views of drainded cathode struture by foreign patent

4.4 Technology of low voltage to save the energy The low voltage technology means not only the low voltage to be applied in busbar, anode, cathode conductor, and contact points between busbars, anodes, cathode conductors but also the further shortened ACD properly applied in a stable operation of the cell by utilizing the existing potentialities of ACD. The low voltage of energy saving technology depending on contraction of ACD requires the high ACD of cell ( > 4.5cm ) and is potentially stable. The 7 aluminium enterprises of China using low voltage technology to save energy are investigated and listed in table 4-2 The effect of this technology shows the normal operation of cell working at 3.85 - 4.01V. The main characters are:

In China, Northeastern University, Yunnan aluminium smelter, Chalieco, etc had applied the patent. But the partial patents of China had been covered as the structures showed in table 4-1. In the meanwhile, the applied achievement had been identified in Tiantail auminium, Yunnan aluminium, and East China aluminium. It indicated that the 1112, 743 and 817kWh was saved. After study on 10 smelters using flow resist techonology of drained cathode, the conclusion was summerized by the author: 1) The flow-resist block applied in drained cathode breaks the circular eddy of aluminium liquid and reduces both velocity of flow and fluctuation of interface of aluminium liquid, which improves the dynamic stability. 2) The cell, adopts insulating lining of drained cathode, could compensate lOOmV voltage around the cell itself during heat radiation. 3) The flow-resist protruding part is neither small nor high. The top of protruding part has to be lower than the aluminium liquid by over 6cm, otherwise the protruding part would be worn badly during unequal current distribution at early stage and reduction at high temperature. 4) The cells with low current density of anode, or big processing Table 4-2 Enterprises

1 1

7 aluminium enterprises of Chalco using low voltage technology Fa Xiang

Long Xiang

Zhong Fu

Lin Feng

Yugang Longquan

WanJi

Qin Yang

I

Capacity , kt

85

60

460

360

860

650

18.6

|

Current, kA

190-215

173

338, 414

221, 400

304, 400

175, 330, 428

162, 320

|

3.80

3.95

—

Setting voltage , V

3.75-3.80

Average votage , V

3.85

3.90

3.85/400kA

3.86

4.01

3.99

3.92, 4.00

Effect coefficient Concentration

0.05

0.05

0.002

0.05

0.05

0.05

0.03

1.5-2.5

2.0-2.5

2.5-3.0

2.5-3.0

27-28

24-26

27-28

42-47

26-28

26-30

20-22

17-21

18-20

17-20

18-20

18-21

18-20

19-21

of alumina , % 1

1) Followed with ACD closed to limitation of stability, the elaborated operation of operators are required highly. 2) To maximize the affection on dynamic stability of cell caused by shortened ACD, the aluminium liquid is generally increased. The current is accordingly intensified if apply the shortened ACD directly onto the operating cell which brings into the insufficient internal heat.

Aluminium liquid level , cm Bath level , cm Temperature of reduction , °C Molecular ratio

2.4-2.6

2.4-2.5

2.4-2.5

2.45-2.5

2.5-2.7

2.6-2.7

2.4-2.5

1

Cathode voltage drop , V

363

400

350

360

320

280

351, 320

|

CE , %

90-91

91-92

92.5

90-91

92

92.5

93

925 - 935

940-950

596

945 - 950

945

|

2.8

1

930 - 940

|

940 - 960

|

935-940 |

I

I Comprehensive A C consumption , kWh

13740

13550

13740

13525

The new drained aluminium reduction cell technology jointly developed by Zhengzhou Research Institute of Chalco, SAM and GAMI was identified that it saved 1220kWh per ton of aluminium by China Non-ferrous Metals Industry Association in April, 2009 The drained cell was initially presented in "Mushroom" shaped TiB2 coating of graphite cathode by American Kaiser Mead smelter. In 1986, Austrialia Comalco conducted the titanium boride coating and drained cathode industry experiments on 92.5kA cell. But the cell was only working for 700 days.Afterwards, there was no any progress in the world. The main feature of new Chalco cell is TiB2-C coating cathode, netted drained grooves for flush flow. As the TiB2-C cathode increase the wettability of aluminium liquid, and the netted drained grooves for flush flow enhance the dynamic stability of

Competitively, the potentiality of new cell developed by Chalco is farther on saving energy. 4.6 CE increase by '3-VariabIe'Control Technology A very important indication of cell having high efficient energy saving is both the material balance and electroheating balance achieving the optimum controllability. There are 2 main techinical features of GAMI's '3-Variable'Control Technology as follows: 1 ) Material "deficiency" control Additional feeding

AE AI9O3 volume

Change of cavity

Reduction °C

Deposition

l·^

Stable electromagnetic

Deposition 2 ) Energy balance as "need"

0

14700

cell, the prospective possibility for having further shortened ACD to save energy become more wide. The pilot cell of Qinyang, and the application in Lanzhou smelter, and Jiaozuo Wanfang, etc indicates that the voltage of cell could be decreased by 3.69 - 3.75V (ACD less than 3.7 cm).

4.5 Drained aluminium reduction cell technology

Electrolyte volume

13900

13700

High efficiency-

O

Energy balai balance

Enerav savina Tendency of °C

Degree of superheat

11η nr Hmvn

l

O

Al liquid level

Votage fluctuation

Variety conductiviety

of

Table 4-2 Optimum control on material balance is aiming at controlling "deficiency" through timely trim the feeding interval to match cell production, in order that the bad cell operation may be encountered due to excessive feeding. And, the electroheating balance is to keep the degree of superheat, regulation of cavity, and increase of CE via regulating ACD, aluminium volume and excessive A1F3. This is the core of '3-Variable'control technology, of which the application brings the increase of CE by 1-2.5% efficiently.

temperature T of molten bath derived from the formula respect to additives against conductivity p of molten bath:

I4J

in

ln/> = 2.0156 - 0.020ºAl203% - 0.005CaF2% - 0M66MgF2% + 0.0178Z/F% + 0M7Ui3ÄlF6% + 0.0063NaCl% + 0.4349C7? 2068.4/T So, there are 3 measure to be taken to increase the CE by decreasing the molecular ratio: D increase the voltage according to the decreased molecular ratio;or D improve the solubility of alumina; or D enhance the conductivity of bath. The affection of additives are versatile and complicated, for instance, LiF could increase the bath conductivity greatly but lead to a big fall of temperature, and the high concentration of LiF ( > 3% ) may also result in bad affection in case of improper control.

From talbe 3-3, accompanied with decrease of molecular ratio but increase of excessive A1F3 volume, CE will be improved. However, the decreased molecular ratio would lead to reduce the conductivety of electrolyte and cell temperature, which may cause heat out of balance and decreased solubility of alumina. The pracitise tells that the voltage drop of electrolyte will be increased about 30mV as molecular ratio decreased by 0.1. The bath conductivity pis of direct ratio for molecular ratio CR and

5 Conclusion

597

Under the multi-pressures at present, aluminium industry of China has to carry out the policies of energy saving at all levels and cost reduction. In respect of principle of energy saving, the cell is potentially to be excuted with energy saving but never only on 1 factor. As the loss of CE by 1% equally to increased voltage drop by lOOmV subject to the cost of primary aluminium, the shortened ACD technology to save energy is favorable to avoid possible bad cell operation and big loss of CE. The newly developed energy saving technology in China have different features and bring about diversly good effect in saving. In despite of a gap in front of the Chinese CE coefficient to world top level, the great contribution had been made by Chinese aluminium industry to the world. By Chinese unremitting effort, the DC coefficient of alumium reduction cell could be 12500KWh and even realize the energy consumption and high efficeint index of 12000kWh in the future. So, the date of Chinese aluminium industry catches up with and surpasses the top level is not too far. References: [1] Zhuxian

Smelting Technology of Cell, III edition, Qiu

[ 2 ] CE of 160kA cells, China non-ferrous metal journal, 2 period, volume 10. [3] study on experiment of drained high amperage cell, Mining Engineering, 3 period, volume 27. [4] Aluminium Smelter Technology , K. Grjotheim and BJ.Welch

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CELL VOLTAGE NOISE REDUCTION BASED ON WAVELET IN ALUMINUM REDUCTION CELL Binchuan Li, Jianshe Chen, Xiujing Zhai, Shuchen Sun, Ganfeng Tu School of Materials and Metallurgy, Northeastern University, Shenyang 110004, China Keywords: Aluminum Reduction, Wavelet, De-noising reduction than time domain [7]. In this paper, it mainly discusses the application of wavelet transform to analyze noise reductions of cell voltage signals.

Abstract For line current fluctuation, cell voltage signals collected in the aluminum electrolysis process have high noise, which has a significant impact on the aluminum reduction cell voltage and precision of the amount of alumina feeding. Based on wavelet denoising theory this paper analyses and compares different wavelet bases for signal de-noising effect by application of MATLAB modeling simulation and field research. Simulation and processing results of field data show that 5-order Haar Wavelet is a good choice for filtering cell voltage signal, with better prospects.

Wavelet De-noising Algorithm The basic method of wavelet de-noising is: Transform signals with noise using multi-scales from time domain to wavelet domain, and extract wavelet coefficients as much as possible in each scale to remove noise in the wavelet coefficients, and then reconstruct signals by using inverse wavelet transform [8]. Selection of Wavelet Base

Introduction

A very important issue for wavelet analysis in engineering fields is the selection of the optimal wavelet base function. How to select the best wavelet function during wavelet analysis is still questionable [9]. The general principle is on the basis of wavelet base function traits, signal characteristics and other specific requirements.

The aluminum electrolysis process has characteristics of nonlinear, time varying and large delay, and strong electric field, magnetic field and thermal field interfere with each other. Thus a number of important parameters in the production process, such as electrolyte temperature and alumina concentration cannot be achieved by online continuous testing [1]. In current times, only the cell voltage signal and the line current are available online, and the resistance calculated by the cell voltage and the current is only online access to reflect the state of the cell signal.

This paper chooses Haar wavelet as wavelet base function. Haar wavelet has orthogonality and symmetry while the features make it a linear phase, and compact support and highest time resolution of the simplest wavelet base function.

Cell voltage signals collected in the aluminum electrolysis process are with noise, which will affect the identification of the signals and the judgment the status of cell [2]. Many scholars work out a wide variety of de-noising methods according to actual signal characteristics, noise statistical features and their spectrum distribution. The most common method is based on one phenomenon that noise energy is generally concentrated in high frequency, while the true signal spectrum is located in a limited range, so low-pass filter approach can be used to de-noising, such as Fourier Transform, Moving Average Window filter, Wiener linearfilters,etc. [3,4].

Choice of Filter Order Mallat (1989) proposed the concept of multi-resolution analysis [10], and calculated the fast wavelet decomposition and reconstruction, which is Mallat Algorithm. Signal x(t) is the orthogonal wavelet decomposition: iv

U)

JAik u)

-Vi, —

r

(1)

xu~l)

\d =Ye

But the traditional method of matched filter wave can only get better results by more signal prior knowledge. As complexity and randomness of reduction cells work environment, cell voltage signal is inevitable non-stationary. Therefore the wavelet theory gets more attention for its time-frequency characteristics of signal processing and noise reduction [5].

0_1)

^"θ(η-2*)Λη

xk®—scale coefficient, dk(j)—wavelet coefficient, hok> gok—multi-resolution analysis filter coefficients, j—decomposition order. Decomposition of wavelet reconstruction process is the inverse operation, and the corresponding reconstruction formula is

The wavelet transform has multi-frequency characteristics, so it can achieve high signal resolution of local orientation in both time domain and frequency domain, which extract effective temporal, steady-state information and waveform characteristics information in non-stationary signals [6]. In turn, it can reduce noise through distinguishing between different rate signal and noise distribution. In addition, the signal can be removed or reduced correlation by wavelet transform, and it has trend of whitening after noise transformation, so wavelet domain is more favorable for noise

x Λ

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Signals can be decomposed at different scales by multi-resolution analysis of wavelet, and the signal intertwined with composition of mixed-signals of different frequency was made into different sub-band signals [11]. The essence of wavelet filtering is to let signals pass through combination of high and low filter, so the signal is decomposed into high and low frequency band, and continue this process until the need is met. As long as we select appropriate decomposition order, we can get the required width and starting and ending frequencies of the frequency band. The concrete steps by using wavelet decomposition and de-noising reconstruction are: decompose a noisy signal into different frequency band scales, set noise frequency band zero, and then reconstruct the wavelet, so as to achieve the purpose of denoising.

Select different order of the Haar wavelet to filter voltage signal, analyze filtering effect of Haar wavelet in different order levels. Filtering effect of different order levels Haar wavelet is shown in Figure 3. 4.140

Π

4.135

ruH] W

4.130

:

4.120

-vj"^^~^Jy 200

300

400

Time (s) (a) 4th Haar Wavelet

Useful information on cell voltage signal spectrum is concentrated in one section of frequencies, and decomposition of appropriate layers using wavelet theory can decompose the signal to a specific scale, which is conducive to the separation of signal and noise, so as to extract signal characteristics to analyze.

4.132 4.130 ^

>gj,

4.128 4.126

•

;> 4.124

Results and Analysis

4.122

This paper is based on noise signal extracted by cell voltage data in a cell production process, which is added voltage step signal, finally establishing cell voltage signal with noise, and analysis of the cell voltage signal with noise. Extracted noise signal is shown in Figure 1, and noisy step voltage signal is shown in Figure 2.

4.120 ,

-1 UN 200

300

1

,

1

,

400

Time (s) (b) 5th Haar Wavelet

4.132 4.130 ,-s 4.128

>

& 4.126

S

> 4.124

4.122 4.120

100

200

300 Time(s)

400

500

600

200

4.118

300

400

Time (s)

Figure 1. Extracted noise signal.

(d)7th Haar Wavelet 4.132

4.20

-

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-

Ì& 4.126 >

4.124

-

4.122 4.120

Time (s) 100

200

300

400

500

600

(c)6th Haar Wavelet

Figure 3. Filtering effect of different order levels Haar wavelet.

Time (s)

Figure 2. Noisy step voltage signal.

From the above diagram of the Haar wavelet filter effects we can see that noise is still confused in the voltage signal by using 4-

600

order Haar wavelet filter, and it has not been filtered, so the filter effect is not obvious; By using 6,7-order Haar wavelet filtering, the filtering effect is obvious, but some low-frequency voltage signal is filtered out. In multiple wavelet filter, 5-order Haar wavelet filter is the best.

Technical Sessions Presented by the TMS Aluminum Committee at the TMS 2009 Annual Meeting and Exhibition, 2009,311-315).

By 5-order Haar wavelet de-noising, wavelet decomposition and reconstruction of signal characteristics are shown in Figure 4. Coefficients of the 5-order low-frequency layer after wavelet decomposition are considered as cell voltage changes, while other factors can be treated as a cell noise, therefore we reconstruct low frequency coefficients of order 5 to get the de-noised cell voltage. 89 til«

£4it οittr Inοwt

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2.

Majid, Nazatul Aini Abd, Young, Brent, Taylor, Mark and Chen, John, "Real-time for monitoring aluminium reduction cells using multi-way PCA (MPCA) and dynamic euclidean distances", (Paper presented at 2009 IEEE International Conference on Control and Automation, ICCA 2009, 454458).

3.

Johnson III, A., and Li, C.-C, "Wavelet packet time series analysis of aluminum electrolytic cells," Proceedings of the SPIE The International Society for Optical Engineering, 2001, no. 4391:228-371.

4.

Liu, Hongfei et al., "The study of unbiased-estimation threshold wavelet de-noising method applied on midwavelength infrared image", (Paper presented at Proceedings-2nd International Conference on Information Technology and Computer Science, ITCS 2010, 202-205).

5.

Huaigang, Zang, Zhibin, Wang, and Ying, Zheng, "Analysis of signal de-noising method based on an improved wavelet thresholding", (Paper presented at ICEMI2009 - Proceedings of 9th International Conference on Electronic Measurement and Instruments, 1987-1990).

6.

Cattani, Carlo, "Sparse representation with harmonic wavelets", (Paper presented at 6th International Conference on Fuzzy Systems and Knowledge Discovery, FSKD 2009, 159-163).

7.

Zhang, Zhen, and Xue, Tao, "Application of a modified algorithm for wavelet threshold de-noising based on the ultrasonic signal of optical fiber defect", (Paper presented at Proceedings of the 2009 2nd International Congress on Image and Signal Processing, CISPO9).

8.

Dupas, Nicolas, "Increasing electrolysis pot performances through new crustbreaking and feeding solutions", (Paper presented at Light Metals 2009 - Proceedings of the Technical Sessions Presented by the TMS Aluminum Committee at the TMS 2009 Annual Meeting and Exhibition, 2009, 337-340).

9.

Brooks, G. et al., "Challenges in light metals production," Transactions of the Institutions of Mining and Metallurgy, Section C: Mineral Processing and Extractive Metallurgy, 116(1) (2007), 25-33.

*j

Figure 6, Floor coefficient after the resolving of Haar wavelet. 5-order Haar wavelet filter works well out of the basic reduction step by waveform after de-noising back cell voltage, and the step signal error is below 4.92 x 10"4 V. Conclusion Wavelet de-noising transform can achieve precise reduction of the voltage step signals, and 5-order Haar wavelet has better noise reduction effect, with higher recognition accuracy. Wavelet noise reduction technology can reduce the monitored cell voltage with characteristic parameters to a number of data, which is conducive to the realization in computer. The program by the computer implements all calculations with high accuracy of correct diagnosis for reduction cell work state, and is conducive to online monitoring and real-time identification. Furthermore we can reduce the error of alumina concentration identification using resistance calculated by the cell voltage and the current.

10. Mallat, Stephane G. and Zhang, Zhifeng, "Matching pursuits with time-frequency dictionaries," IEEE Transactions on Signal Processing, 41(12) (1993), 3397-3415.

Acknowledgements This project was supported by Ministry of Education of the People's Republic of China (N090402016).

11. Davis, Geoffrey, Mallat, Stephane G., and Zhang, Zhifeng, "Adaptive time-frequency decompositions," Optical Engineering, 33(7) (1994), 2183-2191.

References 1.

Stam, Marco A. et al., "Development of a multivariate process control strategy for aluminium reduction cells", (Paper presented at Light Metals 2009 - Proceedings of the

601

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011 ALUMINUM REDUCTION TECHNOLOGY Poster Session SESSION CHAIR

Abdulla Habib Aluminum Bahrain- Alba Manama, Bahrain

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

HUMAN FACTORS IN OPERATIONAL AND CONTROL DECISION MAKING IN ALUMINIUM SMELTERS Yashuang Gao1,2', Mark P. Taylor2', John JJ. Chen1, Michael J. Hautus3 1 Chemical & Materials Engineering 2 Light Metals Research Centre 3 · Department of Psychology The University of Auckland, Private Bag 92019, Auckland 1142, New Zealand Keywords: Human factors, decision making, aluminium smelting, process control Abstract

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The aluminium smelting process involves highly complex mechanisms and has rich information but low observability. To operate such a process at maximum current and energy efficiency while striving for continual improvement requires a set of scientific and systematic approaches to problem solving and decision making along with constant human intervention and interaction based on informed and considered judgment. Understanding the influence of human factors on the process is important to the improvement of the performance of the operational staff and enhancement of positive impact while minimising negative effects on the process. This paper explains some aspects of the human factors and decision making involved in the smelting reduction operations and control process, which includes sensing and monitoring signals, identifying abnormalities and implementing appropriate responses to not only correct the immediate causes of the identified problems but also to reduce the variation and to continuously improve the process over time.

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Figure 1: A simple model for the control process adapted from [1] The aims of the present research are to explore the impact of human factors and decision making on process control as well as the interaction between the systems and human operators. Ultimately, the findings will be integrated into systems, which will improve current process operation and control practices, hence reduce the cost of production, energy consumption and the impact on the environment. The impact of human factors and decision making in process control is significant and pervades every aspect of the operation and control. This paper will use a few examples from Smelters A and B (names cannot be stated for confidentiality reasons) to demonstrate the impact of human factors and some of the missing elements in the current control practice in the aluminium smelting industry, as well as the methodology for potential improvement based on the model in Figure 1.

Introduction Process control is commonly understood as an engineering discipline that deals with computer architecture and algorithms for controlling the output of a specific process. Digital computer control, which is linked to the availability of computers, was brought into the smelting industry in the early 1980s. The aims were and still are to maintain a process operating steadily under the designed specification conditions, achieve the desired quality of the product, and reduce cost through improving the efficiency of the process, while minimizing the human and systems error occurring in the process. However the potential to integrate the power of human reasoning and decision making with the almost unlimited computational capacity now available has not yet been explored. Furthermore, more advanced control concepts are needed in process control, rather than "fiddling" the parameters and/or compensatory actions.

Observing process state - Measurements Aluminium smelting is a multivariate process and involves highly complex mechanisms such as mass and energy balances, electrochemical reactions, supply of raw materials, and maintenance of the composition of the electrolyte [2]. The large amount of information coming in from such a process, some in real-time and others intermittently at varying frequency, is challenging for a human brain to process. This information consists of digital events and analogue signals and signatures sampled from the pots by the computer system, as well as the discrete manual measurements and visually observed signs or features Figure 1.

As Taylor and Chen [1] pointed out, a more robust overall control model is required for industries, such as aluminium smelters, to meet the needs of energy reduction and environmental compliance [1]. Figure 1 shows a simple illustration of such a control model. However, only a fraction of this model can be observed in the current control practices in smelters. Compensatory actions, such as manipulating the input to reach the desired outcome temporarily, without an understanding of the variation, often occur. Many important elements in the control model are still missing. To fully implement this control model requires a higher level of system and human interaction.

To control this complex process to achieve high productivity and efficiency requires day to day (and sometimes minute to minute) monitoring of the variables, and a high level of deductive problem solving and decision making skills using the process information as described above. However, in this information-rich environment, the absence of a strategic procedure or a system to manage the data, visual observations, and verbal communications, some of the valuable information is often diluted or even lost.

605

This is illustrated by the flow chart in Figure 2, which shows the common observations obtained from many modern smelters. Therefore, the efficiency and effectiveness of the operational staffs decision making or problem solving are made based on incomplete information and might not be optimal. Furthermore, implemented decisions cannot be easily tracked and followed up. This could have short term or long term impacts on the process and improvement. This indicates that the first phase in the control model in Figure 1 is not fully implemented in many smelters and it has negative impact on the process in the subsequent phases.

will allow the smelter information to be recorded and analysed systematically and consistently, and provide meaningful information to assist the operational staff to determine the process state. : Awtaaing «ssigr*^ and recogrtftaft :^*?

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Figure 3: A flow chart illustrating a better information management model with the application of an advanced supervisory control system

PuniswwU System

Figure 2: A flow chart illustrating the current information management situation in many smelters

Bath Level Control One example demonstrating the effectiveness of a better designed supervisory system which is able to provide meaningful information to assist the operational staff on process decision making is a case study conducted of bath level control in Smelter A [3]. Low bath level has been a serious issue for this smelter. The percentage of the potline having bath level below the lower control limit every month in the previous two years ranged from 20% to 80%. The root cause was not found for a long time. Meanwhile, the impact of low bath level to the process was detrimental. Eventually it was identified by the external experts as problems with the supply of crushed bath materials. The crusher was broken down so often that the supply of crushed bath material for anode cover dressing was not reliable. Therefore, at times, there was no material to replenish the liquid bath. This root cause was simple and clear, but because of the missing links in the information management within the smelter, it was difficult for the operational staff to pinpoint.

Carbon Dust and Airburn One example of the influence of human factors on process control and the consequence of making decisions based on incomplete information, (which is illustrated on the left of the flow chart in Figure 2) is the presence of excess carbon dust and airburn in Smelter A. When 80% of the pots in the reduction line have excess carbon dust and severe airburn, it is most likely that the carbon plant gets the blame for poor anode quality. That was the situation in Smelter A, until the external experts observed the operational and control practices in the potroom. It was diagnosed and identified that the carbon pieces floating between the anodes and cathode in the pots and the presence of carbon pieces in the bath material process circuit were the root causes of excess carbon dust and airburn. Without full information and the understanding of the cause and effect, the decision made to blame the anode quality (which is controlled by the carbon plant), and not to investigate the situation and take corrective actions, is in agreement with the illustration in Figure 2. It also shows the fact that the concept of the control model was also missing in the practice in this.

There are 6 sections in the potline in Smelter A and the performance of the 6 sections were significantly different [3]. Even though the crushed material was not available for the whole potline, one of the sections managed to control the bath level within the control region consistently. It was found that the leader of this section used a newly implemented supervisory system to monitor the bath level situation, while the leaders of the other sections resisted the use of the new system and instead continued using printed daily reports with only numbers and tables. The new system incorporated some tools such as colour management of data, statistical graphs and pictorial illustrations for operation practice standard. The system was designed to be user friendly and took into account human factors such as perception of risk. For example the user will feel pressure and high level of risk

Improved Information Management What has just been described is only one of the million cases which happen every day in smelters. It has been recognized that an advanced supervisory control system integrated with the robust control model would improve information management and process control by allowing the operators to observe the process with more meaningful information [1, 3]. Figure 3 shows the proposed improved information management model with the implementation of an advanced supervisory control system. This

606

Phüm Et Understanding fise Variation

when the measurements of the pots are highlighted in red (ie indicating it is outside the control region). By using the system, the section leader was aware of the bath level situation of the pots [3]. He asked the operators to tap the bath out of the pots which had high bath level and stored the bath material for anode cover material when the crushed bath material supply was not available. Furthermore, as a leader, he created a reward system to motivate the operators to do the daily operational and control tasks such as bath level control. The performance of this section demonstrates the impact of leadership, management and decision making skill on operation and process control.

Phase $t €ontr®tUnQ the Oirfcortie

Figure 5: A more detailed representation for Phase 2 and 3 of the control model, adapted from [1]

Understanding the variation - Root Cause Diagnosis As shown in the control model in Figure 1, understanding the variation links the observation of the process (ie. the stimulus, what you see) to controlling the outcome (ie. the response, what you do and achieve). Process variation is classified into three generic types: common cause or natural variation, special cause variation and structural cause variation [1]. Without the understanding of variation, the actions taken can often be compensatory or incorrect, thus the variation stays within the process persistently.

Operators as Signal Detection Systems However, as discussed previously, many of the current systems do not even provide complete information to the operators, and furthermore, most of these systems also do not have the ability to guide or assist the operators to understand the variation and identify the root cause. They rely heavily on the operators to make judgments and detect the process problems. In this case, the operators take on the role of signal detection systems for identifying problems. Figure 6 illustrates a basic signal detection system [5]. The input x (ie. process information) is a stimulus, and the stored 'database' is the memory of the operators. The box is the simplified decision making process for the operators to arrive at some function Z which is used to formulate a response based on some decision making criteria. However, without guidelines or the provision of technical constraints, this decision making process model gives the operators an un-reined degree of freedom. Coupled with the influence of a low level of situation awareness (due to incomplete information), perception bias, and selective attention, it can often lead to poor decision making [6, 7]. The consequence is that process abnormalities (special causes) are not identified and left un-attended. Thus such abnormalities keep recurring.

The case study of bath temperature control in Smelter B is an example which demonstrates the importance of understanding the variation and identifying the root cause [4]. Figure 4 shows the thermal cyclic situation observed from a pot. It shows that chemical additions were used to adjust bath temperature without diagnosing the root cause. Soda ash was added when bath temperature was low and outside the control specification box, and a large amount of A1F3 was added to attempt to fix high bath temperature. The consequence was that the temperature cycled and this problem remained in the process. A statistical control tool (control ellipse) was then implemented to guide the operators to understand the variation and diagnose the root cause. The root cause of the continual temperature cycling was the inappropriate use of additions, while the root cause of some of the extreme high temperatures was due to alumina feeding problems [4]. This is a practical example of the implementation of 'understanding the variation' in the control model. A detailed representation of phases 2 and 3 of the control model is shown in Figure 5.

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Stored «lata about \

Stimulus Input X

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Response Yes No

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Figure 6: A basic signal detection system, redrawn from [5] Improved Signal Detection System As suggested by Taylor and Chen, with clear identification and classification of the three types of variation in phase 2 as given in Figure 5, an automatic control system incorporating a robust control model can be designed to assist the operators to not only observe the process but also to detect and diagnose the root cause of the process abnormalities. Figure 7 illustrates a model of automatic system and human operators working together on detecting process abnormalities. In this model, the system detects and classifies the process abnormalities and sends alarms to the operators. The operators then respond to the alarms by investigating the process and diagnosing the root cause. The overall performance is an improvement over the model in Figure 6 [5,8,8,10,11].

Figure 4: Temperature and chemical (A1F3 and soda) addition indicate the thermal cyclic situation of a pot

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Every step in Figure 8 involves a large degree of human reasoning, thinking, decision making, and action taking as demonstrated in the example above. The outcomes and the learning from the process of problem identification, diagnosis and solution achieved should be stored in a system and carried on to be used in solving other problems. A well designed system will not only guide the human operators to implement the scientific methodology but also allow for feedback from the operators; hence a learning process takes place and the system can be continually improved. The system design should also take into account human perception, attention and situation awareness to avoid bias from individual experience or knowledge.

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Figure 7: A detection model with system and human monitors, modified from [5]

Conclusions In conclusion, this paper demonstrates the importance of the control model and the influence of human factors to aluminium smelting process operation and control, by using actual examples from aluminium smelters. The examples have shown the influence of human factors in operation and control decision making. This shows that smelters urgently need an advanced system which incorporates scientific human reasoning functions, tools and guidelines to assist the operators to better observe the process, understand the variation, remove the root causes, and therefore better control the outcome.

Controlling the outcome - Response Once a problem is detected by the system and confirmed by the operators, corrective actions are required as illustrated in phase 3 in Figure 5. Here, human operators can take immediate stabilizing actions to correct the immediate causes, such as removing anode spikes, cleaning the bath built up on the breaker or fixing the malfunctioning feeders. The operators can also implement a scientific problem solving methodology to lower the long term variability of the process [12]. The process of the scientific problem solving process is illustrated in Figure 8.

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MEASURE - Whafresedei

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References 1. M. P. Taylor and J. J. J. Chen, Advances in process control for aluminium smelters. Materials and Manufacturing Processes, 22(7):947-957, 2007 2. J. J. J. Chen and M. P. Taylor, Control of temperature and aluminium fluoride in aluminium reduction. International Journal of Industry, Research and Applications, 81:678-682, 2005 3. Y. Gao, M. P. Taylor, J. J. J. Chen, and M. J. Hautus, Operation decision making in aluminium smelters, Engineering Psychology and Cognitive Ergonomics, Human - Computer Interaction International, pages 167-178, 2009 4. Y. Gao, M. Gustafsson, M. P. Taylor, and J. J. J. Chen. The control ellipse as a decision making support tool to control temperature and aluminium fluoride in aluminium reduction, The 9th Australasian Aluminium Smelting Technology Conference and Workshop, 2007 5. R. D. Sorkin and D. D. Woods, Systems with human monitors: A signal detection analysis. Human-Computer Interaction, 1:4975,1985 6. S. B. Most, D. J. Simons, B. J. Scholl, R. Jimenez, E. Clifford, & C. F. Chabris, How Not to Be Seen: The Contribution of Similarity and Selective Ignoring to Sustained Inattentional Blindness. Psychological Science (Wiley-Blackwell), WileyBlackwell, 2001, 12, 9-17 7. J. A. Swets, & A. B. Kristofferson, ATTENTION, Annual Review of Psychology, Annual Reviews Inc., 1970, 21, 339-366 8. J. A. Swets, R. M. Dawes, and J. Monahan, Better decisions through science. Scientific American, 283(4):82-, October 2000 9. J.A. Swets. The science of choosing the right decision threshold in high-stakes diagnostics, American Psychologist, 47, No.4:522532, April 1992 10. J.A. Swets, RobynM. Dawes, and John Monahan, Psychophysical science can improve diagnostic decisions. American Psychological Society, 1, No. 1, 2000 11. J.A. Swets, David J. Getty, Ronald M. Pickett, and David Gonthier, System operator response to warning of danger: A

BUILD HYPOTHESIS •8 * » ίΐΛ A. οhs S&&2 W b» 8,

TEST Not by aaaeiv* tttorffcitstig of

Figure 8: An illustration of the scientific problem solving methodology, adapted from [12] Referring to the case study of bath temperature control in Smelter B described earlier, once the root cause was identified as an alumina feeding related issue and low metal level situation, the problem was DEFINED. The MEASURE was temperature. Therefore a 10-day response plan was implemented for the hot and sick pot (HYPOTHESIS BUILDING and TEST). The thermocouples were recalibrated and extra temperature and metal level measurements were taken every shift. The cathode was cleaned every day during anode changing. A1F3 addition was maintained at a nominal rate. Along with other corrective actions, the temperature of the pot was brought down from 985 °C to 976 e C at the end of the plan (IMPROVE). A second plan was then implemented and this brought the temperature down and maintained at 960°C (CONTROL). This approach removed the root cause by the scientific problem solving methodology illustrated in Figure 8.

608

laboratory investigation of the effects of the predictive value of a warning on human response time. Journal of Experiemental Psychology: Applied, l,No.l:19-33, 1995 12. M. P Taylor. Scientific problem solving: Getting to grips with complex problems in industry, level 2, Imre smelter training course - reduction module

609

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ALUMINUM ROLLING

ORGANIZER

Kai Karhausen Hydro Aluminium Bonn, Germany

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ALUMINUM ROLLING

ALUMINUM ROLLING Session I SESSION CHAIR

Kai Karhausen Hydro Aluminium Bonn, Germany

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

AN INVESTIGATION OF DEFORMATION BEHAVIOR OF BIMETAL CLAD SHEETS BY ASYMMETRICAL ROLLING AT ROOM TEMPERATURE Xiaobing Li1, Guoyin Zu1, Qiang Deng2 1 School of Materials and Metallurgy, Northeastern University; Shenyang 110004, P. R. China 2 Cold rolling company, Pangang group company ltd.; Panzhihua 617067, P. R. China Keywords: Roll bonding; Roll speed ratio; Deformation behavior; Interface; Bimetal clad sheet complete analysis about the relation of stress state with deformation behavior of asymmetrical cold roll bonding for bimetal clad sheet. In present work, it was intended to analyze the stresses on the interfacial zones of the component metals at the roll gap during an asymmetrical rolling, furthermore, the systemic experimental data of the aluminum/mild steel (Al/St), copper/mild steel (Cu/St) and aluminum/copper (Al/Cu) were attained to investigate the deformation behaviors, bonding conditions and interfacial layer thickness of the clad sheets by the asymmetrical roll bonding.

Abstract The different thickness metals mild steel, aluminum and copper were bonded with each other by means of asymmetrical rolling at room temperature. The deformation behaviors, bonding conditions and interfacial layer thickness of the clad sheets were discussed. According to the slab stress in the plastic deformation region at the roll gap, the relations of bonding condition and metal flow were analyzed. The influence of cross shear on the bonding due to the roll speed mismatch is obvious. The large speed mismatch makes a good bonding and drops the critical reduction. The improvement of bonding is achieved with the increase of the total rolling reduction. The reduction of both layers increases in direct proportion with the total reduction, and the difference between hard and soft metals gradually diminishes. The large initial thickness ratio of hard and soft metal is unhelpful for the bonding due to the inconsistent deformation of bimetal.

Experimental Procedure The materials used in this investigation were commercial pure aluminum 1060, pure copper T2 and mild steel (0.14 wt. % C) of 25 mm width, 200mm length, respectively. The thicknesses of specimens are shown in the table 1. The surface of specimen was degreased using acetone. NaOH (10% in mass) was used to remove the oxide coating of aluminum and copper, and H2S04 (10% in mass) for the mild steel. The specimen was then scratched by a stainless steel circumferential brush with wires 0.3mm across running at a rotational speed of 1400 rpm. After surface preparation, the handling of the strips was performed carefully to avoid renewed contamination. Two pieces of the strips were stacked together by a soft aluminum wire. To investigate the effects of different parameters, a series of cold roll bonding experiments were made with the thickness reductions from 30% to 70% on the four-high reversing mill with roll diameter of 300 mm. The ratios of works roll speed, defined as speed mismatch ratio, were set as 1.06, 1.19 and 1.31. The specimen was fed into the roll gap while the hard component touched with the higher speed work roll. All the parameters of roll bonding are presented in the table 1. Some tests were conducted to try to enhance bonding on the interface by annealing at various temperatures for 20 min. The temperature deviation of the furnace was ±5 D. The cross sections of specimens, transverse to the rolling direction, were mechanically polished and buffed. They were then etched at room temperature using a solution of 5 ml HNO3+100ml C2H5OH for Al/St and Cu/St specimens for 30 s, and of 2 ml HF+3 ml HC1+5 ml HNO3+95 ml H 2 0 for the Al/Cu specimens for 30 s. The microstructures of the interfacial zone and intermetallic compound layers were identified by optical microscope, scanning electron microscope with energy disperse spectroscope.

Introduction Roll bonding technique is a solid-state phase bonding process used to join similar or dissimilar metals. During the processes of multilayer clad sheet and strip production, it is the most economical and productive manufacturing process that can be applied to produce high bonding strength of various materials in continuous rolling and processing lines [1-2]. At present, there are many researches about roll bonding, mostly base on the hot or warm rolling. Some researchers make the experimental study on the cold roll bonding, using the symmetric rolling techniques [3-6]. According to the previous studies, the roll bonding strength is determined mostly by the roll temperature, but the high temperature makes the surface of bimetal clad sheets poor and causes the appearance of the intermetallic compounds on the interface of the two dissimilar metals. For the cold roll bonding, the thickness reduction plays a more important role for the effective bonding [1,7-8]. The bonding of the common metal and alloy can be achieved when the roll reduction reaches 50% [2,7,9], whereas the higher reduction imposes a strict demand for the equipments. Asymmetrical rolling of bimetal clad sheet can significantly reduce the rolling force compared with conventional cold rolling, while still ensure the same primary bond strength[10]. In the asymmetrical roll bonding, there is a shear zone in the central region of roll gap due to the different peripheral speeds of two identical work rolls. The relative sliding on the interface between the two metals is enhanced at the entrance, whereas the plastic flow of the two metals becomes more homogenous at the exit of the roll gap[l 1]. The application of the cross shear roll bonding of the aluminum/stainless steel has shown the advantage of significantly reducing rolling load, while still guarantee equal (or even higher) primary bonding strength[12]. However, there is no

Table 1. Thickness of Specimens and Parameter of Roll Bonding No. 1

615

Com ponent thickness Mild steel Copper Aluminum /mm /mm /mm 1.50 1.00 -

Speed ratio

Thickness reduction/%

1.06

48

2 3 4 5 6 7 1 2 3 4 1 2

1 3

-

0.50

1.00

1.06

48 48 79 53 64 48 30 30 40 30 50 50 50

1.50

1.00

1.19

0.50

1.00

1.19

0.50

1.00

1.19

0.50

1.00

1.19

1.50

1.00

1.31

1.00

1.06

1.00

1.19

1.00

1.31

1.50

-

1.00

1.31

1.20

0.50 0.50

1.50

0.50

-

1.19

1.20

1.50 1.50 1.50

1.31 1.31

zone IV (0<x<xn2) for the bounded regions. The slab stress state of the clad sheet in zone II is shown in Figure 2(b), zone III is the cross shear region where thefrictionalshear stresses are reverse as showed in Figure 2(c), and the frictional shear stresses in zone IV are opposite to that in zone II. The direction of the shear stress on the interface changes along with the rolling direction due to the mismatch speed Zone II, III, and IV for the shear stress on the interface (im) should be determined by the model.

1

y 1

1

Γ *
^ι

H^'^J

«%J

The deformation behavior of roll bonding is difficult to perform. Some researchers develop various theories and establish the mathematical model to perform the experiment [7,13-17]. Figure 1 shows a schematic illustration of asymmetrical roll bonding under constant shear friction. It is assumed that the process of roll bonding is plane strain state and accordant with slab method, and distribution of normal stress at upper and lower work rolls is homogeneous. The material just makes rigid plastic deformation.

r

1 L^Afc

+ q2+dq2 q2 +

•VJ Pi

(a)

i

ini

i—*.q2+dq2

Deformation Stress Analysis

P l

^ 1

(b)

r2

+ ^

^

P2 (c)

(d)

Figure 2. Stress state of distinct regions at roll gap, (a) (b) (c) (d) represents the subzone of I, II, III, IV respectively. The slab stress state presents the metal flow tendency. In zone I and zone II, the shear stress both promotes the hard metal to deform. Unlike layer 2, layer 1 gets little shear effect, because of the same direction of shear stress on roll direction. Simultaneously, both layers take a large thickness reduction, the interfacial zones get enormous metal flow. When the metals come into zone III, the cross shear stress cause both layers to shear deform, an amount of metal flow is made on the interfacial zone. Afterward, cracks appear on the interface, the fresh virgin metals are extruded. In zone IV, the virgin metals achieve a mechanical bonding due to the metallic atoms bonding under load and thermal energy. After that, the component layer is thrown out the roll gap as an integral sheet. It should be noted that this analysis has discussed only the stress direction, not involve the value and relation of each other.

Qai Qo

Qo2

Results and Discussion Effect of Speed Mismatch Ratio on Bonding Figure 3 represents the effect of speed mismatch ratio on the interfacial bonding of bimetal clad sheets. The diffusion thickness on the interface increases clearly with the speed mismatch ratio varying. When the speed ratio increases from 1.06 to 1.31, the thickness of the interfacial zone of Al/Cu clad sheet and Al/St clad sheet grows from 2.8um and 1.8um to 3.5um and 2.1 um, respectively. Similarly, that of Cu/St clad sheets increases from 6.4μπι to 7.5 urn with speed ratio varying from 1.16 to 1.31. The large speed mismatch ratio causes a dramatic cross shear stress state on the interfacial zone, which has been analyzed in part of stress analysis. It is noted that the experimental results agree with the analysis. In the process of asymmetrical bonding, the way that the hard metal touch with the higher speed roll and the soft metal with the low speed roll enables the metal flow of both metals to uniform. Figure 3 shows that the morphology of interfacial zone of bimetal clad sheets has an improvement with increasing speed mismatch ratio.

Figure 1. Schematic illustration of roll bonding. Ri=R2, V2Ü V!, the deformation zone at the roll gap is divided into four distinct subzones I, II, ΙΠ and IV. The Figure 1 reveals the geometry at the roll gap. Unbounded clad sheet is initially bit into the roll gap, the soft sheet (layer 1) is yielded and the hard sheet (layer 2) is not yet yielded. Thus, this region (zone I) belongs to the unbounded region. The slab stress state in zone I is shown in Figure 2(a), where the shear stress on the interface is im. As the harder sheet is yielded, the clad sheet begin to bond, the bonding point (xb) is generated. The plastic deformation region at the roll gap can be divided into four distinct subzones along with rolling direction[13], zone I(Xb<x
616

In addition, the temperature increase induced by deformation and friction, due to the large speed mismatch ratio of asymmetrical rolling, play an important role on interface diffusion. [(a)!

deformation uniformity of both metals is useful for the metal atoms to make metallic bonding and the improvement is corresponding to the higher rolling reduction. !(a)

11»::: .

.3 11

; liai 1 5μ mi

Hbf ·

Figure 4. Micrograph of interfacial zone variation with different reduction, (a), (b), (c) represents 53%, 64% and 79%, respectively, of Al/Cu sheet with annealing at 400 D; (d) (e) is for Al/St sheet with 30% and 40% followed annealing 400 D.

Figure 3. SEM image of the morphology of interfacial zone with different speed mismatch ratio, (a) (b) shows that of Al/Cu sheet with annealing at 400 D; (c) (d) represents the Al/St sheet with annealing at 350 D; (e) (f) is the Cu/St sheet with annealing at 750 D. Effect of Thickness Reduction on Bonding In cold bonding processes like rolling, the total thickness reduction is one of the most important parameters that affect bonding strength. On the other hand the large normal roll pressure on the surfaces leads to produce cracks on the brittle surface layers. Thus extrusion of these virgin metals is convenient to get metallic atoms bonding in the deformation region at the roll gap. Figure 4 shows the interfacial zone variation of the clad sheet with rolling thickness reduction. It is found that morphology of the clad sheet is significantly improved with increasing rolling thickness reduction. As the rolling reduction increases from 53% to 79%, the interfacial zone thickness of Al/Cu clad sheet is, to some extent, improved. The improvement is similar for the Al/St clad sheet, the interfacial zone grows into a continuous layer, and the thickness increases from 4μπι to 7μηι when the reduction increases from 30% to 40%. The deformation of hard metal is delayed to the soft metal because of their different deformation resistance. So normally, the soft metal has a larger deformation accumulation during the whole roll bonding reaching the maximal strain hardening first. Afterward, the deformation rates of soft metal reduce, therefore, the deformation distribution of hard metal increase at the later stage of bonding. Figure 5 represents the deformation distribution of both metals variation with the total deformation. The reduction of both components of Al/Cu as well as Al/St clad sheet increased in direct proportion with the total reduction, and the difference between hard and soft metal gradually diminished. The higher

Total thickness r e d u c t i o n / %

total

thickness

reduction/?

Figure 5. Variation of components thickness reduction with total thickness reduction, (a) represents that of Al/Cu clad sheet, (b) represents that of Al/St clad sheet.

617

4. Reducing the initial thickness of hard metal improves the deformation uniformity of both metals to accomplish a good rolling.

Effect of Initial Thickness on Bonding When the rolling force is constant, the sheet with thin initial thickness has a small contact area with the rolls and suffers a higher rolling stress, ultimately the metal yields rapidly. The couple of metals have an asynchronous deformation owing to their different resistance. Generally, the plastic deformation of soft metal is higher distinctly than that of the hard metal. By decreasing the initial thickness of hard metal strips, the shear stress of the deformation region at roll gas may cause hard metal to yield easily. In this way, the deformation uniformity of both metals is achieved. The bonding is, therefore, improved. Figure 5 shows the experiment results, which are agreement well with above theory. The initial thickness ratio is expressed by the ratio of hard and soft metal, shown as Si/S s . It can be seen that the interfacial layer thickness increases from 3μπι to 7 urn, when the initial thickness ratio of Al/Cu clad sheet changed from 1.5 to 0.5 with the initial thickness of aluminum strip constant. The trend is similar for the Cu/St clad sheet, the initial thickness ratio changed from 3 to 2.4 results in the improved interfacial diffusion.

Acknowledgements The authors gratefully thank the National Natural Science Foundation of China for the financial support given to the research (Grant No. 50971038/E010702). References 1. M. Abbasi and M. R. Toroghinejad, "Effects of Processing Parameters on the Bond Strength of Cu/Cu Roll-Bonded Strips," Journal of Materials Processing Technology, 210 (3) (2010), 560563. 2. L. Li, K. Nagai and F. X. Yin, "Progress in Cold Roll Bonding of Metals," Science and Technology of Advanced Materials, 9 (2) (2008), 1-11. 3. M. Abbasi, A. Karimi Taheri and M. T. Salehi, "Growth Rate of Intermetallic Compounds in Al/Cu Bimetal Produced by Cold Roll Welding Process," Journal of Alloys and Compounds, 319 (1-2) (2001), 233-241. 4. J. An et al., "Hot Roll Bonding of Al-Pb-Bearing Alloy Strips and Steel Sheets Using an Aluminized Interlayer," Materials Characterization, 47 (3-4) (2001), 291-297. 5. N. Bay et al., "Bond Strength in Cold Roll Bonding," CIRP Annals - Manufacturing Technology, 34 (1) (1985), 221-224. 6. G. P. Chaudhari and V. Acoff, "Cold Roll Bonding of MultiLayered Bi-Metal Laminate Composites," Composites Science and Technology, 69 (10) (2009), 1667-1675. 7. S. Mahabunphachai, M. Koc and J. Ni, "Pressure Welding of Thin Sheet Metals: Experimental Investigations and Analytical Modeling," Journal of Manufacturing Science and EngineeringTransactions of the Asme, 131 (4) (2009).

Figure 6. Micrograph of interfacial zone with initial thickness ratio S\JSS varying, (a) (b) is Al/Cu sheet with Si/S s as 1.5 and 0.5 annealed at 500 D, (c) (d) represents Cu/St sheet with Sh/Ss varying from 3 to 2.4 annealed at 600 D.

8. X. K. Peng et al., "Rolling Strain Effects on the Interlaminar Properties of Roll Bonded Copper/Aluminium Metal Laminates," Journal of Materials Science, 35 (17) (2000), 4357-4363.

Conclusions According to the deformation stress analysis and microstructure on the interfacial zone, the deformation behavior and bonding of bimetal clad sheets in the asymmetrical cold roll bonding process were analyzed. The following conclusions may be drawn from the present work: 1. The deformation zone at roll gap is constituted of four districts, and the stress analysis shows that the hard metal touched with the higher speed roll has a beneficial stress field. 2. The larger speed mismatch ratio of work rolls is helpful for the couple metals to achieve a fine interfacial diffusion layer for good bonding. 3. The bonding of clad sheet is improved due to increasing thickness reductions, and the reduction of both layers increases in direct proportion with the total reduction, simultaneously, the difference between hard and soft metal gradually diminishes.

9. H. D. Manesh and A. K. Taheri, "Bond Strength and Formability of an Aluminum-Clad Steel Sheet," Journal of Alloys and Compounds, 361 (1-2) (2003), 138-143. 10. N. Bay et al., "Cross Shear Roll Bonding," Journal of Materials Processing Technology, 45 (1-4) (1994), 1-6. 11. Y. M. Hwang and H. H. Hsu, "An Investigation into the Plastic Deformation Behavior at the Roll Gap During Plate Rolling," Journal of Materials Processing Technology, 88 (1-3) (1999), 97-104. 12. D. Pan, K. Gao and J. Yu, "Cold Roll Bonding of Bimetallic Sheets and Strips," Materials Science and Technology, 5 (9) (1989), 934-939.

618

13. S. C. Pan et al., "Analysis of Asymmetrical Cold and Hot Bond Rolling of Unbounded Clad Sheet under Constant Shear Friction," Journal of Materials Processing Technology, 111 (1-3) (2006),114-120. 14. Y. M. Hwang, T. H. Chen and H. H. Hsu, "Analysis of Asymmetrical Clad Sheet Rolling by Stream Function Method," International Journal of Mechanical Sciences, 38 (4) (1996), 443460. 15. Y. Jiang et al., "Analysis of Clad Sheet Bonding by Cold Rolling," Journal of Materials Processing Technology, 105 (1-2) (2000), 32-37. 16. M. Eizadjou, H. D. Manesh and K. Janghorban, "Mechanism of Warm and Cold Roll Bonding of Aluminum Alloy Strips," Materials & Design, 30 (10) (2009), 4156-4161. 17. M. Salimi and M. Kadkhodaei, "Slab Analysis of Asymmetrical Sheet Rolling," Journal of Materials Processing Technology, 150 (3) (2004), 215-222.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

COIL BUILD UP COMPENSATION DURING COLD ROLLING TO IMPROVE OFF-LINE FLATNESS Lourival Salles de Almeida Neto1, Tuggan Ayhan2 achenbach Buschhütten GmbH; Siegener Straίe 152; Kreuztal, NRW, 57223, Germany 2 Assan Alüminyum; D-100 Karayolu Üzeri 32. km; Tuzla, Istanbul, 34940, Turkey Keywords: coil, build up, compensation, off-line, flatness, improvement understood by the automation system as zones tighter than reference. Immediately after this detection, the flatness control commands its actuators to put back such zones on target, loosing the correspondent strip fibers. The hypothesis elaborated is that releasing the tightness at center area due to coil build up during rolling will create center buckles which will increase center buckles after coil unwinding. The main difficulty is that the buckles created by the flatness control are not seen at the operator's screen. And, worst, sometimes the center buckles are visible during rolling but the operator' screen keeps showing flatness on target. This phenomenon often leads the production team to think the flatness measuring roll is not detecting the buckles, causing the machine to stop for verification. When flatness actuators release the tight center fibers, they will do it until flatness is again on target. However, as coiling keeps going on, the tight center fibers will reappear because of the convex strip profile. Center area fibers again too tight due to coil build up will command flatness actuators to put the strip back to flatness target, increasing center buckles just after the gap. This cycle repeats until end of rolling. The conclusions: a. There is nothing wrong with the flatness measurement device. It simply cannot distinguish if a zone is tight due to rolling or due to coil build up. b. The material is coiled with increasing online center buckles but not seen on flatness screen because of the coil build up stress

Abstract During rolling, overlaps of material with a convex profile at recoiler cause greater tension in strip center fibers than in fibers of the edges, which is known as coil build up. The high tension on central area is understood by the automation system as zones tighter than reference commanding its actuators to put back such zones on target, loosing the correspondent strip fibers. The difficulty is that the buckles created by the flatness controller are not seen at operator's screen. A coil build up compensation based on material profile was developed in order to have the cause of center buckles after unwinding reduced to the coiling process. From 24 coils with thickness ranging between 0.7 and 2.0mm, 100% presented better offline flatness. Flatness carpets related to profile suggested a paradigm breaking: perhaps we have to accept rolling with not so good flatness in order to have the desired offline flatness. Introduction Good off-line flatness has still been a challenge for producers of rolled coils. Although the rolling mills have been improving the ability of controlling flatness, customers are still claiming for bad flatness after unwinding. This is a strong proof that online flatness is not the only variable which determines the off-line flatness. One known effect of the coiling process is the increasing coil stress with diameter due to the positive strip profile leading to center buckles. Understand as positive profile a bigger thickness on the center of the strip compared to its edges. The common recommendation is, then, to roll with reduced recoiler tension. Accepting the profile of the material being rolled causing the effect mentioned and a cold rolling mill with automatic flatness control, one question arises: if the flatness automation system is detecting the increment in stress distribution along the strip width, shall it react with the flatness actuators? The difficulty to answer this question relies on the fact the material can become tighter in a certain measuring zone because it was reduced less in that area. Even if the gap bite does not create any flatness disturbance, measuring zones in the center of the coil will still measure increasing stress due to coil build up. Thus, it was thought about integrating this conclusion to the flatness control strategy, compensating the increment of stress due to coil build up before sending the flatness measurements to the flatness control. This paper presents the results of a coil build up compensation for flatness control during rolling of the last pass before slitter in order to improve the off-line flatness of aluminium strips in the range of 0,7mm to 2,0mm.

Design of the Coil Build Up Compensation The coil build up compensation is simply the integration of the elaborated hypothesis with the flatness control strategy. Before the measured flatness tensions are sent to the flatness controller, they are reduced by the amount of stress caused by the coil build up - characterizing the compensation developed. Doing so, the cause of center buckles after unwind will be reduced to the coiling process only. The amount of stress reduction applied by the compensation must be near or exactly the increment of stress the coil build up causes. If the compensation reduces too much the measured tensions, the flatness control will understand the material has loose center and will commands its actuator to make them tighter. This tightness created by the flatness control allied to the center tightness created by the build up phenomenon will increase the off-line errors. Thus, tuning the compensation correctly is very important. The compensation was designed to actuate without any information about the recoiler tension because, normally, production lines have difficulties to change production parameters without guarantees of a new practice will improve the situation. The hoop stress in a coil during rolling increases as the coil builds up [2], especially in the center area of the coil. This characteristic requires the compensation to be stronger as the recoiler diameter increases. Although the increment on hoop stress is present since the first winding, the coils don't present center buckles on their

Cyclical Creation of Online Center Buckles While the machine is rolling, overlaps of material with a convex profile at recoiler cause greater tension in the strip center fibers than in fibers of the edges [1]. The high tension on central area is

621

complete length when observed at the slitting line. Most of the coils sent to the slitter don't have measurable off-line center buckles at the end of the coil. The situation faced was like described on Figure 1. Technically, this behavior matches the observation made by D.T Oliver [3]. Observation at slitter line revealed that although the hoop stress increases with the diameter, the off-line center buckles has almost the same width along most of the strip length - region "a" at Figure 1. Then, it starts covering fewer surfaces as showed by region "b". Looking at the decoiler diameter at slitter, it was noticed that region "b" lasts for around 200mm, not important the thickness or width of the strip. In order to understand the reason the buckles expansion rate decreases after a certain point, one coil rolled down to 0,4 mm was uncoiled until its half diameter and the profile was measured and compared to the profile measured at the casting line, when it had 6,1mm thickness. Unwind direction

Decoiler

The wideness of the center buckles duplicated with the profile, also defining a linear function for it - Figure 3. The curve suggests that if profile is 0%, there will be no build up compensation as wideness is zero. The measured widths of the buckles were used to determine the wideness behavior of the compensation curve. For all tests made, the initial wideness was 2 measuring zones. The amplitude of the compensation curve is defined in percentage of the amplitude of the flatness reference curve so that the compensation benefits from the customer rolling experience. It starts in zero and reaches its maximum when the coil is finished. To do it, the compensation reads the decoiler diameter as soon as the mill is rolling. The maximum amplitude of the compensation curve and the strip profile were the only fields included on presetting screen for the operators. Remaining decoiler diameter Tmrnl

Recoiler

1200

>V

1000

• 00

920 725 N.

ø

region "b

Start

r

><-

0,6

region a

Buckles spread starts decreasing

Figure 1. Center buckles characteristic at slitter line - post rolling

1

The profile of the rolled coil was 5 1 % smaller than the original profile just after the caster. Relating this result with the rolling process, the explanation is that the constant generation of center buckles during rolling promotes a lower rate of stress increment during coiling and limits the buckles spread towards the edges of the strip. The compensation was designed to actuate around the strip center line, with wideness dynamically varying from a minimum to a maximum width defined experimentally. The amplitude of the generated compensation curve continually increases with the diameter until the end of rolling as the coil build up never stops.

1

1,2 Profile [%]

B l l C k l e S n O t r n A C I C , i r c , ^ i i ar»\/mr»tv»

1

Figure 2. Determination of start point of compensation at slitter line. Maximum wideness [measuring zones] A V 16 8

Compensation Parameters

™>η ^P\ \ 0,6

Decided that the compensation must be as less invasive in the rolling process as possible, it is not activated until the point the center buckles are detectable at slitter line, see "Start" at Figure 1. The first parameters to determine were when the compensation shall start actuating and how wide it must be. The production team experience is that the buckles width expansion and point of start vary with material profile. Therefore, three coils with 1,2% profile and three with 0,6% were cast with 2080mm width. These coils were rolled down to 0,8mm and sent to the slitter. Results, accepting a linear behavior, are showed by Figure 2. At any case, the compensation curve will reach its maximum wideness when recoiler diameter is 200mm bigger than the diameter when the compensation started to actuate.

1,2 Profile [%]

Figure 3. Maximum wideness depending on material profile Tests and Measurements Tests were made to prove the hypothesis elaborated and to find the best tuning for the compensation parameters. Although the compensation curve was designed to increase its amplitude and width of action as recoiler diameter increases, tests were made in opposite directions, and even with constant amplitudes and wideness, to be sure the hypothesis was correct. The study could not disturb the normal production line. To overcome this limitation, the tests were made in batches of coils

622

cast together, at the same caster, for the same final product and with same profile. The batches were composed of at least two coils because one coil was taken as base coil and, thus, rolled without compensation. The remaining coils of each batch were always compared with the base coil in order to determine if the compensation gave good results. Doing so, it was not necessary to wait always for the same target thickness to be rolled. During the test of a coils batch, all coils received the same pass before the next pass was applied to any of them. When a test batch was in progress, coils not belonging to the batch were not rolled until the test had finished. Also, it was taken care that the set of work rolls in use had rolled coils for at least two hours before tests started, suffering interruptions only due to the normal coil change procedure. To reduce the amount of variables to analyze, the compensation was activated only for the last pass before the slitter, which was considered final product. At the slitter, the off-line flatness was measured in five points indicated by Figure 4. After threading the material and cutting the scrap edge, 100 meters were unwinded. Then, machine was stopped and tensions released for the measurement to take place. The next four measuring points were made when decoiler diameter was down to 1400mm, 1200mm, 1000mm and 800mm, following the same procedure: machine stop, tensions released. On the inspection table of the slitter, the wave amplitude and length were measured and the off-line flatness calculated in IUnits.

compared to thick materials, although coils had better flatness than the base coil - Figures 6. The off-line center buckles problem increases as thickness decreases and, within the range of study, 2,0mm strips had almost not detectable center buckles at slitter when the decoiler diameter was 1400mm or less, using low recoiler tension.

Center buckles y reduction[%] 69,5 56

1 1

0,71

^ Thickness | LmmJ

2,0 '

Figure 5. Same amplitude test

Center buckles [I-Unitsl

0,71mm 2,00 mm

100 th meter of coil Outer Diameter: 1400 mm Outer Diameter: 1200 mm

Amplitude [%]

Outer Diameter: 1000 mm Outer Diameter: 800 mm

Figure 6. Same thickness test To evaluate the effect of the compensation on this thickness, a special test was made with very high maximum amplitudes for the compensation curve: three coils were rolled with only 9,5N/mm2 as recoiler tension. One coil was taken as base coil. One was preseted with 200% maximum amplitude and the other 350%. The coil with the biggest amplitude revealed heavy edge buckles at the beginning of unwinding procedure but at the first measuring point it already had lessflatnesserror than the others. One more evaluation about the behavior of the compensation curve amplitude was to roll coils without an automatic increment of the amplitude of the compensation curve. The coils didn't present good improvement compared to the base coil, reassuring the amplitude must increase during rolling. The coil with constant 60% of amplitude was over compensated in the beginning of rolling, causing online center buckles as flatness controllers understood the strip was too loose. Then, at the end of rolling when it really had tight center due to coil build up, the compensation was not enough. The coil with constant 100% of amplitude also over compensated the effect of the coiling process in the beginning of rolling. Online buckles were seen from the operator's room before the flatness measuring roll. However, the average off-line flatness was better because the amplitude of the compensation was big enough at the end of rolling.

Figure 4. Decoiler diameter points at slitter for off-line flatness measurement Results Seven batches of tests were done manipulating wideness and amplitude. To observe the effect of the amplitude of the compensation curve, the wideness behaviour was set to increasing mode. One coil was rolled with increasing amplitude and another with decreasing amplitude. Although both coils had better off-line flatness than its base coil, the coil rolled with increasing compensation amplitude gave the smallest average center buckles. This result was observed in all 3 batches tested: 0,84% ; 0,89% and 0,93% profile. Although the center buckles width increases with rolling, the same test was performed to observe wideness, setting the amplitude to increasing mode. Both coils of each test had better off-line flatness than the base coil but coils rolled with increasing wideness presented undoubted better result. Wideness tests were done with batches of 0,84% and 0,89% profile. The tests suggest the maximum amplitude shall increase with thickness - Figure 5. Within the same thickness, more amplitude doesn't necessarily mean better result. Thin material has an opposite behavior

623

Also, the flatness carpet observation is suggesting breaking a paradigm: perhaps we have to accept rolling with not so good online flatness in order to have the desired off-line flatness.

The Recoiler Tension Observation Finally, the study confronted rolling with common recommendation from coiling process studies, which is the use of reduced recoiler tension, and the compensation created. For this test the operators could not use the recoiler tension values from the plant process experience but defined low and high values. From the same batch, two coils were rolled with recoiler tension at 14N/mm2 and two at 23N/mm2. On each coil pairs with same tension, one was rolled without compensation to serve as base coil. Results at Figure 7.

References 1 J. Mignon, C. Counhaye, H. Uijtdebroeks and C. Stolz, "Improvement of Coil Flatness During Unwinding" (The Control of Profile and Flatness, Birmingham, UK, 1996) 241-249. 2. D.B. Miller and D. Nardini, "Prediction of Plastic Strain During Coiling of Sheet" (Alcan International Limited, B anbury Laboratory, Southam Road, Banbury, Oxon., U.K., OX16 7SP, 1996)

Center buckles V [I-Units] 3 4

3. D.T. Oliver, "Off-Line Flatness Problems in Aluminium Strips" (1st International Conference on Modelling of Metal Rolling Processes, London, UK, 1993) 518-524.

27,2 14

4. M. Falk, "Fiction and Reality of Aluminum Strip Tolerances" (Aluminium, Issue 74, Kreuztal, Germany, October 1998) 731738.

7,8 14

25

Recoiler tension rN/mm2l No compenination H With compe nsation Figure 7: Compensation and reduced recoiler tension The investigation point was to be sure that the coil build up compensation would improve the off-line flatness for a coil being rolled with reduced recoiler tension. And the result was positive. The test also proved that the recommendation of rolling with reduced recoiler tension is correct and has an impact on off-line flatness bigger than the coil build up compensation itself. The Flatness Carpet Observation The flatness carpets issued by the automation system of the mill were always present during the off-line analysis. It is interesting to register here that if the flatness carpet of a coil using the compensation is better than the flatness carpet of the base coil, it doesn't mean it will present better off-line flatness. It was found that for strips with convex profiles greater equal to 0,89%, if carpet was better than the carpet of the base coil, the coil after unwinding had fewer center buckles than the base coil. On the other hand, for profiles smaller than 0,89% the behavior was mostly the opposite. Conclusion The coil build up compensation developed improves the off-line flatness of aluminium strips using minimum efforts from the production team. The best combination, however, is the use of the coil build up compensation with the smallest possible profiles and recoiler tensions. The adoption of a linear characteristic for varying wideness and amplitude of the compensation curve showed well the control strategy is in therightdirection.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Through process effects on final Al-sheet flatness S. Neumann, K.F.Karhausen Hydro Aluminium Rolled Products GmbH, R&D, Georg-von-Boeselagerstr.21, 53117 Bonn, Germany Keywords: Through process modelling, rolling, winding, creep, flatness cold rolling winding and unwinding coil storage coil annealing

Abstract To achieve the desired flatness of rolled aluminium products in ever decreasing tolerances becomes more and more a key challenge. The typical process chain of strip material consists of hot rolling, cold rolling, annealing as well as coil winding and unwinding at various stages. Primarily the product flatness is dominated by the profile and internal stress generation during the rolling steps. However, in the coil some aluminium alloys show a strong tendency to reduce internal stresses by creep and relaxation. This transformation of elastic into plastic strains frequently leads to considerable changes of shape and flatness of the product over the different steps of the process chain. This paper introduces a model based through process analysis with emphasis to the evolution of strip flatness via the most decisive production stages. A case study from aluminium strip production is used to illustrate the impact of different process parameters. Introduction The rolling production process of Al-sheet involves a large number of processing steps like hot and cold rolling, annealing, slitting, surface treatments and further downstream processing. Inbetween many of these stages, the material is coiled and de-coiled again. Besides the impact of the original rolling passes on shape and flatness, the coiling has a large impact, since considerable stresses are induced throughout the coil build up. Due to the nature of this particular processing chain in conjunction with the specific material behavior of aluminium there is a strong interaction between the sequential steps with cumulative effects which may reveal at the final product.

The models need to be coupled to enable the tracing of profile, shape and flatness through a particular production chain, i.e. a Through Process Model (TPM) for matching the complete process sequence is demanded. Through process analysis The objective of this work is to present results of a through process analysis with emphasis on final flatness quality. In particular it is the objective to study the impact of rolling parameters to the flatness of a plate that has been cut of the final coil. To this end our recent ambitions have been concentrated on filling the gap between the various in-house developments as well as commercial tools towards a TPM for flat rolled aluminium products. The analysis is based on a stepwise simulation of: 1. Hot and cold rolling to calculate the thickness profile and residual stresses of the strip 2. Coiling at the final cold rolling step to predict the wound in stresses 3. The evolution of creep strains in the coil when being stored 4. The flatness in the final strip when the residual stresses are released on unwinding and cutting Note that the first simulation steps, i.e. rolling and winding simulations, are based on software tools that have been developed in-house whereas the two final steps involve the application of the commercial FE-package ABAQUS®.

A typical example in this context is the final flatness quality when a plate is cut out of a coil. Generally speaking, the flatness of a strip has to be regarded as evolving with respect to the particular stage of the process chain. The condition of the strip is strongly dependent on: material properties of the product, e.g. temper and alloy thickness profile as produced and evolving throughout the rolling winding conditions at the various steps set-up of the cutting tools temperature history of the coil

Hot rolling The hot rolling operation starts with an ingot thickness of typically 350-600 mm, which is rolled down to a hot strip of a thickness in-between 3-6 mm in several reversing passes. Depending on the mill set-up, the final two to four passes can also be performed in a continuous tandem line. While the ingot is usually cast into a rectangular cross profile, the strip gradually assumes a "cigar" like profile. To some extent this form is desired, since it is beneficial for cold rolling and following processing steps. However it needs to be controlled within a desired window. The profile essentially results of the bending of the rolls under load, work roll flattening, thermal expansion, tensile stresses, material properties, etc. Conventionally it can be controlled by choosing an appropriate work roll crown, by active work roll bending and local cooling. Furthermore, there are a number of more advanced actuators available on modern mills.

In industrial practice a given target shape is usually attained during rolling due to the available modern shape control systems. But practice shows that considerable flatness or shape deviations may still occur in the final processing stages at the slitter or in finishing lines. Accordingly a precise knowledge of the process mechanisms and their impact on the final product condition is required to detect and to avoid the nucleus of such problems. Due to the large number of process steps with an even larger number of processing parameters, it is time consuming and costly to optimize the process chain by trial and error. The development of simulation tools is required to model theflatnessand shape effects of the consecutive process stages hot rolling

A coupled model for predicting profile and shape has been developed in-house and is used for a sample analysis. The principles are described in [1]. Most notable is the necessity of coupling the longitudinal stresses in the strip, generated by profile changes, with the roll gap model [2], which takes place iteratively in this model without introducing further empirical factors.

625

Furthermore a thermal roll model is implemented to trace the temperature development over larger pass schedules. In this case a simple reversing mill is assumed for a pass schedule of alloy AA1200 of 1400 mm width. A work roll crown of 0.4 μπι has been set and no active work roll bending was applied. Since at high enough temperatures, the material usually does not develop residual stresses, only the profile of the final strip, in this case of 5 mm centre thickness, is the relevant characteristic for the following cold rolling. In this case the development of temperature at the mill has a great impact. For the model example, the mill was started without any initial thermal profile and simulation results are presented in fig. 1 for the first strip in this condition as well as for the following 5th, 10th and 20th strip, where the thermal profile of the rolls has developed.

-700

-500

-300

-100

100

300

500

2.22 |

-700

1 -500

, -300

r-2:2-J -100

1 100

. 300

1 500

1 700

distance from strip centre in mm

Figure 2: Thickness profile of hot strip 1 after following cold pass Winding The winding in-between passes has no impact on final strip flatness, since the complete plastification in the subsequent pass will remove all pre-existing residual stresses. However it can influence the performance of this pass. But the final coiling operation is decisive for the strip quality. Throughout the coil built up large stresses can be generated. This stress state has to be regarded since the coil stability as well as the strip quality is affected. Simulation tools derived to model the winding process and to calculate the wound in stresses due to the different winding parameters proved to be an effective assistant for process optimization, cf. [3]. In this study a winding software is applied that was initially derived by [4] and was further enhanced as outlined in [5]. The tool treats the winding process basically under the following assumptions: The coil is a coherent body, i.e. no contact conditions between single laps The coil material is non-linear elastic and transversal orthotropic The coil is an assembly of stacks of layers, each with "equivalent material behavior" as the (macroscopic) coil 2D axisymmetric geometrical and loading conditions

700

distance from strip centre in mm

Figure 1 : Thickness profile development during hot rolling If no further measures are taken with eventually available mill actuators, it can be seen that the thermal profile on the work rolls strongly influences the strip profile. While the first strip has a desired cigar shape, the magnitude of 2% is very high. The profile heights is greatly diminished towards the 20th strip, where it only amounts to 0.2%, but now undesired up-coming edges are observed. The impact of these two extreme cases is analyzed in the following.

It accounts for the most important process parameters of centre drive winding, namely the incoming residual stresses of the strip, the strip thickness profile, the spool material and geometry and the winding tension profile.

Cold rolling In the present case study, the strip is cold rolled to 0.28mm. Essentially only the first pass can introduce further small profile changes. With decreasing strip thickness, a profile deformation will lead to shape problems, essentially making a stable cold rolling impossible. Thus, in the simulation the cold rolling conditions (work roll crown, bending) were set in order not to introduce great profile changes, as done in rolling practice.

Concerning the two case studies under consideration, the final flatness of the strip is mostly affected by the conditions of the coiling at the last cold rolling step. It is assumed that the thickness profile remains unchanged after cold rolling to 2.3 mm. The thickness profiles can then be downscaled to the final nominal thickness of 0.28 mm to perform the winding simulation after the final cold rolling. The following parameters have been assumed for this winding step: Constant winding tension of 30 MPa Steel spool with 480 mm inner diameter and thickness of 35 mm Coil outer diameter of 1740 mm Nominal strip thickness of 0.28 mm with profiles according to fig. 1, 1st and 20th strip Constant strip and spool temperature of 140°C

The profile change of the first cold pass from 5 to 2.3 mm is presented in fig. 2. Here the true calculated profile is compared to the proportional profile which would result if the initial profile remained unchanged. The total profile deviation is already very small in this pass and will be even smaller in subsequent passes. Consequently, it is not useful to simulate all further cold rolling passes which do not affect the profile anymore.

With a view on flatness effects it is the hoop stress which is most relevant. Fig. 3 illustrates the wound in hoop stress in a cross

626

prove to be detrimental and outweigh the benefit of the large bearing fraction. Due to the upcoming edges the outer wraps of coil 2 develop an additional tension peak with a maximum value of 43 MPa at the vicinity of the edges.

section of coil 1. The stress distribution is typical for a wound strip with a cigar shaped thickness profile which cumulates to a barrel. According to the larger diameter at the centre of coil 1 large tensile stresses at the outer wraps are formed, indicated by the bright domain. This stress component will be discussed in more detail in figs. 4 and 5, which present cross sectional cuts of the hoop stress for both coils according to the positions shown in fig 2.

70.

coil 1

60.

Cross sectional cut in outer wrap

0-

¢

50

-

(0

2 40 to g

*

♦

.

*\

30

coil 2

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V

ί

* 1

♦ ♦ ♦ ♦ ♦

10.

r

100.

200.

300.

400.

500.

600.

70C

distance from strip centre [mm]

Figure 5: Hoop stress for cross sectional cuts in the outer wrap after winding (see fig. 3) Storage The winding simulation of the previous section assumes pure elastic material behavior and neglects thermal aspects. The calculations reveal that the winding step produces considerable stresses inside the coils with significant differences between both coils. Certain aluminium alloys show a tendency to level internal stresses by creep and relaxation already at room temperature and at comparative low stress levels. The corresponding mechanisms are strongly temperature dependent and can cause severe changes of the flatness inside the coils during storage. To account for the effects related to creep and relaxation a consecutive FE-analysis is applied using the commercial the FE software ABAQUS. Based on the results of the winding simulation a 2D-axisymmetric FEmodel is prepared for the coils. As in the previous section the coil is modeled as a coherent body whereas contact between spool and coil is involved in the following. A phenomenological approach for modeling creep is applied where the resulting strain rate ιcr follows a power law:

Figure 3: Hoop stress for a cross section of coil 1. The difference in hoop stress on a radial cut in both coils (fig. 4) is quite drastic. Due to the large crown of the thickness profile of coil 1 the barrel shape is more pronounced as compared to coil 2. Accordingly, large tensions up to a maximum of 80 MPa act on the strip centre towards the outer wraps. The thickness profile of coil 2 has a comparatively large bearingfraction.In general this is regarded to be positive since the tensions in the outer wraps at the coil centre are minimized. The present results show, that the large bearing fraction drops the hoop stress by -35% to a value of 53 MPa. In both cases the results reveal a compressive hoop stress at the vicinity of the spool. This indicates a considerable necking of the spool and, accordingly, the inner wraps are compressed.

=(^»[(m+i)ecrrr

where Gv denotes the von Mises stress and ε cr the equivalent creep strain, cf. [6]. Constants A, n and m are experimentally predicted for the alloy by isothermal tests. Fig. 6 shows the results of creep tests on the present alloy conducted with applied tensions of 10, 20, and 60 MPa, being a representative range for the wound in stresses.

*

coil 1 | ^ ^ ^ . *

^^< t***^***

(1)

coil 2

300.

Winding height [mm]

Figure 4: Hoop stress of cross sectional cuts in radial direction after winding (see fig. 3) Fig. 5 illustrates the hoop stress distribution of both coils in strip width direction at the outer wrap. The upcoming edges of coil 2

627

2000η 1800-1

1600 ο 8 HÖH ï J î 120oJ 2 \ ■a ßïïï-\ 800

0

24

48

72

96

time [h]

Figure 6: Stress relaxation measurements at room temperature

IOO.

200.

300.

400.

500.

- 11

distance from strip centre in width direction [mm]

The stress relaxation simulations where conducted under the conditions, that the coils were stored for 5 days. During storage the coils cool down from an initial temperature of 140°C of the last rolling pass to room temperature. The initial stress distribution is transferred from the previous simulations.

Figure 8: Hoop stress for cross sectional cuts in the outer wrap after storage (see fig. 3) De-coiling and cutting Following the coil storage, the resulting flatness of the strips after de-coiling needs to be determined, since this is a decisive quality parameter for the product. During storage, stresses inside the coil are re-distributed due to the transformation of elastic strains into creep strains. This induces local length changes in the strip. Differential lengths of single longitudinal strip fibres cause the formation of residual stresses since the strip must remain a continuous body. By unwinding and cutting of a strip segment, the stresses are partially released and may cause unflatness of a strip resting on a table without further external tensions. The flatness changes of both coils are assessed by considering single wraps of different radial positions of the coils and performing a buckling analysis for afinitesection of the wraps. The previously simulated creep strains are interpreted as a differential lengths deviation and used to calculate the local residual stress tensor ores in the decoiled strip as follows:

Fig. 7 shows the hoop stress of the radial centre cut of both coils after storage. The stress distributions of both coils are very similar and the maximum difference in hoop stress has dropped to 5 MPa. The cooling effect of the coils becomes visible at the vicinity of the spool. Due to the different thermal expansion coefficients of spool and coil, shrinkage stresses in a range of 30 MPa occur at the inner wraps.

o^fe^-kfcO-eJ·^ (2)

200.

30Ö'.'

with the local creep strain ecr and the Youngs modulus of aluminium EAi. The quantity Λcr denotes the average of the creep strain over the strip width.

4ÖÖ".'"

Winding height [mm]

Figure 7: Hoop stress of cross sectional cuts in radial direction after storage (see fig. 3) The effect of relaxation is similar in the outer wraps in width direction, as illustrated in fig. 8. There are smoothened stress progressions for coil 1 and 2 after storage and the hoop stress is drastically reduced, e.g. the maximum value of coil 1 drops from 80 MPa to -20 MPa.

628

Fig. 9 visualizes the residual stress in coil 1 after the simulated period of storage calculated according to eqn. (2). The respective residual stress field of a certain wrap of the coil is transferred to a FE-model, mapping afinitesection of the wrap,freefromexternal tensions. In particular the FE-analysis considers a finite plate with a dimension of 1400 mm x 2500 mm. Standard shell elements, S4R, are used for discretization. In a first step of the analysis the redistribution of the initial residual stress fields in the released plate is predicted and thereafter buckling of the strip is initiated and simulated. Note that dead loads have been neglected throughout the buckling analysis.

simulations demonstrate that centre buckles are to be expected most likely on the outer part of coil 1, i.e. in wrap 1. 15

A-"""

10 „ 5 E Κ 0

V \J

centre

4

\\

N

CO 3

\ A" x

-5

edge

-10 -15

500

1000

1500

2000

2500

Position in length direction [mm]

Figure 11 : Out-of-plane deformation for longitudinal sections of coil 1, wrap 1, at the centre (straight line) and edge (dotted line) Wrap 2 is more or less dominated by a homogeneous long-wave deflection, with no tendency of short-wave bag formation. Figure 9: Residual tangential stress in coil 1 before de-coiling after storage 10

For each coil the buckling analysis is conducted for two wraps, as indicated in fig. 9. The buckling pattern of the final state of coil 1, wrap 1, is presented in fig. 10. A considerable change of the initial flatness can be observed, which originally had been produced at the cold rolling mill. The tendency of long edges is completely removed and a newly developed waviness appears after decoiling. The buckling pattern of wrap 1 is a superposition of longand short-wave fractions. The long-wave fraction is a global deflection of the complete plate and acts homogeneously over the strip width. The short-wave fraction of the buckling pattern is localized in the centre of wrap 1 and forms the bags.

5

Io CO

13

centre edge

-5 -10

\A

-..

-15

500

1000

1500

2000

2500

Position in length direction [mm]

Figure 12: Out-of-plane deformation for longitudinal sections of coil 2, wrap 1, at the centre (straight line) and edge (dotted line) The buckling analysis of coil 2 is presented in fig. 13 for wrap 1. The simulation reflects the impact of the up-coming edges in the thickness profile. The corresponding formation of a hoop stress peak (see fig. 5) creates bags, rsp. waviness, at the edges of wrap 1. As observed for coil 1 the general trend of a positive crown in the profile forms centre buckles in the final flatness of the outer wraps of the coil. But the lower total profile magnitude and the large bearing fraction in the profile of coil 2 creates a better flatness quality of the plate. This is confirmed by the maximum values of out-of-plane deformation u3. The amplitude of u3 of coil 2 amounts to 15 mm in contrast to coil 1 where the corresponding value is in a range of 26 mm. Figure 10: Out-of-plane deformation u3 [mm] for coil 1, wrap 1 (seefig.9) The transition of flatness with respect to the wound height of coil 1 is demonstrated by the out-of-plane deformation u3 for longitudinal sections as indicated in fig. 10. The results plotted in fig. 11 concentrate on a centre and edge section of wrap 1, whereas fig. 12 presents the respective graphs for wrap 2. The

629

Conclusion The application of various models along the processing chain of aluminium strips has been demonstrated with a view of predicting the flatness of the final product. A special focus has been put on the through process character of theflatnessdevelopment.

O, U3

Figure 13: Out-of-plane deformation u3 [mm] for coil 2, wrap 1 (seefig.9)

It could be shown that the flatness of a 0.28 mm thick sheet is already determined by the strip condition exiting the hot mill at a gauge of 5 mm. It was demonstrated, that the profile in this condition can take effect in the final product by causing inhomogeneous stress distributions inside the wound coil, which can partially relax, depending on alloy, winding parameters and storage conditions. Upon unwinding and cutting, the remaining residual stresses are set free and the plate deforms elastically on a flat table without further external loads. The plate develops characteristic wave patterns in different positions over the width which also vary along the total length upon unwinding.

The flatness of coil 2, wrap 1 plotted on longitudinal sections in fig. 14 shows less long-wave fractions. This is most likely the consequence of the lower total profile and the superior bearing fraction resulting in lower residual stresses to be released.

The use of the through process model allows the detection of flatness relevant mechanisms in the processing chain. Thus it is possible to develop adequate rolling and winding strategies in order to improve final quality or to minimize leveling operations. References

0

500 1000 1500 2000 Position in length direction [mm]

[1] K.F.Karhausen, M.Wimmer: Walzen von Flachprodukten, Hrsg. J.Hirsch, Wiley-VCH (2001) p.89-98 [2] R.B. Cresdee, W.J.Edwards, P.J.Thomas, Iron and Steel Engineer, Vol.68, No.10,1991, 41-51 [3] James K. Good, David R. Roissum, Winding: Machines, Mechanics and Measurements, Norcross: Tappi Press, 2007. [4] M. Wimmer, "Simulation des Wickeins von Bandhalbzeugen aus Aluminium", PhD, thesis RWTH Aachen, 2005. [5] M. Wimmer, A. Cremer, S. Neumann, K. Karhausen, "Optimisation the winding of Al- strip", Proc. TMS 2009. [6] NN: Abaqus Users Manual, Version 6.10, 2010.

2500

Figure 14: Out-of-plane deformation for longitudinal sections of coil 2, wrap 1, at the centre (straight line) and edge (dotted line) In fig. 15 the same analysis is plotted for the inner wrap 2. The flatness is dominated by a long-wave deflection of the complete plate. The centre buckles nearly completely disappear. The effect of the upcoming edges seems to persist even at the interior parts of the coil which show lowfrequencywaves.

0

500 1000 1500 2000 Position in length direction [mm]

2500

Figure 15: Out-of-plane deformation for longitudinal sections of coil 2, wrap 2, at the centre (straight line) and edge (dotted line)

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S CAST SHOP for ALUMINUM PRODUCTION

ORGANIZERS

Geoffrey Brooks Swinburne University of Technology Hawthorn VIC, Australia John Grandfield Grandfield Technology Pty Ltd. Brunswick, Australia

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S CAST SHOP for ALUMINUM PRODUCTION

Casthouse Productivity and Safety SESSION CHAIRS

Leonard S. Aubery SELEE Corporation HendersonviUe, North Carolina, USA David De Young Alcoa Pittsburgh Pennsylvania, USA

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

NEW CASTHOUSE SMELTER LAYOUT FOR THE PRODUCTION OF SMALL NONALLOYED INGOTS: THREE FURNACES/TWO LINES Jacques Berlioux, Jean Louis Baudrenghien, Arnaud Bourgier Rio Tinto Alcan-Aluminium Pechiney-Aluval-BP07-38341 Voreppe Cedex-France Keywords: Casthouse, ingot line, furnace, modelling The casthouse is composed of groups of pairs operating independently of each other (other than in certain exceptional downgraded operating conditions such as: prolonged failure of an item of equipment, major potline backlog, etc.). The capacity of the furnaces is geared to that of the ingot line and the arrival of metal from the potline (size and number of metal ladles). This ensures that casting is practically continuous: the second furnace is ready to cast when the first one is empty. Figure 2 shows the new installation with Trio arrangement.

Abstract This paper deals with a layout for a smelter casthouse producing small non-alloyed ingots. The layout of aluminium casthouses has important consequences with respect to safety, investment and operating cost. Traditional solution for casthouse producing non-alloyed ingots is two furnaces per one ingot line. In order to improve the economics and logistics of the casthouse a detailed study was done, which required: • Detailed performance analysis, in particular relating to furnace/ingot line connections, impact on maintenance, operations, etc. • Hazard study • Modelling of metal flow from tapping in the potline through to the end-product, including breakdowns and downgraded operation • Thermal and flow modelling in the launder, in steady state and temporary operation • Detailed determination of the launder. It was demonstrated that a layout based on such a configuration is superior in terms of full economic cost. This design is now part of our casthouse basic engineering package.

Figure 2: 3D view of a trio: 3 furnaces-2 ingot lines

Introduction

Features of the trio arrangement are:

The study to reduce the number of furnaces in a casthouse producing non-alloyed aluminium ingots was conducted in the framework of a larger project aimed at reducing overall smelter costs (CAPEX and OPEX). The new arrangement consists in installing a group comprising three furnaces/two ingot lines (referred to as a "trio" in this document).instead of two groups of one ingot line connected to two furnaces (referred to as a "pair" in this document) as it is conventionally done. This new arrangement remains to be implemented but is adopted in new Rio Tinto projects.

• • •

Presentation of layout Figure 1 shows a conventional installation with a pair of furnaces.

•

•

• Figure 1: 3D view furnace pair: 2 furnaces-1 ingot line

635

A launder connects the three furnaces and the two ingot lines. While a single furnace is casting, the two others are being filled and in preparation. Alternatively, each furnace casts either on one or two ingot lines. Safety: a risk analysis at the design stage was carried out with the participation of operational experts from Rio Tinto Alcan smelters. This led in particular to: - giving preference to a symmetrical layout of the ingot lines with the operators' working side being separated from the machine circulation side, - physically separating the operators* working zone on the launder from the working zone on the furnaces, - defining pedestrian routes and mobile equipment routes, - analysing detail drawings of the launder. The position of the furnaces and ingot lines is such that the differences in distance between furnaces and ingot lines are kept to a minimum. Molten metal temperature loss therefore varies only slightly, regardless of the furnace/ingot line configuration during casting. The building floor area is optimised, in particular thanks to inclination of the ingot lines with respect to the casthouse centreline, and thanks to the location of the ingot lines and superpack (assembly of 24 stacks) formation equipment in the shadow of the furnaces. The symmetrical layout of the ingot lines makes it possible to keep the operators' working area completely separate from the area in which vehicles circulate. The casting stations are also close to each other, enabling a single operator to supervise both lines (refer Figure 3).

• • • •

Distance in and between shops Duration of each operation Sequencing of operations Process values: e.g.: targeted casting temperature range, metal cooling rate in the ladles and in the furnaces, reject rate. This data comes from measurements, information collected in the various Rio Tinto Alcan smelters and from suppliers. They are introduced into the model in the form of an algebraic function depending on the type of equipment. The sequencing of operations was determined with the participation of operation experts from Rio Tinto Alcan smelters. The simulations performed on the model are used to check the consistency and respect of operating criteria of a casthouse. 2. Methodology and acceptance criteria

saw «

The casthouse is modelled: • In normal mode: all equipment is available (breakdowns and short maintenance periods are taken into account) Two criteria are analysed: - The percentage of shifts and maximum number of backlogged ladles, - The over-capacity of the furnace which is estimated by the times required to accept a one shift potline tapping backlog. • In downgraded mode (long-term maintenance on an ingot line or a furnace): The criteria analysed are: - for the furnace maintenance, the percentage of shifts with backlog, - for the ingot line maintenance, the number of shifts with backlog built up during the stoppage.

Figure 3: vehicle/pedestrian traffic plan As in the case of a pair, the capacity of the furnaces is determined so as to keep the lines in almost continuous operation. In a complete casthouse, the number of lines naturally depends on production. If the number required is odd, a pair will be added to complement the trio modules. For example, a casthouse at a smelter with a production of 378 000 kt/year comprises a pair and a trio (refer figure 4).

All these criteria and their acceptance values are based on Rio Tinto Alcan's experience and are shown in Table 1.

T :*·?

Table 1 : Criteria and acceptances Criteria During normal operation, shifts with backlog

unit

%

Acceptable

TüCfcKfc

< 10%

10- 15%

"""^η > 15%

one Electrolysis shift caught up capabality

itlB

i ,«i *yri

II

d

"

1 (around 10 % over capacity of the system) shifts with backlog during furnace | maintenance (4 weeks) Generated backlog during maintenance of | an ingot line

l

r w î ! .f"

shift

% shift

<6

6-9

>9

< 40 %

40 - 50 %

>50%

<0.85

0.85 - 1

> 1

|

Accepta bi ! Validation criteria of a configuration

ς

5 greens | 1 to 2 orange(s)| 3 1 0 ^ 9 ° â 8 |

3. Basic data for the modelling

Figure 4: Casthouse layout for 360 AP37 pots

To be representative, the model must be studied in the configuration of a complete smelter. The following configuration was studied. Potline: • AP37 potline with 360 pots • Output: 378,000t/yr • Ladle capacity: 13t Casthouse: • Furnaces filled by ladle siphoning • Ingot line flow rate: 30t/h • Ingot weight: 24kg • Scrap ratio: 3% • Constitution of superpacks on leaving the line.

Production performance 1. Presentation of modelling principles The design of the system presented is based on complete modelling of the metal flow, from pot tapping to formation of the superpack. The model takes numerous parameters into account: • MTBF and MTTR of all the equipment: pot tending machines, metal ladle transport vehicles, ladle skimming stations, installation of metal siphoning system in the furnaces, ingot lines, installation of superpack formation units, etc.

636

4. Modelling results. For a pair casthouse: • five pairs each comprising two 80t furnaces and one 30t/h line. For a trio casthouse: • one trio comprising three 100t furnaces and two 30t/h ingot lines, • one pair comprising two 80t furnaces and a 30t/h ingot line. The table 2 summarizes the results from modelling the two layouts according to the validation criteria used by AP technology

Figure 5: 3D drawing of launder: furnace side

Table 2: Flow model results

|"**i

Configurations

3 Duos Figure 6: 3D drawing of launder : ingot line side

ANNUAL PRODUCTION PER LINE Average T 1 (kt)

PH

Average D1 (kt)

D3

NORMAL MODE Maximum Backlog Vo of shifts with Backlog 1 shift catch-up capability (shift) DOWNGRADED MODES Furnace maintenance 4 weeks Ingot line maintenance 1 week

122

126

A screen protects the operators in the vicinity of the furnaces: risks of splattering, etc. A monorail is used to handle launder sections. Bins collect the metal in the event that it must be drained quickly during an incident. Preheating is electrical, by hot air blasting, and covers are fitted to enhance its efficiency.

133

2 1.0%

I 3,5

12% 0,2

3 1.7% 3,9

|

42% I 0,6

2. Motorized gates The motorized gates, shown in Figures 7 and 8, perform two functions: - Shutting off or connecting the furnaces and the required lines and the drainage bins. - Limiting level disturbance when starting up or stopping a line. Excessive disturbance would result in scrap being produced on the other line currently casting. They are controlled automatically by the ingot line PLC. The gates connected with the lines are opened sufficiently slowly to limit disturbance on the line currently casting and hence avoid producing scrap.

This table shows that the output from a trio is equivalent to that of two pairs in both normal and downgraded operation. In the rest of the study, we will therefore compare one trio with two pairs capable of producing 250,000t/yr. 5. Remarks on operating mode As indicated in the section on layout, the operating mode proposed is to have only a single furnace casting. Depending on their availability, this furnace casts on one or two ingot lines. A line can be started up or shut down without any problem while the other line is casting. With this design another operating mode can also be adopted: each end furnace is assigned to one line, with the middle furnace being capable of casting on either one or the other ingot line. This operating mode is not recommended. Modelling shows that the production capacity is lower (5% on average) and that the operating mode lacks flexibility in the event of equipment break down. Shared launder This new, key piece of equipment required a detailed study of its design, a thermal modelling study and a study of the metal level disturbance caused by shutting down or starting up a line. 1. Launder design Figures 5 and 6 show the furnace and ingot line sides of the launder respectively.

Figure 7: 3D drawing of a gate

637

Steady-state operation The same model used in transient operation shows the temperature distribution on a section (refer on Figure 9 at the outlet of the furnace). The temperatures of the insulating material (about 500°C) and metal casing (20°C) comply with the requirements.

º-600

Figure 8: 2D drawing of layout of the gates Thermal modelling Aim Compared to that of a furnace pair layout, the launder has several major differences in characteristics: • very long (32m maximum compared to a maximum of 10m in a furnace pair layout) • a difference in length of metal route varying from 32 to 24m depending on the service configuration • a difference in flow rate from 30 to 60t/h depending on the number of lines in service. This modelling procedure calculates the thermal losses in the extreme configurations during the unsteady flow period and the steady flow period, in order to check whether they are compatible with the process. The aim is to optimize the thickness and nature of the insulating material and concrete in contact with the aluminium by minimizing heat losses.

Figure 9: Thermal section from modelling The model also shows that that: • The smallest loss (short length: 24m, flow rate 60t/h) is 5°C • The largest loss (long length: 32m, flow rate 30t/h) is 15°C. This difference has no impact on casting wheel operation. Conclusion By indexing the furnace temperature within a range of values between 725 and 745°C, start-up, casting, the furnace sequence and shutdown can be controlled irrespective of the configuration in service. This 20°C range does not cause problems for operation.

Modelling principle

CAPEX and OPEX costs

The model used is known as Cosmos and is based on a finiteelement calculation. The thickness and characteristics of the components (concrete, insulating material, metal casing) are input data. The calculation is run over the entire length of the launder in order to calculate heat loss. By studying the crosssection at the furnace outlet (the hottest point), it is possible to check that the conditions of use comply with the characteristics of the materials. The heat equation from the furnace spout to the nozzle is solved, taking into account heat transport, conduction in the launders, the initial state of the launders (cold/preheated), the materials and thickness of the launders.

Assumptions are: The costs are direct costs. The costing is in US$, base early 2010. The smelter is located in the Middle East. The factors differentiating between the two versions relate to the replacement of four 80t furnaces with three 100t furnaces and the installation of a shared launder. 1. CAPEX

Transient operation The model shows, that with the preheating and the use of the covers, the temperature drop on start-up is about 25 °C at the maximum and 8°C at the minimum. These values are managed without difficulty. The target temperature at the casting wheel is generally between 700 and 740°C, so the furnace should be indexed at a value between 725 and 748°C.

Gain with one trio versus two pairs Quantity Casthouse building: m2 900 Furnace civil engineering: m3 30 80t furnaces including siphoning 4 Simple launders 2 Total gain

kUS$ 1120 30 14,500 50 15,700

Losses with one trio versus two pairs Quantity 100t furnaces including siphoning 3 Launder 1 Gate automation 1 Total loss

US$ 12,300 250 150 12,700

Result

638

The CAPEX gain in direct costs of the one trio version compared with two trios therefore amounts to US$3 000k for an annual production of 250 000t 2. OPEX Furnace maintenance The study is based on an analysis of the costs of maintaining the various Rio Tinto casthouses. It reveals that the difference in maintenance costs between three 100t furnaces and four 80t furnaces amounts to US$200k per year, in favour of the 100t furnaces. Energy consumed by the furnaces The reference value in the casthouse with 80t pairs is 450MJ/t. Impact of the number and capacity of furnaces Casthouse furnaces are kept permanently hot. In a casthouse where there is little scrap (in the order of 3%) and the process does not require preparation, we adopt the hypothesis that the energy consumed only serves to keep the furnaces at the required temperature and that it is proportional to furnace capacity. 3xl00t in the trio casthouse corresponds to 4x80t in the pair casthouse, i.e. a consumption of (3x100/4x80) x450 = 420MJ/t i.e. a gain of 30 MJ/t. The gain, for an output of 250,000t and a gas cost per MJ of US$0.01 is therefore equal to US$75k per year. Result The OPEX gain of one trio compared with two pairs is therefore US$275k per year.

Conclusion In an ingot casthouse, the trio furnace arrangement (three furnaces supplying two ingot lines) offers substantial gains in comparison with the traditional solution using pairs (two furnaces supplying one ingot line). In the particular case studied in this document, for a production of 250 000 t/year, a trio of three 100t furnaces supplying two 30t/h ingot lines has an output equivalent to two pairs of two 80t furnaces supplying one 30t/h ingot line. While at the same time meeting the HSE criteria, the CAPEX gain amounts to US$3 000k and the OPEX gain to US$275k per year. This represents a NPV of US$4 000k This arrangement forms part of the AP technology package and is retained for Rio Tinto Alcan projects.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

USE OF PROCESS SIMULATION TO DESIGN A BILLET CASTHOUSE Gwenola Jaouen Rio Tinto Alcan - Smelter Technology - Centr'Alp - BP 7 - 38341 - Voreppe Cedex - France Keywords: Casthouse, Modeling, Simulation, Billet It was clear that process simulation modeling would be required to allow effective casthouse design. Process simulation is a modelDesigning a casthouse for billets is a complex activity. It must be based representation of physical, chemical or other processes and operations in software. The goal of process simulation, which is a correctly sized to optimize metal flow and continuously feed the well established technique [1], is to find optimal conditions for an homogenization shop. It must also be borne in mind that overexamined process. Within a few months a new process simulation sizing adds no value and is costly. To solve this problem, tool had been developed. A shop model was built, calibrated and casthouse operations must be analyzed accurately in real time using a discrete simulation model. Such a model was developed, optimized so as to have a well sized billet homogenization shop. with a library containing all the necessary equipment: conveyors, Figure 1 shows the billet homogenization shop designed with the new tool. continuous and batch furnaces, andfinishing stations. Abstract

The model was used to design a fully integrated casthouse, requiring flexible, robust facilities and management of complex product mixes. By combining this model with the metal-flow sizing model, design execution and its impact on casthouse operation was optimized. The result is a shop designed for customer needs, at optimal cost. Introduction High value-added products are well rated in the aluminium market for their specialty and their manufacturing complexity. For marketing strategy reasons, a project to build a billet casthouse to cope with the increase in capacity of the Isal smelter in Iceland was launched in 2009. The aim of this project is to build a fully integrated billet casthouse: a billet casting shop complemented by a billet homogenization shop.

Figure 1 - Billet homogenization shop design.

As-cast billets need to be homogenized to improve extrudability and ensure specific mechanical properties in the extruded profiles. Billets are homogenized in either batch or continuous homogenizing furnaces. They are heated and held at a temperature slightly below their melting point for several hours. They are then rapidly quenched in air to below 300°C in dedicated coolers, or in a cooling chamber in the case of a continuous furnace, and further cooled down to below 80°C.

Metal flow modeling The casthouse study was separated into two parts: • •

In order to complete the studies as quickly as possible, the teams were divided into two. The first team was in charge of building a metal flow model on the basis of an existing library. This library contains modules representing potline tapping groups, metal ladle transport vehicles, ladle cleaning, furnaces and casting pits. It covers the required product mix, furnace charging times, metal preparation times, equipment reliability, etc. The resulting model is shown in Figure 2. It was first calibrated using current plant data, then tested in both nominal operation and downgraded mode (four-week furnace shutdown for refining, one-week maintenance work on a casting unit).

The product mix to be considered for the casthouse was broad: • • •

Metal flow between the potline and billet casting. Billet homogenization.

Commodity (6060 and 6063 alloys), Specialty 1 (6xxx alloys) and Specialty 2 (3xxx alloys) alloy billets. Diameters from 7" to 10". Lengths varying between 6 m and 8 m.

Sizing casthouses is a regular activity for the Rio Tinto Alcan engineering site in France. Tools are used for modeling the metal flow between the potline and the casthouse furnaces, the metal preparation area and billet casting. All dimensional designs are analyzed and validated using flow simulation models. The challenge of this project was that there were no tools available for modeling a billet homogenization shop. It was therefore difficult to design the casthouse in its entirety.

The model was used to optimize casthouse dimensional design and metal ladle flow logistics. It integrated process and potline operating constraints.

641

Figure 2 - Schematic diagram of modeling metal flow between the potline and casthouse • • • •

Development of a new tool The second part of the casthouse study included construction of a tool for modeling a billet homogenization shop and then modeling the shop required for the study. We did not have this modeling tool at our disposal. Some equipment suppliers used simulation models, but they did not take the whole workshop into account. It was more prudent to incorporate all shop management rules, conduct the tests and then be in a position to check if results tallied with those of suppliers.

Conveyors Each conveyor operates independently. It moves the billets one by one according to a predefined cycle and its own breakdown rules. Inspection, sawing andfinishingstations

We also wanted to connect the two models together, taking billet casting times given in the metal flow model results and importing them into the homogenization shop model. The tool had to cater for the following needs: • • • •

Stacker De-stacker Automatic handling trolley Billet finishing and marking station

Each device processes the billets one by one and is fully configurable (cycle times, breakdowns, etc.). Figure 3 shows the modeled billet sawing station.

Model a complete shop, from billet stripping to the finishing and bundle-marking stations, Launch standard production plans corresponding to the required product mix and other non-standard plans, Visualize bottlenecks, Validate shop dimensional design.

The tool was developed using ExtendSim software, selected for its robustness, its integrated database and easy interfacing with the model, its three-dimensional development capabilities and the fact that the team was already familiar with it. All equipment items were modeled, resulting in a library dedicated to this type of shop, including: • • • • • • • •

Conveyors Visual inspection Saws Helical ultrasonic inspection Continuous homogenization furnaces Batch homogenization furnaces Coolers Bundle storage areas

Figure 3 - Model of billet sawing station

642

Continuous furnace Temperature

The continuous furnace is intended for homogenizing Commodity and Specialty 1 type billets. Billets enter the furnace one by one on a conveyor set at a rate that varies according to the required homogenization time. Figure 4 describes the continuous homogenization cycles. The furnace is divided into three zones for heating, temperature-holding and cooling. The model of the continuous furnace is shown in Figure 5.

\

CONTINUOUS FURNACE

x € { 7 , 8 , 9 , 10}

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χÈ{7,8,9,10}

Upstream of the batch furnaces, as shown in Figure 7, a machine automatically stacks billets from a given cast to form bundles. These bundles are then taken by charging car to one of the batch furnaces. The homogenization cycle starts up in the batch furnace. Once completed, and if a cooler is available, the charge is removed automatically by the charging car. If no cooler is available, the charge is held in the furnace.

. ^ ^ ^ ^ ^ ^ ^ " Hip»ii^^ja^a

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Figure 6 - Batch homogenization cycles

Figure 4 - Continuous homogenization cycles [ — —

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After the cooling cycle, the charge is finally moved to the de-stacker. The bundles are de-stacked automatically and placed on a conveyor. Temporary storage areas are provided between the furnaces in order to absorb variations in charges.

^ Ρ £ £ 2 |

^ — ^ -4-1

Figure 5 - Model of continuous homogenization furnace Batch furnace The batch furnace is intended for homogenizing all types of billet. Process times, as shown in Figure 6, are calculated according to billet type and diameter.

Stacker

Charging car

Charging car

Cooler

Charging car

De-stacker

Figure 7 - The stages of batch homogenization process

643

Shop modeling Once the new tool was developed, the next stage was to assemble the modules representing the equipment, just like assembling Lego® bricks. This provided the main features of the shop: two billet feed conveyors, two visual inspection stations, one helical ultrasonic inspection station, two saws, four batch homogenization furnaces + two batch coolers, one continuous homogenization furnace, and two finishing stations.

Each intersection was therefore programmed at a high level of complexity. On leaving the billet sawing station, Commodity billets are preferably sent to the continuous furnace, and Specialty 1 and 2 billets preferably to the batch furnaces. Equipment overloading always needs to be considered. For example, it is sometimes preferable to occupy a batch furnace with a Commodity billet cast. Otherwise, the helical ultrasonic inspection station may become blocked by a Commodity billet cast waiting for the continuous furnace to become available. A shop power limit also had to be incorporated. For reasons of available smelter-power levels, no more than one batch furnace can function in operating phase 1 at any given time. Before operating a batch furnace it therefore has to be checked that no other furnace is already in this initial phase of its heating cycle. Figure 9 shows a bundle waiting in front of a batch furnace.

Figure 9 - Model of the bath furnace showing, in the foreground, a bundle waiting to be loaded. Production plans were validated with experts in the subject. The plans had to guarantee that sales requirements be met and installation overloading avoided as much as possible. Product changes, which are always detrimental to production yields, had to be limited. Each equipment item was configured (cycle time, availabilities, etc.) by comparing supplier data with internal plant records. Parameters were adapted whenever considered necessary. The modeling of the shop is shown in Figure 10.

Figure 8 - Flow diagram of billet flow through the workshop The billet flow through the workshop is shown in Figure 8. Once the billets are stripped, they are laid on one of the two conveyors at the shop entrance. Specialty 1 billets must be moved to a helical ultrasonic inspection station, while Commodity and Specialty 2 billets go directly to the sawing station. After sawing, Commodity billets are preferably sent to the continuous furnace, while Specialty billets go to the batch furnaces. On leaving the furnaces, all billets are sent to thefinishingand marking stations. The difficulty at this stage was to correctly define the system rules, in particular at conveyor intersections. A major constraint was to avoid splitting casts, even in the event of equipment failure or conveyor overloading.

Figure 10 - Plan showing simulation of shop operation in progress

Results obtained

Bottlenecks concerning:

Rules

Data

One year of production was simulated. Using the software it is possible to track variations on graphs in real time, view animations and see the shop load factor (utilization rate, number of waiting casts, etc.). Simulations quickly showed that some rules had to be adjusted. For example, it was preferable to send all Specialty 1 billets to the batch furnaces because the continuous furnace was a critical item of equipment and was not to be slowed down. In billet casthouses, it is generally recognized that the major criterion is waiting time at the casting pits, otherwise the casthouse becomes the smelter bottleneck. A threshold was therefore set at: • •

A delay of one potline shift at any time (i.e. 2 casts waiting in our case) over a year in normal configuration. A delay of two potline shifts in downgraded configuration.

Figure 13 - Plan of the facility showing locations of bottlenecks Examples of bottlenecks:

The downgraded configuration adopted corresponds to exclusive production of 6 m long billets for two consecutive weeks. This is equivalent to increasing the number of billets by 14% during this period. Figures 11 and 12 show respectively, acceptable and unacceptable levels of casts on hold at pits, from simulation results over a one year period.

• • •

• •

Layout: The buffer stock upstream of the continuous furnace is very important for smoothing the charge, and hence maximizing the continuous furnace filling rate. Layout: The number of waiting spaces at the batch furnaces helps absorb production variations and allows a degree of flexibility in the production plan. Layout: It is vital to set aside room to store casts upstream and downstream of the helical ultrasonic inspection station. This disconnects the station from the rest of the facilities and avoids blockages (namely in cases of breakdown). Rules: In production plan, billet diameters have to be smoothed. Data: The shop power limit has to be respected.

Different strategies were tested for managing the facilities. Impacts of the various products (type, diameter and length) were analyzed and quantified, and recommendations formulated for scheduling production:

300

Figure 11 - Graph of simulation results for casts on hold at a pit over a 1 year period (acceptable level)

•

•

365 days

Figure 12 - Graph of simulation results for casts on hold at a pit over a 1 year period (unacceptable level) The simulations identified the bottlenecks at the facilities. They were classified according to three types: layout, rules and data as shown in Figure 13.

Impact of 7" billets - essentially Commodities: given the large number of billets to be processed one by one in the continuous furnace (112 billets per cast), 7" Commodities casts are critical in the installation. Conversely, 10" Commodities (56 billets per cast) enable the shop to empty the buffers. The recommended sequence used for the first pit production plan is shown in Figure 14.

7" Commodities

7 successive casts

Buffers saturated Continuous furnace full

10" Commodities

14 successive casts

Buffers and continuous furnace "empty"

8" Commodities

21 successive casts

Buffers and continuous furnace moderately saturated

9" Commodities

7 successive casts

Buffers and continuous furnace relatively "empty"

Figure 14 - Recommended production plan to minimise bottlenecks

645

Impact of Specialty 2 billets: due to long homogenization time (10" billets are almost twice as long as Specialty 1 billets of the same diameter); it is preferable to smooth Specialty 2 production as much as possible.

Conclusion The Isal smelter project was launched in September 2009. Its billet casthouse will be completed in April 2012 and it will reach full capacity in July 2014. Casthouse operators, equipment suppliers and the engineering department collaborated throughout this project, combining optimum technological design with best operating practices.

The recommended production plan adopted in the simulations is a Specialty 2 - Specialty 1 - Commodities sequence lasting 18 days on the second pit, followed by production of Specialty 1 for 22 days. This production plan matched our dimensional design criteria. One of the results thus obtained was the utilization rate of the various equipment items, as shown in Figure 15. Delay on the stripping on one cast: Utilization rate of saw 1 Utilization rate of saw 2 Utilization rate of batch furnace 1 Utilization rate of batch furnace 2 Utilization rate of batch furnace 3 Utilization rate of batch furnace 4 Utilization rate of batch cooler 1 Utilization rate of batch cooler 2 Utilization rate of finishing station 1 Utilization rate of finishing station 2

Future operators are now aware of the impact of key parameters and suppliers have had the chance to test and calibrate their simulation models. The Rio Tinto Alcan engineering department has at its disposal a new tool capable of simulating the operation of a billet homogenization shop. The models developed are available for future shop optimization studies and can be used for other billet casthouse projects.

Max 2 casts 55% 0% 91% 89% 90% 87% 48% 15% 73% 0%

References [1] Professor Stewart Robinson, "Simulation: The Practice of Model Development and Use", Warwick Business School, University of Warwick. [2] Jean-Franηois Claver, Jacqueline Gιlinier, Dominique Pitt, "Gestion de flux en entreprise, Modιlisation et simulation".

Figure 15 - Predicted equipment utilization rates The second saw is only used as a backup when the first saw is down for long periods. Both coolers are required to operate batch furnaces at maximum capacity. The second finishing station is only used in downgraded configurations (temporary shop overproduction). During these critical periods, putting the second station into operation avoids slowing down the furnaces and keeping products waiting at the casting pits. Product mix changes The model was used to track changes in the product mix required by customers, assess different layouts and define buffer needs. In total, about one hundred configurations were tested. These were used to fine-tune the layout, define the project phases and optimize the shop according to the space available in the buildings and the required product mix. With the aim of maximizing value, the starting hypotheses were changed and the product mix modified to include more Commodities and only Specialty 1 billets. The batch furnaces, which are much more costly than continuous furnaces, could therefore be eliminated. The model was adapted to meet the needs of this new homogenization shop comprising, at most, three continuous furnaces. Furnace capacity was tested, billet distribution challenged, billet circulation in the shop reviewed and buffer sizes studied in detail. This resulted in a solution compliant with our dimensional design criteria: a casthouse that is neither oversized nor a bottleneck for the entire smelter.

646

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

OPTIMIZING SCRAP REUSE AS A KEY ELEMENT IN EFFICIENT ALUMINIUM CAST HOUSES Ir. Jan Migchielsen, Dipl. Ing. Hans-Walter Grab, Tom Schmidt Otto Junker GmbH, Germany Keywords: Stimulus, cycle, durability, garbage to garbage, separation, fluxing, dilution, twin chamber furnace separately collected. With a recycle rate of 65%, many tons of aluminium, paper and plastic is recovered from these packages, while in the rest of the world these packages, that consists of valuable components surely will end up in the incinerators [5]. Worldwide only about 45% of the aluminium packaging material is recycled. In tonnage aluminium, there is still a treasure to be recovered.

Abstract In the last decades, the market share of secondary aluminum is slowly replacing primary aluminium. However, as the demand for aluminum is still considerable larger than the collection of used aluminium, primary aluminium will also be needed for the coming decades. This paper will address the current position of this development and the possibility to further increase the part of used aluminium scrap as the source for production of new aluminium to come to a durable production. The second part of the paper will concern the melting facilities for recycling used scrap. Introduction Remelting of used aluminum takes only 5 percent of the energy that is needed for producing primary aluminium [5]. Recycling is therefore not only interesting from an environmental point of view, but also commercially attractive. The growth of the worlds recycling capacity has led to a structural shortage of aluminium scrap, which even in today's economic crises still exists. The structural shortage led to a high scrap price that is an important incentive for the collection of used aluminum scrap [2]. Together with the stimulus from many governments, the collection of used aluminum is steadily growing. Aluminum can be recycled indefinitely and is a durable solution for many applications.

Figure 2: UBC scrap Cycle For recycling of used aluminum there are two main approaches, referred to as the "closed cycle" and the "open cycle". A typical example of a closed cycle is producing can stock from UBC (Used Beverage Cans). The scrap has already the right composition and is supplied in a uniform shape. In the closed cycle the material can be re-used endlessly without change of alloy or quality. This is a durable way of operation. Can stock is a material with a short life-cycle. Only about eight weeks after production, the material is back in the factory in the form of UBC (see Figure 2). Aluminium from the automotive sector takes about twelve years to return, while it takes more than a human generation for material from the building sector to return. The open cycle is using a mixture of all kinds of aluminium scrap to produce new aluminum products. The recycling process, frequently referred to as refining, makes use of low grade, low cost scrap with the purpose to produce high quality aluminium alloy out of it. Low grade aluminum scrap is not only contaminated with all kinds of unknown matter, but in many cases it will consist out of a mix of alloys. For making a useful metal out of it, it will be needed to dilute the metal with primary material or with a higher grade scrap with known composition.

Figure 1: Aluminum use. Source: Aluminium Association Stimulus The high scrap price is an important incentive to collect used aluminium from buildings, automotive and used appliances. However, with regard to packing materials, additional stimulus is required. With improved production techniques, there are too many cans in a kilogram of UBC to make it worthwhile to collect all of them [5]. A nickel refund per can makes the difference and leads to almost full recycling. In Germany used Tetra pack is

The two types of recycle facilities have different needs with respect to the remelt facilities. The UBC recycler needs an installation that is specially designed for UBC processing with

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respect to highest recovery and minimum operating cost. The cost price of UBC scrap and can stock is both determined by the market. The recycler must operate a cost effective process, inbetween these prices to keep in business.

foreign matter and various alloying elements. Recycling is more than just melting. It is the modern variant of alchemy, a true art of transferring a dirty hot soup into precious metal. There are various means to enhance this operation: Separation Fluxing Dilution Holding furnace with porous plugs Degas ser Metal filter

Looking at UBC, the recycle rate is approximately 45% [5]. This means that still most can stock must be made from primary aluminium. All available UBC should be preferably used for making new can stock as it has the ideal composition for that. Using UBC for other purposes, by changing the composition, is a waste of valuable alloying elements and will be less profitable than re-using it for can stock.

Separation

Figure 3: Various types of scrap: Packaging and old coolers The recycler in the open cycle needs a melting furnace that eats anything. The different types of scrap come in different sizes and shapes. The recycler tries to make an optimum mix with regard to alloys, but the furnace must be versatile to handle any type, composition and shape of scrap (Figure 3). The target of this operation is to produce a useful quality metal with lowest cost input materials. The success of the enterprise depends strongly on the experience of the operators and control of the remelt process.

The first aspect is build into the design of the twin chamber furnace. It has a heating chamber, referred to as clean chamber and a melting chamber called dirty chamber. The furnace melts contaminated scrap in the dirty chamber with superheated liquid metal from the heating chamber. The concept of the furnace is that the dirt and dross remain in the dirty chamber while the metal that is tapped from the clean chamber is relatively clean. The dirty chamber needs to be skimmed every one or two hours, while the clean chamber needs to be skimmed only once a day. Fluxing With fluxing salt, the metal is cleaned from part of the oxides and foreign matter. The reactants will float on the surface of the bath as dross that can be skimmed off. Fluxing is a pre-treatment step in increasing the metal quality. Fluxing salt has negative side effects on the refractory and the filter plant. Many operators today decide to operate the furnace without fluxing salt. The quality of the metal can be compensated with selecting a high end degassing system.

Durability The closed cycle recycling is really durable. This is only partially the case for the open cycle. Here are still possibilities for improvement, not only for the melting facilities, but also for the packaging industries. In the filter plant of the remelt facilities, substantial amounts of Chlorine and other noxious substances are found in the additives. The origin of these substances must be in the coatings used on the scrap. Apparently producers of aluminium appliances and even of packaging materials seem still to have little concern for the end of the lifecycle of their products. Noxious components in the used filter dust make the dust useless for any further application such as building materials.

Dilution Adding primary metal into a twin chamber furnace is in many cases needed to get the metal on spec. As it is virtually impossible to get most alloying elements from the melt, the only way to reduce concentration is to dilute the melt with primary metal. Dilution is a relatively expensive way for increasing the quality. Analyses and pre-selection of the scrap should always have the preference. Controlling the concentrations at the entranced of the process is far more economic than diluting the large content of metal of a twin chamber furnace.

Low grade scrap contains many alloying components that are virtually wasted when making a new alloy out of the scrap. For making the alloy on spec, elements can be leached out with Chlorine or concentrations can be lowered by diluting with primary material. This makes that only part of the material is actually recycled. Another part needs to be supplemented with primary metal. Garbage to garbage In scrap remelting a general saying is that one makes garbage from garbage. Remelting low grade aluminum scrap makes low quality aluminium, low quality with regard to high level of oxides,

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Secondary cast house For recycling, the larger the amount of metal that is processed, the more economic the operation will be. However, it must be noted that melting and casting 200 MT per day is no big deal, obtaining 200 MT of suitable scrap, every single day, is! The secondary cast house must therefore be designed for having maximum flexibility in handling various types of aluminium scrap, to respond to the fluctuations on the scrap market. Both rotary type furnaces as twin chamber furnaces can handle scrap in various shapes, different alloys and are flexible in handling contamination. Rotary furnaces, as shown in Figure 5, are more suitable for small and varying quantities and qualities of scrap. Apart from contaminated scrap, also aluminum dross may be processed in these furnaces. Figure 4: Twin chamber furnace (on back) with tilting holding furnace (on front) Holding furnace

Twin chamber furnaces are designed for large amounts of scrap with a metal content of 90% and more [3]. Although flexible in type and consistency of scrap, the twin chamber furnace should be used at full capacity most of the time and with not too much variation in alloy. Due to its size, operation at part load results in relatively high fuel consumption per ton of scrap. Twin chamber furnace

For casting metal that is produced with a twin chamber furnace, it is in general required using a separate holding/casting furnace (Figure 4). The metal transfer to the holding furnace results in cleaner metal in the holding furnace by retaining dross and sludge in the twin chamber furnace. Equipped with porous plugs, the holding furnace can directly contribute to an increased metal quality. The gas bubbles from the porous plug system not only take out hydrogen from the melt, but the bubbles take oxides to the surface of the melt as well. Together with the inline degasser and the filter, the metal quality at the casting machine is sufficient for casting extrusion billets or rolling slabs. Developments to purify aluminium with the partial crystallization process are promising but still in a test phase.

Figure 6: Twin chamber furnace for Hindalco Aluminum has always been a valuable material. As the production of aluminium product goes with quite some production of internal scrap, there was always a need for furnaces that remelts the internal scrap. Internal scrap is mostly clean and pure material that can be processed easily. Further the amount of internal scrap is the smaller part of the total production. Part of internal scrap is formed by cuttings and thin foil. Processing this type of scrap is best with an (open) side well furnace by indirect melting. For processing contaminated scrap, the open side well type of furnace was further developed into the closed side well furnace with closed (electromagnetic) pump, also referred to as twin chamber furnace. A bottleneck in the development of these furnaces, during a long time, was the pumping capacity of both the electromagnetic as the mechanical pump. Only since a couple of years, larger pumps exist on the market that allowed scaling up twin chamber furnaces to melting capacities beyond 10 tons per hour [3]. The types of twin chamber furnaces that were developed by Otto Junker mainly

Figure 5:10 Tons rotary furnace

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serve the "open" cycle remelt facilities. It makes no sense to optimize these furnaces for a certain type of scrap as they will be fed with different scraps from day to day. Hot circulation fan

Description of the furnace The furnace has two chambers, connected via a metal pump with LOTUSS bowl [3]. The main heat source is formed by regenerative burners. Heat for pre-heating is supplied by a hot gas circulation system that transports contaminated fumes to the thermal oxidizer and brings back hot flue gas to the melting chamber. A docking type charging machine charges the scrap into the melting chamber with minimum emission of fumes to the cast house environment. Finally a bag house filter plant cleans the flue gasses from dust and dirt. Due to volatility of the scrap market, the furnace is designed in a way that has maximum flexibility with regard to alloy, scrap shape and contamination.

Metal pump + charge well

Figure 7: Simplified schematic of a twin chamber furnace The scheme shows a twin chamber furnace with a thermal oxidizer for combusting the volatile contaminants that are released during the melting process. Alternatively these contaminants can be combusted in the heating chamber of the installation. The latter has both advantages as disadvantages. The major advantage is a reduction in fuel consumption. Keeping the external thermal oxidizer on operating temperature requires fuel while combusting the fumes in the heating chamber make optimum use of the energy content of these fumes.

Large volume scrap recyclers mostly used baled scrap as these have logistical advantages. Loose scrap has such volume that it becomes impossible to charge it into a modern furnace in the required production rate of the furnace. Flexibility has its price with regard to the downstream processing of the fumes. Thermal oxidizer and filter plant needs to be prepared for unknown fume supply.

Major disadvantage of combustion in the heating chamber is that with the combustible fumes an amount of dirt and noxious components come into the heating chamber. The clean chamber now gets polluted with dirt and dust, which also may end up in the metal and have to be removed in the downstream processes. The dirt also collects in the regenerators of the burners of the heating chamber, which results in increased maintenance time and cost. Melting process

Figure 6: Twin chamber furnace with docking type charging machine

Objective of the recycle facility is to have a high as possible metal recovery. In addition to separating flames from thin gauge scrap, it was learned from observations that the highest recovery is obtained with quick melting of the scrap. In the first generation of twin chamber furnaces, especially melting of baled contaminated scrap took a long time. It seemed that the longer the melting time took, the higher the metal loss. Temperature profile bales

The twin chamber furnace separates the burner flames from the combustible scrap. In a large heating chamber, the heat from the burners is transferred into a bath of liquid metal. The superheated liquid metal is pumped into the relatively cold melting chamber, where the thin gauged scrap is molten by submerging into the liquid metal. Before the scrap is submerged, it is preheated with almost inert flue gas in order to dry the metal from organics.

Figure 8: Temperature profile in submerged bale

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For improving the furnace performance, the melting process had to be improved. We needed to have better understanding of the physics of the melting process. Mathematical models of heat and mass transfer in combination with temperature measurement of full sized bales in a test furnace showed that in the melting process in a twin chamber furnace are governed by metal velocity rather than by heat transfer. Figure 8 shows the metal temperature in a bale, calculated with the method of Euler [3]. From the figure it may be evident that the outer skin is already melting while the inner part of the bale is still relatively cold. In a stream of superheated liquid aluminum, the viscous forces will take the melting outer skin and dissolve the material. Further calculations resulted in a quantification of the dissolving of the bales, as shown in Figure 9.

References [1] Richard C. Chandler, "Options for Optimizing Recoveries and Energy Consumption in Light-Gauge Scrap Recycling", Metaullics Systems Co. L.P. [2] A.L. Steward, J.G. McCubbin "Melting Aluminum and Aluminum Alloys", Aluminium Company of Canada Ltd. [3] J. de Groot, J. Migchielsen, "Multi Chamber Melting Furnaces for Recycling of Aluminium Scrap", Casthouse Conference Brisbane 2003. [4] W. Trinks e.a. Industrial Furnaces; sixth edition, Wiley, 2004 [5] A.L. Steward, J.G. McCubbin "Melting Aluminum and Aluminum Alloys", Aluminium Company of Canada Ltd.

By melting with high metal velocity, the heat transfer value goes up and the bale will be peeled off quickly. The blocks quickly dissolve, rather than slowly melt [3].

[6] Werner Heiligenstaedt, Wärmetechnische Rechnungen für Industrieφfen, Stahleisen-Bücher, Düsseldorf, 1966.

Dissolving cube for different metal velocities (Re for 350 mm cube)

[7] Rosenow, Warren M., Hartnet, James P. Cho, Young I., Handbook of Heat transfer, McGraw Hill, 1998

200

300

400

Time in seconds

Figure 9: Dissolving time for different metal velocities In the new generation twin chamber furnaces, we keep the metal velocity in the melting chamber at a level that all metal is molten before the next charge. The highest recovery is achieved with high liquid metal temperature and high metal velocity. For further improving the melting velocity and - at the same time - recovering heat, the bales are preheated on the charging ramp of the furnace. By keeping a charge of relatively cold metal on the ramp all the time, waste heat from the flue gasses is transferred into the metal. Improvements It has been proven that governmental stimulus has a major success on the recycling rate of used aluminium packaging materials. The recycling of used aluminum scrap is not only beneficial for energy consumption and the associated carbon dioxide production, but at the same time it reduces the amount of waste that is produced and will result in a durable economy. When, for packaging materials, only coating materials would be used that contain no noxious substances, the recycling process would be more durable. In that case the dust from the filter plants would have a further application as building materials.

65

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

IMPLEMENTATION OF AN EFECTIVE ENERGY MANAGEMENT PROGRAM SUPPORTED BY A CASE STUDY Roger Courchιe TMS (The Minerals, Metals & Materials Society); 184 Thorn Hill Rd.; Warrendale, PA 15086, USA KB Alloys, LLC: 2208 Quarry Drive, Suite 201, Reading PA 19609, USA

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Keywords: Energy, Management, Program •

Work with the Working Party to identify and provide training, review data and set and monitor targets This paper contends that an Energy Management program not driven by Upper Management has a high risk of failure.

Abstract Effective energy management programs require commitment and leadership from Upper Management supported by an experienced Energy Manager. The vehicle for change is the Energy Management Working Party. The Working Party will be led by a Senior Manager and include other line managers under the technical guidance the Energy Manager. The first step is to use historical data to calculate an energy management performance indicator called Specific Energy Consumption [SEC] and then continue the calculations using real time data. The Working Party will initiate programs to achieve a continuous improvement in the SEC and in conjunction with Upper Management agree long term targets. Techniques will include evaluating base load data (energy consumed when not in production), targeted projects on energy inefficient equipment and a review of the various manufacturing processes to develop low energy solutions. This presentation is supported by a case study of the KB Alloys facility at Wenatchee WA

Energy Management Working Party This is the team charged with developing and implementing the Energy Management Program. The following are considered to be key appointments: Chairman • This role should be undertaken by a Senior Operational Manager for example the Plant Manager Energy Manager • Clearly an important appointment - in the absences of anybody with Energy Manager skills and experience the Engineering Manager or Maintenance Manager would be good choice although some training may be appropriate Other Members • Depending on the size of the company it would be appropriate to add two or three members who should be managers or supervisors who run energy consuming departments - for example Production Manager • Additional members can be added as required as the Working Party begins to make progress

Introduction This paper outlines a method of implementing an effective Energy Management Program in a manufacturing facility. The paper focuses on the need for leadership and commitment from Upper Management and its appointment of the Energy Management Working Party which is the vehicle for driving the process forward. The paper addresses the organization of the Working Party and the work programs to be undertaken with the first priority being the setting up of a metric to calculate an Energy Performance Index which is called Specific Energy Consumption.

Meetings should be held at least monthly and they must be formal with appropriate notes or minutes reporting progress, results and actions planned (by whom and by when) Energy Management Working Party - Work Program 1 - Measure Energy Management Performance

The paper concludes with a progress report on the results achieved at KB Alloys, LLC facility Wenatchee WA using the Energy Management Working Party as the method of implementing an Energy Management Program

The first task of the Working Party is to measure the existing energy performance of the facility and then put in place programs to continuously improve that performance. To achieve this we need a meaningful metric in the form of an Energy Performance Index and as energy is used to produce things we are looking for a metric that relates the energy consumed to the saleable production produced. This Energy Performance Index is typically called Specific Energy Consumption or SEC for short and is calculated by dividing the total energy consumed by the total saleable production Typical units are kWh/#

Management Commitment For an Energy Management Program to be effective it requires leadership and commitment from Upper Management with the following being crucial: • Publish an Energy Management Policy demonstrating the Company commitment to energy management and the methodology to be implemented • Provide resources (both people and finance) • Appoint the Energy Management Working Party

Calculating SEC Step 1 - Identify the types of energy to measure Step 2 - Identify means of recording the energy used [probably from invoices and or meters]

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When action is taken the benefits will be seen as a reduction in Specific Energy Consumption. Once you have a monitoring system in place keep it going as it is all too easy to let things drift and the benefit gained will be lost.

Step 3 - Determine a method of measuring saleable production Step 4 - Record both sets of data regularly [probably monthly] Step 5 - Calculate SEC [total energy -r total saleable production] Step 6 - Develop an effective reporting system [graphical is recommended]

3 - Specific Projects This is where the specialized knowledge of the Energy Management Working Party will come to the fore. They clearly know your plant, equipment and infrastructure and should be able to target areas where energy management programs can be implemented. Nevertheless anything that uses energy can become the focus of an energy management project and the following may be relevant:

Once the SEC reporting system is in place you have achieved the following: • A base or reference performance • A means of measuring the effects of your improvement programs • Aframework against which targets can be set and monitored

• • • • • • • •

2 - Minimize Base Load Energy Having established a measure of a site's Specific Energy Consumption the Working Party can now turns its attention to identifying opportunities to reduce energy consumption. The first Project is to investigate Base Load which is the energy being used by each energy stream when there is no production for example over a weekend or between shifts. The costs can be minimal but the rewards can be significant.

For each of the subjects above you will probably need a measure and monitor program with the objective of producing a dedicated SEC for the equipment or process. This will almost inevitably lead to the need for sub-metering which are essential tools in the fight to eliminate waste. As each subject will require a slightly different approach the exact way forward is best addressed by the Energy Management Working Party under the leadership of the Energy Manager.

Each plant will have to assess the best way of monitoring its consumption but if energy usage can be reported on say a half hourly basis such data can make some interesting reading particularly during non productive periods. As an example the following is a chart showing Vz hour electrical consumption for a site operating a 24 hour shift system 5 days per week. The weekend can clearly be seen.

There are, however, other issues that are worth considering. Downtime • by reducing downtime you reduce the amount of energy wasted and improve your SEC Reject Rate • When you produce a reject its energy is wasted. Reducing reject rate will improve your SEC Capital Investment • Every capital investment should consider energy implications and where practical opt for low energy opportunities.

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In any typical plot of half hour data don't initially look at the peaks but look at the energy being consumed when not in production and ask some questions • • •

Air compressors and compressed air Furnaces Dedicated process equipment Lighting Large motors Hydraulic systems Air handling equipment HVAC systems

4 - Review Manufacturing Processes So far the Working Party has been looking at opportunities to improve the energy performance of the infrastructure. It is now time to look at your manufacturing methods and processes to identify or develop lower energy solutions. As a starting point you need to produce an energy profile of your processes so that you know exactly the energy consumed for a particular process. Thereafter it's a case of investigation, process engineering and development.

What is running? Why is it running? What can be done minimize energy consumption?

Some issues worth considering: • Make sure that anything that is left running is necessary. • Design well thought through Start-up and Shutdown procedures and make somebody is responsible for supervising and monitoring.

This is likely to be a significant challenge but the rewards could be significant. It will also be the time to consider adding additional specialists to join the Energy Management Working Party:

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The blue dashed line is the target for the year of 1.37kWh/#. The red line reports monthly SEC and the green reports the Year to Date Figure. It can be seen that the performance is better than target.

KB Alloys, LLC - Case Study of Wenatchee Facility 1 - Management Commitment The Company published a Company Policy on Energy Management and in September of 2008 set up Energy Management Working Parties at its manufacturing facilities.

4 - Base Load The chart below shows Vi hour electrical consumption for a week in 2008 and a corresponding week in 2010. The area of interest is the weekend consumption. In 2008 the typical energy consumption over a weekend was 300kWh per Vi hour. In 2010 that figure had been reduced to about lOOkWh per Vi hour. This came about as a result of an investigation into what was running and what could be done to minimize the energy consumed. There is still work to do to drive that figure lower and the engineering work is in hand.

The Working Party at Wenatchee followed the structured approach as outlined above and within 2 months and using historical data produced a Base Year SEC. 2 - The Energy Management Working Party - The team The Wenatchee Plant Manager was appointed Chairman of the Working Party and the Production Manager and the Maintenance Manager were added to the team. I was appointed as Energy Manager but in the fullness of time that role will be transferred to the Maintenance Manager. Subsequently the Technical Manager was added to the team as the work programs required specialist knowledge. 3 - SEC Measurement and Reporting 3.1 - SEC Base Year [2008]

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- Specific Projects 5.1 - Reverb Furnace Relining Program In 2009 we completed a relining program in which we had worked with a specialist refractory vendor to improve the energy performance of the furnace. The program was successful and we achieved a 10% improvement. We support this program with regular measurements of heat loss.

Jan-08 Feb-08 Mar-08 ApM)8 May-08 Jun-08 Jul-08 Aug-08 Sep-08 Oct-08 Nov-08 Dec-08

The red line reports the Monthly SEC and the Green reports the Year to Date Figure. The reason for the sharp upturn in the red line at the end of the year is the effect of the recession.

Operational Monitoring SEC data is reviewed weekly and action taken if trend indicates a deterioration of performance. We also monitor gas consumption at weekends when the furnace is full of molten metal but in an idle state. In comparison with other similar furnaces in the Company its performance is good and has become a benchmark.

At the end of the year the average SEC was 1.52kWh/# of net product. The CEO reviewed the data published and set a target for the Working Party to achieve a 10% improvement in SEC for 2010 [target for 2010 is 1.37kWh/#]. 3.2-SEC 2010 2010-SEC[kWhper#]

Future Projects Initial investigations indicate that a burner up-grade could deliver a 30% improvement which is clearly an opportunity for the future.

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May-10

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Volume It is almost inevitable that volume will be a factor in setting and monitoring targets. This relationship will need to be understood which was achieved at Wenatchee by plotting monthly volume against monthly SEC as follows:

5.2 - Compressed Air Leaks With the plant shut down we measured a leakage rate of 16.4% of compressor output Following repair of leaks reduced leakage rate to 4.6%. Regular monitoring is carried out.

VOLUMEvSEC ALGORITHM

Weekend demand There is minimal demand over weekends and we installed a small compressor to meet demand and shut down the main compressor Compressed Air Demand We have a program to minimize demand and look for alternative lower energy technologies Size of compressor The compressor is 60% bigger than we need. Perhaps an opportunity for the future

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As can be seen there is a natural reduction in SEC as volume increases. When that curve is understood it can be used in target setting and monitoring with actual target set being adjusted relative to volume.

5.3-Rod Mill The DC drive system was due for an up-grade. As a part of the project we included in the specification the objective to reduce to reduce the mills' SEC. Following successful completion of the project we have reduced SEC by 14.5%

Involving Others The impression may have been given that that the Energy Management Working Party is an exclusive club working behind closed doors. This could not be further from the truth.

5.4 - Lighting Mindful of the advances in low energy lighting we took advantage of an opportunity to have a survey of both our lighting and the associated control. We now have a new project in the pipe line that will reduce our total electrical energy by 7% and plan to complete this project in the next 6 to 9 months

The Working Party needs to be open and transparent and to engage with and encourage everybody working in the Company to get involved. To achieve this it needs to communicate and the following are subjects that could well be included: • • •

Factors outside the Control of the Working Party The Working Party is charged with delivering the Energy Management Program and meeting jointly agreed targets. However not all factors that can effect SEC are under the control of the Working Party and such issues may need to be taken into account when setting and monitoring targets.

Company Energy Policy Energy Management Program Progress report and future plans

In addition it should consider: • Seek ideas for energy management opportunities from the workforce at large • Establish dedicated working groups to address specific issues • Regular Publication of energy management data relevant to the audience

Weather This is not the difference between summer and winter but the differences between summer and summer and winter and winter when the energy used to heat and or cool could be significantly different year to year. To date this has not been an issue at Wenatchee but if it is a problem it may be necessary to consider this when targets are set.

In Finality The plan outlined above works for KB Alloys and it will work for you. However as reported above it is unlikely to be successful unless the driving force comes from the most senior management.

Product Mix If you make a range of products with widely differing energy demands the mix of products will affect you ability to achieved target. This relationship will need to be understood. To date this has not been an issue at Wenatchee but if it is a problem it may be necessary to consider this when targets are set.

Projections indicate that energy prices will increase significantly - if you are not running an energy management program to-day there is no better time to start. The concept outlined above is an effective method of implementing energy. Maybe now is the time to embrace it and reap the benefits.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

MOLTEN METAL SAFETY APPROACH THROUGH A NETWORK C. Pluchon, JP. Riquet, F. Fehrenbach, GM. Raynaud1 B. Hannart, L. Jouet-Pastrι, M. Bertherat, J. Hennings, J. Mathieu, R. Wood2 V. Brockhagen3 1 Alcan EP Research Centre, 725 rue Aristide Briand, 38340 Voreppe, France 2 Alcan EP Casthouse managers 3 Alcan EP Safety and EHS Audit manager, Singen, Germany Keywords: Casthouse, aluminum, safety, explosion, approach

into the furnace. Several witnesses indicated an enormous explosion with large flames of burning gas coming out of the furnace. Immediately afterwards, the atmosphere was filled with dust, leading to poor visibility. The explosion was heard also outside of the building. A lot of dust, insulation plates and other material fell down from the roof of the casthouse. No metal or any other material was ejected out of the furnace. The SECURIT glass disk of the loader burst and the employee was hit by the heat wave of the explosion. He climbed out of the vehicle and called for help. The driver of the fork truck had been on the cold-side and was not injured. An operator had been working about 8 meters away from the furnace door. He just fell down and was not injured. Both started immediately to help the injured employee.

Abstract Molten Metal explosion or splash is a major risk encountered in the aluminium casthouses. Alcan EP has had to face an accident in 2006, an explosion in one of its cathouses. Unfortunately an employee suffered 2nd degree burns on the face. After expertise of the accident and implementation of corrective action in the plant, it appeared that we had a potential of improvement in management of Molten Metal risk by sharing experience coming from the merged companies, Alcan, Pechiney and Algroup. The EP management decided to create a network constituted of experienced managers of most of the casthouses. It was key to look deeply into the details and to think about the practical aspects of every decision. A new approach of the management of Molten Metal risk was born in the company. The system was very powerful to align the whole organization on the objective to minimize risks. Introduction Environment, Health and Safety First (EHS First) is the EP management system for Environment, Health and Safety. When a big explosion occurred in one of a major casthouses, it was obvious that we had to make further progress on molten metal safety to keep our employees safe in our casthouses. An explosion in a casthouse During winter 2006 a big explosion occurred in a major casthouse of the company. Root causes were identified by a comprehensive incident investigation.

Figure 1: After the explosion in the casthouse.

Description

Conditions

A strong explosion occurred while loading plates and cuts from plate edges into a 50t furnace. One employee, who was operating the loading vehicle suffered 2nd degree burns at his face and injuries to his eyes. Fortunately he regained his eyesight. The ceiling of the furnace was completely destroyed and the roof and cladding of the casthouse building was seriously damaged. The main door of the 50 t furnace had already been opened for at least 10 minutes. The furnace had been emptied half an hour before by the previous cast. A forklift truck operator had filled the bucket of the loading machine with aluminium scrap plates and cuts (alloy 7212) of the fourth container, which belonged to the designated furnace load. The driver of the loading machine was sitting inside the cabin of the loader and pushed the content of the filled bucket

The furnace had been emptied half an hour before the explosion. The tilting system was in maximum position. The inside of the furnace was controlled and there was only a pool of molten metal of usual size. The load was made of tubs of plates, of thickness about 8 cm. These tubs had been brought in the shop 48 h before. There was no instruction giving a minimum storage time inside the shop before loading. The outside temperature was -10°C.The first load was made of one plate of big dimensions. Then two tubs were loaded over it. The explosion occurred just after the fourth loading of plates. No water was seen when the tub went through the loading funnel.The inspection of the furnace showed that the first loads had not melt. That was normal, as only the holding burner was on. The molten metal depth was estimated at 7 cm when the explosion took place. So, taking into account the

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surfaces, the quantity of metal can be estimated at about 1500 kg when the explosion occurred. No alumina trace was found in the furnace. The weather was very cold with a lot of snow outside. The temperature scarcely got above - 5°C the days before. The inspection of the fifth tub, that was not loaded, showed ice and snow on the plates. The inspection of the inside of the box showed a few snow in an angle. All the inspected tubs were pierced of holes in the bottom. Like the other operator who loaded the loading vehicle, the injured operator (on the vehicle) attested that the load was not particularly wet, and no water fell down from the tub when he loaded it. Also he attested there was only one explosion.

This casthouse was initially in a smelter plant. The smelting part had closed and the casthouse was operating by recycling scraps. It appears that it had not learnt enough from the other casting plants of the company to completely avoid the risk of explosion. At this time it was decided by the Alcan EP management to create a Molten Metal Safety network which was constituted of the casthouse managers, experts and EHS managers. Its mission was to write a Directive and to ensure its application at all cathouses. By sharing their casthouse experience and using existing documents the members of the network developed a molten metal safety directive and a guideline for audits.

Explanation

EHS Directive

The further investigation on the site confirmed the hypothesis of a physical-type reaction between water (ice) and molten aluminum in a confined space and the with the generated vapour put under pressure. The explosion intensity may be explained by the fact that there was water in the form of ice. The lower the water temperature is, the more violent the explosion is (Fig 2). The mechanical impacts due to the loading operation could be a factor that trigged off or aggravated the explosion intensity.

This directive establishes EP's program requirements for molten metal safety [1]. It applies to all Alcan Engineered Products sites. It takes into account a series of existing reference documents, [2], [3], [4], [5], [6], [7]. Definitions Primary Protection PPE is personal protective equipment designed to be worn for work activities during which significant exposure to molten substance splash, radiant heat, flame, and hot surfaces is likely to occur. Secondary Protection PPE is designed for continuous wear. This applies to all work activities in designated locations in which infrequent exposure to molten substance splash, radiant heat and flames sources is possible.

Intensity of Explosion

Risk Management

10

20

30

40

50

60

70

80

Water Temperature (°C)

Figure 2 : Intensity of explosion related to the water temperature What we learnt from this accident. Referring to the Aluminum Association statistics, this accident belongs to the too numerous family of serious injury cases in rolling plants during the melting operation. (Table 1).

The site shall identify potential molten metal hazards utilizing the hazard identification, risk assessment and risk control process (HIRARC).For each job task the current and potential molten metal explosion hazards must be identified and documented. Identified molten metal hazards are assessed initially and reviewed on an ongoing basis, at a minimum annually (severity and probability). Risks are prioritized. Molten metal requirements identified during the assessment are documented, addressed by the V EHS Corrective and Preventive Action Plan', or integrated in the Standard Operating Procedures (SOP) and communicated to the appropriate parties. Material related hazards have to be considered in the risk assessment. They are related to the nature of scrap. They cannot be fully described here. Here is only a short list: Trapped water, surface moisture oxides on product surface, foreign and hollow objects, reagents, product shape. Procedures of the EHS Directive To illustrate the accident described before, the procedures related to the loading of scrap, including the upstream process, are presented here.

Table 1 : Statistics of accidents in the casthouses, from the AA. AA 1980-2006

total

melting pasting transfer

|

|Smelters

1177

273

k43

303

1

Material Purchase

Rolling / extrusion 867

274

^82

195

1

Total

2482

931

b09

564

1

[Fatality

65

40

19

6

Serious injury KFat+Ser)/Total

305 15%

97 15%

144 144

^0% b o%

64 64 12%

All materials (e.g. ingots, alloying materials, aluminum scraps) must be purchased with a specification which includes requirements regarding safety, packaging, storage conditions and transportation. The specifications must be integrated in each order. External and internal suppliers must be audited.

|

658

Material Reception and Inspection Inspection: All shipments of materials to be charged into furnaces and/or in contact with molten metal must be inspected prior to acceptance of the material. Step 1 inspection "Before unloading": Materials, shipping containers and delivery vehicles (trucks, railcars) must be inspected for contamination. Step 2 inspection "After unloading": After the materials have been unloaded they must be inspected. Quarantine: The quarantine process of a hazardous material must be described in a procedure. The load has to be tagged (do not charge tag ...), isolated and thoroughly investigated if any of the following conditions are detected: Residual fertilizers, dry fire extinguishing powders, reactive chemicals (nitrates, sulfates, oxidizing material, etc.) if found on the metal and/or delivery vehicle; Water or other volatile substances (liquid or solid). Heavy grease or oil; Garbage and trash that may contain cans, bottles or other objects that can trap moisture; Salt fluxes that contain nitrates, sulfates, and oxidizing chemicals; Corroded or oxidized material; Crimped or closed end pieces of tubing, extrusions or containers that may contain water; Scrap contaminated with hazardous or toxic materials (PCB's, selenium, lead, cadmium, and radioactive materials); Miscellaneous contaminants such as batteries, butane lighters, live ammunition, medical waste and aerosol cans. If white powder is visible on the surface of magnesium, the material must be either rejected or brushed and dried. Other requirements: For checking the load a check-list must be used. Receiving and inspection of materials must be assigned to well trained, experienced persons. These persons must be formally authorized to reject any hazardous material. Hardener batches must be used in FIFO sequence (First In, First Out). It is recommended for each plant to visualize the standards of acceptance. See more details in the AA document: "Guidelines for aluminum scrap receiving and inspection based on safety and health considerations" December 2002. Ref. [4] Material Storage Storage has a significant influence on major hazards, such as condensation, rain, snow and ice. Ice often is not even visible. The storage of a material determines its dryness level. The dryness level determines the permitted furnace configuration. Condensation occurs when the air meets a cold point either because of contact with a cold surface or when the temperature of the air decreases. Condensation occurs when the surface temperature is below the dew point temperature of the adjacent air. The dew point temperature is the temperature at which the air must be cooled to reach saturation. By measuring the temperature of the air, the relative humidity (also called hygromιtrie degree) and the temperature of the metal, the risk of condensation can be determined using the following figure. From a starting point defined by temperature and relative humidity, there will be condensation if the temperature drops below the saturation curve (Fig 3). One of the factors to prevent condensation is good ventilation in storage facilities. This prevents stagnant air remaining in contact with cold metal. We can consider that there is no risk of condensation if the product is stored indoors in a place where the

temperature and humidity are controlled and the temperature remains above the corresponding dew point.

metal

Material must be stored where direct contact with snow or rain is prevented (e.g. under a roof). Storage building structures, such as roofs must be kept rainproof. Roof, gutters and down pipes must be inspected, at least annually. Deformed tubs, which may accumulate water, must be eliminated. Tubs must be inspected, at least annually. The plastic foil over magnesium must always be removed to avoid the accumulation of condensed moisture. If material has got in contact with rain, snow or ice, strict charging procedure have to be followed, described in chapter "charging". (E.g. charging in cold furnace, prohibition of dilution etc.). Hardener storage conditions: Suppliers must be asked to pack hardeners in closed and covered buildings. Suppliers must also be asked load and transport hardeners under cover. Hardeners must be stored under a cover.

Figure 3: Risk of condensation HR % is the relative humidity. Case 1: Outdoor storage under a roof Point A: Air at 15°C with 75% HR. Then over night the air temperature suddenly drops to 2°C —> Point B Condensation! Case 2: Storage inside the cast house Point C: Air at 20°C with 40% HR. Cold metal with a temperature of-10°C is carried into the cast house —> Point D point D Frozen condensed water! Case 3: Storage inside the cast house Point E: Air at 20°C with 75% HR. Metal with a temperature of 5°C is moved into the cast house —> Point F Condensation! Case 4: Storage inside the cast house Point G: Air at 15°C with 25% HR. Metal with a temperature of5°C is moved into the cast house —> point H No condensation Best practices are: To store all materials inside under controlled conditions preventing condensation, to store alloying materials in a heated room, to unload hardeners under a cover. Material over-drying The site has defined for each oven temperature setting and corresponding drying time for each material ensuring trapped water gets vaporized. The definition of oven temperature and drying time has to be based on experiments with measurement of the core temperature. Here are some examples. These examples are by no means an alternative to real measurements and

experiments, but give an indication of the necessary temperatures and times to vaporize trapped moisture:

the surface level of the molten metal in the furnace. The preparation of a "bed" will be explained later on.

Material Sows T-Ingots Divided Scrap

Core Temperature 200°C 150°C 100°C

Time 4 hours 4 hours 1 hour

High volume liquid metal: The volume of liquid metal is high. Even with a "bed" it is not possible to prevent that materials charged into the furnace "dive" into the liquid. Any liquid metal level beyond 30 cm is considered as "high volume liquid metal".

Material Aluminium Magnesium

Chamber Temperature 400-450°C 150-200°C

Time 6 hours

Rotary furnaces We distinguish three furnace configurations of rotary furnaces: hot furnace emptied, hot furnace with salt or black dross, hot furnace with salt or black dross + molten metal.

For magnesium the temperature must be lower to prevent excessive oxidation and to avoid spontaneous combustion at a temperature of about 560°C. Care must be taken with regard to hot spots in ovens. It is preferable to use an oven with an even temperature. Electrical ovens are recommended. Before ovendrying the plastic strapping around ingot stacks must be removed. It does not withstand the oven temperature. For example ingot may be packed in bins. After oven-drying it can only be considered oven-dried as long as the material is continuously kept completely dry and kept under conditions that prevent condensation. It is generally not recommended and material cannot be considered oven-dried, if kept on a ledge of an operating furnace for drying.

Channel induction furnaces We distinguish two furnace configurations of channel induction furnaces: Heel operation with max. 30 cm heel and charged on a bed, high volume liquid metal. Crucible induction furnaces We distinguish two furnace configurations of crucible induction furnaces: Hot and completely emptied furnace, high volume liquid metal. Material dryness levels If water is charged into a furnace, the probability of a molten metal explosion depends much on the kind of materials, the type of furnace and the furnace configuration. In some cases even a big amount of water cannot cause any explosion, in other cases a small quantity can cause a huge explosion. The following concept of dryness levels is related to the probability that water is part of the charge. These dryness levels focus on water as the most frequent cause of molten metal explosions. Other kind of contaminations must be considered equally. There are three dryness levels. These levels represent three different risk levels, are applicable to all materials and are an essential element of the charging rules.

Furnace charging To easily understand the charging rules in this chapter, two concepts have to be introduced before: furnace configurations and dryness levels. Furnace configurations There are four different furnace configurations distinguishable. These configurations represent four different risk levels. They are applicable to all cast houses and structure the set of charging rules. It's obvious that any furnace can change "state" or configuration during the different stages of melting. When we look at charging we look at the furnace configuration at the moment when the material is charged into the furnace. The following configurations are defined for reverbatory furnaces, rotary furnaces, and induction furnaces.

Level 1- Sure dry: This level represents the optimum in terms of safety. A material can only be considered sure dry, if: The material is oven-dried or equivalently heat-treated or homogenized, or if the material is inside scrap that has never been stored or moved outside of the cast house building. In both cases and for all times before Reverbatory furnaces charging the following requirements must be met to consider a Cold furnace - empty or solidified: The furnace has been emptied material sure dry: The material was always been stored in a and the low temperature ensures that any metal that remaining closed building and controlled conditions preventing inside can only exist in a solidified state condensation; the material was always transported under a cover (e.g. roof). The pre-charging inspection has indicated no visible Hot emptied furnace - dry hearth operation: The furnace has been wetness. Materials all can only be classified sure dry if all emptied. But the furnace may still contain remaining metal which mentioned processes (drying, storage, transportation, preis liquid because the furnace is hot. In this furnace configuration it charging inspection, and correct classification) regularly checked is possible to solidify any remaining liquid metal by adding a bed by planned inspections and Leadership Safety Tours. of dry divided scrap (see below, description of a bed). Heel max. 30 cm high - heel operation: There is liquid metal in the furnace; it has not been completely emptied. The amount of liquid metal may vary between a few centimeters up to 30 cm. Depending on the surface area of the furnace; this can be equivalent to 10 or even up to 30 tons for the largest furnaces. In this furnace configuration it is no longer possible to dry the liquid metal by divided scraps. But it is still possible to prepare a "bed" to prevent that materials charged into the furnace can drop below

660

Level 2 - Probably dry This level can be achieved if there is no visible wetness, but it can not absolutely be excluded that water is hiding in cavities, cracks or in areas which are difficult to be inspected. A material can only be considered probably dry, if: Case 1 : The material is stored under a roof and conditions where condensation is unlikely and the receiving inspection has not indicated any visible wetness. The inspections during storage period have never indicated any visible wetness of the material or

the ground under the stored material. The pre-charging inspection has indicated no visible wetness. Case 2: The material has been stored outside and the following additional requirements are met: There was no rain or snowfall during the whole storage time. The temperature was always above 0°C during the last week of storage.

6:closed building, reduced condensation risk, plastic foil removed, defined site-specific minimum storage duration and temperature Preparation of a bed The bed is a first layer, made up of particular components of the load, with at least one of the following or often all objectives simultaneously: to protect the earth of the furnace from the shocks caused by the massive components of the load; to prevent that the loading of solid elements of the load does not cause splashes; to allow some other components to be charged on a solid material heap, to avoid (or limit) the risk of immediate contact of these components with molten metal; in case of a dry earth operation, to solidify all the remaining liquid metal.

Level 3 - Wetness must be taken under consideration A material has to be considered wet, if it does not meet all requirements of sure dry and probably dry. Another case is forbidden. In this case a certain combination of a material and furnace configuration is forbidden independent of the dryness level of the material. Charging rules for the combination: material - dryness levels furnace configurations

In case of a dry earth operation a bed must ensure that all remaining liquid material is solidified. To achieve uniform coverage of the earth of the furnace the bed must consist of materials easy to spread.

The following section is a key element of this directive. It defines the charging rules for the different types of furnaces as a function of the material, of the furnace type and configuration and of the dryness level. The charging charts (Fig 4) concern safety and environmental matters. However, in some cases charging might not be possible or meaningful for other reasons too, e.g. due to productivity or design. Hct emptied t u n a c e - d r y

Heel max 30cm - heel operation

Cold empty» scarified

1

Used t o dry the Charged on t r e

i

S

i

I

V

Y

V

Compacted Chips

Y

Y

Y

Y Y

Hardeners ether t i a n Mg Sheet s c r a p - side tn m

Y

Y

Y

y

Clean compacted ; o a p

Y

Y

Ertrustonbutts

Y

Y

Ejdr. sections open et least at one side

ã

Y

Y

â " * * * YATSUT cracks

Y

Y

Y

V

Y

Discs or stamping skeletons Loose Chips

Y

-■

V 1

■.

Y

Y

Y

Y

Shape casting scrap

Y

Y

Erfr. sections hoBovActeed both sides

Y

Y

■'

Plate scrap

Y

V

Y

\

«m

ã

Loose Cote

■.

■

■:

y

Y

Y

i

i

Y.

Y N

Tigrt C d s w t i cardboard mandrel

Lacquered material Material v#h at, grease, plastic or paper a ass 1 can stock scrap Loose Sawdust

K

Y

Y

Y

H

Y

Y

Y Y

Y

Y Y

·,'

Y

[:

¥

']

•

·

Y,..

,

Y

'

Y

Y

Y'

Y

Y Y

* 1 Y » Y H | .

Y

Y

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Y. \ H 1 Y Y «

Y'

Y

M j

Y

Y | Y ]

Y Y

Y

H

n

Y Y

H

v

n 8

«

t)

s

! Y j

Y

VM Y**

» V « -,' « H

Y

Y

Y

H

y Yj

Y i Y

Y B H Y. H Y Y Y ] ■', i Y Y Y ' | v, N N Y» Y a H H Y Y Y ' .-,.. N Y H Y H V Y V '. : Il H Y Y H H H Y Y Y H H H ti II « M H H H N M H Γ»ΤΗ N tl tl

P i s Cornfakes/pope cms

Y

Y

Ingots 1 2 - 22kg

Magnesi u n seve

Y

«! v Y HA "Il I m Y » i . H

1

Y

H

Y

£ Y

H

Y

bed

Y

Y Y

|

Charged on a

(bed)

Y

Y

*

A s a
&

i

Y

EdemalScMe

ß f.

Y

Y

Y.

Y

Y Y

Tight Cces j g h o i t car aboard mandrel

Irtemal Scwe

Y

ãΊβ - ?"

Slabs rfhcU cracks Slabs y * h cracks

B i e t s v s t l cracks

Y

Y,

Y Y

Compacted Sawdust

T-lngDts(exriuc«ng ends)

s

Y

Wire rod

T-Ingots ends Magnesiun Ingots

«

5 S

Y Y Y

V»

N

N

N

N

N

H

H N

H N

The ideal components of the bed are chips and rolling edge trims. They are practically essential "to completely dry" a earth with molten metal. The compacted packages are also good constituents of a bed, but they do not bring the same guarantee of uniform cover of the residual molten metal. Other types of divided waste can also be used. Some materials, due to their shape, are completely unsuitable (e.g. large rolling coils). The loading practice should start with the more divided materials to increasingly massive materials. For example sows or T-ingots should not be put directly on a bed of chips.

Y Y Y

Y.

>

"ft-'

Y,.a

Y. V, Y,

it y

» » » M w H

«

tt

Ys-« Y,.*. Y»OM

N Y, H

N M

y

« » M

M N 1 Y H H H Y H N ■m H 1 Yj

N N H H H H H VfM N H H H H N H

n

Y Y Y Y Y Y

V

Y

Y

v

Y

Y

Y

Y

Y

Y Y;

Y Y Y Y y Y

*N

Y

s

Y3 H Y

Y

Y'

Y

Y

Y

Y

','

M H

« ·,', « H

11

X H M H

m H «i

1 H

N

'.'*.<->', Y,

Y,

Y,

*M

Y

Y

Y,

Y3 Y,

v N

Y

Y,

{

;

S

Y,

H

«quid metal

Y

11 II Y*.«

H H 11

Y

¥

Y,

H

tl H

H H

'M H H 1 I

Y;

Y: Y3

H N tl

: «1 » ft« «S II Y.·;

Y.·.-

Yi-Î Y.·;

H

;

Y,.r

Yfl 1 H Y*j M H tl

*

H li HJL H

| H

M

II

Y|

y. H H H tl N H H tl N M ti ti ti VW« H tl H H H H

The bed is not a mandatory practice. To charge more massive components directly is possible as long as they comply with the rules described in the charging tables. The bed in dry earth operation configuration offers real advantages for the safety for the materials charged afterwards. In the case of a heel a bed is much less effective. This is the reason for the more restrictive rules for the remainder of the load in this case.

8 tt;

•

«I

Ø. ·■

n

y·. ■

■

* * « H

Other charging rules Sequence: The correct charging sequence starts with light materials and ends with heavy and problematic scrap: 1. Light scrap (e.g. dry or scalping chips ). 2. Medium weight scrap.3.Heavy scrap. 4. Sows, coils, ingots. It is recommended to record the sequence in which a furnace has been charged.

M

II

n

« M: n H

H H

H

':

Pre-charging inspection: The charge must be inspected before being charged: research of moisture, oxidation, foreign substance, dust. The best practice for inspection is to make it twice: a first time when preparing the load and a second time when charging the material into the furnace.

Figure 4: Charging chart of a reverbatory furnace. Explanation of abbreviations: Y = yes = Charging recommended if all requirements are met N = No = Do not charge! Yi = Index i referring to additional material specific charging requirements

Vehicles and people: Cabs of charging vehicles must be completely closed when operating in an area potentially exposed to a metal explosion or to splashes. During charging operations, only people directly involved in charging are permitted to stay close to the furnace.

Footnotes - specific requirements: 1: only oven-dried 2: smooth introduction important 3: only in small blocks 4: as the only load, not as a bed 5: brushing needed, if more than 10 % surface oxidized

Dilution and addition: Dilution must be considered as charging in a furnace full of liquid metal. Addition material may be added in a trough under the

661

following restrictions: Material must be at a minimum "probably dry" (see definition); documented trials prove that form, quantity, and liquid flow rate do not result in splashes. Magnesium: For metallurgical reasons (melting loss, risk of spinels, etc.), the best practice is charging magnesium in a furnace that is full or has a heel.

pareto of cat2 findings

Big-bags and other products: Compacted products with an unknown or dubious origin, must be crushed and sorted before charging, as they may contain foreign bodies. Charging of big bags is prohibited because their content cannot be inspected properly. Pallet charging should be avoided. This can be done for example by previous removal of the pallet or by keeping the pallet on rotating forks. If the charging of a pallet or any other packaging item (cardboard, plastic film...) occurs, it does not change the charging restrictions for all materials as long as the package is dry and as long as the packaged material has been stored inside and properly inspected.

Figure 5: Pareto of Category 2 findings Need of complementary work for progress However, in spite of this highly involving program, several explosions - smaller than the explosion described at the beginning - occurred in 2009, leading to complementary work of the network: For example a new explosion occurred in a casthouse due to an unexpected descent of the lift at the end of a cast. Cooling water covered the sump not yet solidified, creating projection of molten metal with no injury. An addition to the procedures of the Directive was made on the design and the maintenance of the hydraulic circuits of the lifts.On another side, some recurrent small explosions occurred at the starting phase of casts in the same casthouse. It appeared that some operators did not understand the procedures and the reason why they had to be applied. A special training session was established, dedicated to this specific need. It can be adapted to the specificity of other plants. Finally, for a better understanding of the Directive, it had been translated into the three main languages of EP.

Induction furnace: A procedure must be developed how to prevent arch formation while loading the furnace. All the process is described in the Directive as it is done and presented for the material purchase, for its inspection, for its storage and for its loading. Application of the Directive Carry-out of the audits Since the first year, all the Alcan EP casthouses were audited once a year by two auditors on the base of the guideline. Few major "Category 1 non-conformances" were found. 6 "Category 2" per plant were found in average, leading to have an action plan to control the risk (Table 2) .The following years no "Category 1 non-conformance", but the same number of "Category 2" was found. While the standards became increasingly better known by sites and auditors, the requirements were handled more and more strictly.

Conclusion An accident in a plant of Alcan Engineered Products had been the trigger of a comprehensive cooperative involvement of casting experts of the company to prevent all the casting personnel from molten metal explosions. A Directive and its application was the result of this involvement. During this work a new shared culture of EHS regarding the molten metal explosion risk was born. Same vocabulary, same standards, same request were also shared. EHS results in the casthouses were in progress. To keep the system efficient, it is currently up-dated by the network.

Table 2: Findings within the plants

tot 2009 [average/plant 2009 average/plant 2008 [average/plant 2007

cat1

cat2

cat3

Best Pract

Tobe studied

0

84

27

46

37

0 0 0,3

8 7 6

3 3 5

5 5 5

4 3 3

I |

References [1] "EHS Directive for Molten Metal Safety". Internal paper 2010 [2] "AA guidelines for Handling Molten Aluminum" (Third Edition July 2002) [3] "Heavy Lifting and Handling Equipment- Rules for the design of hoisting appliances" FEM1001 (8 booklets) [4] "Guidelines for Aluminum, Scrap Receiving and Inspection" [5] "Based on Safety and Health Considerations". Publication GSR 2002 [6] "Lightning risks and prevention". Internal paper DRTF. 2001 [7] "Fire Protection Engineering Standards: Clean up of metal dust - 28". March 2009 EHS for molten metal explosion prevention, Alcan EHS 4.8.6

|

An analysis of the type of findings was made (Fig 5). Many findings were related to the PPE wearing, and to the procedures and their compliance. Many investments were also made to respect the standards, especially of roofs to protect the material and of furnaces to dry it.

662

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

IMPROVED MONOLITHIC MATERIALS FOR LINING ALUMINUM HOLDING & MELTING FURNACES Andy Wynn1, John Coppack1, Tom Steele1, Ken Moody2 rhermal Ceramics UK Ltd.; Tebay Rd.; Bromborough, Merseyside, CH62 3PH, UK 2 Thermal Ceramics Inc.; 2102 Old Savannah Rd.; Augusta, Georgia, 30903, USA

lr

Keywords: Aluminum, Monolithics, Castable, Refractory Background

Abstract To remain competitive, aluminum producers continue to increase productivity through their Melt-Hold furnaces. Increasing heat input to the furnace using more powerful burners is common practice. But faster melting leads to increased metal losses from surface oxidation and to segregation from large heat gradients. These effects are countered by increased use of fluxes and increased stirring. Given the increasingly challenging environment within which the refractory lining has to work, traditional lining solutions can no longer be relied upon to provide the service lives that were previously achieved. Therefore, a new generation of furnace lining materials is required to cope with today's aluminum furnace. This work reports on a new monolithic material with improved performance, compared to existing materials, designed for use in the ramp/hearth area of aluminum furnaces. Improved behavior against the critical performance criteria in this furnace region are demonstrated in the laboratory using industry standard test methods.

Over the last 30 years, a group of Monolithic technologies has emerged which have been designed specifically to perform within the unique environment of Al melt-hold furnaces. These Alresistant grades often contain 'non-wetting' additives, particularly in the metal contact areas, to minimize interaction between the refractory and the melt to suppress damage to the lining from 'corundum growth' [2]. As aluminum producers strive to increase productivity, the environment within the Melt-Hold furnace is becoming more arduous. Chamber temperatures are increasing and more aggressive fluxes are being used, necessitating more frequent and severe cleaning operations of the refractory wall. A key requirement for maintaining high levels of productivity is the need to minimize thefrequencyand duration of furnace downtime. The more aggressive conditions within which the refractory lining has to work today means that the Al-resistant lining materials developed in the past to cope with these applications are now being used beyond their original intended design boundaries and their service performance is under threat, leading to more frequent lining repairs. In order to minimize the frequency of furnace downtime a new breed of Al-resistant products are needed by Al producers, specifically designed to perform within today's more aggressive operating environment. This paper describes the development and behavior of once such newly developed material.

Introduction The refractory lining of a typical furnace used for holding and melting aluminum has to withstand a wide variety of physical and chemical environments. The different areas within the furnace are shown in Figure 1. Each of these areas presents a different set of operating conditions, in terms of peak temperature, temperature fluctuation, metal contact, salt contact, impact from ingot loading, etc. Therefore, in order for a monolithic material to successfully perform in a particular area of the furnace, it needs to be able to cope with the specific environmental conditions in that region of the furnace. This is why furnace linings are complex arrangements, with different materials installed in different locations [1].

Walls

Ram

P

Hearth

Approach An Al producer will make the decision to take a furnace off-line for repair once a critical lining area has degraded to the point of affecting the efficiency &/or safety of the operation. At this stage, not all the lining will have degraded to the point that it is in need of replacement or repair. Therefore, the frequency of furnace lining repairs and furnace downtime is determined by the area of the furnace that is most quickly and frequently degraded during operation. Therefore, in order to increase campaign times and decrease frequency of repair stoppages, we need to improve the service life of this weak link in the lining arrangement. To identify the region mostfrequently& quickly degraded, we worked with a number of Al producers. Their feedback suggested that the most common area that was the cause of repair downtime was the ramp/hearth area. The failure mechanisms within the furnace environment, that limit refractory service life, are of two main types, chemical attack (corundum growth, corrosion from flux addition) and mechanical damage (e.g. ingot loading, cleaning practices, thermal shock) [1]. Since our target is to improve refractory performance in the ramp/hearth region, we need to understand which of these failure modes are most critical to lining performance in this region.

Sub Hearth

Figure 1. Furnace lining zones in a typical Aluminum Melt-Hold furnace.

663

Performance Targets

Secondary Performance Parameters

A study of working practices and furnace operating conditions at a number of Al producers revealed that the ramp/hearth region of an Al melt-hold furnace is subjected to severe mechanical and thermal stress during the loading of large ingot down the ramp. Frequent loading of heavy ingot to feed the furnace, often by fork lift truck, subjects the ramp to severe abrasive forces. As the ingot is usually at room temperature, there is also considerable thermal shock on the ramp/hearth refractory, which is at furnace operating temperature. As the bottom of the ramp and the complete hearth are in contact with molten metal, the refractory is also subject to chemical attack from the alloy, alloying elements and flux additions. A study of ramp/hearth degradation of Al-resistant materials containing 'non-wetting' additives suggested that damage leading to furnace downtime is mostly due to the mechanical action of the erosion and thermal shock from ingot loading. We therefore focused our work on developing a new Alresistant material with improved abrasion & thermal shock resistance. To achieve significant improvements in performance we set out to increase abrasion and thermal shock resistance by 20% compared to existing materials. As metal and alkali resistance are secondary performance parameters in this furnace region, we also had to ensure that any changes we made to the materials did not degrade chemical resistance.

1.

2.

Aluminum Resistance 'Cup' Test; 'Cup' samples are prepared (Figure 2) and filled with 7075 alloy. Samples are ramped up to 1000°C and held for 100 hours. After cooling, the samples are sectioned vertically and visually assessed for the degree of metal penetration and corundum growth. The alloy is then analyzed for any pick up of critical elements from the refractory during the test. Maximum allowable is 0.5% Si and 0.1% Fe. The test method is described in more detail in the literature [3]. Alkali Resistance 'Cup' Test; Sample preparation is the same as for the Al resistance cup test. Instead of Al, the samples are filled with mixtures of K 2 C0 3 and Na 2 C0 3 and fired to either 900°C, 1000°C or 1100°C for 5 hours. After sectioning, samples are analyzed by visual inspection for cracks, bulges, depth of penetration and color change.

Experimental Two existing, industry leading Al resistant monolithic materials used by many Al producers in the ramp/hearth area of melt-hold furnaces were selected as baseline materials for the study. The in service performance of both these materials is well known and so serve as useful benchmarks against which to compare new developments. A detailed analytical investigation of the baseline materials was undertaken in order to identify those aspects of the materials technology that were considered to be constraining performance in terms of abrasion and thermal shock behavior and thus leading to premature mechanical failure. The bond chemistry and aggregate granulometry were then re-engineered through several iterations to find the optimum balance of material types and grain size, shape and distribution that produced the maximum improvement in abrasion and thermal shock performance without negatively affecting other important properties. This paper presents the results of performance and property measurements of the final, optimized development composition compared to the baseline standards. All materials in the study were tested against the four key performance parameters using industry standard test methods;

Figure 2. Mold & test sample for Al Contact & Alkali Tests. Results & Discussion The physical characteristics and chemical composition of the optimized new material compared to the two standard baseline materials are displayed in Tables 1 & 2. Table 1. Physical Properties of Materials Studied. Standard Standard New 1 Material 2 Water (%) 5.5-6.5 5.7 5.3 2840 2630 2640 110°C Bulk 2800 2590 815°C Density 1000°C 2790 2580 (kg/m3) 2570 2510 1300°C 815°C -0.29 -0.43 PLC -0.32 1000°C -0.26 (%) 1300°C -0.35 0.38 0.95 110°C 128 122 147 815°C 99 163 ces 1000°C 129 95 (MPa) 144 1300°C 138 119

Primary Performance Parameters 1.

2.

Abrasion Resistance Test (ASTM C704); Pre-fired samples are blasted with a stream of SiC grit of specified grain size for a set time. Samples are cross-sectioned and the amount of material abraded across the section is measured and reported in cm3. Thermal Shock Resistance Test (ASTM C1100 - Ribbon Test); Pre-fired samples are subjected to alternating heating and cooling cycles on one face using a ribbon burner. The Modulus of Elasticity (E-modulus) of samples is measured non-destructively by ultrasonics before and after testing. The percentage of retained E-modulus is used as a measure of retained strength.

Table 2. Chemical Analysis of Materials Studied. Standard 1 Standard 2 New Material 80.6 65.8 66.6 %A1203 11.2 % Si0 2 26.7 25.6 3.6 3.2 % CaO 1.8 % Ti02 2.0 2.1 2.2 1.2 1.0 1.0 %Fe 2 0 3 %MgO 0.2 0.1 0.2 0.2 0.2 0.2 % Alkalis

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Secondary Performance Parameters

In order to be of practical use to the Al producer, it is important that any new material developed not only meets the performance targets, but also can be installed easily. Both baseline materials are low cement, vibrocast grades. The optimized new material in Table 1 could be cast at 5.3% water, lower than the baseline grades, and gave free flow of 125 mm and tapped flow of 160 mm.

As the baseline materials 1 & 2 are commonly used in service around the world, we fully expected them to pass the Aluminum resistance 'cup' testing. Both these materials, and all of our new development formulations, contain well proven 'non-wetting' additives. Our final, optimized new composition passed all Al contact testing and performed identically to Standards 1 & 2 in the visual assessment of Al 'cup' test samples after testing, both dried at 110°C (e.g. Figure 5) and pre-fired to 1200°C (e.g. Figure 6).

Primary Performance Parameters 4.5

4

Figure 5. AI 'Cup' testing - Standard 2 (left) & New Material (right) - dried samples. Standard 1

Standard 2

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New Material

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Abrasion resistance test results of the materials, one of the primary performance parameters in the ramp/hearth region, are presented in Figure 3. As dried, the new optimized material was observed to deliver 16% better resistance to abrasion compared to Standard 1 and 20% better resistance compared to Standard 2. When pre-fired to operating temperatures, the new material delivered a 30% improvement on abrasion resistance compared to Standard 1 and 20% compared to Standard 2.

1

2

3 Number of Cycles

4

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Figure 3. Abrasion Loss Resistance of Test Materials.

0

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Figure 6. Al 'Cup' testing - Standard 2 (left) & New Material (right) - pre-fired samples. However, subsequent analysis of the alloy after testing in the prefired state revealed subtle differences in interaction between the alloy and refractory. Table 3 reveals that although all materials pass the test (target pick up <0.5% Si, <0.1% Fe), Si pick up is much reduced in the new material compared to the two standards. Since 'Cup' test failures are normally accompanied by increased concentrations of Si & Fe in the alloy after testing, this result may be an indication of a much reduced interaction between the new material and the test alloy compared to the standards and thus may indicate superior 'non-wetting' behavior. Similar behavior has been noted in the literature [3]. Table 3. Alloy analysis after AI 'Cup' Testing. New Material Standard 2 Standard 1 0.314 0.093 0.011 % Si pick up 0.04 0.052 % Fe pick up ao4

5

As with the Al contact testing, for the alkali resistance testing, we expected Standards 1 & 2 to possess good resistance to alkalis as they are used commonly in service and so should have already passed Al producers' approval testing. Our final, optimized new composition passed all Alkali contact testing with K2C03 and Na2C03 and performed identically to Standards 1 & 2 in the visual assessment of Alkali 'cup' test samples after testing, at all test temperatures (e.g. Figures 7-12; samples tested at 900°C (left), 1000°C (middle) & 1100°C (right)).

Figure 4. Thermal Shock Resistance of Test Materials. Thermal shock resistance test results of the materials, the other primary performance parameter in the ramp/hearth region, are presented in Figure 4. After 5 test cycles, Standard 1 lost 42% of its E-modulus and Standard 2 lost 32%, compared to only 20% loss for the new optimized material. These results suggest that the new material is capable of delivering 52% improvement on thermal shock resistance compared to Standard 1 and 38% compared to Standard 2.

665

Conclusions 1.

2. Figure 7. Standard 2 after alkali testing with K2C03

3.

4.

5.

Figure 8. Standard 2 after alkali testing with Na2C03.

6.

7.

8. 9. Figure 10. New Material after alkali testing with K2C03.

By working closely with Aluminum producers, the most frequent cause of melt-hold furnace downtime has been identified as mechanical damage in the ramp/hearth region of the refractory lining. The main factors leading to mechanical damage in this region have been identified as severe abrasion and thermal shockfromthefrequentloading of heavy, cold ingot. Through re-engineering of the bond chemistry and aggregate granulometry of a series of Monolithic formulations, significant improvements have been achieved in abrasion and thermal shock resistance for material in the ramp/hearth region of Al melt-hold furnaces. An optimized formulation has been developed which has been shown to deliver 20-30% improvement in abrasion resistance and 40-50% improvement in thermal shock resistance compared to existing materials. The new material has been shown to pass industry standard Aluminum contact and alkali resistance tests. More detailed investigation has indicated that the new material interacts less with the industry standard test alloy and therefore may possess superior 'non-wetting' characteristics compared to existing materials in the ramp/hearth area. The results of our development program and subsequent laboratory analysis work suggest that the new material should be capable of surviving the unique set of service conditions in the ramp/hearth region of aluminum melt-hold furnaces, better than the existing materials used in the industry and thus deliver longer service life. Extended service life in the ramp/hearth area is expected to reduce thefrequencyof furnace downtime and thus allow Al producers to run longer production campaigns, increasing productivity and minimizing the need for expensive repairs. The new Monolithic material has been designed as a vibrocast grade, with a degree of free flow, for improved ease of installation at low water content. This new material is now on trial in the ramp/hearth area of melt-hold furnaces at several Aluminum producers around the world. References

1. A.M. Wynn, T.J. Coppack, T. Steele, K.J. Moody, and L. Caspersen, "Monolithic Material Selection for the Lining of Aluminum Holding & Melting Furnaces" TMS 2010, Seattle, USA, Feb. 14-18,2010. 2. D. Jones, A.M. Wynn, and T.J. Coppack, "The Development and Application of an Aluminium Resistant Castable" UNITECR '93, Sao Paulo, Brasil, Oct 31-Nov 3, 1993.

Figure 11. New Material after alkali testing with Na2C03.

3. A.M. Wynn, T.J. Coppack, and T. Steele, "Methods of Assessing Monolithic Refractories for Material Selection in Aluminium Melt-Hold Furnaces" 53rd International Refractories Colloquium, Aachen, Germany, Sept. 8-9, 2010.

Figure 12. New Material after alkali testing with K2C03/Na2C(>

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S CAST SHOP for ALUMINUM PRODUCTION Direct Chill Casting SESSION CHAIRS

Dmitry G. Eskin Delft University of Technology Delft, Netherlands Arild Hβkonsen Hycast AS Sunndals0ra, Norway

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

COLD CRACKING DURING DIRECT-CHILL CASTING D.G. Eskin1'2, M. Lalpoor2, L. Katgerman3 Materials innovation institute, Mekelweg 2, 2628 CD Delft, The Netherlands 2 Brunel University, BCAST, Uxbridge, UB8 3PH, U.K. 3 Delft University of Technology, Department of Materials Science and Engineering, Mekelweg 2, 2628 CD Delft, The Netherlands Keywords: aluminum alloys, DC casting, cold cracking,fracture,thermomechanical simulation and stages of casting over which high tensile stresses appear and trigger cold cracking [9, 10]. The next step was the application of fracture mechanics to assess the critical crack/void size required for the occurrence of catastrophic failure. Boender et al. [11] and O. Ludwig et al. [12] applied thefracturemechanics to the results of the thermomechanical simulations with the aim to find the critical locations in the ingots where the brittle fracture is more probable to occur. Although precious from technical point of view, these results were not validated upon casting trials to check how closely they relate to the actual crack sizes. The computer simulation results may be more reliable if the constitutive parameters, mechanical properties and the plane strain fracture toughness of the material are gained from the as-cast material. As the 7xxx series aluminum alloys are mainly used in the fully hardened state, very little research has been performed on their properties in the genuine as-cast condition and the properties of the material in the heat treated states would not represent the as-cast conditions [6]. Even in the homogenized [13] or stress relieved [14] states, the constitutive properties are considerably different than the as-cast ones, which may lead to unreliable calculated stress values and eventually critical crack size estimations. In order to prevent all these problems, the constitutive parameters, mechanical properties and the plane strain fracture toughness of the material in the genuine as-cast condition are required. In this study, we report the results of the thermomechanical tests for an AA7050 alloy, which were used as input data for computer simulations. Having applied the fracture mechanics to the results of the thermomechanical simulations, the critical crack sizes required for catastrophic failure were calculated for the entire billets conditions. The simulation results were eventually validated upon pilot scale casting trials on another typical 7xxx alloy which is highly prone to both hot and cold cracking.

Abstract Cold cracking phenomenon is the least studied, yet very important defect occurring during direct chill casting. The spontaneous nature of this defect makes its systematic study almost impossible, and the computer simulation of the thermomechanical behavior of the ingot during its cooling after the end of solidification requires constitutive parameters of high-strength aluminum alloys in the as-cast condition, which are not readily available. In this paper we describe constitutive behavior of high strength 7xxx series aluminum alloys in the as-cast condition based on experimentally measured tensile properties at different strain rates and temperatures, plane strain fracture toughness at different temperatures, and thermal contraction. In addition, fracture and structure of the specimens and real cold-cracked billets are examined. As a result afracture-mechanics-basedcriterion of cold cracking is suggested based on the critical crack length, and is validated upon pilot-scale billet casting. Introduction 7xxx series aluminum alloys are highly prone to cracking at different stages of solidification and cooling during the direct chill casting process. The susceptibility to cracking is on the one hand due to the precipitation of low melting point brittle intermetallics on grain boundaries and interdendritic spaces [1, 2], and on the other hand to the development of high residual thermal stresses [3]. Above the solidus, the low strength of the low melting point nonequilibrium eutectics and intermetallics provide potential sites for hot tearing [4]. Further propagation of such cracks under the tensile stress fields that develop to some critical levels below the solidus may result in catastrophic failure of the ingot in the solid state; cold cracking [5]. What makes the material even more susceptible to cold cracking is the sever loss in ductility upon cooling after the end of solidification [6]. Although the specific microstructure of the material in the as-cast condition provides potential crack sites for hot and cold cracking, the stress development in the ingots during the casting is brought about by the non-homogenous cooling conditions and the poor thermophysical properties of the material. Relatively higher coefficient of thermal expansion and the lower thermal conductivity of some of 7xxx series alloys compared to other aluminum alloys [7] result in steep temperature gradients and consequently high thermal stresses appear. In order to predict the cold cracking phenomenon in high strength aluminum alloys, a true understanding of the residual thermal stress development during the casting is required. In the 1950s, thermal stresses were calculated analytically by Livanov [8] for ingots with various geometries. He even established a criterion based on numerous experimental trials under different casting conditions. With the development of computer simulations, however the costly casting trials were replaced by numerical calculations. Numerical simulations helped to reveal the critical locations in the casting

Experimental procedure The material used in this research work was extracted from an AA7050 billet produced by DC casting with a conventional mold from the melt that was degassed in the furnace, and supplied by Corus-Netherlands (Umuiden). Chemical composition of the alloy in terms of wt% is as follows: Zn: 6.3, Mg: 2.42, Cu: 2.49, Zr: 0.098, Ti: 0.03, Fι: 0.07, Si: 0.04, Mn: 0.04 and Cr: <0.01. Tensile mechanical properties of the as-cast samples were measured using a Gleeble-1500 thermomechanical simulator. Samples were cut from a cylindrical billet with a diameter of 255 mm along its radial direction. Tensile specimens were heated through the Joule effect at a rate of 10 K/s, kept for 10 seconds and then uniaxially deformed at three strain rates of IO-2, 10"3 and 10-4 s"1. The range of strain rates was chosen to resemble those typical of DC casting [15]. Mechanical properties were measured from room temperature to 400 °C at 100 °C steps. Four samples were tested for each specific temperature and strain rate combination, and the average values are reported.

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In order to assess the critical crack size leading to catastrophic failure, plane strain fracture toughness tests were performed from room temperature to 200 °C following ASTM-E399 regulations. The samples prepared based on the dimensional proportions mentioned in the standard were locally polished around the chevron notch to be able to follow and accurately measure the fatigue crack length. Samples were then fatigue pre-cracked at room temperature to reach the total crack length of 15 mm (including the chevron notch length of 13 mm). Loading rate was chosen in such a way to keep the stress intensity within the range of 0.55-2.75 MPa m1/2 s"1. At higher temperatures, samples were first covered with a ceramic coating for insulation and preventing a sudden temperature fall (temperature tolerance was ±10°C) and then kept at desired temperatures in an oven for 20 minutes.

The effect of strain rate on the flow stress of the AA7050 alloy at the true strain 0.002 is shown in Figure 3. At high temperatures (300 and 400 °C) the flow stress increases as the deformation rate increases, but at lower temperatures (beginning at 200 °C) the material behavior becomes strain-rate independent. 1

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Figures 1 and 2 show the effect of temperature on the frature strength and ductility (% reduction in area). Due to very low ductility the fracture strength can be accepted as the yield strength. The fracture strength reaches 266 MPa at room temperature and falls to 23 MPa as the temperature increases to 400 °C. The fracture strength increases as the temperature decreases in a linear manner. As temperature falls below 400 °C, the material starts to behave even more brittle and shows less ductility. The alloy loses its ductility completely below 200 °C, which is accompanied by a higherfracturestrength.

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The comparison of the Kic results gained for the genuine as-cast material here with the minimum Kic for the material in the precipitation hardened state (20.9-27.5 MPa-m1/2, at room temperature [7]) shows how brittle the material is under as-cast conditions.

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Figure 2. Ductility of the AA7050 samples at different temperatures and strain rates [16].

In order to simulate the DC-casting process from solidus temperature to room temperature suitable equations describing the

670

material's mechanical behavior are required. One approach is the extended Ludwik equation [17]:

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equal to 0.001 [12]. The true stress-strain curves of the material were fitted to Equation 1 to determine K(T), n(T) and w(7), and the results are presented in Figures 5 through 7. In the same graphs, the reference data gained from the stress relieved samples [14] and the homogenized samples [18] are also reported. As can be seen in Figure 5, the consistency of the genuine as-cast material falls in a continous manner with temperature. In homogenized and stress relieved samples however, it passes through a plateau and then falls with increasing the temperature. It is noticeable that the main difference occurs at lower temperatures were the material is more prone to cracking due to its extreme brittleness. The same trend is observed for strain hardening coefficient of the alloy (Figure 6), although the difference in the results appears to be higher at elevated temperatures especially in the stress relieved material. Stress rate sensitivity values are in a relatively good agreement at lower temperatures, but the difference becomes larger at higher temperatures (Figure 7). m values turn to zero below 200 °C in agreement with the strain rate independent behavior shown in Figure 3. 1000

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Figure 7. Strain rate sensitivity of the AA7050 alloy gained from various samples: as-cast (this study), and homogenized [18]. Computer simulations ALSIM5 was used for the computation of temperature profile and stress/strain fields in the round AA7050 billet. Detailed description of the model can be found elsewhere [19, 20]. Geometry of the setup consisted of hot top, mold, water jet, bottom block, and the cast domain (Figure 8). New elements with the size of 0.75 mm are added to the geometry at the casting speed to simulate the continuous casting conditions. So during casting, the bottom block moves downwards while new elements are added to the system, and the mold, hot top and molten metal retain their initial position. Due to axial symmetry, only half of the billet is considered. Time-dependent thermal boundary conditions are defined to account for filling time, the air gap formation between the billet and the bottom block as well as at the billet surface, and for different heat extraction in different parts of the casting system [19]. The process parameters are listed in Table 1.

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Figure 6. Strain hardening coefficient of the AA7050 alloy gained from various samples: as-cast (this study), stress relieved [14], and homogenized [18].

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trend is observed for the circumferential stress, but it turns to compressive in the vicinity of the surface (71 MPa in the center, and -86 MPa at the surface; Figure 9c). Contour map of the axial stress (along "y" axis) follows the same trend in the lower part of the billet, i.e. tensile stresses in the center and compressive stresses at the surface. This, changes in the upper part of the billet where we see compressive stresses below the high temperature zone of the billet and tensile stresses around the water impingement area (Figure 9d). It is important to note that the circumferential stress at the surface passes through a transition from tensile to compressive as soon as the surface leaves the water impingement zone (WIZ). The tensile stresses in the WIZ reach the highest value (100 MPa) among the stresses formed in the billet. To be able to discuss the failure probability in the billet the contour maps of the three components of the principal stress tensor (σ33<σ22<σιι) are shown in Figures 10a, b, and c. In agreement with the results shown in Figure 9 the principal stress components appear to be tensile in the center of the billet and compressive at the surface.

Table I. Description of casting process parameters. Process parameter Ingot diameter (mm) Final length of the billet (mm) Casting speed (mm/s) Melt temperature (°C) Water flow rate (1/min) Water temperature (°C) Start temperature of bottom block (°C)

Value 200 380 1 680 80 15 20

1

Thermophysical properties of the alloy were gained from the thermodynamics database JMat-Pro provided by CorusNetherlands (Umuiden). The temperature dependence of the coefficient of thermal expansion, specific heat, heat conductivity, fraction solid, density and kinematic viscosity were extracted and embedded in the model. The solidification range as well as the liquidus and the non-equilibrium solidus were determined using DSC tests. Tensile mechanical properties, constitutive parameters and the Kic values of the as-cast AA7050 samples obtained from the experiments mentioned in the previous section were utilized in the model. A new module was introduced to ALSIM5 by means of which critical crack sizes could be calculated for the desired crack type and geometry based on the given Kic values and the maximum principal stress component. Simulation results and discussion Casting simulation was performed for 380 seconds to make sure that the steady-state conditions are gained and stress analysis can be done. After this moment, the stresses remained more or less unchanged until changes in the thermal boundary conditions were applied or casting ceased. As can be seen in Figure 9a, the lower part of the billet has reached temperatures below 80 °C.

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As discussed by Boender et al [11], although having three tensile principal stress components in the center results in a nearly zero equivalent von Misses stress, it does not imply that no failure may occur. In other words, such a stress state facilitates the occurrence of a brittle fracture in the center of the billet. According to Rankine's theory which is more applicable to brittle materials, failure occurs when either the maximum principal stress reaches the tensile strength or the minimum principal component reaches the uniaxial compressive strength [21]. As the stresses computed by ALSIM are far below the tensile strength of the material (225 MPa at 200°C to 266 MPa at room temperature), the effect of stress raisers (cracks and flaws) should be taken into account. Similarly, in a brittle material and under a triaxial state of stress cracks mainly orient themselves normal to the largest component of the principal stresses (óη)- Hence, the maximum principal stress component was selected for calculation of the critical crack sizes. A penny shaped crack was chosen as a 3D crack resembling the actual void shape in billets. For such a crack the critical crack size (radius of the penny) can be calculated as follows [5]:

WÊÊÊÊÊ (b) ft„ (MPa)

(e

Figure 9. Computer simulation results showing the contour maps of: (a) temperature, and normal residual thermal stresses in b) radial, c) circumferential and d) axial direction after 380 s of casting at speed 1 mm/s. Figures 9b through 9d show the contour maps of the normal residual thermal stresses generated under steady-state casting conditions. As can be seen, the residual radial stresses displayed in the lower part of the billet (on the top of the bottom block) are compressive which turn to tensile as we move in the positive "y" direction. Along the radial axis (x-axis), radial stress diminishes from 71 MPa to 3 MPa as we move towards the surface. Similar

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center of the billet, which means higher failure probability (Figures 11a through c). In agreement with our results in Figure lid, the water flow rate has a negligible effect on the crack size (Figure 12d).

The calculated critical crack sizes are shown in Figure 10c, with most critical locations in the center as well as the WIZ. Validation of the simulation predictions In order to check the validity of the cold cracking criterion, a newly developed high strength 7xxx aluminum alloy, which is highly prone to both hot and cold cracking, was selected. The material was cast in the form of a 260-mm diameter billet at Corus-Netherlands (IJmuiden) through DC casting with a conventional mold (without hot top) from the melt that was degassed in the furnace. Constitutive parameters, mechanical properties and the plane strain fracture toughness of the material were determined following the same procedures mentioned for AA7050. The mechanical properties of the new alloy resemble that of AA7050 except for the fact that the new alloy exhibits more plasticity at 200°C and the Klc values are higher compared to AA7050. The thermophysical properties of the material were obtained from the JMat-Pro database provided by CorusNetherlands. Finally, mechanical as well physical properties databases were prepared to be implemented in ALSIM5 for thermomechanical simulations and calculation of critical crack sizes. A 260-mm diameter billet, cast at 1 mm/s with water flow rate of 35 1/min was taken as the standard case. Casting speed and water flow rate were varied to study the effect of these variables on the cracking propensity (Table II).

Figure 11. Simulation results showing the óη (MPa) in the billets cast at various conditions mentioned in Table II. Arrows show the transition between casting regimes [23].

Table II. Description of the casting process parameters for the 0 260-mm billets. case 1 2 3 4

Casting speed (mm/min) 60 (1 mm/s) 80 (1.3 mm/s) 110 (1.8 mm/s) 60 (1 mm/s)

Water flow rate (1/min) 35 35 90 70

As in many cases the billets fracture at the end of casting when they are in the complete solid state, it is wise to adjust the simulation conditions in such a way that resemble the ones during the failure of the billets in practical situations. To achieve this, the casting conditions were set as follows: 1 - the first 500 s of all simulations was performed under standard conditions (case 1), 2 changes in casting speed or water flow rate were applied afterwards by ramping up from case 1, 3 - the billet was cast then with new parameters for another 500 s, 4 - eventually, the casting speed was ramped down to zero and the billet was cooled down to room temperature over 200 s. Contour maps of the óη are shown in Figure 11 for the cases described in Table II. In Figures 1 lb,c and d, the lower parts of the billets correspond to the initial casting conditions and the upper parts to the new conditions. With increasing the casting speed, the óη increases in the centre of the billet (Figures. 11a through lie). The increase in the magnitude of the au is due to the increased heat input mainly in the axial direction (-y) resulting in higher temperature gradients in that direction [22]. Water flow rate has a negligible effect on the magnitude of the óη and the simulation results of the billet cast with 70 1/min water flow rate (Figure 1 Id) are very similar to Figure 11a. Figure 12 shows the contour maps of the CCS distribution in billets cast under conditions described in Table II. As expected, increasing the casting speed leads to smaller CCS values in the

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Figure 12. Simulation results showing the critical crack size distribution (mm) in the billets cast at various conditions mentioned in Table II. Crack sizes larger than 30 mm are neglected and appear as gray between the billet surface and the black area at mid-radius [23]. Arrows show the transition between casting regimes and the star in (c) indicates the location of the inclusion which triggered the catastrophic failure. DC-casting trials were performed for the cases mentioned in Table II to check the simulation results. No cracks were observed in cases 1, 2 and 4. A real cold crack however occurred in case 3 during the casting with an audible bang resulting in the failure of the billet. According to Figure 12c, in the center of the billet cracks or defects with a critical diameter 6 to 12 mm (3 to 6 mm in radius) may lead to catastrophic failure. Further investigation of the fracture surface in case 3 revealed an inclusion with a length of 7 mm, which was located 20 mm away from the center of the billet and 730 mm above the bottom block (Figure 13). The predicted critical crack size (the diameter of the penny) for the coordinate mentioned above is 7.5 mm, which is 0.5 mm longer than the actual observed void. The reason for such a deviation

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might be the fact that the actual crack shape is more complicated and irregular than the simplified penny shaped crack. In reality, cracks have sharper edges that may approach that of an ideally sharp crack.

Figure 13. The cold crack surface in the billet cast at 110 mm/min and water flow rate of 90 1/min. The dark point shown by an arrow triggered thefracture.The chevron markings radiating from this oxide inclusion indicate its role in the fracture. Conclusions Cold cracking propensity was studied using the simulation results and casting trials on several billets under various conditions. The simulation results showed that as the casting speed increases, the CCS decreases leading to a higher failure probability of the billets. This was proven by casting trials. The critical crack sizes were validated upon experiments, where a 7 mm inclusion in the centre of the billet triggered the catastrophic failure. Acknowledgments This research was carried out under the project number MC4.05237 in the framework of the Research Program of the Materials innovation institute M2i (www.m2i.nl). Support of Corns (Tata Steel Europe) and Aleris is greatly appreciated. Authors would like to especially thank Dr. A. Ten Cate for the preparation of the geometry and mesh simulation files. References [1] J.B. Hess, "Physical Metallurgy of Recycling Wrought Aluminum Alloys", Metallurgical Transactions A, 14 (1983), 323-327. [2] M. Lalpoor, D.G. Eskin, and L. Katgerman, "Microstructural Features of Intergranular Brittle Fracture and Cold Cracking in High Strength Aluminum Alloys", Materials Science and Engineering A, 527 (2010), 1828-1834. [3] B. Hannart, F. Cialti, R. Van Schalkwijk, "Thermal Stresses in DC-Casting of Aluminum Slabs: Application of a Finite Element Model", ed. U. Mannweiler, (TMS, Warrendlae, PA, 1994), 879887. [4] J. Campbell, Castings (Oxford, United Kingdom: ButterworthHeinemann, 1991), 242-245. [5] M. Lalpoor, D.G. Eskin, and L. Katgerman, "Cold Cracking Assessment in AA7050 Billets during Direct-Chill Casting by Thermomechanical Simulation of Residual Thermal Stresses and Application of Fracture Mechanics", Metallurgical and Materials Transactions A, 40 (13) (2009), 3304-3313.

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[6] M. Lalpoor, D.G. Eskin, and L. Katgerman, "Fracture Behavior and Mechanical Properties of High Strength Aluminum Alloys in the As-Cast Condition", Materials Science and Engineering A, 497 (2008), 186-194. [7] J.R. Davis, ASM Specialty Handbook: Aluminium and aluminium alloys, (Materials Park, Ohio: ASM international, 1994), 68-75. [8] V. A. Livanov, "Casting of Large Ingots for Sheet Production from Aluminum Alloys", A.F. Belov and G.D. Agarkov Eds., (Paper presented at Aluminum Alloys, Oborongiz, Moscow, 1955), 128-168. [9] J.-M. Drezet, M. Rappaz, and Y. Krähenbühl, "Thermomechanical Effects during Direct Chill and Electromagnetic Casting of Aluminum Alloys, Part II: Numerical Simulation", ed. J. Evans, (TMS, Warrendale, PA, 1995), 941950. [10] K,-M. Chang, B. Kang, "Cracking Control in DC-Casting of High Strength Aluminum Alloys", Journal of the Chinese Institute ofEngineers, 22 (1) (1999) 27-42. [11] W. Boender et al., "Numerical Simulation of DC-Casting; Interpreting the Results of a Thermo-mechanical Model", ed. A.T. Tabereaux, (TMS, Warrendale, PA, 2004), 679-684. [12] O. Ludwig et al., "Modeling of Internal Stresses in DCCasting and Sawing of High Strength Aluminum Alloys Slabs", ed. C.-A. Gandin and M. Bellet, (TMS, Warrendale, PA, 2006), 185-192. [13] J. Wan et al., "As-Cast Mechanical Properties of High Strength Aluminum Alloy", ed. B. Welch, (TMS, Warrendale, PA, 1998), 1065-1070. [14] K.-M. Chang et al., "Computer Simulation of Solidification Cracking in High Strength Aluminum Alloys: Basic Concepts and Approach" (Paper presented at the 2nd International Conference on Quenching and the Control of Distortion, Cleveland, Ohio 4-7 November 1996), 341-345. [15] W.M. van Haaften, "Constitutive Behavior and Hot Tearing during Aluminum DC Casting", (Ph.D. thesis, Delft University of Technology, 2002), 5. [16] M. Lalpoor, D.G. Eskin, and L. Katgerman, "Constitutive Parameters, Mechanical Properties and Failure Mechanism in DCCast AA7050 Billets" (Paper presented at the 12th Intern. Conf. on Fracture, Ottawa, Canada, 12-17 July 2009). [17] B. Magnin, L. Katgerman, and B. Hannart, "Physical and Numerical Modeling of Thermal Stress Generation during DC Casting of Aluminum Alloys", eds. M. Cross and J. Campbell, (Paper presented at MCWASP VII, TMS, Warrendale, PA, 1995), 303-310. [18] W. M. van Haaften, (Internal report, Corus Ijmuiden, The Netherlands, 2002). [19] H.G. Fjaer, A. Mo, "ALSPEN-A Mathematical Model for Thermal Stresses in Direct Chill Casting of Aluminum Billets", Metallurgical Transactions £21 (6) (1990) 1049-1061. [20] D. Mortensen, "A Mathematical Model of the Heat and Fluid Flows in Direct Chill Casting of Aluminum Sheet Ingots and Billets", Metallurgical and Materials Transactions B, 30 (1) (1999)119-133. [21] J.H. Faupel and F.E. Fisher, Engineering Design, (New York (NY): John Wiley & Sons, Inc., 1981), 242-252. [22] J.F. Grandfield and P.T. McGlade, "DC casting of Aluminum: Process Behavior and technology", Materials Forum 20(1996)29-51. [23] M. Lalpoor et al. "Application of a Criterion for Cold Cracking to Casting High Strength Aluminum Alloys", Materials Science Forum 654-656 (2010) 1432-1435.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SURFACE DEFECTS ON DIRECT CHILL AS-CAST 6XXX ALUMINUM BILLETS Mikael Erdegren, Torbjφrn Carlberg Mid Sweden University, Department of Natural Sciences, Engineering and Mathematics, 851 70 Sundsvall, Sweden Keywords: Aluminum, DC casting, Segregation, Surface defects, Vertical drags the rest, and the studied defect was very large and less usual than the rest. According to statistics 6082 billets do not usually have VD defects going along the whole billet. They rather have a characteristic rough surface with small random ridges a few millimeters long, which for this type of alloy, is considered normal. That type of surface is represented in sample G.

Abstract Surface defects found during air-slip direct chill casting are today a major quality factor for extruded aluminum, as they can cause increased segregation, pores and unwanted precipitations. The surface zone in billets of the aluminium alloys 6063, 6005 and 6082 have been analysed by metallographic methods and by chemical analysis. Surface defects, of the type vertical drags, were investigated and compared to defect free surfaces. Inverse segregation to the surface was quantitatively analysed. The concentration profiles were coupled to the appearance of the defects and to microstructures from corresponding areas. It was shown, for vertical drags on 6005 billet surfaces, that either the segregation depth or the precipitated particles were different from surface areas without defects. For the 6063 alloys on the other hand the vertical drag zones contained different particles than surfaces without defects and the segregation had noticeably increased.

Samples C, E and G were selected for a segregation analysis, and comparisons were made between areas with and without defects on the surface of the same billets. Those billets were homogenized, but also two not homogenized ingots H and I were analyzed in the same way. All samples mentioned were also used to identify intermetallic phase particles. Table 1. The samples designated letters for respective billet type. VD = Vertical drags and UH = Unhomogenized. Sample name A B C D E F G H I

Introduction In the production of aluminum billets a major problem has been identified, which is addressed in this project, namely surface defects. This problem is of large economic importance, and is today the main cause to scrap during vertical direct chill (DC) casting. Billet surface quality is for example evaluated by measuring surface appearance, segregate zone thickness, large Mg2Siparticle near the surface and area fraction of pores. Surface defects is said to increase the segregation zone, but also the size and number of the Mg2Si-particles. The presence at the surface of many Mg2Si- and ί-particles can lead to pick-up defects on the aluminum profiles during extrusion [1]. Vertical drags (VD) are one of the most common defects found at DC casting. A possible cause for formation of VD at air-slip hot top casting could be that, due to uneven and/or low air-pressure in the graphite ring, liquid aluminum sticks to the ring in the DC casting mould [2]. The purpose of this article is to analyze various surface defect structures, the segregation zones and to identify the intermetallic phase particles in the surface region. Most of the previous works connected to this research were done on a macroscopic level [3], but this paper involves studies of both microstructure and macrosegregation.

Al alloy

Type

6005 6005 6005 6063 6063 6063 6082 6005 6063

VD VD VD VD VD VD UH+VD UH+VD

Billet diameter [mm] 178 178 228 178 203 178 203 178 178

The surface defects were first studied at lower magnifications in a stereo light microscope (SLM). Before further microscope studies all samples were mounted in a polymer with hardness similar to that of the alloys to avoid edge effects during grinding. They were then ground with wet silicon paper and polished with a polishingsystem. No water was used in the last step of polishing, as water etches the water sensitive Mg2Si particles. After polishing the samples were studied with a light optic microscope (LOM) axially and radially. Later a scanning electron microscope (SEM), with an energy dispersive x-ray spectrometer (EDX), was used to identify intermetallic phase particles. This was done by calculating the Fe:Mn:Si:Mg ratio for particles with different shape and color [3]. After the phase identification the samples with surface defects were etched in a solution of 0.5% hydrofluoric acid (HF) for 10 seconds, restudied in a LOM and the defects and segregation zone depth were measured with the software Kappa Imagebase Metero 2.7.

Experimental Procedures The 6xxx series aluminum alloys are primary made for extrusion. Three of the most commonly extruded alloys were therefore chosen for this study. These three were 6005, 6063 and 6082. From the aluminum industry 30 mm thick discs were cut from the billets, as-cast by air-slip technology (Wagstaff). From the discs different samples were extracted and table 1 shows all the samples presented in this article. Samples A-G were chosen to show different types of defects on homogenized billets. They consisted of three VD defects from 6005 billets and three from 6063 billets. Sample F was taken from a considerably older billet compared to

The grain sizes at different depths were also measured, at both defects and at defect free areas. To do this the samples were anodized to improve the grain measurements. They were placed as

675

of 400 μιη was recorded. The second observation was that the aAl8Fe2Si particles in the defects followed the meniscus lines. In samples E α-particles could sometimes be seen 1 mm along these lines.

anode in a solution, consisting of 24 ml ethanol, 1 ml HF (50%), 1 ml fluoroboric acid (HBF4)(35%) and 74 ml tap water. The cathode was a platinum basket. The voltage was 50 V and the anodization time was 90 seconds. When inspecting the anodized samples in a LOM polarized light was applied. Grain size variations on the macro scale for these three alloys have also been presented in previous work [4]. To properly study the segregation in a sample the composition at different depths from the surface had to be measured. This was accomplished with a method that was used in [5]. The billet discs were put in a milling-machine so the surfaces of the billets could be cut out like ring segments, 3-4 mm thick. These samples were then straightened with a 20-ton press and placed in an optical emission spectrometer (OES) for a composition measurement. After each run in the OES the samples were ground and again put in the OES, until the depth from the original surface was 1000 μπι. First five measurement points were done with 10 μπι steps as the segregation shifts fast near the surface.

200 ìçé

Results Looking at the billets surface, the VD defects could have varying length, location and "track" appearance. Sometimes they spanned across the whole length of the billet and other times just short sections. Common VD appearances were named: tire tracks (figure la), wave tracks (figure lb), zipper tracks (figure Id) and fish bone tracks (figure If). The width and depth of the VD's also varied along the billets.

(a) 6005 VD, sample A.

The defect profiles for most of the VD defects tended to go both in and out in relation to the cast surface. When following the billets axis saw-tooth or wave shaped patterns were observed. These types of patterns created corresponding variations also in the structure at the defects, i.e. the amount of precipitations, the types of precipitations and the influenced depth. For 6005 and 6063 billets, both beside and at defects, the structure variations were coupled to the meniscus lines. At positions where the meniscus lines went deep into the billets, or just before the start of the lines, the zone depths were large but just after the lines started at the surface the segregation zones were thin.

400 ìçé

(b) 6005 VD, sample B.

For two of the three 6005 samples, A and B, the grains seemed to have started coming apart and move away from the surface (figure la and lb), thus creating an increase of the porosity. In none of the 6005 VD cases did the radial segregation zone depth, seen as a zone with increased amount of precipitations, change at the defects compared to at defect free surfaces. In figure 3a the surface zones of a homogenized 6005 alloy is shown. The segregation zone depth had an average of 90 μπι and the segregation consisted of thick plate- or rod-shaped ί-Al5FeSi particles. For the 6063 samples the analyzed particles were found to be of different types in the defect zones compared to defect free surfaces. The defect free homogenized 6063 surfaces contained mostly thin needle-shaped ί-particles, figure 3b, but at the defect zones Chinese script-shaped a-AlgFe2Si particles dominated, figure 3c. At the surface they can have the appearance of small cracks as in figure Id. Following the axial direction two observations could be made. Firstly, the segregation zone depths varied a lot more at the defects compared to at a normal surface. In the 6063 samples studied, the average normal segregation zone depth was 70 μπι, but for defects the variations were strong, and a maximum depth

(c) 6005 VD, sample C. Figure 1. LOM surface pictures of samples A-G with their respective radial structure.

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The 6082 sample, G in figure lg, had a segregation zone that, continuously throughout the billet surface, randomly changed in depth. When viewed radial and axially, the segregation zone looked the same and the depth of the zone varied between 50 μπι and 500 μπι. The zone contained many large pores and a high concentration of them often occur at the bottom of the segregation zone creating a porous line parallel to the surface. This area was clearly visible by a color difference between the segregation area and the bulk of the sample, caused by the etching. The particles found were mainly a-Al15(FeMn)3Si2 (figure 3d), probably due to the higher Mn concentration in this alloys. Mg2Si particles appear as black areas when studied with a LOM or SEM. Black areas could sometimes be pores, but pores are generally larger than Mg2Si particles. When the samples are observed in a LOM instantly after polishing with no water, the particles remain visible and appear as slightly darker gray areas than a-and ί-particles. It was hard to compare the amount of Mg2Si between the samples, however, they tend to lie near the a- and ίparticles. Therefore an increase of these particles normally implied an increase of the number of Mg2Si particles.

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(f) 6063 VD, sample F. 6005 and 6063 unhomogenized billets have rather similar structures, but the fraction eutectic is larger in 6005 billets, figure 2. Comparison to the homogenized billet in figure 3a shows that the particles thicken while the amount of eutectic areas decreased during homogenization. Both alloys contain some needle-shaped particles, which remain to a larger extent in the 6063 alloy after

677

4

The phase particles mentioned in this section were confirmed by EDX measurements and by comparison with previous work in the same field [6-8]. Composition ratios and shape of a particle revealed the phase type. All the spot-analyzes gave higher than expected Si concentration, which made some of the ratios deviate from expected values, but not enough to make identification impossible. The thick and thin ί-Al5FeSi particles shown had a Fe:Si ratio of 1:1.1 and 1:1.3 respectively, the Chinese scriptshaped <x-Al8Fe2Si particle had a ratio of 2:1, and the aAl15(FeMn)3Si2 particle had the Fe:Mn:Si ratio of 3.3:3:3.6. The grain size measurements in the surface zones were rather inconclusive. The average grain size was random. Neither at defects nor at regular surfaces could a clear pattern be obtained. Grain sizes ranged between 80-130 μπι for all samples. A rough pattern was that for the VD samples, regardless of alloying type, smaller grains appeared near the surface, but at 200-300 μπι from the surface some larger grains were found. Even further away from the surface, the grain size started to decrease again. The grains near the surface had grown slightly more columnar towards the surface hence making them appear smaller when measuring along lines parallel to the surface, which was the method used. The measurements were performed to the maximum defect depth, which was at 500 μπι.

(a) Plate- or rod-shaped ί-Al5FeSi particles found in 6005 billets.

(b) Needle-shaped ί-Al5FeSi particles found in 6063 billets.

Results from the composition measurements are presented in figure 4. First measurement points are at 10 μιη from the surfaces as the spark patterns obtained on the not grinded surfaces were not always reproducible. The concentration profiles compared well with most of the segregation zones as they were observed in LOM, i.e. the thicknesses of the zones with increased particle precipitation coincided with the starting points of strongly increased concentrations. For the 6005 samples (C and H), when profiles beside a defect and at a defect are compared, no significant differences could be distinguished. Comparing the different billet slices C and H, the homogenized and not homogenized, respectively, clear differences can be seen. The former goes deeper into the billet, to about 400 μπι, while the latter has an increased concentration starting about 200 μπι from the surface. For the 6063 samples (E and I) the VD defects have caused segregation to go further into the billets causing deeper segregation zones, especially in the homogenized case. Near the surfaces of the VD defects the Fe and Si contents have been lowered compared to the defect free surface areas. For the 6063 and 6082 samples the defects had varying segregation depths, which meant different concentration profiles for selected measurement locations, but the total area for the OES measurement points at each depth was considered large enough to take this into account and give an average value. All surfaces except for billet (I) seemed to have a Fe content of about 1.5 wt% close to the surface. When looking near the surface the differences between a 6005 VD and a 6063 VD were the content of Mg and Si. 6005 VD had -1.5 wt% Mg and -2.5 wt% Si, while 6063 VD had -0.8 wt% Mg and -1.5 wt% Si.

(c) Rod- and Chinese script-shaped a-Al8Fe2Si particles found at VD defect in 6063 billets.

(d) a-Al15(FeMn)3Si2 particles found in 6082 billets. Figure 3. SEM and LOM pictures of different intermetallic particles from surfaces of homogenized billets. The darker pictures to the left are taken from SEM.

Discussion The VD defects can have very different appearances as can be seen in figure 1, which implies that it is not possible to give a general explanation to the formation of the defects or to the struc tures and segregations related to them. The inverse segregation to

the surface defects is dependent on defect type, but the differences between defects and smooth surfaces sometimes are smaller than differences between different billets, indicating that the casting parameters are important for controlling the surface segregation.

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Figure 4. Diagrams showing Fe, Mg, Mn, and Si concentration profiles as a function of depth from surface. The letters C,E,H,I and G refers to different billets listed in table 1. VD notes that the profile has been taken at a VD defect, and UH notes that the ingot is not homogenized. Note the scale difference on the x-axis for sample 6082 (G). 200

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679

One of the reasons for formation of VD defects, mentioned in the introduction, is that aluminum sticks to the graphite ring due to drop in air pressure, and that metal is coming out of the oxide skin at theses point where the skin is strained. The sticking to the ring can explain the grain cracking for the larger of the 6005 VD defects shown in figure 1, A and B. The strong segregation, figure 4 C and H, and thus large fraction eutectic formed at the surface of 6005, figure 2a, decrease the coherency between the dendrites which, together with sticking, can open up the structure at the surface with oxides and pores as a result. This can be the reason why 6005 alloys have a high sensitivity to formation of VD defects. The inverse segregation to the surface in this alloy creates a local composition with relatively large areas with low melting point, and thus a loose structure, which becomes sensitive to sticking at the graphite ring. This can be the reason why 6005 alloys have a high sensitivity to formation of VD defects. The inverse segregation to the surface in this alloy creates a local composition with relatively large areas with low melting point, and thus a loose structure, which becomes sensitive to sticking at the graphite ring.

probably are of less importance than the movements of the shell. If the shell, dependent of defect type, is stretched out or pressed together the enriched liquid will be sucked in or pressed out, respectively, giving different types of macro-segregations in the surface zones. Conclusions The surface segregation in the studied 6005 billets was not changed by VD defects, while the segregation in a 6063 billet showed an increased depth and a high addition of a-particles. The segregation zone for 6082 varies randomly in depth and contains large pores. No defect in this research caused the segregation zone to go deeper than 500 μπι. All analyzed as-cast billets have different concentration profiles at the surface, implying that the dimension of the billets as well as the casting process are as important as the defects for the surface structure and composition. Acknowledgments This project has been financed by the Knowledge Foundation and EU Mβl-2. Dr Majed M.R. Jaradeh, MSc Mohammad W. Ullah and MSc Par Âslund at Mid Sweden University are acknowledged for help with the experimental equipments. Lena Evertsson and Henrik Oscarson at Sapa Technology are thanked for sharing knowledge about metallographic preparation of aluminum alloys.

A slightly different defect type can be seen in figure lc. Here the defect profiles mostly go inwards and the structure is not loose or porous as in figure la and lb. A possible cause for this type of defect can be that oxides or refractory particles are stuck between the graphite ring and the billet surface [9]. An old graphite ring can sometimes have very clear indentions, with a scratch below probably caused by the release of the particle.

References 1. T. Minoda et al., "The Mechanism of Pick-up Formation on 6063 Aluminum Alloy Extrusions," Sumitomo Light Metal Technical Report, 40(1999), 22-27. 2. R.J. Collins, "Consistent Casting Ring Performance Measuring and Tracking," (Form WP103-1, R. J. Collins Incorporated, Spokane, Washington, USA, June 2006). 3. D.G. Eskin: Physical Metallurgy of Direct Chill Casting of Aluminium Alloys (CRC press, USA, 2008), 47-49. 4. Jin Hu, Majed Jaradeh, and Torbjφrn Carlberg, "Solidification Studies of 6xxx Alloys with Different Mg and Si Contents," Light Metals, TMS, 2005, 773-779. 5. D. Mortensen et al., "Coupled Modelling of Air-gap Formation and Surface Exudation during Extrusion ingot DC-casting," Light Metals, TMS, 2008, 773-779. 6. S.R. Claves, D.L. Elias and W.Z. Misiolek: Materials Science Forum, vols. 396-402, 2002, 667-674. 7. R. M. Kelly et al., "Predictive Metallographic Assessment of Extrudability and Comparative Testing of 6xxx Aluminum Alloys Billets," (Paper presented at the 8th International Aluminum Extrusion Technology Seminar ET, Orlando, FL, May 2004), 3950. 8. M.W. Ullah, "Solidification Studies of Aluminium 6000 Series Alloys to Increase the Understanding of Surface Defect Formation during DC casting of Aluminium Billets," (Master thesis, Royal institute of technology, Sweden, 2009). 9. J.M. Ekenes, "Visual Observation Inside an Airslip Mould during Casting," Light Metals, TMS, 1990, 957-961. 10.1.L. Ferreira et al., "Analytical, Numerical, and Experimentaal Analysis of Inverse Macrosegregation during Upward Unidirectional Solidification of Al-Cu Alloys," Metallurgical and Materials Transactions B, vol 35, 2004, 285-297.

The different phases precipitated in the segregation zones are dependent on the local alloy composition, and the concentration profiles shown in figure 4 shows that there are completely different alloy contents at the surface than in the bulk of the billets. The ratio between the components controls the type of precipitation, and for the AlFeSi phases the Fe:Si ratio is important. For the 6005 alloys, C and H in figure 4, the Si content is higher in the whole surface zone, and the dominating phase is as expected ίphase in the shape of relatively thick plates, figure 3a. For the 6063 sample E in figure 4, the Fe:Si is about 1 at the very surface, and in this alloy, and especially at defects, a-particles were found, figure 3c. Different cooling conditions can also influence the type of precipitates, and for some of the defects in the 6063 alloy a slower cooling rate, combined with a smaller Fe:Si ratio, can explain the appearance of a-particles. The surface segregation, both at smooth surfaces and at VD defects, is basically a result of shrinkage in the solidified shell, which causes interdendritic liquid of high alloy content to be sucked to the surface. This results in the typical increased concentrations at the surface and sometimes in exudation of enriched liquid out of the surface. The shape of the composition profiles shown in figure 4 shows a large variety, which could be confirmed also by analysis of more samples not included in this presentation. Both defects and smooth surfaces can have different types of surface segregation. In an analysis of inverse segregation [10] the influence of heat transfer at the surface was discussed, and it was found that high heat transfer gave steeper profiles at the surface and less penetration. High concentration of a certain element, in that case Cu, could also increase the steepness of the profiles. The present investigation indicates that, in cases where a difference can be seen, the defects rather give a deeper penetration of the segregation, i.e. higher contents in a thicker layer inside of the surface. Here the contributions from cooling rate differences

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECT OF COOLING WATER QUALITY ON DENDRITE ARM SPACING OF DC CAST BILLETS S.P.Mohapatra, S.Nanda, A.Palchowdhury NALCO (National Aluminium Company Ltd), Smelter Plant, Angui, Orissa, 759145, India Keywords: Direct Chill, Dendrite Arm Spacing, Solidification

attempted to quantify the interfacial heat transfer during solidification in terms of heat transfer coefficients [4] or heat flux [5].They have highlighted several factors that affect heat transfer during DC casting process and in turn the microstructure. Influence of process variables such as mold surface roughness, mold material, superheat, alloy composition & lubricant on heat transfer & cast structure has been studied by C.A. Moujekwu et al [6]. They have proposed an empirical transient heat flux model to take the above variables into account. Dominique Brochord et al [7] have tested the general predictability of dendrite arm spacing for unidirectional solidified binary alloys through experiments. They have concluded that for prediction of DAS, the exponent 1/3 in the exponential relationship of cooling rate with DAS is adequately represented in both steady & unsteady state heat flow.

Abstract For a given alloy chemistry & casting technology, the quality of a billet produced through DC ( Direct Chill) casting route is largely defined by the DAS (Dendrite arm spacing). DAS in turn, is determined by the cooling rate and local solidification time. The cooling rate is influenced by the quality & quantity of water besides casting speed and melt temperature. This paper focuses on the relationship between the quality of cooling water & the resultant DAS during production of DC cast billets. An empirical relation is derived to indicate the influence of cooling water characteristics on the DAS, based on a mathematical model supplemented by experimental investigations. The model has been validated in plant scale trials.

Using empirical relationships, the dendrite arm spacing & degree of homogenization have been predicted by C.Devdas et al [8]. The effect of casting speed, super heat & metal height on the heat transfer has also been analyzed by them. The study of Ivan Todaro et al [9] also focuses on the relationship between DAS and local cooling rate. They have also been able to correlate DAS with the mechanical properties of the cast. The work of Bomberger et al [10 ] have demonstrated that increase in silicon content of an aluminium alloy reduces cooling rate and hence has an influence on DAS. Mark Easton et al [11] have been able to predict SD AS (Secondary Dendrite Arm Spacing) for multi-component aluminium alloys and have inferred that grain refiner has little effect on SDAS; rather composition of liquid at the temperature at which SDAS is determined is the most influencing factor.

Introduction The quality & properties of cast products strongly depend on microstructure development during solidification. The DAS is a fundamental characteristic of the microstructure and has been used over the years as a means offinenessand, hence is a measure of the quality of cast products. It has also been established that DAS exercises a significant influence on extrudability. Therefore it is of great interest to analyze the inter dependence of the DAS and process parameters for improving product quality & development of superior methods for quality castings. Cooling rate and solidification time have a dominant impact on DAS. A quantitative understanding of the cooling rate experienced by the component during solidification is necessary. Heat transfer rate of an alloy during DC casting depends on the speed of casting, casting temperature, quantity & quality of cooling media. During steady-state stage of casting, the shape and dimensions of the solidification region remain constant. Shape & dimensions determine thermal gradient which is responsible for structural homogeneity in the castings. Thus, cooling water used in DC casting process plays an important role in controlling the cooling rate during the solidification process.

The aim of the present work is to determine the effect of the quality of water on heat transfer rate during direct chill casting of billets through hot top Air slip casting system and to correlate the water quality variables with SDAS.

The influence of water quality has been analyzed by many researchers. In one of the first papers on the subject, Ho Yu [1] has shown from his missile test that dissolved air, surfactant, cationic poly electrolytes & oil content, all have an adverse effect on the heat transfer rate. Langlais et al [2] have also inferred that oil content & flocculants beyond 10 & lOOOppm respectively, reduce cooling rate. The effect of composition, temperature and flow rate of cooling water on heat transfer has been studied by Grandfield et al [3] through experiments.

A. Laboratory experiment

Experimental Procedure Experiments were carried out in two stages.

A cylindrical sample of diameter 2" & length 12" was machined from the middle part of an AA6063, 7" billet. It was heated up to a predetermined temperature in a muffle furnace (equipped with a temperature controller) & soaked for 30 minutes to ensure uniform temperature distribution through out. Thereafter, the heated sample was quickly removed from the furnace and immersed in a fixed quantity of cooling water. The instantaneous temperature of the sample was obtained by a K-type thermocouple located at the axial & radial centre of the sample. A data logger

Numerous solidification studies have been developed with a view to predict microstructural parameters. Many scientists have

681

system (Yokogawa-DX2008) recorded the temperature at a sampling rate of 20 Hz.

B. Shop floor experiment

Water quality was varied by changing temperature, hardness, and oil content. Several trials were carried out by changing the value of one quality variable at a time and the corresponding cooling curves were recorded. In all the trials, constant water quantity was maintained.

AA6063, 7" diameter billets were cast using 3 different parameters as described in table II. The casting recipe of a 7" billet was chosen as base case. Case-2 and 3 conditions were achieved by fitting one 7" mold with its starting block in a 6" diameter billet casting table and again another one in an 8" casting table.

DM (demineralised) water at 30°C was used as the base case. For the purpose of the experiment, the quality of DM water was changed by I. Raising the temperature. II. Addition of salts to raise hardness III. Addition of lube oil The matrix of the different experimental conditions are shown in table I. Table I. Experimental cases Cases 1 2(basecase) 3 4 5 6 7 8 9 10

Water temperature (WT), °C 20 30 40 60 30 30 30 30 30 30

Water hardness (HD) ,mg/L 0 0 0 0 70 150 300 0 0 0

Table II. Different casting parameters Cases

1 (base case) 2 3

Oil(OL), ppm 1

Water flow rate, 1pm 78

Pouring temp, °C

126

705

38

68 130

145 102

708 706

38 38

Casting speed, mm/min

Water temp,

°C

Cooling water of hardness 130 mg/L and oil ppm 40 ppm was used for all the three cases. The surface temperature of the billet during steady state (i.e. after 1 meter of billet casting) was measured at three fixed points at a particular time. The temperature readings were also confirmed through thermographs. Pool depth was measured after steady state regime during the three cases.

0 0 0 0 0 0 0 50 100 150

Metallographic studies were conducted on mid radius regions of samples obtained from the central portion of the billets. SD AS was measured by use of Olysia Image Analyzer Software.

Results & Discussion The results obtained from experiments were utilized & analyzed in four stages. A. Correlation Between Water Quality Variables & Surface Heat Flux The cooling curves recorded by the data logger are shown in figure 1. 500.0 i

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The relatively constant temp plateau at the beginning (depicted in the cooling curves (fig 1)) can be attributed to film boiling, in which a vapor blanket insulates the surface from cooling water. It is evident from the fig 2(a) that higher water temp decreases critical heat flux value and promotes film boiling at lower surface temp. The effect of rise in water temperature from 20°C to 30°C had a very little influence on heat flux. This observation also matches with the findings of Langlais et al [2] & Grandfield et al [3].

seen in fig 2(c). In general critical heat flux decreases when oil content in the water is raised because of diffusion in an oil enriched region surrounding the growing vapor film. The rise in water temp because of insufficient cooling and seasonal variation can be seen to affect cooling rate significantly. The scale formation in the water pipe lines because of repeated recycling and inadequate water treatment contributes to increase in hardness in the casting station. Oil content rises when there is leakage into the water system. All these factors promote film boiling & diminishes heat transfer rate during DC casting thereby affecting solidification.

In a salt solution, cations and anions play a major role in formation of bubbles in a pool boiling. Cations reduce bubble coalescence whereas anions favour it thereby promoting film boiling. The reduction in critical heat flux in fig 2(b) indicates the pronounced effect of anions in the water. Cations are sometimes added in the water to promote nucleate boiling overfilmboiling.

In order to assess the extent of influence of these factors on microstructural parameter (DAS) of castings, correlations have been derived. From the experimental & model data, correlations could be formed between water variables & heat flux at various surface temperatures (fig 3).

The oil in water forms a water-oil emulsion that was found to have lowered the leidenfrost point thus promoting film boiling as I.00E+05 9.0OE+O4 8.00E+04 -| 7.00E+04 e.OOE+04 CÎ.OOE+04 g*.OOE+04 X3.QOE+G4 2.00E+04 1.00E+04 Q.00E+O0

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683

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Table III. Heat flux at different surface temp Surface temp,°C

Combined equations for HF 100 200 300 350 400

HF = 8.6e3 -128.71 *(T- 30) -3.17*HD - 6*OL HF = 4.27e4 - 340*(T - 30) - 6.1847* HD - 27.6*OL HF = 7.01 e4 - 679.43*(T - 30) - 24.237* HD - 37.03*OL HF = 7.85e4 -1038.9*(T - 30) - 45.562* HD - 108*OL HF = 7.66È4 - 2002.7*(T - 30) - 37.367* HD - 51.29*OL

The strength of the equations were tested with heat flux values predicted by lab model (from temperature measurement) in two different water trials (T= 36 & 34 °C, HD=260 & 50 mg/L , OL= 10 & 50 ppm) in fig 4. The empirical relation found to be closely matching with predictions for combination of all variables.

validate the model, center temperature of billet at three different depths in the sump was measured by dipped in thermocouple and found to be in close agreement with model predictions (fig 5(b)).

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C. SPAS & Cooling Rate:

Figure 4. Comparison of formula output with model

Secondary dendrite arm spacing (ë) which has an impact on the billet properties, is known to vary inversely with cooling rate during solidification as per formula [14, 15],

Absolute heat flux values obtained in the experiment pertains to the sample geometry and experimental conditions, which are different from actual billet casting. The geometry of the experimental set up possesses cylindrical symmetry with billet casting. The rate of change of heat flux due to change of water quality is considered here to be independent of size & process parameters.

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(2)

Where, Ά ' is a constant for a particular casting technology, alloy and size of casting. 'Tc' is the cooling rate. The exponent 'n' lies in between 0.2 to 0.4 for most of the aluminium alloys [14, 15].

The rate of change of heat flux values were obtained from the relationship of various slope values of water quality (noted in table III) and surface temperature, for each variables. These were used as correction factors to the heat flux input table of a billet model to get changed set of heat flux data for different water variables. These would be put as boundary conditions for the billet model to generate cooling profile in the billet and would predict cooling rate to be used in stage *C\

The average cooling rate in the solidification range at mid radius region was computed by the billet model for the three cases mentioned in table II. Measurements were carried out with the samples representing the corresponding three cases for SD AS. Computed Tc & measured SDAS for the corresponding cases are tabulated in table IV. Table IV. Computed cooling rate and measured SDAS

B. Development of Billet Casting Model

Case

A 2-D transient thermal model based on finite element method was developed by using ANSYS software for the 7" (AA6063) billet casting. The casting parameters were as per the base case shown in table II. The model was simulated for case 2 and 3 of table II for validation. Material property was taken from Ref [13] and fluid flow aspect was not considered. Equivalent specific heat approach has been adopted.

1 2 3

The measured pool depth (average of 5 measurements during steady state) was found to be 64 mm and matched well with the predicted solidification profile (figure 5 (a)). In order to further

Model predicted Tc at mid radius point, (°C/ sec) 0.75 1.05 1.5

SDAS measurement 1 (Average of 10 measurements) μπι 20 18 16

The correlation derived from the measurement values through curve fitting (Fig.6) is expressed by the equation,

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Figure 6. Cooling rate Vs SDAS The micrograph of the two 7" billets (mid radius region) are shown in figure 8. tst«!

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(4)

This equation was reliable for standard casting recipe of 7" billet as mentioned in case-1 of table I. The equation was verified during five shop floor castings with different water quality parameters. The calculated SDAS was found to be comparable to the values obtained from metallographic measurements. The comparison is shown in table V.

Water Temp °C

Hardness Mg/1

36 34 30 38 40

260 50 60 130 120

Oil ppm

10 50 40 40 40

SDAS Calculation μπι 19.05 17.8 17.6 18.13 18.25

SDAS (Average of 10 measurements) μπι 20 19 18 19 21

A 16% variation in SDAS could be observed during the trials in the shop floor. This would have an impact on homogenization as well as extrudability of billets. Conclusion The effects of water temperature, hardness, oil content of the cooling water used in a DC casting process on the dendrite arm spacing of billets was investigated through experiment and modeling. An empirical relation was derived. This correlation has

Table V. Comparison of calculated SDAS with measurements

685

been tested in a plant scale DC casting station in order to closely predict dendrite arm spacing. The empirical relation serves as a useful tool for fine tuning casting recipe for different conditions of water quality. SDAS at mid radius of billet is specified by billet manufacturers to indicate the quality. Low and consistent SDAS helps the extruders to run the press with uniform parameters so as to maximize productivity. The variability of SDAS because of different cooling water quality during different seasons like summer, winter or rainy seasons can be judiciously taken care of by this tool.

SDAS increases with increase in water temperature. ii. Hardness and oil content in cooling water also increases SDAS in DC casting billets, but the effect is minimal. Acknowledgments The authors are grateful to Nalco for giving permission to publish this paper. The authors would like to thank Laboratory and Billet Casting personnel for their help and support during the experiment.

From the results of the study following conclusion can be made: i. The water temperature has a very strong influence on the cooling rate of the casting and has maximum impact on SDAS. References

10. M. Bamberger, B.Z. Weiss, and M.M. Stupel, " Heat Flow and Dendritre arm spacing in chill-cast Al-Si alloys", Material Science and Technology, January 1987, Vol. 3, 49-56.

1. Ho Yu, "The Effect of Cooling Water Quality on Aluminium Ingot Casting", Light Metals 1985, TMS,1985, 1331-1347. 2. J. Langlais et al, "Measuring the Heat Extraction Capacity of DC Casting Cooling Water", Light Metals 1995, J.W.Evans eds, TMS 979-986.

11. M. Easton, C. Davidson, and D. St John, "Effect of Alloy Composition on the Dendrite Arm Spacing of Multicomponent Aluminum Alloys", Metallurgical and Materials Transaction A, 41 A, June 2010, 1528-1538.

3. J.F. Grandfield, H. Hoadley & S. Instone, " Water Cooling in Direct Chill Casting : Part-1, Briling Theory and control" Light Metals 1997, R.Huglen ed, TMS-691-699.

12. D.Li, M.A. Wells & G. Lockhant," Effect of surface on morphology on boiling water heat transfer during secondary cooling of the DC casting process", Light Metals 2001, J.L.Anjier ed, TMS,2001,865-871.

4. K. Ho and R.D. Pehlke, "Metal- Mold Interfacial Heat Transfer", Metallurgical and Materials Transaction B, 16B, No3, (1985), 585-594.

13. T.S. El. Raghy et al "Modelling of the Transient & Steady State Periods during Aluminium DC Casting" , Light Metals 7995, J.Evansed,TMS, 1995, 925-929.

5. T. S. Prasanna and K. Narayan Prabhu, "Heat Flux Transients at the Casting/Chill Interface During Solidification of Aluminum Base Alloys", Metallurgical and Materials Transaction B, 22B, No-5, (1991), 717-727

14. D.G. Eskin, Physical Metallurgy of Direct Chill Casting of Aluminium Alloys, (New York, CRC press, Taylor & Francis Group, 2008), 19-24.

6. C.A. Muojekwu, I.V. Samarasekera, and J.K. Brimacombe, "Heat Transfer and Microstructure during the Early Stages of Metal Solidification," Metallurgical and Materials Transaction B, 26B,April (1995), 361-382.

15. M.C Flemings, Solidification Processing (Mc Graw-Hill, USA, 1974), 146-150.

7. D. Bouchard & J.S .Kirkaldy, "Prediction of DAS In Unsteady & Steady State Heat Flow of Unidirectionally Solidified Binary Alloys", Metallurgical and Materials Transaction B, 28B,Aug (1997), 651-663. 8. C. Devdas, I. Musulin & O. Celliers, "Prediction of Microsrtucture of DC Cast 6063 Billets & its Effect on Extrusion Process", in 5 th International Aluminium Extrusion Technology Seminar-1992, Aluminium Association and Aluminium Extruders Council 1992, 121-128. 9. Ivan. Todaro et al., "Effects of Cooling Rate on Microstructure in En-Ac43000 Gravity Castings and Related T6 Mechanical Properties", Light Metals 2010, J. A. Johnson ed., TMS, 2010, 619-624.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

MOULD WALL HEAT FLOW MECHANISM IN A DC CASTING MOULD A. Prasad1,1.F.Bainbridge1. *CAST CRC, University of Queensland, Brisbane, QLD 4072, Australia Keywords: DC casting, Mould-wall, Heat Flow, Radiation Knowledge of the actual heat flow from metal to the mould wall and the factors that influence this are thought to be important for both the design and operation of a DC casting mould [9, 10]. Cast product quality has been directly linked to the amount of heat removed through the mould wall [4, 11]. Moreover, accurate simulation of the process by mathematical modeling is dependent upon reliable data for this factor.

Abstract Experiments have been performed to study the effect of the mode of heat transfer on the heat flow in the wall of a DC casting mould. Billet casting uses graphite as inner lining material within the mould-wall. Graphite being a black-body, provides a possible pathway for radiation to play a role in the mould-wall heat transfer. As such, experiments were performed with a graphite probe on the experimental apparatus presented at TMS 2010. An overview of these results will be presented together with the implications of the results for the DC casting process.

CAST CRC initiated research in this direction to endeavor to quantify some of the important factors that can possibly influence the mould-wall heat flow in a DC casting. However, the measurement of the mould wall heat flow under controlled conditions during an actual DC cast is extremely difficult and costly. Determination of the effect of individual factors that may have an influence on the mould wall heat flow is not possible under such conditions. Hence, at CAST, the focus has been to develop a laboratory simulation of an operating mould and to use this to accurately measure the heat transfer coefficient between the molten metal and the simulated mould wall under a range of controlled conditions known to occur within a DC casting mould.

Introduction The practice of DC casting of aluminium and aluminium alloys requires the removal of heat from the surface of the solidifying metal at two points within the system: the sub-mould cooling impingement area (often referred to as the primary cooling zone) and the mould wall area. Heat flow in the sub-mould cooling area has been reasonably well studied and confirmed values are available for the heat transfer coefficient for a range of casting conditions [1-4]. Several estimates of the mould wall heat transfer have been made, together with some attempts to measure the heat flow during an actual cast. The amount of heat flow is characterized by the heat transfer coefficient and Table I summarizes the current data available.

The essentials of the experimental apparatus were described in the previous work [12]. It is also briefly described here under "Equipment design and operation" section. Initial experiments were done with an AA601 probe on two sample types - 99.85% Al sample (molten metal at -700 °C) and stainless steel samples (solid sample at 700 °C). Note that the AA601 alloy probe is the same material from which industrial moulds are made. The first set of experiments was performed by passing dry air @ 1.0 1pm between the probe and the sample. Further experiments were performed with different gases at the same flow rate namely, nitrogen, argon, and carbon-dioxide with stainless steel samples. The reason for using a stainless steel sample was to mitigate the difficulty of obtaining a flat sample surface with molten metal samples and hence a more accurate sample/mould gap [12]. The experimental results obtained for experiments with different gases on the stainless steel samples are repeated here in Figure 1 for convenience.

Table I. Published values for mould wall heat flow Parameter Heat transfer coefficient molten metal to mould Heat transfer coefficient air gap formed in mould Heatflux- molten metal to mould wallHot top casting Air pressurized casting Rolling block castingSumitomo Rate of heat removal through the mould (for 152mm dia. 6063 billet cast at 101,152 & 228 mm/min.)

Value 2000 3000 2400 1000 1000-2000 650-1600 200 250-350

Units Wm^K1

Wm^K1

Ref. [1] [2] [3] [4] [5] [6] [1] [2]

1270 420 -800

kWm2

[4] [4] [7]

7,600 (9.1%) 11,000 (9.0%) 18,000 (10.4%)

W (%of total)

[8]

The results clearly showed a linear correlation between the heat transfer coefficient and the gas thermal conductivity. Thus the conduction mechanism was suggested to be dominant. However, since a billet mould for a DC casting has a graphite lining, graphite being a black-body (emissivity = 1 [13]), potentially provides a pathway for radiation to play a role in the heat transfer. From the previous experiments with an AA601 alloy probe, the effect of radiation in the metal-probe gap heat transfer was not clear. To study the effect of radiation, further experiments were planned with a graphite probe with the same design features as the AA601 probe.

687

the sample surface and the probe. This feature makes this steadystate experiment unique in that the other experiments reported in the literature only estimate transient values of hg [5,14], whereas our values were for steady-state conditions. The effect of the metal-probe gap on hg for the present work is therefore unequivocal.

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The experimental procedure details were the same as in the previous experiments [12]. Thus the sample temperature was kept constant at -700 C, the gas flow rate was fixed at -1.0 1pm etc. The choice of sample type used for the current experiments is described later. Also, the data gathering and recording of different parameters viz., ΔΤ1? ΔΤ2, L! etc. mentioned in equations 1 and 2, were performed in the same manner. The probe features were also kept the same as before (Figure 2). A Tokai grade G384 graphite was used for the probe with a thermal conductivity of 128 Wm^K" *. However, because of the brittleness of graphite, the probe broke at several thin sectioned regions during installation. It was decided to redesign a slightly bigger probe with thicker sections for structural rigidity. The essential features of embedded thermocouples, water cavity etc., from the original probe was kept the same.

Dry Air

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50

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Gas Thermal Conductivity x 10 , [Wm^K ]

Figure 1. Heat transfer coefficient versus mould gas thermal conductivity for a constant 0.5 mm gap [12]. Sample = stainless steel, probe material = AA601 Equipment design and operation The basic concepts of the experiments are the same as reported earlier [12]. The experiment attempts to simulate the heat flow during an actual DC casting operation. During DC casting, as molten metal is poured into the mould, a liquid meniscus is formed which has a variable gap between it and the mould wall. Concurrently, the molten metal at higher temperature loses heat through this gap to the mould. This heat flow translates to a temperature gradient within the mould. This situation can be simulated in a laboratory experiment by using a thermal probe as analogous to a mould and subsequently measuring the temperature gradient within the probe. This was explained in the previous publication [12].

Figure 2 shows the schematic of the previous probe. For the present set of experiments, the flange section of the probe attaching to the top plate was made thicker since this section was the most vulnerable. The region in the probe where the thermocouples were embedded was also lengthened slightly to enable embedding 3 thermocouples instead of 2 thermocouples previously. This allowed a check on the thermocouple data since the temperature gradient between thermocouples 1-2 should be the same as that between thermocouples 2-3. The new probe was therefore slightly longer in length (from the top plate to the face) compared to the earlier version. The rest of the design features were kept the same.

The experimental apparatus consists of a water cooled probe (AA601 or graphite) with thermocouples embedded facing a sample (pool of molten 99.85%Al or solid steel sample - both at 700 C). The gap between the probe and the sample is continuously changed and the resulting heat flow through the probe is captured by the thermocouples which record the proportional temperature gradient within the probe.

Water ly/cooling^Nk I 1 Top plate"

Knowing the temperature difference (ΔΤ2) across the probe (from the sample facing front face to the water cooled rear section, i.e. distance L2) and the temperature difference between the front face of the probe and the sample (ΔΤ^, the heat transfer coefficient (hg) can be calculated using Fourier's law, by first determining the quantity of heat, Q, flowing through the probe.

20mm

(1)

β = ^(ÄΓ2)

Figure 2. Drawing of the probe showing the positions of the three thermocouples and water cooling.

This is then converted to a heat flux ( q = — ) and then

There were some changes in the experimental plans which were deemed important. Firstly, the previous results were quoted for stainless steel samples. However, it was considered important to perform experiments with an aluminum based sample, simply because of its direct relevance to the actual DC casting situation. Secondly, since the probe dimensions were changed, albeit slightly, it was considered important to perform experiments with a similarly modified AA601 probe so that a direct comparison between probe material types could be made. Thus experiments

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=

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is used to determine the heat transfer coefficient, hg. In these equations, k is the thermal conductivity of the mould material and A is the mould surface area through which heat is transferred. In the reported experimental work, these equations were used to determine hg across a fixed (but adjustable) distance between a

688

were performed with molten 99.85-wt%Al in dry air and argon with both AA601 and graphite probes to the modified design. Results The results for each particular set of experimental conditions (probe material type and mould gas) were plotted and the trend line equations then used to calculate hg at a particular mould gap. It was mentioned in the previous work that the heat transfer coefficient varied non-linearly with the gap, but had a linear corelation with the inverse of the gap [12]. As the inverse relationship (hg vs 1/gap) is a straight line this was used for the data analysis. The equation thus obtained was used to calculate the heat transfer coefficient, hg, at 0.5mm probe-metal gap. The following figures show the plots between hg vs 1/gap from which equations and R2 values were generated using linear regression analysis. These equations were then used to estimate the hg values at 0.5mm gap (reported in Table Π).

0.5

120 100 *

80 -

1

60 ~ M

20 U "

()

Jt

O Dry air + Argon

For both the figures (3 and 4) there is a clear difference in slopes for the two gases tested. This replicates the effect of gas type as seen in Figure 1 for the steel samples with the older probe. Furthermore, considering the effect of probe material types, the yaxis scales are different for the two figures. The highest value in the scale shows that the graphite probe gives slightly lesser values of hg as compared to the AA601 probe. This difference is seen more clearly in Table II. Table II compares the results from the AA601 and graphite probe for the two gases. The average values of hg at a mould gap of 0.5mm are quoted with all the data points taken into the calculations. The hg values were obtained by using the linear least square regression lines for each set of data points i.e. [AA601 probe, 99.85-wt%Al, dry air] as stated before. Also included in the table is the percentage difference in conductivity between the two gas types and the probe material types. The gas conductivity values were evaluated at 600K, which is approximately the average of the sample temperature (700 C = 975 K) and the probe surface temperature (-40 C maximum = 31 OK).

^ W^

^ L

ί* t***^ 1

0.5

i

1

i

1.5

1/Gap, [mm1]

1

2

1.5

Figure 4. Plot showing the heat transfer coefficient as a function of inverse of mould gap - Graphite probe, 99.85%Al.

Figure 3 shows the results from experiments with molten 99.85wt%Al sample at 700 C facing a AA601 alloy probe in dry air and argon. The figure plots the heat transfer coefficient as a function of the inverse of the gap. The heat transfer coefficient between the molten metal and the mould wall in a DC casting mould was found to be dependent upon the size of the gap between the molten metal and the mould wall [12]. The same result is seen here, which is a validation for the modified probe working correctly (similar R2 values approaching 1 were obtained with both the older and the modified version of the probe). 140 "

1 1/Gap, [mm 1 ]

Table II. Heat transfer coefficient, hg for different combinations of probe material and gas types.

Λ

2.5

Figure 3. Plot showing the heat transfer coefficient as a function of inverse of mould gap - AA601 probe, 99.85%Al. Likewise, Figure 4 shows the same results for the 99.85-wt% Al in dry air and argon respectively with the graphite probe. A similar trend is seen as with the AA601 probe. In each of the sets of plots in Figures 3-4, there are approximately 5000 data points. A single run produces approximately 2250 data points of heat transfer coefficient vs gap (or the inverse of the gap). And for each combination of probe material type and gas type, at least 6 runs were performed. Therefore, the plots show only a fraction of the data points actually obtained. A tight band of data points for a given set of experimental conditions shows the consistency of the experiments. Note that all the successful experiments showed a similar plot with closely bound data points within a given set.

Gas type / Probe material

hg, [Win 2K_1] (at 0.5mm) AA601

Graphite

Dry air

148.3+18.9

Ar % diff. h s (gas)

88.3+9.8 40.5%

%

% diff Gas cond., kg

% diff Probe cond. k

125±10.2

diff. hg (probe mtl.) 15.7%

33.0%

19.5%

86.9+8.0

1.5%

33.0%

19.5%

30.5%

-

-

-

The effect of radiation may be manifested in two ways - one due to the radiative properties (absorption, reflection etc.) of the mould material (probe material in our case) and the other due to the emissivity of the molten metal (sample in our case). Therefore, the results from the four sets of experiments were also compared with the steel sample results in Figure 1. Note that Figure 1 was based on an earlier version of the probe, and also, had data for four different gas types. Nevertheless, a comparison is valid for two reasons. Firstly, the modified probe gives similar trends as the older probe, therefore the working of the two probes were not

689

encountered within a DC casting mould. These runs showed that the hg (at 0.5mm) values were not affected within this range of flow rates. This suggests that the convection effects (forced or natural) were negligible. Hence any difference in the hg values between various combinations of sample type, gas type or probe type must have been a result of major contributions from conductivity and radiation mechanisms only. The effect of gas conductivity on hg values has already been demonstrated.

vastly different. Moreover, only a small change was made in the new probe. Secondly, given that the old and the new probe yield similar results, the comparison provides an opportunity to compare two different sample types, steel and aluminum, whose oxides are very different. Thus the comparison provides more information to compare the radiation component of heat transfer via differences in the emissivity of the two oxides. Figure 5 shows this comparison. A theoretical set of data points are also plotted. The theoretical data points were based on the HoPehlke [15] prediction of hg for steady state conditions. Their model is simply given by: hg = kg/gap, where kg is the gas conductivity. For our case, the gap was fixed at 0.5mm.

For a given gas type, when the results between different sample types are compared, the following comparisons can be made. Within the Al-based samples, the results can be compared between the two different types of probe. For the two probes, the hg values were similar, particularly for Ar. For this case, as discussed earlier, it transpires that the gas conductivity was the dominant mechanism. This suggests that the material of the probe did not play a major role in the heat transfer. In other words, the absorptivity and reflectivity properties of the mould material may be of minor consequence in a DC casting operation.

Discussion and Conclusions From Figures 3-4 and Table II it is clear that there was a difference in the amount of heat flow across the gap when the gas type and/or the probe material was changed. In Figure 3, as the gas type was changed, the slope of the data points changed, showing a smaller hg value for Ar. The same trend was seen in Figure 4. There seems to be a larger change in the slope for the AA601 probe as compared to the graphite probe. However, from Table II it is clear that for either of the probe types, the change in hg values due to a change in gas type was similar (30 - 40%). This was close to the difference in the gas conductivity between dry air and Ar (33%). Note that the percent difference was based on the average hg values. 190

J

»SS-AAoOl

«99.85%A1-AA601

170 1 A99.85%A1-Graphite

Ý πï i ^

90

50 -i 20

XTheoretical

I

70 i

X 25

The second comparison is that between the steel and Al sample for the AA 601 probe for the two different gases. It can be seen that for a given gas type, the steel samples had higher values than the Al samples. Since the conductivity is the same for the gas type, by elimination, the difference must be attributed to the radiation. Literature [16] suggests that the emissivity of an oxide layer between stainless steel and molten aluminium will be different. However, this is a function of the temperature and the degree of oxidation. A higher degree of oxidation in steel would result in a higher emissivity of the oxide layer. In our experiments, the molten metal was skimmed prior to the start of each run, but the steel sample surface was left untouched between runs. Note that the literature search on emissivity was performed much later and therefore at the time of performing steel runs, the authors were unaware of the effect of oxidation on the emissivity of oxide of the steel. The literature [16] also suggests that beyond -800K, highly oxidized steel would have higher emissivity than aluminium oxide. Our experiments were performed at -975K, which is well over the 800K limit suggested in the literature.

30

χ

i 35

40

3

XX 1 45

-1

Gas thermal conductivity x 10 , [Wm^K ]

. 50

It is interesting to note that for dry air, the hg values between Al and steel were not vastly different. However, for Ar, this difference increased markedly. While difference in emissivities can be attributed to explaining the difference in hg values between Al and steel for dry air, differences in emissivities fail to explain the large difference in hg values for Ar.

Figure 5. Comparison of hg values versus mould gas thermal conductivity for a constant 0.5 mm gap with different probe/sample combinations.

The overall mechanism seems to be gas conduction dominated. On the other hand, radiation can affect the gas molecules by transferring energy to the molecules, thereby heating them. This in turn would affect the conduction since conduction through a gaseous medium relies on molecular movement. Since both the probe face and the sample had a finite area (φ = 20mm each), the view factor could be an issue as well. The view factor will change with a change in gap (increases as the metal-probe gap decreases). Thus at smaller gaps the effect of radiation on Argon's conductivity could be higher than that for dry air. For our experiments with dry air and Ar, such effects of radiation on conduction have not been determined.

On the other hand, the effect of difference in probe material was not as straightforward. For example, for dry air, the difference between hg values from AA601 and graphite probe was -16%, but it dropped to 1.6% for the Ar gas. Note that the difference in the probe material conductivity was -20%. In other words, the difference in probe material conductivity had been overridden by the gas type. Thus the gas conductivity seems to have been the more dominant parameter. Effect of radiation: Before discussing the effect of radiation, it should be mentioned that there had been a limited number of experiments on the steel sample with different gas flow rates (0-3 1pm). The flow rates were so chosen as to simulate the range of flow rates normally

This would require an in-depth study, perhaps a modeling effort to quantify such an effect of radiation. Also, a more rigorous quantitative analysis can be performed by doing radiation-focused

690

experiments in a high vacuum. In such a situation, with absence of gas molecules, the conduction and convection components will be negligible with radiation the major mechanism. However, a high vacuum is not easy to obtain and this would also require major changes in the experimental set-up. As such, the vacuum based experiments have not been performed at this time. Thus at this stage it is only possible to conclude that radiation has a definite effect on the heat transfer via differences in emissivities of the sample type.

2. D. Mortensen, "Mathematical model of the heat and fluid flows in direct-chill casting of aluminum sheet ingots and billets", Metallurgical Transactions B, 30B (1999), 119-133. 3. J. M. Drezet et al, "Determination of thermophysical properties and boundary conditions direct-chill cast aluminium alloys using inverse methods", Metallurgical Transactions A, 31A, (2000) 1627-1634. 4. J. F. Grandfield, and P. T. McGlade, "DC casting of aluminium: Process behaviour and technology", Materials Forum, 20 (1996), 29-51.

Implications for DC casting: The aluminium sample results with different probe types showed a small difference in hg, which was not large enough to cause a major difference in the heat extraction through the mould wall. Recall that the majority of the heat extraction in DC casting is through the sub-mould water jets. However, the effect of the gas conductivity has been shown to be quite pronounced. Hence it is suggested that in the situation where vapors from the oil, steam from sub-mould cooling etc. are present in the metal-mould gap, the heat extraction through the mould gap will be affected. Finally, the type of metal cast may or may not have an effect depending upon the emissivity of the oxide skin formed. The effect of oxide skin on radiation emissivity also requires in-depth study in terms of the variables, namely: amount of oxidation, the chemical composition, crystal structure of the oxide etc. In another aspect of the present work, the hg for different aluminium alloys in dry air had been determined. In cases where high levels of Mg and or Zn are present, e.g., 5xxx or 7xxx series alloys, a thicker, more adherent oxide film was formed, along with some difference in the hg values. The results of this work are yet to be reported.

5. K. Ho, and R. D. Pehke, "Metal mould interfacial heat transfer", Metallurgical Transactions B, 16B (1985), 585-594. 6. M. Trovant, and S. Argyropoulos, "A Technique for the Estimation of Instantaneous Heat Transfer at the Mould/Metal Interface During Casting", ed. R. Huglen (Orlando, Florida: Light Metals, TMS, 1997), 927-931. 7. N. Muto, N. Hayashi, T. Uno, Sumitomo Light Metals Technical Reports 37 (1996) 180-184. 8. D. C. Weckman, and P. Niessen, "Numerical simulation of the DC continuous casting process including nucleate boiling heat transfer", Metallurgical Transactions B, 13B (1982), 593-602. 9. P. W. Baker, J. F. Grandfield, "Mould Wall Heat Transfer in Air-Assisted DC Casting", Solidification Processing, (London UK: 3rd Conference, The Institute of Metals, 1987), 257-260. 10. M. Ekenes, W. S. Peterson, "Visual Observations Inside an Airslip Mould During Casting", ed. C. Bickert (Anaheim, California: Light Metals, TMS, 1990), 957 - 961.

In summary, a new technique has been developed that enables the measurement of heat transfer across a simulated metal-mould gap for a DC casting. In the present work, this technique has been used to study the effect of radiation on the heat transfer across the gap. This was done by performing experiments on Al-samples under two different gases with two different probes - AA 601 and graphite. It was found that radiation had a small effect on the heat transfer coefficient via emissivity of the sample. The absorption/reflection of the probe body did not have any noticeable effect. Overall, the effect of radiation resulted in a small difference. However, it is suggested that this difference is not sufficient to cause a significant change in mould wall heat transfer in an actual DC casting. To accurately quantify the effect of radiation, further work (experimental/theoretical or both) is required.

11. S. Instone, W. Schneider, and M. Langen, 'Improved VDC Billet Casting Mould for Al-Sn Alloys", ed. P. N. Crepeau (San Diego, California: Light Metals, TMS, 2003), 725-731. 12. A. Prasad, J. Taylor, I. Bainbridge, The Measurement of Heat Flow within a DC Casting Mould, ed. J. A. Jensen, (Seattle, WA: Light Metals, TMS, 2010), 765-770. 13. Michael F. Modest, Radiative Heat Transfer (Academic Press, 2003), 746-752. 14. M. Trovant, and S. Argyropoulos, "Finding Boundary Conditions: A Coupling Strategy for the Modeling of Metal Casting Processes: Parts I and II", Metallurgical Transactions B, 31B (2000), 75-96.

Acknowledgements CAST CRC was established under, and is funded in part by the Australian Federal Government's Cooperative Research Centre scheme.

15. K. Ho, and R. D. Pehlke, "Transient methods for determination of metal-mould interfacial heat transfer", AFS Transactions, 92 (1984), 587-598.

References

16. Frank P. Incropera, and DavidP.Dewitt, Fundamentals of Heat and Mass Transfer (John Wiley and Sons, 1990), 722-723.

1. A. Sabau et al, "Heat Transfer Boundary Conditions for the Numerical Simulation of the DC Casting Process", ed. A.T.Taberaux (Charlotte, NC: Light Metals, TMS, 2004), pp.667-672.

691

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Productivity Improvements at Direct Chill Casting unit in Aluminium Bahrain (ALBA) , Aref Ahmed Mohd Noor1, Sukanta Chatterjee2A. Rasool Ahmed1 1 Casthouse, Aluminium Bahrain B.S.C, (e), Kingdom of Bahrain 2 R & D, Aluminium Bahrain B.S.C, (e), Kingdom of Bahrain Keywords: Casthouse, Rolling Ingot, DC, OEE, Levers, TAC, SOP, Steady State, KPI Casthouse 1 is left with only production of Rolling Ingot at DC1. This resulted in increase in cost of production at DC1, Casthouse 1 due to high fixed cost to maintain all the infrastructural facilities. Around this time the price of Aluminium started to dip as a result of worldwide financial crisis . This forced ALBA to take a stock of situation and optimize cost of production.

Abstract ALBA has been producing Rolling Ingots utilizing two DC casting units viz., DC1 & DC6. Out of total annual production of 140,000 MT of Rolling Ingot, DC6 production was 100,000 MT (lxxx, 3xxx & 8xxx series alloys) while DC1 production was 40,000 MT of 5xxx series alloy. Due to higher cost of production & low productivity at DC1, ALBA management decided to close down DC1 as a part of restructuring exercise and maximizing production at DC6 to meet customers demand. An extensive study was carried out towards improving productivity. Four levers of improvement which has the maximum impact were identified e.g., furnace preparation practices, preparatory activities between casts, increasing casting speed etc. Various action plans were drawn up & were implemented. These actions were closely monitored by DC operators themselves by using trend charts. As a result a 30% improvement in productivity was achieved at DC6 on a consistent basis.

It was found that cost of Rolling Ingot production per ton at DC1 is almost 2 times than that at DC6 which makes DC1 unviable. After lot of deliberations ALBA executive management took a decision to close down operations at DC1. It was expected that by doing this total Rolling Ingot production capacity will be reduced by almost 30% as per existing levels of production. ALBA took a decision look into feasibilities of improving productivities at DC6. A core team comprising of representatives from operation, maintenance, technical was formed to carry out this study by applying principles of lean manufacturing process [1].

Introduction Aluminium Bahrain (ALBA) started its operation in the year 1971. The first Casthouse was commissioned in the same year with the production of Rolling Ingots (DC1) & Tee Ingots. Subsequently over the year Casthouse 1 expanded the operation by adding facilities for casting Extrusion Ingots. During 1992 ALBA went for major expansion by adding Polline 4 which necessitated setting up of additional casting facility. As a result Casthouse 2 was commissioned with Standard Ingot Casting Lines (PI020 & Foundry grade) and Rolling Ingot Casting facility as DC6. ALBA has been producing Rolling Ingots from 2 DC units viz., DC1 at Casthouse 1 and DC6 at Casthouse 2. Over the years Rolling Slab casting facility at ALBA underwent many upgrades like top control to bottom control system for metal distribution, change over from 420 mm to 460 mm thickness slabs along with increase in furnace capacities. Total Rolling Ingot production capacity at ALBA stood at 140,000 MT per annum.

Figure 1 DC6 Casting Unit with automatic metal level control system

Background ALBA has been producing Rolling Ingots of various alloy types, viz., lxxx, 3xxx, 5xxx & 8xxx between 2 DCs, viz., DC1 & DC6. As a part of convenience & logistics DC1 has been producing 5xxx series alloys while DC6 has been producing lxxx, 3xxx & 8xxx series alloys. Total annual production of Rolling Ingot is around 140,000 MT. DC1 has been producing 40,000 MT while DC6 has been producing 100,000 MT per annum. During the year 2005 ALBA went for another major expansion by commissioning Pot Line 5 and Casthouse 3. New Casthouse 3 was commissioned with state of the art Extrusion Ingot casting facility. This expansion has made Extrusion Ingot casting facility at Casthouse 1 redundant. Once the operations at Casthouse 3 stabilized, total production at Casthouse 1 reduced drastically and

Figure 2 DC6 Casting Unit with 460 mm thickness Rolling Ingots

693

Study for Productivity Improvement

With the closure of DC1 entire 5xxx series alloy production is required to be produced at DC6. This would necessitate additional furnace washes after each campaign which has an adverse impact on the productivity. Addition of casting 5xxx series alloy in DC6 range of alloys posed a challenge for the team.

To get maximum benefit out of the study, an external consultant was appointed give exposure to techniques of lean management. Interactive training sessions were organized for all the supervisors & operators to buy-in the lean concept.

The present study was initiated in October 2009 and the production figures of year 2008 was taken as a reference. However it was made adequately clear to the team that although productivity improvement is the prime objective of this exercise, one cannot compromise on safety & quality.

The core team along with supporting members studied all the activities / processes related to production of Rolling Ingot at DC6, Metal receiptfromPot lines Pouring metal into furnaces Furnace batch preparation DC preparation In between cast activities Actual casting process Waiting time before ingots are ready for discharge Drying of ingots Time to discharge ingots Station change activities Net yield per cast

i î

ACD

CFF

w

Initially OEE (Overall Equipment Effectiveness) study was carried out for DC6 operations. This involved contribution of the following factors, Planned maintenance and shutdown Pit cleaning Machine breakdown Furnace washes for alloy changes Delays due to operational & technical reasons Preparation time between casts Station (size) change time Casting time Time loss due to abort casts DC utilization for maximum number of ingots Scrap & defects (planned & unplanned) After a thorough measurement & analysis of the above the factors which can contribute to maximize the OEE were identified.

DC6 DC Preparation time

TKR

Furnace Preparation Size change ind. Wash

FURNACES Figure 3 DC6 Layout (Schematic)

34% WΚΚΚΚΚΚΚΚΚΚΚΚ 1 2 % WΚΚΚΚΚΚΚΚΚΚ 10%

Casting speed losses m M, Shutdown, Pit dean

As per ALBA's definition lxxx & 8xxx series alloys are considered to be soft alloys while 3xxx & 5xxx series alloys are considered as hard alloys.

DC utilization

6%

M i 6% i 4X

Breakdowns I l i 4% Defects

Alloy wise Production

i 4%

Abort casts 1 1% OEE

WΚΚΚΚΚΚΚΚΚΚΚΚΚΚΚΚΚΚΚΚ 1 9 %

Figure 5 DC6 OEE (%) Analysis based on year 2008 production [3]

Based on the study following 4 major levers were identified where there potential to rninimise losses and improve OEE. 1. 2. 3. 4.

Figure 4 Alloy wise Production trend of Rolling Ingot [2]

694

DC Preparation time Furnace preparation PM, Shutdown, Pit Clean Casting speed losses

on action list and suitable corrective actions for these deviations were derived.

DC preparation time Activities from stopping the platen at the end of cast until starting of the next cast were mapped, as shown below, • • • • • • • • • •

Furnace Preparation

Shift Casting launder to parking position Cooling the ingots for complete solidification Tilting mouldframeup Drying the ingot head Discharging ingotsfromDC pit Tilting mouldframedown Cleaning the mould Engaging starting head into mould Drying starting head including water drain Shift casting launder to casting position

Following activities of furnace preparation were looked at for idea generation, Crucible pouring Alloying Stirring Sampling Skimming Settling

Brainstorming sessions were carried out with all the casting crew to generate ideas for minimizing preparation time. Going through this exercise it could be established that there is potential to improve DC preparation time by 40% by doing the following, 1. 2. 3. 4. 5. 6.

During brainstorming sessions some of the ideas were, 1.

Standardizing cooling time ingot size wise Use 2 operators to dry ingot head Establishing & marking reference points on the ground as well as on the crane for hoist positions for discharging ingots at various locations Follow the best path from DC to ingot laydown conveyor for fastest turn around time Coaching of all the operators by most skilled / senior operators Carrying out several activities in parallel like moving starting head upward while discharging last ingot from the DC pit

2. 3. 4. 5. 6.

Changes in sequences of crucible pouring & alloy and other cold charge addition to furnace to ensure faster melting Doing some activities in parallel Ensuring liquid metal availability as & when required by close communication with TAC plant Focus on getting furnace "on spec." in the 1st attempt During shutdown of 1 of the 3 furnaces only soft alloys to be planned to lower no. of attempts required for on specification metal Train / retrain operators for "best practice" procedures

The saving potential is envisaged to be 25%. The major emphasis has been to prepare the furnace in 1st attempt.

Subsequently all the ideas generated were validated in the shop floor and were integrated into the Standard Work Instructions for DC6. The implementation was easy since most of the ideas were generated by shopfloorpersonnel only.

PM, Shutdown & Pit Clean Maintenance activities play a vital role in productivity of the equipment. The objective for the team was improve maintenance performance without compromising productivity of DC6. During the brainstorming exercise 5 areas were identified. 1. 2. 3. 4. 5.

Planning & Scheduling Improved Execution Equipment Strategy Autonomous Maintenance Spare Parts Management

The action plan evolved out of this brain storming activities are as follows, Preprocess all the requested jobs to screen real needs (Gate Keeping) Define fixed accountability for job execution Improved workplace organization Improved remote work place organization Create SOPs for most recurrent complex jobs Equipment failure analysis for corrective/preventive action

Figure 6 Standard Work Instructions displayed in the work area During implementation period shift wise performances were closely being monitored by operators with the help of trend charts. Performance review meeting was being held at the end of each shift. Any deviations from target was discussed & recorded

695

• • •

Execution of simple clean & check activities by production operators and develop SOP Optimization of spare store room layout to reduce handling time Carrying out some activities in parallel like overhauling actuators/sensors in between casts, pit cleaning during PM etc.

φ*&>«km: Tim« men to pmp*m tm DC im easting. Startingfrom*w* «t » ea*t ~ p§t*t*a mp ~ tHI statks« ί* being tmó? fer cmtrng D*t» mates: »«nior operator m#Δ*yram*nt

8y whsm: Soaier Qp»**«or fttrηafcney s>* «a&üteüo«: Bv«ty «tiift Tara«! So Mifls, Toi*r*nc»; *S«*î«*

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Casting Speed Losses

Senior 0$»r*to*

Casting productivity is one of the area where work was already initiated from the beginning of year 2009. The existing steady state casting speed for various alloys and sizes were reviewed. The following aspects were looked into during this exercise,

3. 4. 5. 6. 7.

OmfcO*t*C»onf w « » : IMi«8 m * waîc* measofmg tfc* j*r#pars8o8 mm ir<m pimn #m tJ8 »μfcfioo te fwι? fef mm ami

Timing of ©enaction; 6««ί of shift

Based on the above it was possible to change over from bi weekly PM schedule to four weekly schedule. As a result it was found that about 200 hr of production downtime could be eliminated.

1. 2.

L«v*i e* wNen « is U**Τ; %& cm,

Figure 7 Procedure for monitoring KPI as displayed in shop floor

Maximum design speed for the type mould Maximum metal flow rate for the existing down spout without causing turbulence Maximum metal rate for the existing distribution bag Dimensional tolerances for the output ingot viz., concavity Grain size (macro) for the ingots after etching Shell zone depth for the ingot Any other obvious surface defect

Individual shifts were conducting performance dialog at the end of every shift. During initial period of implementation these meetings were being moderated by members of the implementation core team and subsequently these meeting were handled independently. During the discussions any deviations from the targets were highlighted, possible reasons were identified and corrective actions were recorded in the board for subsequent follow up. These activities generated lot of enthusiasms among the operators and they have been driving the process.

It was found that it is possible to reach maximum design cast speed for majority of size / alloy combination. This would result in an increase of casting productivity by 7%. For implementations some of the parameters needed to be adjusted e.g., steady state water flow, steady state metal level in the mould etc. The implementation was carried out in incremental steps of cast speed increase prior to reaching the target speed. The quality of the output ingots were very closely monitored and only after ensuring that it meets the customer specifications, they were shipped.

It was also agreed that targets for various activities are dynamic and maximum validity of targets to be 1 year maximum and beyond which it must be set at higher levels. Results The effect of the implementation started to be felt from March 2010 onwards. The net production in March 2010 was 11363 MT of rolling ingot. To sustain this achievement ALBA's rolling ingot customer was also taken into board. A joint task force was formed with members comprising of production, R & D and marketing from ALBA and relevant personnel from ALBA's customer. Various issues which can improve production and delivery of rolling ingots were discussed in details. Some of the issues are [4],

Implementation & Monitoring Inline with ALBA's strategy towards process optimization operations at DC1 was completely stopped 31 December 2009 onwards.

1.

Implementation of the agreed points for DC6 started to be implemented January 2010 onwards in steps and real effect of implementation was being felt March 2010 onwards.

2. 3.

KPIs for various activities were agreed upon and were displayed in the shopfloorwith clear responsibility and accountability.

4.

Performances of the individual shifts were being closely monitored by Performance dashboard.

Switching over from bi weekly placing of orders to monthly order placement Agreement on sequence of casting various alloys to minimise requirement of furnace washes Optimizing length of ingots for maximizing furnace output Close communication between the companies for better alignment and fast response

The understanding is to arrive at win-win situation which benefits both the stake holders and the results, as given below, adequately demonstrates this.

696

NET PRODUCTION, MT 12000

2008 Jan-10 avg/month

Feb-10 Mar-10 Apr-10 May-10 Jun-10

Figure 8 Trend of Net Production at DC6 [5]

Conclusion As an out come of implementation of "lean" process and recommendations from the task force ALBA is able to achieve higher levels of production over a period. ALBA is able to record a 32% improvement in production (March - June 2010 period) on a sustained basis as compared to year 2008 levels of production at DC6.

References 1.

Womack, James P.; Daniel T. Jones, and Daniel Roos (1990). The Machine That Changed the World.

2.

Internal Production Data (2009), Aluminium Bahrain

3.

Al Faseela Team Report (2010), Aluminium Bahrain

4.

ALBA-GARMCO Task Aluminium Bahrain

5.

Internal Production Data (2010), Aluminium Bahrain

Force

Report

(2010),

697

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

THE COUPLING OF MACROSEGREGATION WITH GRAIN NUCLEATION, GROWTH AND MOTION IN DC CAST ALUMINUM ALLOY INGOTS Miha Zaloznik 1 , Arvind Kumar 1 , Hervι Combeau 1 , Marie Bedel 1 ' 2 , Philippe Jarry 2 , Emmanuel Waz 2 1

Institut Jean Lamour, CNRS - Nancy-Universitι - UPV-Metz, Ecole des Mines de Nancy, Parc de Saurupt CS 14234, F-54042 Nancy cedex, France 2 Alcan CRV, 725 Rue Aristide Berges, BP 27, F-38341 Voreppe cedex, France Keywords: Direct chill casting, Aluminum alloys, Solidification, Macrosegregation, Microstructure Abstract The phenomena responsible for the formation of macrosegregations, and grain structures during solidification are closely intertwined. We present a model study of the formation of macrosegregation and grain structure in an industrial sized (350 mm thick) direct chill (DC) cast aluminum alloy slab. The modeling of these phenomena in DC casting is a challenging problem mainly due to the size of the products, the variety of the phenomena to be accounted for, and the non-linearities involved. We used a volume-averaged multiscale model that describes nucleation on grain refiner particles and grain growth, coupled with macroscopic transport: fluid flow driven by natural convection and shrinkage, transport of free-floating globular equiaxed grains, heat transfer, and solute transport. We analyze the heat and mass transfer in the slurry moving-grain zone that is a result of the coupling of the fluid flow and of the grain nucleation, growth and motion. We discuss the impact of the flow structure in the slurry zone and of the grain packing fraction on the macrosegregation. Introduction The macrosegregation in the DC casting process is governed mainly by two mechanisms: by the melt flow induced by thermosolutal natural convection, shrinkage and pouring, and by the transport of solute-lean free-floating grains [1-6]. A commonly observed, surface-to-surface distribution of alloying elements at a transverse crosssection of a DC cast ingot reveals distinct regions of positive (solute-rich) and negative (solute-depleted) segregation [1]. A solute-depleted region is present in the ingot center, adjoined by a positive segregation zone spreading into the outward direction, an adjacent thin negative segregation zone and another positive segregation layer at the surface. Experimental investigations were published on macrosegregation and macrostructure in grain refined and non-grain refined ingots [2,7]. It was reported

that macrosegregation generally increases with grain refinement and linked this to the increased transport of free-floating coarse (slowly growing) grains, formed either by the fragmentation of dendrites or by nucleation on grain refiner particles. Eskin et al. [7] presented a systematic experimental investigation of the dependence of macrosegregation and structure on process parameters. A strongly supported hypothesis states that the negative center line segregation is caused by the transport of solutelean free-floating grains to the center of the casting. First attempts to model the influence of free-floating grains were made by Reddy and Beckerman [3]. They considered nucleation of grains at a fixed temperature and the transport and growth of spherical globular grains in a slurry zone. The solid phase was assumed to form a connected rigid porous structure (packing) at a solid volume fraction of 0.637 (packing fraction). In the case of simulations accounting for grain motion, a significant negative segregation at the center of the billet was found. Vreeman et al. [4,5] proposed a simplified model with regard to grain nucleation and growth, assuming local equilibrium (lever rule) and a constant imposed characteristic grain diameter. They calculated the grain velocity directly from the grain diameter and the solid fraction. Macrosegregation distributions in DC cast billets were calculated [5] and a parametric study of two key model parameters, the packing fraction and the grain diameter, was performed for Al-4.5wt%Cu and Al-6wt%Mg billets with a diameter of 400 mm. The results were qualitatively consistent with commonly observed macrosegregation trends. The study revealed a large degree of dependence on both the packing fraction and grain diameter. It has to be noted that their imposed grain diameter has to represent an actual grain size distribution and that the packing fraction is not well known and, moreover, might not be uniform throughout the mushy zone. In a later work Vreeman et al. [4] compared model predictions to measurements on industrial-scale 450 mm diameter DC cast billets of an Al-6 wt%Cu alloy and tried to determine the value of the packing fraction to obtain the best fit. They found rea-

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sonably good agreement between the experimental and model results and they estimated the packing fraction to be in the range of 0.2-0.3. For a more detailed overview of macrosegregation modeling the reader is directed to the recent extensive review [1]. We recently conducted a systematic study of the influence of the individual transport mechanisms, viz. shrinkage, natural convection and grain motion, and their interactions, on the macrosegregation formation [10]. In the present paper we go further and analyze in more detail the heat and mass transfer in the slurry moving-grain zone that is a result of the coupling of the fluid flow and the grain nucleation, growth and motion. We discuss the impact of the flow structure in the slurry zone and of the grain packing fraction on the grain growth, motion and the resulting macrosegregation. Model Physical Model and Solution Procedure The multiscale two-phase model SOLID is presented in entirety in [8]. We here therefore provide only a brief description and point out model extensions with respect to ref. [8]: the consideration of nucleation on inoculant particles and their transport, and the consideration of shrinkage-induced fluid flow in the mushy zone. The model is based on a volume-averaged Euler-Euler two-phase model that consists of two parts: a macroscopic part with momentum, mass, heat, solute mass, and grain population conservation equations, and a microscopic part that describes the nucleation and growth of grains. At the macroscopic level, the model accounts for heat and solute transport coupled with flow driven by thermal and solutal buoyancy and by solidification shrinkage (assuming no strain in the solid). Depending on the behavior of the solid phase, we consider two flow regimes in the mushy zone. Where the solid volume fraction gs is larger than the packing limit (gs > p| )ack ) the solid is considered to be blocked and moving at the casting velocity. The flow of intergranular liquid through the porous solid matrix is described by a momentum equation including a Darcy term for the drag interactions, with the permeability modeled by the Kozeny-Carman law. The density of the solid phase is assumed to be constant and for the liquid density the Boussinesq assumption is employed. The flow due to solidification shrinkage is induced via the enforcement of mass conservation, by considering different densities of the solid and liquid phases. At solid fractions smaller than the packing limit (gs < ^ a c k ) the solid phase is considered to be in the form of free-

floating grains. Their motion is described by a balance of buoyancy, drag and pressure forces acting on a grain. In this way, the solid and liquid have locally different velocities. The interfacial particle drag is considered dependent on the grain size, which produces the tendency that the larger the grains are, the stronger their tendency to settle; contrarily, smaller grains are more easily entrained by the liquid motion. The microscopic level is treated locally; within SOLID, this means within each discrete volume element. The nucleation is controlled by the addition of inoculants. According to the theory of Greer et al. [9] an inoculant particle is activated as a nucleation (or better, growth-onset) site at a critical undercooling that is inversely proportional to its size: ATuc(d) oc d~l. A typical particle size distribution in the active particle range in a commercial inoculant is exponential. The largest particles nucleate first and afterwards a nucleation-growth competition takes place. As long as the number of nucleated grains is still small, the solidification kinetics is too slow and the constitutional undercooling will continue to increase, triggering the nucleation on smaller particles. As the grain density increases, the growth kinetics of the nucleated grains approaches equilibrium and the undercooling decreases; further nucleation is blocked. This physics is modeled by considering a distribution of inoculant particle size, discretized into 10 classes. Each class has its own activation undercooling, depending on the mean particle size in the class, and an initial density, calculated from the known distribution density. Moreover, the transport of nuclei is considered, assuming that they move at the velocity of the liquid. The conservation equation for nuclei of class i is ^ ί + ν . ( ^ φι

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ί

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where iV£ucl is the volume density of nuclei of class i, v\ is the intrinsic velocity of the liquid, Φ1 is the nucleation source term, δ is the Dirac delta function, ATUC = WIL(C* — C\) is the local undercooling, m^ is the liquidus slope, C* is the concentration of liquid at the solid-liquid interface, C\ is the local average concentration of the liquid, and ΔΤ^ υο1 is the activation undercooling for the nuclei particles of class i. At the same time the conservation equation for grains is

700

— +ν·(^Λ0 = -]Γ.Φ 1

(3)

where N is the local volume density of grains and vs is the velocity of the solid grains. The source term accounts for nucleation of grains from the grain refiner particles. The nucleation is solved coupled with the macroscopic transport and the likewise local (microscopic) phase-change. The phase change (solidification and melting) is controlled by solute diffusion in both phases at the grain scale, assuming local thermal equilibrium and thermodynamic equilibrium at the solid-liquid interface. In the present work the grains are considered spherical with a fully globular morphology, which in many cases can be considered as a realistic assumption for inoculated DC cast aluminum alloys. We solve the macroscopic equations with a finitevolume method. The local microscopic growth model is integrated implicitly and is coupled with the microscopic nucleation and the macroscopic model via a three-step operator-splitting integration method that separates the integration of the macroscopic terms and the microscopic nucleation and growth terms into three separate stages. This method is an extension of the algorithm [8] that introduces a separate solution step for the nucleation. Process Model We studied the solidification in a 7449 alloy slab of 350 mm thickness. We consider a simplified 2D geometry with symmetry, i.e. a domain of 175 x 800 mm. The computational grid consists of 40 x 115 rectangular cells, refined in the solidifying and liquid zones. The solidified metal leaves the domain at a casting speed of 75 mm/min at the bottom. Note that this casting speed is much higher than the industrial practice, which we did to amplify the macrosegregation. The feeding of the liquid metal with the nominal composition, inoculant particle density distribution and casting temperature is at the top across the whole cross section. The feeding velocity is uniform across the inlet and is given by a mass balance accounting for the solidification shrinkage. The heat extraction in the mold is described by three zones with different heat transfer coefficients: a meniscus at the top, a contact zone, and an air-gap zone. The heat extraction in the water-chill under the mold is described by the classical Weckman-Niessen correlation. The 7449 alloy was modeled in a simplified way, as an equivalent pseudo binary alloy, the approximation done to match the solidification path of the multicomponent 7449 alloy, calculated with a Calphad model. This gave a partition coefficient of 0.257, a linear liquidus slope of —6.05K/wt%, and a melting temperature of pure Al at 677.8 °C (a projection of the liquidus to C= 0). The alloy

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0.05 0.1 x[m]

0.15

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0.05 0.1 x[m]

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Figure 1: Segregation and streamlines of relative liquid velocity (Q -vcast). Left: # s pack = 0.3. Center: g%ack = 0. is further modeled with constant density of the solid and a Boussinesq approximation with constant thermal and solutal expansion coefficients for the liquid. The liquid density is thus variable in the buoyancy terms, for both the liquid and the solid force balance (the grain buoyancy depends on (ps — pi)), but is constant in the mass balance. The shrinkage coefficient ίs\ — (ps/pi,ref — 1) is thus constant; we used ίs\ = 0.057. The thermal and solutal expansion coefficients were modeled to fit the variation of liquid density along the solidification path of the 7449 alloy. For this alloy thermal and solutal expansion are cooperating and the influence of heavier solutes rejected into the liquid upon solidification dominates. Results and Discussion The Role of Grain Motion in Segregation Formation When a part of the grains growing in the mushy zone of a DC cast ingot are free to move, these grains, as they are usually heavier than the surrounding liquid, have a tendency to settle. They settle along the inclined mushy zone towards the center, where they accumulate at the bottom of the sump. Upon settling of the solute-lean grains the solute-rich liquid is expelled upwards and this creates a negative macrosegregation tendency in the center of the ingot (Fig. 1). In addition to the grains, the

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liquid is also set in motion. In the slurry zone the principal driving force for fluid motion is the entrainment by the moving grains; additionally, thermal and solutal buoyancy forces can induce motion too. This flow is globally downwards with the fast current of settling grains along the packing front, and recirculating slowly back upwards in the center (Figs. 1 and 4). The upward flow entrains some grains, which creates the extended slurry zone. Just below the packing front the thermosolutal buoyancy is the prime driving force for the intergranular fluid flow through the packed porous solid matrix. A little deeper into the mushy zone at smaller liquid fractions the permeability strongly decreases, so the high flow resistance completely blocks the natural convection flow. The flow is now controlled by the solidification shrinkage that creates a high pressure drop and orients the flow towards the solidification front (Fig. 1).

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Figure 2: Dependence of the centerline segregation on the packing fraction.

fraction region of the mushy zone is reduced due to a lower velocity difference vs — v\ (the grains move). In the packed Corresponding to this flow situation we can distinguish region, on the other hand, the segregation is weaker than four zones for segregation formation: the slurry zone, the before since the liquid flow is less intense due to a lower packing front, the moderately permeable packed layer permeability of the packed region. As the packing frac(moderate g\), and the impermeable packed layer (low tion is increased further, this effect becomes stronger and g\). We investigated these interactions in [10], where stronger. At the same time the solute transport by grain we have shown that the grain transport is not only di- settling starts to become important. We now find a negrectly responsible for the creation of the negative center- ative segregation in the center (at gfack ~ 0.05). From line segregation, but also changes the sump shape and this point on, a large slurry zone develops with a particthus modifies the action of the natural convection and ular flow structure (Fig. 4), where the flow descends in a shrinkage flow and the segregaton they cause. While it strong current limited to the close vicinity of the packing is generally supported that the transport of solute-lean front and the liquid ascends slowly in the almost stagnant free-floating grains causes a negative segregation at the core of the slurry zone. The grains that are carried into centerline, where they settle, the modification that comes the core settle downwards very slowly in a countercurrent with the intensity of grain transport is not clear a-priori. motion, some are even entrained by the liquid and leave We would expect that as more grains are free to move and and remelt into the fully liquid zone. settle, the negative centerline segregation will be ampliHow is this flow set up? We can observe that the liquid fied. A numerical parameter study of the dependence of in the core of the slurry zone is close to thermodynamic ack the centerline segregation on the packing fraction (#J ) equilibrium. As we will see later, this happens because of shows a quite nonlinear image (Fig. 2). We can see that the slow phase change (solidification/melting) of the floatthe centerline segregation in a columnar ingot is strongly ing grains. The slow phase change means that the solute positive (Figs. 1, 2). When a part of the grains is free exchange of the interface of a grain is slow and the diffuto move the centerline segregation drops sharply at low sion in the liquid surrounding a grain has enough time to gpack a n ( j q U i c kiy reaches a negative centerline segrega- maintain the liquid close to the thermodynamic equilibtion (at gfack ~ 0.05). As the proportion of free grains rium concentration at the interface. This means that the increases, the centerline segregation continues to drop, temperature and the liquid concentration are closely couuntil it reaches a minimum (at # s pack - 0.25 in our exam- pled, which can be described by the equilibrium relation ple) and then increases again for higher pfack. T = Tf -{-m^Cx for the liquidus temperature. As the soluLet us look in more detail at what happens in Fig. 2. As the packing fraction is increased, more grains are free to settle at the bottom. At low packing fractions the impact of the solute transport by solute-lean grains remains relatively fable. The grain transport by itself does not provoke the sharp decrease of the positive centerline segregation. However, the segregation caused in the low-solid-

tal effect on the buoyancy force is much stronger than the thermal one, and the Schmidt number Sc = D/v ~ 100 of the solution is very high, the dominating solutal buoyancy creates a stable solutal stratification. The coupling of the temperature and the concentration maintained by the phase change then induces a corresponding thermal stratification, stable as well (Fig. 4). The solid fraction is

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(a) Stratification controlling the flow in the slurry zone. Left: average density p m (color map) and liquid fraction. Right: temperature (color map) and liquid concentration (contour lines in mass fraction).

Figure 3: Segregation and streamlines (v\ — vcast) for different packing fractions. Left: #fack = 0.03. Center: ^pack = o 20. Right: g^ck = 0.50.

stably stratified too. In this situation the driving force for the flow comes from the lateral gradients of temperature, concentration and, most notably, solid fraction, which are localized next to the inclined packing front in the growth region and drive the descending current. The relative grain settling velocity vs — v\ is essentially driven by a balance of buoyancy and drag forces acting on the grains. If we simplify this force balance, we can show that essentially, the settling velocity of a globular grain is proportional to the density difference between solid and liquid and the square of the grain size: va—v\ oc (ps—p\)d2. While the grain size can differ and is dependent on a rather complex nucleation-growth competition in the nucleation zone as well as the grain transport, we did not observe a fundamental variation with grain size as a function of the packing fraction. The decisive factor is the density difference (ps - pi). The density of the liquid phase depends on its temperature and composition. In the present case, where the alloying elements are heavier than aluminum, the density of the liquid increases as solidification progresses. This can be demonstrated by applying a Scheil solidification path to the density function PX = a r e f ( l - ίT(T - Tri) - ίciPx - Cref)). The density of the primary solid phase, on the other hand, is approximately constant. At the onset of solidification the solid density is greater, however in many alloys the liquid density becomes larger at a certain solid fraction. In the modeled alloy this happens at approximately gs = 0.30. The grain settling velocity slows down and even reverses

(b) Grain nucleation, growth, motion and accumulation. Left: grain population density and solid velocity streamlines (an approximation for the grain trajectories due to a qualitatively steady state). Right: undercooling (—0.684 °C is the activation undercooling for the largest inoculant particles) and liquid velocity vectors. The liquidus, packing {gi*C = 0 . 3 ) and solidus fronts are marked.

(c) Grain motion and growth. Left: relative grain velocity (vs — vi)) magnitude and streamlines. Right: solidification rate (dgs/dt > 0 for solidification and dgs/dt < 0 for melting).

Figure 4: Flow, motion, nucleation, growth and coalescence of free-floating grains - conditions in the slurry zone for #P ack = 0.3. (i.e. the grains start to fractions, # s pack > 0.30 grain transport reduces tion. At the same time

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float upwards) for high packing (Figs. 3 and 4). The decreased the negative centerline segregathe shape of the sump is modi-

fied as more grains settle to the center with an increasing gpack rpkg p a c k e c j p ar ^. Qf fae mushy zone becomes shallower and the streamlines of the shrinkage-induced flow in this region are less divergent. The shrinkage induced segregation in the center is thus reduced.

References [1] R. Nadella et al., "Macrosegregation in direct-chill casting of aluminium alloys," Prog. Mater. Sci., 53 (2008), 421-480. [2] G. Lesoult et al., "Equi-axed growth and related segregations in cast metallic alloys," Sci. Technol. Adv. Mat, 2 (2001), 285-291.

Grain Nucleation, Growth and Motion in the Slurry Zone The flow enters the slurry zone exclusively in the stream flowing downward along the packing front. This is shown in Fig. 4. Analyzing the grain trajectories we can see that the grains then either attach to the front, settle to the bottom of the slurry zone or enter the core of the slurry zone. In the core the grains slowly settle downwards before rejoining the stream and settling to the bottom of the sump. The residence time of these latter grains in the core is rather long and they have enough time to regain equilibrium. We can see in Fig. 4 that the core is a zone of very small undercoolings/superheats and the phase change rate is correspondingly small. This is a zone of very slow growth and partial remelting of the freefloating grains, which is also shown in Fig. 4. The undercoolings ATUC in the core are also not enough to trigger nucleation in this zone. This means that the grains here all originated elsewhere and were transported here. The zone of higher undercoolings is located along the packing front in the main stream (Fig. 4). This is the region of nucleation and fast growth. The free floating-grains thus nucleate mainly in the stream entering the slurry zone. They first grow fast while descending along the packing front and then their growth slows down as they float around in the stagnant slurry zone.

[3] A. V. Reddy and C. Beckermann, "Modeling of macrosegregation due to thermosolutal convection and contraction-driven flow in direct chill continuous casting of an Al-Cu round ingot," Metall. Mater. Trans. B, 28B (1997), 479-489. [4] C. J. Vreeman, J. D. Schloz, and M. J. M. Krane, "Direct chill casting of aluminum alloys: Modeling and experiments on industrial scale ingots," J. Heat Trans.-T. ASME, 124 (2002), 947-953. [5] C. J. Vreeman and F. R Incropera, "The effect of free-floating dendrites and convection on macrosegregation in direct chill cast aluminum alloys, part II: Predictions for Al-Cu and Al-Mg alloys," Int. J. Heat Mass Tran., 43 (2000), 687-704. [6] M. Zaloznik and B. Sarler, "Modeling of macrosegregation in DC casting of aluminum alloys: Estimating the influence of casting parameters," Mater. Sci. Eng. A, 413-414 (2005), 85-91. [7] D. G. Eskin et al., "Structure formation and macrosegregation under different process conditions during DC casting," Mater. Sei. Eng. A, 384 (2004), 232-244. [8] M. Zaloznik and H. Combeau, "An operator splitting scheme for coupling macroscopic transport and grain growth in a two-phase multiscale solidification model: Part I - model and solution scheme," Comp. Mater. Sci., 48 (2010), 1-10.

Conclusions We analyzed the coupling of the flow structure in the slurry zone, grain growth and motion, and macrosegregation. The grain settling dynamics strongly depends on the packing fraction and is closely coupled with the segregation. We could show that the grains initially nucleate and grow fast in a zone of high undercooling, while settling along the packing front. They then continue with a phase of slow growth while floating in the core of the slurry zone.

Acknowledgements

[9] A. L. Greer et al., "Modelling of inoculation of metallic melts: Application to grain refinement of aluminium by Al-Ti-B," Acta Mater., 48 (2000), 28232835. [10] M. Zaloznik et al., "Influence of transport mechanisms on macrosegregation formation in direct chill cast industrial scale aluminum alloy ingots," Matιriaux 2010, submitted to Adv. Eng. Mater. (2010).

This work was supported by Alcan CRV.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INVESTMENT CASTING OF SURFACES WITH MICROHOLES AND THEIR POSSIBLE APPLICATIONS Todor Ivanov1, Andreas Biihrig-Polaczek1, Uwe Vroomen1, Claudia Hartmann2, Jens Holtkamp2, Arnold Gillner2, Kirsten Bobzin3, Nazlim Bagcivan3, Sebastian Theiss3 ^WTH Aachen University, Foundry Institute, Intzestrasse 5, 52056 Aachen, Germany Fraunhofer-Institute for Laser Technology, Steinbachstrasse 15, 52074 Aachen, Germany 3 RWTH Aachen University, Surface Engineering Institute, Kackertstrasse 15, 52072 Aachen Germany 2

Keywords: Investment casting, microstructured surface, lotus leaf, microcasting, gypsum-bonded investment roughness of between 0.5 μπι and 1 μπι. The possibility of casting microstructured surfaces on a cast part such as a dummy turbine blade with a structural size of 50 μηι and an aspect ratio of 1 were demonstrated for an aluminum alloy and a Ni-based superalloy [6]. Other analyses focused on the molding accuracy achievable for casting of low melting Bismuth-Tin and Bismuth-Lead alloys. These were directly cast into polymer as well as quartz molds to avoid the process steps of wax fabrication and ceramic mold making. Owing to the very high surface quality of the used molds, grooves of 4 μτη width and 200 nm depth [6] as well as ridges with a periodicity of 400 nm [7] could be cast. The objective of the work presented here is the characterization of the investment casting process in order to cast a microstructured hydrophobic surface with a common AlSi-cast alloy. Therefore the relevant influencing parameters will be identified.

Abstract The usual way of realizing microstructured features on metallic surfaces is to generate the designated pattern on each single part by means of laser ablation, electro discharge machining or micro milling. A disadvantage of these process chains is the limited productivity due to the additional processing of each part. The approach taken by this project is to replicate microstructured surfaces via investment casting. The main research objective deals with the investigation of single process steps of the investment casting process with regard to the molding accuracy. To demonstrate the potential of microcast surfaces, current results for the casting of a microstructured hydrophobic surface will be shown. Introduction

Experimental Details

Many microstructured surfaces are known in nature such as shark skin, lotus leaf and insect feet structures. By understanding these effects and using this knowledge for technical applications, several material properties such as drag, friction, adhesion, hydrophobia can be adapted. Possible ultra precision technologies for creating microstructured surfaces on metal parts are laser ablation, electro discharge machining and micro milling. The shortcoming of these processes is their limited productivity since each part has to be separately microstructured [1]. Furthermore, micro structuring inner areas of technical parts is limited. A promising reproduction process that meets all requirements is the investment casting process. By molding from a microstructured grand master pattern, parts of arbitrary geometry can be manufactured at relatively low cost compared to common ultra precision technologies. As described above, laser structuring provides one possibility for creating the functional surfaces on the grand master pattern. Laser ablation has already been proven for the production of micro replication tools. In principle there are two technologies which can be employed for generating microscaled structures on tool surfaces: Direct microscale structuring by laser evaporation which is applied in the process described here - and indirect micro structuring by laser ablation of surface layers and then subsequent etching. The grand master pattern is then used for replicating the microstructured surface on metal parts via the lost wax investment casting process. Several analyses have been conducted to examine the molding accuracy of this process. The accuracy of casting microparts were investigated using aluminum/zinc alloys [2, 3] and stainless steel [4] as well as low melting Au-Ag-based precious metal alloys [5]. According to these results, microparts with geometrical features in a range of 700 μπι to 50 urn can be cast with acceptable geometrical variations and an average surface

In order to examine the molding accuracy of each investment casting process step (Figure 1), an inverse lotus leaf like surface was used. The grand master pattern was fabricated by laser structuring a microhole geometry into a steel plate, Figure 2.

Sμ!icone inlay in wax pattern die

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Figure 1. Micro casting process chain for producing a hole structure in aluminum parts A silicone inlay with nep structure was molded from the grand master pattern, which served as an inlay in the wax pattern die. Following this, the microstructured surface could be replicated by

705

manufacturing wax patterns. These were embedded into gypsumbonded block molds and used to cast metal.

hydraulic wax injection press, Modtech Machine Pvt. Ltd., Ahmedabad - India) was used. This was equipped with an adjustable heating device at the injection nozzle to precisely adjust the injection temperature of the wax. For the wax injection process, the wax temperature and the injection pressure was set to 69 °C and 16 bar, respectively. The resulting speed of the wax through the ingate could be estimated at 0.7 m/s. After injection, the pressure was maintained for 1 min in order to compensate for the shrinkage of the wax.

~XJür\a< W

Figure 2. Laser structured geometry for the grand master pattern, 0 = 22μιη, h = ΙΟμπι, w = ΙΟμιη

Table 1. Properties of Straight Pattern Wax A7-11, Blayson Olefins Ltd., Cambridge - UK. Value Property Ash content 0.05 % Max Congealing point 69-73 °C Drop melt point 73-76 °C Free linear contraction 0.9-1.1% Viscosity 0.8-1.3 Pas

Structuring of Grand Master Pattern The laser structuring experiments were carried out a fine drilled steel plate (X8CrNil8-9, AISI 303) with a diode pumped Nd:YV04 MOP A laser (Master Oscillator Power Amplifier; Rapid, LumeraLaser) at a wavelength of ë = 355 nm. The laser operates with repetition rates of up to v = 500 kHz and a pulse duration of τ= 12 ps. For the micro structuring experiments, the laser focus was positioned on the surface of the samples by a galvanometer scanning system with a focal length of f = 53 mm (Fig.3). The ablated geometry consisted of holes with a diameter of 22 μπι. The holes were created by cross-hatching at a line separation of 1.5 μπι and 10 parallel lines in each direction. The hole depth is adjusted by repetition of the hatching. Micro structuring experiments were performed at a repetition rate v = 250 kHz and a laser power of P = 600 mW. The mark speed was set to 50 mm/s, the cross hatching was repeated twice.

Embedding the Gvpsum-Bonded Mold Typical mold materials used in foundry technology are not suitable for casting of microstructured surfaces. Foundry sand is, in general, too coarse. Ceramic shell molding materials used in investment casting as well as phosphate bonded investments are fine-grained but possess very high hardness values which impede their ejection. To remove of the ceramic layer from the cast part, sand blasting is usually used. This would damage the microstructured surface of aluminum part. For this reason, only gypsum-bonded investments fulfill the main requirements concerning moulding accuracy and shake out behavior and are used for the experiments. The gypsum-bonded investment material Gold Star XXX (Goodwin Refractory Services Ltd., Staffordshire, UK) was used for the production of the mold. The material is characterized by a high concentration of small filler particles and has an average grain size of 18μπι (manufacturer's information). For the preparation of the block mold, the plaster was mixed with water in a ratio of 100:40 according to the manufacturer's instructions. This mixture was poured into a mold flask and then degassed in a vacuum system at about 50 mbar residual pressure. The mold was subsequently dried at room temperature for about 12 h. Following drying, the mold was dewaxed in an oven at 150 °C for about 4 hours. The subsequent burning of the mold was carried out at a heating rate of 100 °C/h up to 700 °C with a soak time of 4 h. After the burning process, the mold was cooled down in the oven to the preheating temperature of 420 °C for subsequent casting.

Fabrication of the Wax Pattern The wax pattern die shown in Figure 3 was used to produce the microstructured wax patterns. The material for the silicone inlay was Elastosil M4643 A, B (Drawin Vertriebs GmbH, Riemerling Germany) with a Shore hardness A of 48 [8]. Silicone material is often used for wax pattern molds in investment casting foundries. The advantages of using silicone over metal dies are, on the one hand, rapid manufacturing when the microstructured surface is contaminated with wax. On the other hand, the silicone material is elastic: Therefore the risk of tearing the microfeatures is reduced during ejection of the wax pattern. The lower thermal conductivity of silicone also leads to a slower solidification of the wax: Thus the microstructured features are molded more precisely since the material penetrates into the small features before solidification. The wax pattern die was equipped with a deaeration system in order to scavenge the air out from the cavity during wax injection. Valve foi optional evacuation

Counter Pressure Casting and Demolding An AlSi7Mg0.3 cast alloy which is characterized by a good mold filling capacity [9] was used for casting. The chemical composition is given in Table 2. The melt was neither grainrefined nor modified. The casting experiments were conducted using a counter pressure system. Thus the block mold was placed into a steel sleeve. Directly before the metal was poured into the mold, the air from the sleeve was drawn into a low pressure boiler. When the cast metal sealed the inlet of the mold, a pressure of about 50 mbar was applied to the cavity; this increased the casting pressure by 0.95 bar. A resulting casting pressure of 4 bar could be estimated at the beginning of the mold filling. The melt

Deaeration channels Microstructured silicone inlay Ingate for wax injection

Figure 3. Wax pattern die with microstructured inlay. The wax patterns were made of an unfilled wax (Straight Pattern Wax A7-11, Blayson Olefins Ltd, Cambridge - UK). The wax properties are shown in Table 1. To fabricate the wax pattern, an electronically controlled wax injection machine (35t C-frame

706

was overheated and cast into the mold at 740°C (liquidus temperature: 613°C) in order to achieve a very good mold filling.

Grand Master Pattern Using the parameters described above, laser structuring of a grand master pattern was performed. The reproducibility of the structure can be seen in figure 5. The difference in height can be ascribed to the rough surface of the grand master pattern's base plate. If the surface is not sufficiently smooth and flat, the laser beam is differently focused for each hole. Thus, the ablation depth varies. To increase the reproducibility and therefore the quality of the laser drilled pattern, the grand master pattern used here has to possess a higher quality regarding its roughness (e.g. polished surface) and flatness in order to guarantee a correct focal position.

Table 2. Chemical composition in wt-% of aluminum cast alloy AlSi7Mg0.3 (samples were analyzed with spark spectrometer Spectromax, Spectro GmbH&Co.KG, Kleve, Germany), Ti Si Mg Fe Cu B Sr 6.98 0.339 0.111 0.0053 0.118 0.0015 0.0002 Analysis and Testing In the individual process steps, the reproduction of the microstructured surface has been studied on different samples since some measurements involve the destruction of the samples. To measures the micro feature's height, a white light interferometer (WIM) was used. An error of 1 μπι was estimated for these measurements. This error results from slightly different points of measurement because the identical point could not always be found on the specimen's surface. Furthermore, the grand master pattern's surface had a roughness (Ra) of 0.53 μιη, measured in the unstructured field, which superposes the microstructured surface. SEM analyses of the gypsum-bonded mold and the casting were performed in order to get an impression of the spatial image and the reproducibility of the microstructured surface. The structural investigations of the micro patterned surface were carried out on metallographic polished samples of the cast part using a reflecting light microscope. Finally, contact angle measurements with ultra pure water were carried out to determine the effect of the hydrophobic surface structure. A DSA 10, Krüss GmbH, Hamburg, was used to measure the contact angle. During the measurement of one drop, 60 measurements of the contact angle were performed. Six drops for every field were analyzed to calculate a standard deviation.

Figure 5. White light microscopy of the grand master pattern with lasered microholes. Dimensions: h = 10 μπι, 0 = 22 μπι.

Results and Discussion The microstructured surface could be reproduced over the whole of the investment casting's process chain. However, some loss in molding accuracy of the micro features could be observed at different process steps. The highest loss in feature height occurred on injecting wax into the wax pattern die, Figure 4. The detailed results for each process step are discussed below. Figure 6. White light microscopy of the wax pattern. Dimension: h = 5 μπι, 0 = 22 μπι (scale for height varies fromfigure5). Wax Pattern

Grand master pattern

Silicone inlay Wax pattern Investment (nep) (hole) (nep)

Starting the replication process from the grand master pattern, the hole structure's dimensions have been completely molded in silicone as neps over the entire area. Using a silicone inlay which was placed in the wax pattern die, the microhole structure was molded into the wax pattern. Within this process step, the largest loss of molding accuracy occurred. Only half of the original micro feature's height could be achieved in the wax pattern, Figure 4 and Figure 6. The results of the white light interferometer measurements showed that the grooves, which originated by fine turning the grand master pattern before laser structuring, were transferred in the wax pattern. This indicates that the microstructured surface of the silicone inlay was completely molded into the wax pattern.

Casting (hole)

Figure 4. Moldability of microfeature height in each process step.

707

Considering this excellent moldability as well as results from other examinations with similar completely molded microstructures in wax patterns, it is very unlikely that entrapped air or the material properties of the wax (surface tension, shrinkage, etc.) or the injection parameters caused the micro holes to be too shallow. The height of the micro neps in the silicon inlay did not change even after several injection cycles. For this reason, it can be excluded that the peaks of the ribs, which form the hole structure, adhered in the wax pattern die and tore off during ejection. The most likely reason for the shallow holes in the wax pattern is that the neps of the soft and flexible silicone inlay were compressed by the high wax injection and squeezing pressure of 16bar. If the high pressure changed the contour of the micro features could not be measured because the sides of the holes in the wax pattern did not reflect enough light back towards the white light interferometer.

the holes was much smaller. This may be due to the rough surface of the neps in the gypsum-bonded mold. Moreover, it could be that the strength of the gypsum-bonded investment was too low: This leads to erosion during mold filling or deformation after filling e.g. by shrinkage of the metal. In general, it can be said that the wetting properties of the molten metal on the microstructured gypsum-surface were sufficient as the well molded areas on the cast part demonstrate. Hence, the casting and mold temperature could be lowered in order to reduce the risk of erosion caused by thermal and mechanical loads. Another possibility is to increase the mechanical properties of the gypsum-bonded investment. Therefore, the water content could be initially lowered. Other factors which have an influence are the grain size distribution, thefiller/binderratio and the shape of the filler grains. Also, the chemical composition either of the filler (fraction of quartz and cristobalite) or of the gypsum plaster (modification of calcium sulfate and adding of set-up agents) can be adjusted.

Gvpsum-Bonded Mold To investigate the quality of the gypsum-bonded block mold, the hole structure was measured using WIM and SEM. The height of the micro feature (4μιη) differed from the height of the wax pattern by Ιμπι. This difference was caused by the above mentioned measuring inaccuracy and by the roughness of the mold surface, which depends primarily on the particel size of the investment material. As can be seen in the SEM image in Figure 7, the structure of the entire area of the microstructured field has been reproduced without any major defects. On the other hand, the surface is rugged and porous which is caused by the growth of thin gypsum needles. In order to achieve a smoother surface, it can be assumed that a finer grain size distribution of the mold material will not lead to a better molding accuracy. Rather, it is necessary to achieve smaller gypsum needles and less porosity. Therefore the water content of the gypsum mixture could be reduced, so that the gypsum oversaturated faster and a finer gypsum-needle network develops [10].

Figure 8. SEM image of cast part (the grains in the holes are the remains of investment powder) Nevertheless, the wettability of the four structures was analyzed using contact angle measurements with ultra pure water. The contact angle of 75° on the microstructured surface demonstrates the improvement compared to the unstructured part which has a contact angle of 60°. A further increase of the contact angle is possible by reducing the roughness of the unstructured grand master patterns as well as of the mold material and an increase in number of holes per cm2. Additionally, it is possible that an increase of the roughness by micro structuring leads to a higher influence of the slip-stick effect which results in higher contact angles. However, based on the low standard deviations (3°) of the measured contact angels compared to the significant effect of the structures, the slip-stick effect can be excluded. Most studies concerning hydrophobic effects used non-metallic materials with low surface energies [11, 12]. If it is possible to achieve a hydrophobic effect (contact angle > 90 °) on a common Al-alloy surface by merely improving the microstructure geometry, then this is a prospect for future examinations.

Figure 7. SEM image of gypsum-bonded investment mold Casting The micro structure was molded on the surface of the cast part and the feature height of the gypsum-bonded mold could be achieved on the casting. However, the hole depth and diameter were not transferred reproducibly, Figure 8. In some areas the diameter of

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References

Conclusions, Prospects and Possible Applications It has been shown that a laser structured surface consisting of microholes with a diameter of 22μπι, a depth of ΙΟμπι and a separation of ΙΟμπι can be replicated in the successive investment casting process. Micro features with the above mentioned dimensions can be completely reproduced in silicone. With such an elastic silicone inlay in the wax pattern die, the demolding of the wax pattern can be improved. On the other hand, if the silicone is too soft, the micro features of the silicone inlay can be deformed when high wax injection pressures are applied. In all probability, this was the reason for the insufficient moldability of the micro feature's heights which were 5μπι in the wax pattern. Having the microstructured wax pattern, it is possible to reproducibly mold it in the gypsum-bonded investment over the entire area. However, the gypsum-bonded mold had a porous needle-like surface which reduced the accuracy of the micro features even though the heights of 5μπι given by the wax pattern could be reached. In the last process step, the microstructured surface was cast and 4μπι deep micro holes were achieved. However a few micro features could not be reproduced. This was probably caused by erosion from the liquid metal on the mold surface. Options to reduce the risk of erosion include lower casting temperatures, pressures and increased mechanical properties of the gypsumbonded investment material. Due to the needle like surface of the gypsum-bonded mold, the roundness and diameter of the micro holes varied within a small range of sizes. In order to determine the effect of the used hydrophobic surface, contact angle measurements were carried out. The contact angle of ultrapure water on the microstructured surface could be increased to 75° compared to 60°C on the smooth casting surface. If it is possible to achieve a hydrophobic effect by merely producing microstructured surfaces on aluminum parts or other cast parts without changing the chemical composition of the metallic surface, then this needs to be examined in future research.

[I]

Possible applications

[12]

[2]

[3] [4]

[5] [6] [7]

[81 [9] [10] [II]

Possible applications include cast parts where a reduced wettability or a self-cleaning-effect is desired. Non-stick surfaces can be used on premium life style products such as cast bathroom fittings or cooking equipment. Also, cast parts which become rapidly dirty; such as wheel rims or the underbody of cars and motorbikes, represent very interesting applications for microstructured surfaces, especially in terms of corrosion protection. Another big field is industrial facilities such as chemical or power plants with complex piping and pumping systems which are very difficult to maintain and clean. Also, equipment in the food industry with high cleanliness requirements provides potential applications. Furthermore, micro holes can be used for tribological applications, for example as reservoirs for lubricant films in cylinder-piston- pairings or bearing blocks in engines [13].

[13]

Acknowledgement: The authors thank the German Research Foundation DFG for supporting the research within the Cluster of Excellence "Integrative Production Technology for High-Wage Countries".

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Uriarte, L.; Herrero, A.; Ivanov, A.; Oosterling, H.; Staemmler, L.; Tang, P.T.; Allen, D.: Comparison between microfabrication technologies for metal tooling. J. Mech. Eng. Science, Vol. 220 Part C, 2006 Baumeister, G.; Brando, O.; Rogner, J.: Microcasting of Al bronze: influence of casting parameters on the microstructure and the mechanical properties. Microsystems Technology, Springer-Verlag, 2008 Chuang Yang, B. Sheng Li, M. Xing Ren, H. Zhi Fu: Studies of microstructures made of Zn-Al alloys using microcasting. Int. J. of Adv. Manuf. Technol. (2010) Charmeux, et al.: Benchmarking of three processes for producing castings incorporating micro/mesoscale features with a high aspect ratio. J. Eng. Manufacture, Vol. 221, Part B, 2006 Wφllmer, H.: Untersuchung zum Präzisionsgieίen metallischer Mikroteile. Dissertation, University of Freiburg, Germany, 2000 Schmitz, G. J.; Grohn, M.; Biihrig-Polaczek, A.: Fabrication of micropatterned surfaces by improved investment casting. Advanced Engineering Materials 2007, 9, No. 4 Cannon, A. H.; King, W. P.: Casting metal microstructures from a flexible and reusable mold. Journal of Micromechanics and Microengineering, No. 19, IOP Publishing, 2009 ISO 868 "Testing of rubber - Shore A and Shore D hardness test" Drossel, G.; Friedrich, S.; Huppatz, W.: Aluminiumtaschenbuch, 15th Edition, Aluminium-Verlag, Düsseldorf, Germany, 1999 Kappert, H. F.; Eichner, K: Zahnmedizinische Werkstoffkunde und ihre Verarbeitung. 6th Edition, Georg Thieme Verlag, Stuttgart, Germany, 2008 Kietzig, A.-M.; Hatzikiriakos, S. G.; Englezos, P.: Patterned Superhydrophobic Metallic Surfaces. Langmuir, Volume 25, Issue 8, 2009 Guo, Zhiguang; Fang, Jian; Wang, Libo; Liu, Weimin: Fabrication of superhydrophobic copper by wet chemical reaction. Thin Solid Films, Volume 515, Issue 18, 2007 Schmitz, G.J.; Grohn, M.; Nominikat, J.: Primary shaping method for a component comprising a microstructured functional element. International patent, publication no. WO 2004/087350 A2, RWTH Aachen University, 2004

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

USING SEM AND EDX FOR A SIMPLE DIFFERENTIATION OF a- AND ί-ALFESI-PHASES IN WROUGHT ALUMINUM BILLETS Marcel Rosefort, Christiane Matthies, Hinrich Buck, Hubert Koch1 n TRIMET ALUMINIUM AG, Aluminiumallee l; 45356 Essen; Germany Keywords: aluminum alloys, wrought alloys, intermetallic phases, AlFeSi At first alloy compositions with only a- and only ί-AlFeSi-phases respectively were selected via phase simulations and literature. After a microscopy check of the phases the EDX-measurements were carried out

Abstract Aluminum 6xxx extrusions have considerable potential to make cars lighter. Thus there is a growing demand for high quality aluminum billets. Additionally there is a requirement to improve the aluminum properties. An important issue to fulfill these increasing requirements is the optimization of the microstructure. While utilizing 6xxx-alloys the formation and optimization of the AlFeSi-phase is important. The determination of a- and ί-AlFeSiphases is a challenge, because ί-AlFeSi-phases affect the mechanical properties negatively. A determination in wrought alloys via microscopy is often complicated and leads to questionable results due to the low amount of the phases. TRIMET has investigated one possibility of determining a- and ί-AlFeSi-phases with the help of a scanning electron microscope and energy dispersive X-ray. The paper describes the first tests, the phase simulations and the casting experiments. Finally it presents the tests of such phase determinations in the production of 6xxx-alloys like the new TRIMAL 52. Introduction As with cast aluminum alloys, intermetallic compounds can form from the melt and from the supersaturated solid solution of the solidified aluminum of wrought alloys. This paper deals with the development and application of a new and simple method for determining the intermetallic AlFeSi-phases and differentiating between a- and ί-AlFeSi-phases in wrought aluminum alloys. Because of deterioration of the mechanical properties the plate like structured ί-AlFeSi-phase should be avoided or minimized in extrusion billets as well as in most of the other aluminum products. Instead the curved crystals ("Chinese script") of the ocAlFeSi-phase is preferred, Figure 1 and Figure 2.

mm ' Figure 1. Curved structured oCc-AlFeSi-phase

A quick determination of the different AlFeSi-phases by considering the structure shape via microscopy as in Figure 1 and Figure 2 is often impossible due to the low amount of AlFeSiphases and the small dimension of these phases in typical 6xxx alloys. Consultations of universities and institutes yield agreement in principle that the most accurate possibility to differentiate between a- and ί-AlFeSi-phases is an EBSD-measuring (EBSD: Electron Backscatter Diffraction). The results are very reliable, but a scanning electron microscope (SEM) with an EBSD is expensive and not always available. Additionally, this method is relatively complex. Therefore the target of the project described here was to develop a simple and fast method to differentiate between a- and ί-AlFeSiphases with an EDX-measurement of element concentrations (EDX: energy dispersive X-ray) via the relation of silicon and iron content in the AlFeSi-phases.

711

Table 1: The variance of the prevalent intermetallic phases in 6xxx series [1] Phase

Structure

Stoichiometry

occ(a)-AlFeSi

cubic

Al12Fe3Si, Al12. isFesSil^

cubic

Al12(FeMn)3Si, Al15(FeMn)3Si2

cubic

Al12Mn3Si, Al15Mn3Si2, Al9Mn2Si

ah(oO-AlFeSi

hexagonal

Al8Fe2Si

ί-AlFeSi

monoclinic

Al45FeSi

Figure 2. Plate-like ί-AlFeSi-phases a- and ί-AIFeSi-Phases The typical difference between the a- and ί-AlFeSi-phases is the structure of the phases as visible in Figure 1 and Figure 2. This is the most convenient method for differentiating between the phases. One of the most important disadvantages of the ί-AlFeSiphase is the sharp and plate-like structure of the ί-AlFeSi-phase (Figure 2). The EBSD-measurement uses the different crystal structures of the phases. In the following table the variance of the prevalent intermetallic phases in the 6xxx series is shown. These different phases have to be differentiated [1]. The stoichiometry of the different phases and especially the relation between silicon and iron in the different AlFeSi-phases are described for wrought alloys in various papers [2, 3]. The different Si/Fe-relations for fv and β-AlFeSi-phases ΐ.Griger found out are visible in the diagram in Figure 3. Because the hexagonal och-AlFeSi-phase crystallizes only in high purity alloys the investigations focused on the ac- and β-AlFeSi-phases [2].

Figure 3: The typical Si/Fe-relation of etc- and ί-AlFeSi-phases [2].

24 Fe+Mnwt.%

Figure 4: The simplified diagram with the targeted separator between a- and ί-AlFeSi-phases The target of this project was to show that a simplified diagram as in Figure 4 can be used with the EDX-method for the determination of a- and ί-AlFeSi-phases. This diagram shows the designated separator between the a- and ί-AlFeSi-phases which is expected to be found via EDX-measurements of a- and ίAlFeSi-phases in wrought aluminum alloys. For an accurate analysis the manganese content in the phases is added on the iron

712

The specimens for the a-AlFeSi-reference were produced with the parameters shown in Table 2. At 200°C the simulation predicted only a-AlFeSi-phases. At least the second version (V4) should yield the phase constitution, which is given by the simulation. It was chosen considering the theoretical cooling conditions in Pandat. The predicted forming of the AlFeSi-phases can only be reached with a very slow cooling in the solid state. Therefore a very slow temperature reduction to 200°C was chosen to give exclusively aAlFeSi-phases in the specimens. Additionally a long holding time at 200°C gives time for forming the predicted a-AlFeSi-phases. After the holding time the phases are frozen by water cooling.

content due to the manganese content of the ac-AlFeSi-phases, which could be an additional indication to differentiate between the a- and ί-AlFeSi-phase. Investigations All simulations in this project were carried out at TRIMET with Pandat 8.1 (Version 2009) with the database 8 (Version 2009) from CompuTherm LLC. The main aspects for the alloy composition simulation and development are: • only a- or ί-AlFeSi-phases are present in the alloy to get reference alloys for microscopy, EDX- and EBSDmeasuring • the phases are large enough for a microscopic determination of the different phases. The first alloy ideas for the starting simulations were taken from different papers [1, 2] and former projects at TRIMET. These alloys are reported or proved to form only a- or only ί-AlFeSiphases. The simulations yielded a couple of alloy compositions with promising simulation results for each phase type (a- and ίAlFeSi-phase respectively). But due to problems with homogenization or other production and casting parameter not all of these alloy composition fulfilled the targets in reality. Below only the best alloy compositions for the a- and ί-AlFeSi-phase respectively are described. All simulation diagrams show the last simulation results for the real cast alloy composition. All important alloying elements (11 elements) were considered.

700 600 500 400

Ü

Γ

300 200 100J TAO+AL13CRi+AL3TI+FCC_Al+BEIA_ilfbm+ALPHA_ > lft«+AL13_FEMN3SI2+M02SI

All casting specimens were cast in steel moulds for cylindrical specimens with the following dimensions: • Diameter: 40mm • Height: 35mm The small dimensions make it easy to control the temperatures during the casting and cooling process. The target of the casting experiments was to get the reference specimens, which were simulated in the first step. These specimens allow the EDX-measurement method to be proved for the differentiation of a- and ί-AlFeSi-phases.

0.8

0.9

1.1

1.2

1.3

1.4

1.5

w%(FE) Element wt%

Si 0,8299

Fe 0,9264

Mn 0,1222

Mg 0,8403

Cr 0,2736

Zn 0,0996

Cu 0,3446

Figure 5: Pandat simulation and composition of the alloy with the target of only a-AlFeSi-phases in the specimen Table 2: Casting parameters for the α-AlFeSi-reference specimens

All specimens were analyzed via microscopy and SEM. The phases were analyzed manually via microscopy and with EDX. As additional specimens, two 6xxx-alloys from TRIMET production were taken. In one alloy predominantly a-AlFeSiphases were expected, but a minor fraction of ί-AlFeSi-phases should be present in some samples. In the other a majority of ίAlFeSi-phases are assumed.

Parameter

Value V3

Value V4

Melt temperature

730°C

730°C

Mould temperature

RT

730°C

Cooling

free air cooling

cooling down in 12h: from 730 to 200°C, Holding time:

q-AlFeSi-Phase

12h, 200°C,

The alloy presented in Figure 5 is based on previous TRIMET projects. The addition of manganese should favor the crystallization of the a-AlFeSi-phase. After some variations the simulation result concerning the alloy listed below shows a narrow area with only ct-AlFeSi. Between approx. 300°C and 120°C only α-AlFeSi is displayed. Based on the theoretical cooling conditions in Pandat, a very fast temperature reduction to 200°C-250°C should give exclusively a-AlFeSi-phases in the specimen.

Water cooling

ί-AlFeSi-phase: The synthetic alloy with only Si and Fe in 99.95% aluminum described in Figure 6 is based on Â.Griger [2]. This alloy is not so complex both concerning alloy composition and simulation result. With the high Si/Fe ratio a-AlFeSi-phases should not be present. The simulation result shows that between 615°C and room temperature, there will be no a-AlFeSi, detail to 570°C in Figure 6.

713

especially against the (Fe+Mn) values, Figure 9. The following facts can be observed: • the oc-AlFeSi points group very well and very closely to the Si/Fe-ratio given by Δ.Griger [2] • the correlation between the ί-AlFeSi measurements and the Si/Fe-ratio given by Δ.Griger is reasonably visible. • the spread of the ί-AlFeSi measurements is significantly higher • between the a- and the ί-AlFeSi measurements a free area can be found

Summarizing the simulation results we can presume that an exact alloy compositions, adequate casting and cooling parameters should result in reference specimens with only a- and ί-AlFeSiphases respectively. The casting parameters from Table 3 were used while casting the specimens with the alloy given in Figure 6. Both versions allow a very fast cooling below 610°C. Below this temperature the simulation gives only ί-AlFeSi-phases. Therefore a-AlFeSi-phases should be avoided.

.„ OoU-j

LIQUID+FCC_A1

620 J

LIQUID+FCC_Al+ALPHA_alfesi

610-j O l·-

600 ΐ

1

FCC_Al+BETA_alfesi+LIQUID

590 J 580 J 570 -I 0.9

FCC Al+DIAMOND A4+BETA alfesi

, 1

-T 1.1

, 1.2

Figure 7: Plate like phases in the ί-AlFeSi reference specimen.

1 1.3

w%(FE) Element wt%

Si 3,0914

The spread of the ί-AlFeSi measurements can be explained with the relatively thin phases and the alloy composition. EDXmeasurements of element concentrations are imprecise when the target phase is thin. An additional explanation is the free silicon, which could be found in the ί-AlFeSi reference especially near ίAlFeSi-phases. Silicon can be measured as a background contamination.

Fe 0,9802

Figure 6: Pandat simulation and alloy composition of the alloy with the target of only ί-AlFeSi-phases in the specimen. Table 3: Casting parameters for the ί-AlFeSi-reference specimens Parameter

Value VI

Value V2

Melt temperature

730°C

730°C

Mould temperature

RT

590°C

Cooling

immediate water cooling

free air

Results A brief look at the results found in this project, during the inspection and analysis of the specimens, showed a very good outcome. The first target was to produce reference specimens with only a- and ί-AlFeSi-phases respectively. This point was checked via simulation and microscopic analysis. In Figure 7 and Figure 8 it is shown that in the reference specimens only the desired phases are present. Next important step are the EDX measurements of the reference specimens. In these measurements the Si, Fe and Mn-content was determined. The Si values were plotted against the Fe values and

714

•

Determine whether the measured phases are a- or ίAlFeSi-phases The number of measuring points and measured phases can be low (3 measuring areas each with 3 measured phases) if there is only a- and ί-AlFeSi-phases respectively. If the specimen contains ocand ί-AlFeSi-phases and a determination of the fraction of phases is desired more measurements are needed. The determination normally takes less than one hour. Table 4: Detail of a matrix of EDX-measurements (alloy: TRIMAL 52) 0

8

16

24 Fι + Mn wt.%

32

40

4*

Figure 9: Si/(Fe+Mn) ratio from both reference specimens. Additionally plotted: the Si/Fe ratio for the a- and ί-AlFeSi-phase as described by Â.Griger [2], and the targeted separator between a- and ί-AlFeSi-phases. However, of utmost importance is the free and well defined separating area between the a- and ί-AlFeSi measurements. Therefore a separator can be defined for the decision, if AlFeSiphases are the a- or the ί-type. The final step was to verify the new method with the production samples. EDX-measurements were carried out. These measurements are displayed in Figure 10. In alloy A and B predominantly cc-AlFeSi-phases were expected. The measurements correlate for most of the AlFeSi-phases. These measurements are on the oc-side of the separator. But one can see that a lower content of ί-AlFeSi-phases is present in this sample. In alloy C all AlFeSi-phases are from the ί-type. For all alloys the measurements concentrates around a straight line. The measurements of alloy C are distributed in higher silicon values, analogue to the reference specimen.

Figure 10: Si/(Fe+Mn) ratio from the 2 production samples. Additionally plotted: the targeted separator between a- and ίAlFeSi-phases. In summary the method described in this paper is as follows: • EDX-measurements of the specimen (3-10 measuring points each with 3-10 EDX measurements) resulting in a matrix of points like visible in Table 4 • Plot the Si/(Fe+Mn) ratios of the matrix into a diagram including the separator line

TRIMAL 52 114-0153 Mn[wt°/<j SI [wt%] Fe [wt°/<j Fe+Mn [wtttj 20,27 6,69 1,19 21,46 6,7 16,83 2,36 19,19 8,55 21,93 3,52 25,45 4,23 9.75 1,31 11,06 8.07 22,82 2.31 25,13 6,85 17,94 2,17 20,11 3,66 1,17 8,77 7,6 6,09: 14,74 2,2 16,94 7,57: 20,95 2,53 23,48 4,89! 14,91 1,36 16,27 0,67: 0,44 0,44 0 2,47 6,88? 18,93 21,4 5,67 15,8 2,16 17,96 2.02 3,61 0.9 4,51

This new determination method for a- and ί-AlFeSi-phases was used recently for the development of the new TRIMAL 52 alloy [4]. This alloy for high performance space frame constructions was developed to fulfill the demands of the AUDI TL116-C28. Due to the demand for highest mechanical properties ί-AlFeSiphases had to be avoided in the TRIMAL 52, Figure 12. The amount of the ί-AlFeSi-phases is one part of the alloy properties, which helps to attain the best mechanical properties. The main mechanical data are: • Rp0,2: 281-330MPa • Rm: > 305 MPa • A5: > 10%

Figure 11 : Excellent crush behavior of extrusions with the new high strength aluminum alloy TRIMAL®-52 [4]

20

S5 12

i

^r^

'* 8

yf

4

* TRIMAL 52

«

« V

8

y/øú*

0 i

0

8

16

24 Fé + Mn wt.%

32

40

48

Figure 12: Si/(Fe+Mn) ratio from the new TRIMAL 52 alloy. Additionally plotted: the targeted separator between a- and ίAlFeSi-phases. Discussion and Conclusions The EDX-measurement of silicon and iron/manganese concentration is a sure and inexpensive method for a determination of a- and ί-AlFeSi-phases in wrought alloys. The method is verified via the simulation results and the microscopic determination of the phases in the reference specimens. Additionally EBSD-measurements of the reference specimens and the production samples are planned for an additional validation. Limitations of the EDX-measurement principle have to be considered. Because of the relatively inexact element measurement and the penetration of the measurement in the specimens, a statistically relevant number of measurements must be carried out. Measurements with a very low amount of silicon and iron have to be eliminated due to measuring errors. With a higher number of measurements, the straight line of the element relation becomes visible and a determination of the phase character is possible. Considering these points the EDX-method is an excellent tool for differentiating between a- and ί-AlFeSi-phases in wrought alloys. References 1. N.C.W.Kuijpers, "Kinetics of the ί-AlFeSi to a-Al(FeMn)Si transformation in Al-Mg-Si alloys" PhD Thesis, ISBN 90-7717207-6, Delft, 2004 2. Â.Griger, V.Stefβnia, A.Lendvai, T.Turmezey, "Possible modification of cast structure by continuous casting technology in AlFeSi alloys Part III: Intermetallic phases" Aluminium, 10, 1989, Giesel Verlag GmbH, Isernhagen, 1989 3. A.L.Dons, "AlFeSi-particles in Commercial Pure Aluminium" Zeitschrift für Metallkunde 75 2, pp 170-174,1984 4. M.Rosefort, H.Gers, T.Koehler, J.Ehrke, D.Schnapp, H.Koch: TRIMAL 52® - A New Aluminum Alloy for High Performance Spaceframe Construction. Proc. TMS 2010, The Minerals, Metals and Materials Society, Seattle, 2010

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S CAST SHOP for ALUMINUM PRODUCTION

Dross Formation, Control and Handling SESSION CHAIR

Pierre Le Brun Alcan CRV Voreppe, France

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

OXIDATION OF AlMg IN DRY AND HUMID ATMOSPHERES Darcy Stevens1, Anne Kvithyld1, Thorvald Abel Engh2, Shawn Wilson1 ^INTEF Materials and Chemistry, Trondheim, N-7465, Norway Norwegian University of Science and Technology, Trondheim, NO-7491, Norway

2

Keywords: aluminum, oxidation, carbon residue, humid atmospheres machined surface as compared to a surface polished with 600 grit emery paper because of a larger effective surface area.

Abstract Thermal oxidation of A15%Mg was carried out in dry and humid atmospheres with a thermogravimetric analyzer (TGA), and the composition of off-gases was measured with a mass spectrometer. It was found that when water vapor participates in oxidation, hydrogen gas is produced. A humid atmosphere (7 mol% H20) does not affect total oxidation mass, but seems to catalyze oxidation in both air and air containing 5 mol% C0 2 . In an air atmosphere containing 5 mol% CO2, total oxidation is approximately 50-75 % less than in air alone. Some samples in humid air (7 mol% H20) containing 5 % C0 2 did not oxidize significantly, which may be due to a strong oxide layer which forms around the liquid aluminum sample. A carbon powder coating was shown to protect the aluminum from oxidation for an extended period of time compared to an unprotected sample and humidity extended this protection. This work has potential applications in industrial thermal decoating units.

Material and Sample Preparation The composition of the alloy oxidized (Table I) was determined by inductively coupled plasma optical emission spectroscopy (ICP-OES). Table I. Composition of Sample Material. Element Quantity [wt9c] Mg 4.62 Si 0.07 Fe 0.25 Mn 0.33 Ti 0.02 The samples were machined into discs from a large ingot of the alloy. Table II contains the average dimensions of the samples after final preparation. It is assumed in this work that the effective surface area is the same as the geometric surface area because 1 micron polish gives a flat surface. This assumption is validated by experimental data gathered by Cochran and Sleppy [10].

Introduction In aluminum (re)melting, the formation of oxide i.e., dross, reduces the yield. The formation of dross is higher in a hotter melt, due to increased oxidation. The effect of humidity and carbon residue on oxidation is examined. Aluminum is lost not only by oxidation, but also from the large metal content of the dross [1,2]. The high oxide strength [3] allows for pure liquid aluminum to be trapped within the dross during skimming. If the formation of dross can be minimized, recycling yields increase.

Table II. Important Sample Dimensions. Dimension Average Value Standard Deviation Diameter 7.98 mm 0.02 Thickness 1.93 mm 0.03 Surface Area 148.5 mm2 1.0 Mass 257.4 mg 1.60

Some recyclers use a thermal decoating process, which consists of two stages; (i) evaporation of the volatile organic compound leaving carbon residue behind and (ii) reaction of this carbon with oxygen in the atmosphere [4]. Carbon powder is used to simulate carbon residue during oxidation.

Electric discharge machining was used to form the sample discs. This technique was necessary as the material was too soft for mechanical machining at the small size required. To clean the samples and homogenize the surface, manual grinding and polishing down to 1 micron was carried out. Between each grinding and polishing stage, the samples were cleaned with soapy water and rinsed with water, ethanol, and acetone, then dried. In order to keep the samples cylindrical, only the top and bottom surfaces were polished; the sides were left asmachined.

Aluminum-magnesium alloys are known to oxidise more readily than pure aluminum [5,6] and are investigated in a TGA with humidity control and a mass spectrometer. Attention has been given to the work of Impey, whose thesis contains an impressive literature review, including work by Thiele on the nature of aluminum oxide formation and growth [7,8]. Thiele's review gives an overview of the fundamental research which has been conducted, as well as more complex systems of oxidation. Thiele points to the inherent difficulty of studying aluminum oxidation.

The samples were wrapped in lens paper to protect the polished surface, rather than isolate the samples from humidity. The time between sample preparation and experimentation varied between several hours and several days. Samples were photographed, measured, and weighed before each experiment, to determine the change which occurred during

Surface finish has an effect on the oxidation of aluminum, as reported by Aziz and Godard [9]. Oxidation increases on a

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oxidation. Figure 1 shows a sample as machined, after polishing, and after oxidation.

A schematic of the TGA apparatus is shown in Figure 2. The carrier and auxiliary gases come from a common source and have identical compositions. The dry carrier gas flows through the balance head (10 ml/min) to protect the balance from exposure to humidity and mixes with the auxiliary gas before passing through the furnace tube and around the sample. The auxiliary gas passes through the WetSys apparatus (50 ml/min) and flows into the TGA above the furnace tube, giving identical flow patterns for dry and humid atmospheres. The mass spectrometer samples gases from just below the sample crucible in the furnace tube. 100 ml/min argon is introduced at the furnace outlet, diluting the humid gas to prevent condensation.

Figure 1. Sample progression; as machined, as polished, and after oxidation. The preparation of the samples is a source of error in this work. The samples were ground using SiC abrasive paper, so SiC particles were most likely embedded in the soft aluminum. It is unknown whether the SiC particles affected oxidation of the aluminum, but SiC oxidation was assumed to be insignificant at the temperatures of this work. The ethanol and acetone used to clean the sample may have affected oxidation behavior.

The interior of the furnace and gas flow patterns is shown in Figure 3. Note that sizes are approximate.

For experiments conducted with a carbon coating, Thermax® Powder Ultra-pure N991 was used. The powder was applied manually by rubbing the surface with small amounts of the powder using tweezers. The powder did not adhere well to the polished surfaces, so unpolished samples cut with a band saw were used. It was important to use only carbon powder without an adhesive agent, such as ethanol, to avoid contamination of the off-gases.

Furnace Tube

Apparatus To study the effect of atmosphere on the oxidation of aluminum, a Setaram Setsys 2400 TGA, coupled with a Setaram Wetsys humidity generator and Pfeiffer Quadrupole mass spectrometer was used. The TGA measured the mass change of a sample with respect to temperature and time. More in-depth descriptions of the fundamentals of TGA can be found in Brown [11] and Speyer [12].

Auxiliary Gas

Figure 3. Diagram of the sample within furnace tube.

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The top surface of the sample was assumed to be the most active surface, but the sides of the sample are clearly exposed and oxidation was also present on the underside of the sample.

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Experimental Procedure

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The prepared samples were measured and weighed before oxidation. Three measurements of the diameter and thickness are taken with a Vernier caliper and averaged. These measurements were used to calculate the surface area. The initial and final masses of the sample were obtained by weighing the sample crucible with the sample and subtracting the mass of the empty sample crucible. It was assumed that the crucible mass would not change during an experiment.

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Figure 2. Schematic of TGA Apparatus.

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The solid line shows the temperature profile, while the dashed lines indicate when the humidity is switched on and off. Humidity reached the desired value shortly after being switched on, effectively instantaneous for the time scale of the experiments. Time zero was chosen at the beginning of the 800°C isotherm.

For carbon coated samples, the amount of carbon used was determined by weighing the sample before and after coating, as well as the carbon powder itself. The two values were averaged, giving an approximate mass of carbon on the sample. Alumina crucibles, 10 mm tall with a 12 mm inner diameter and a 0.75 mm wall thickness were used, connected to the balance with platinum suspension wires. With each run, some oxidized material remained in the crucible that could not be cleaned out. This material accumulated in the crucible and fused to the sample which prevented it from being removed, approximately every four runs. New crucibles were used when the sample could not be removed from the old crucible.

At -8 minutes, when the sample was at approximately 550°C, the humidity was switched on. This was done before the start of the 800°C isotherm because it was found that oxidation occurred before the holding temperature was reached. Humidity was not introduced at the beginning of the experiment as the water would condense, damaging the delicate balance head. The humidity was switched off during cool-down of the furnace before the dew point of the gas was reached.

The sample and furnace chamber was evacuated and backfilled with argon before beginning each experiment.

Baselines were run for each atmosphere with empty crucibles, and were subtracted from sample runs to isolate sample oxidation data.

Samples were heated in the atmospheres shown in Table III. The air used was synthetic air, not ambient air.

Results and Discussion

Table III. Oxidation Gas
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Figures 6 and 7 show the results of two dry and two humid oxidation experiments, respectively.

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There was some difficulty in oxidizing the aluminum in air containing carbon dioxide. It can be seen in the figures above that several curves show minimal oxidation. Indeed, out of 11 experiments, only 3 oxidized significantly. It is tempting to conclude that the limited oxidation is due to the presence of the 5 mol% carbon dioxide.

Significant hydrogen signals were not obtained for oxidation in humid air, therefore water did not react. The aluminum reacted preferentially with oxygen, forming an oxide layer which the water vapor could not react with. Out of two experiments in humid air, only one oxidized significantly. Samples were oxidized in air containing 5 % carbon dioxide. Dry and humid conditions were used, Figure 8 and 9, respectively. Figure 8 contains data from 4 experiments and Figure 9 contains 7 experimental sets of data.

An oxide layer forms on the surface of the melted aluminum, trapping liquid aluminum inside a solid oxide layer. This is similar to what occurs during dross skimming [1]. The layer must crack or break, exposing fresh aluminum, in order to obtain significant oxidation. The curves in Figures 8 and 9 which show only a slight mass gain correspond to the formation and thickening of this oxide layer.

Samples oxidized in air containing 5 % carbon dioxide gained less mass that in air, for both dry and humid conditions. Hydrogen production was not observed for oxidation in humid air and air containing carbon dioxide. While the water did not react, it appears to have catalyzed the reaction as the mass gain for humid conditions occurs sooner than in dry conditions.

Figure 10 shows two oxidized samples. The sample on the left did not oxidize significantly, while the sample on the right gained significant mass.

722

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Future Work Future work will focus on oxidation of industrial samples with an unmodified surface. Samples with a well understood surface modification, such as anodizing, are of interest. More experiments of carbon coated samples in various atmospheres will help with understanding mechanisms. A longer holding time will be used to see how long carbon protects the surface.

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To improve understanding of surface oxidation, a detailed comparison of pre and post surface oxides will be carried out. Characterization of the surface before and after will help to describe the mechanisms of oxidation. Acknowledgement The authors would like to thank Hydro Aluminum, Alcoa, SAPA Heat Transfer and the Research Council of Norway for funding through the BIP RIRA project.

Conclusions Mass gain by oxidation and composition of off-gases has been measured simultaneously in a TGA. When water vapor is the only source of oxygen available for oxidation, the water reacts with the sample and hydrogen gas is produced. Oxidation in air neither increases nor decreases with the presence of water. Oxidation is delayed in a dry environment compared to a humid environment. An atmosphere containing carbon dioxide, with and without humidity resulted in approximately 50-75 % less oxidation than in air. Only 2 out of 10 experiments resulted in significant oxidation, which takes place only after the oxide layer has cracked or broken. Humidity may prevent the layer from breaking. In both air and air containing carbon dioxide, humidity seems to initiate oxidation earlier. The off-gas measurements indicate that water most likely did not participate in oxidation in air or air with C0 2 . Water may act as a catalyst, initiating oxidation but not participating in the reactions. A carbon coating delays oxidation of the sample. Oxygen preferentially reacts with carbon powder to form carbon dioxide. Humidity seems to extend this carbon protection compared to a dry environment. More experiments are needed to confirm this. In industry, the formation of dross reduces yield by entrapment of molten aluminum [1,2]. The presented results suggest that metallic aluminum droplets are enclosed in alumina oxide layers, when, upon cracking or bursting, fresh aluminum is oxidized. Furthermore, aluminum is often coated with organic substances. Decoating units are used to remove the coating before thermal melting. The results here indicate that a carbon residue should be left on the surface of the scrap, protecting it from oxidation. A specially designed decoating procedure could achieve this after the organics of the coating have decomposed. This would reduce dross formation and increase recycling yield.

References I. O. Manfredi, W. Wuth, and I. Bohlinger, "Characterizing the Physical and Chemical Properties of Aluminum Dross**, JOM, 11 (1997), 48-51. 2 Q. Han et al., "Dross formatin during remelting of aluminum 5182 remelt secondary ingot (RSI)", Mat. Sci. and Eng., A363 (2003), 9-14. 3. Martin Syvertsen, "Oxide Skin Strength on Molten Aluminum", Metallurgical and Materials Transactions B, 38B (2006), 495-504. 4. A. Kvithyld, "Thermal Decoating of Aluminium Scrap" (Doctoral Thesis, Dept. of Mat. Tech., Norwegian University of Science and Technology, July, 2003). 5. A. Kanti et al., "Numerical simulation of early stages of oxide formation in molten aluminium-magnesium alloys in a reverberatory furnace", Modelling Simul. Mater. Sei. Eng., 12 (2004), 389-405. 6. E. Bergsmark, C. J. Simensen, and P. Kofstad, "The Oxidation of Molten Aluminium", Mat. Sci. and Eng., A120 (1989), 91-95. 7. S. A. Impey, "The Mechanism of Dross Formation on Aluminium and Aluminium-Magnesium Alloys", (Doctoral Thesis, School of Industrial Science, Dec. 1989). 8. W. Thiele: Aluminum, 38 (12) (1962), 707-715. 9. Philip M. Aziz, and Hugh P. Godard, "The Influence of Surface Pretreatment on the Atmospheric Oxidation of 2S (U.S. Alloy 1100)". Aluminum, Journal of the Electrochemical Society, 104, (12) (1958) 738-739. 10. C. N. Cochran, and W. C. Sleppy, Oxidation of High-Purity Aluminum and 5052 Aluminum-Magnesium Alloy at Elevated Temperatures, Journal of the Electrochemical Society, 108 (4) (1962)322. II. Michael E. Brown, Introduction to Thermal Analysis, 2nd edition (Kluwer Academic Publishers, 2001). 12. Robert F. Speyer, Thermal Analysis of Materials (CRC Press 1993). 13. T. Abel Engh, Principles of Metal Refining, (Oxford University Press, 1992) 332. 14.1. Akagwu, and R. Brooks, "Environmental Reaction of Liquid Aluminium", (Paper presented at Conference for the Engineering Doctorate in Environmental Technology, 2003).

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Study of Early Stage Interaction of Oxygen with Al; Methods, Challenges and Difficulties B. Fateh(a-1,3), G.A. Brooks(al), M. A. Rhamdhani^^ J. A. Taylor(a'2), J. Davis(3) and M. Lowe(3) (a) CAST Cooperative Research Centre (CAST CRC) (1) HTP Group, Faculty of Engineering and Industrial Science, Swinburne University of Technology (2) School of Mechanical & Mining Engineering, University of Queensland (3) The Centre for Atom Optics and Ultrafast Spectroscopy (CAOUS), Swinburne University of Technology Keywords: Al oxidation, chemisorption, laser ablation, fast imaging (iii) Matter transports through the oxide layer by solid sate diffusion of one or both of the reactant (in this case metal or oxygen)

Abstract Aluminum is among the metals with the greatest affinity for oxygen. Aluminum oxidation occurs easily and rapidly; a surface film can be formed in order of milliseconds. Study of initial oxidation is very important as it concerns the first step of a metal's oxidation and corrosion behavior. Characteristics of the kinetics of early-stage oxidation are believed to have an important influence on the later steadystate growth. Modern sophisticated experimental studies such as scanning tunneling microscopy (STM), X-ray photoelectron spectroscopy (XPS), low energy electron diffractions (LEED), combining ellipsometry and auger electron microscopy have clarified some of the details of initial stage of both oxygen chemisorption and oxidation kinetics of Al. However, the experimental data are not consistent. Furthermore, the dynamics of absorption and oxidation process are much less understood. This paper reviews the methods, techniques and the challenges of Al initial stage oxidation process measurements and also introduces a new method for studying the oxide layer formation using ablation of the surface with an ultra-fast laser beam and fast imaging techniques.

(iv) Matter exchanges between diffused matter and the interfaced ions (internal semi-reaction [4]). In this case, the thickness of the oxide layer increases with time but the oxidation decelerates and finally reaches to stable plateau behaviour. This explanation reveals that the first step of oxidation is expected to be governed by chemisorption but the last steps are expected to be controlled by diffusion process. Previous experimental work indicate that oxidation of metals are initially very rapid and follow linear first order kinetics. The initial fast linear oxidation is thought to be controlled by chemisorptions [6] .Through further exposure to oxygen, the initial fast linear oxidation changes to slower logarithmic or parabolic growth rate behaviour [6]. Aluminium has a high affinity for oxygen and its oxidation occurs easily and rapidly. At room temperature Al is always covered with an oxide film of 2-3 nm in thickness [23]. At high temperature, the oxidation process is extremely fast and thus, it is very difficult to predict or measure the mechanism of oxidation at the initial stages.

Introduction

According to the collision and kinetics theory of gases, for 1 cm2 of a clean surface (which has about 1015 atoms) at room temperature and at the pressure of 1 atm., it will take nanoseconds to form a monolayer of the oxide, if we assume that all the oxygen that arrives at the surface reacts instantaneously. Therefore, a clean oxide-free surface is seldom found in nature and must be prepared. Defining a clean surface is a challenging step in studying the initial stage of oxidation. To prepare a clean surface and to remove surface scale, techniques such as in-situ sputtering or heating in hydrogen are employed [1]. Since normal weight gain methods are not sensitive enough to detect any changes that occur in time frames of less than a second, in-situ fast response surface analytical techniques should be employed for the measurement of Al initial oxidation. This phenomena has been extensively studied by different methods and techniques using mono-crystalline surfaces and in ultra high vacuum (UHV) chambers: such experimental techniques are low energy electron diffraction (LEED)[2,3], ellipsometery techniques [7-11], auger electron spectroscopy (AES) [12], scanning tunnelling microscopy (STM) [13-17], electron energy loss spectroscopy (EELS) [18-20], photoelectron spectroscopy (XPS) [21] and also medium energy ion scattering [22]. Results of these investigations show that oxygen is chemisorbed and then gradually converts to oxide. However, the details of the process are still unknown and

There are many aspects of Al melt handling and cast process that result in oxidation of Al and dross formation. Dross is an inhomogeneous complex mixture of aluminium oxides along with entrapped molten metal. Since dross contains 10 to 80 wt% of entrapped Al, poor melt handing practices can result in significant profit loss. Of particular importance is the generation of dross during pouring operations, when "fresh" metal is exposed to oxygen as a free surface is formed. Opportunities for dross reduction may be enhanced by a greater understanding of the kinetics of aluminium oxidation especially in molten phase. When a clean surface of a metal is exposed to oxygen, the following processes are considered to occur: (i) Oxygen is adsorbed on the surface and forms an adsorbed layer (chemisorption). (ii) Matter exchanges between the adsorbed layer and the surface. Thus, the first layer of oxide appears through formation of islets of oxide nuclei. This process is also named as external semi-reaction by some authors [4]. It is considered in the majority of low temperature oxidation theories that both adsorption and oxidation occurs predominantly at grain boundaries and then spreads over the whole surface [5].

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also there are great discrepancies among reported data for the observed onset of oxidation, oxygen sticking coefficient and coverage. We will discuss this issue latter in this paper.

Figure 3 depicts the plot of changing thickness of Al and Al-oxide in time, based on the Lindmark et al [7] measured ø and Δ data.

In this paper, we review the several published papers on chemisorption and oxidation of Al surface, discuss some of the inconsistencies and finally introduce a new technique for visualizing the evaluation of surface change after being exposed to oxygen.

Two regimes can be seen in their data: a sharp transition from near zero to several angstrom thicknesses and then a slow logarithmic growth. The authors claim that the first regime describes the diffusion of metal ions to the surface and in the second regime the oxidation is limited by the tunnelling of electrons.

Literature review on early stage Al oxidation The initial interaction of oxygen with Al solid surface has been studied by a number of groups in the past few decades. Lindmark et al employed an in situ ellipsometer for measurement of thin film aluminium oxidation [7]. Their experimental set up included three chambers: sputter, robot handler and oxidation chambers. In the sputter chamber, aluminium layers, with 3000 A thickness, were deposited on top of silicon wafers. Then, the robot arm handler took a wafer and moved it through the robot handler chamber into the growth chamber. The pressure in the robot handler chamber was 8xl0"8 torr. After the growth of the film, the wafer from the sputter chamber was moved by the robot arm, through the robot handler chamber and into oxidation chamber which had a base pressure 2xl0"8 torr. This movement took about 1 minute.

-10

-5

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5

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Tune (minutes)

Figure 2. Plot of raw ø and Δ data versus oxidation time, zero minute is the time when oxygen was let into the chamber [7]

The highly vacuumed oxidation chamber was equipped with a spectroscopic ellipsometer with a high intensity xenon light source lamp. The lamp produced light with wavelengths from 280-760 nm (a schematic of the experimental layout is shown in Fig. 1). In order to measure growth rate, oxygen was introduced into the oxidation chamber with the ultimate oxygen pressure of 20 torr and the ellipsometer was set to take data at 1 or 5 second intervals. 0

5

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Figure 3. Plot of aluminum oxide and aluminum thickness versus oxidation time [7] Kotenev [5, 9, 10] combined ellipsometry and electron microscopy for identifying the initial stages of Al oxidation (see Figure 4). A vacuum chamber with an electron spectrometer and an ellipsometer were combined. The experimental set up has been designed for complex investigation of the initial stages of gas-phase oxidation by means of ellipsometry, electron auger spectroscopy and lowenergy electron diffraction with the usage of layer-by-layer etching of the surface with a beam of argon ions. In detail, his apparatus was set up as follows: a monochromatic radiation beam of laser (1) passed through a polarizer (2) and an electro-optical phase modulator (3). The phase shift between the orthogonally polarized radiation components with a frequency from a generator (6) was modulated. After a semi reflecting mirror, through the window (11), the radiation beam was incident on the specimen (12) placed in a high-vacuum analytical chamber (10) of the electron spectrometer. Then the beam being reflected by a concave mirror (13), it returned upon itself to a beam splitter (4), was reflected and passed through an analyzer (7), to a photoreceiver (8). The output signal of the receiver was sent to a control processing unit (9).

Figure 1. Schematic of the oxidation chamber outfitted with ellipsometer and UV lamp [7] This set up has a certain disadvantage. The aluminium film suffers some oxidation during its journey from the sputter chamber to the oxidation chamber. Both Fig (1) and Fig (2) show oxide thickness prior to the time zero, the time when oxygen introduced to the chamber. As the authors pointed out even without the introduction of oxygen into the chamber, oxide growth was still seen. Thus, to measure oxide growth from a zero thickness on clean aluminium, an ultra-high vacuum system is required, which was not apparently the case in their study. Their report shows that during the exposure ø decrease and Δ (ø and Δ are two important ellipsometeric parameters which tan^) is the amplitude ratio and Δ is the phase shift upon reflection) increases sharply with time sharply (Fig. 2).

726

In the first chamber (16), an Al film of thickness 2000 ΐ was sprayed from a 99.99% Al source on a glass substrate. Then the sample moved to an analytical chamber (10) by means of a magnetic system. The chamber was under a vacuum of 10"10 torr. In order to make a fresh clean surface of Al, immediately before the experiment, the surface was cleaned by an argon beam. This is the advantage of Kotenev design compare to the previously described Lindmark set-up.

(Fig. 8). Figure 8 depicts that the initial fast stages end at L < 100 with a thickness of around 10-12 nm. The slower stage corresponds to three dimensional growth of the oxide film. Hayden et al [8] also studied oxygen chemisorption and oxidation of single crystal aluminium under different temperature and pressure conditions (range of 295-800 K and 10"9-102 torr, respectively) using an ellipsometry technique.

Auger-spectrometer

MP*

Figure 4. The set up for combining ellipsometric and electron-spectral measurements [9] Figure 6. The stages of early stage oxidation (Ellipsometric data) [9]

Figure 5 shows the experimental change of the ø and Δ as a function of Langmuir value, L (1 L = 10 "6 torr. s) in a constant pressure of 3xl0"7 torr. Despite of Lindmark results, both ø and Δ decrease with time. The ø curve has an explicit minimum and all changes in Δ are negative.

dAη/dL, arb. units

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Figure 5. Change in the ellipsometric parameters ø and Δ during the interaction between aluminum and oxygen of the rarefied atmosphere at room temperature vs. the dose of oxygen (1 L = 10 "6 torr. s) [10]

to

i

i. »,..» i.l.U..

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1. ..».,.1 ÁËËË, 1000 L

Figure 8. The change in the thickness of oxide layer at room temperature [10] The authors reported that before each measurement, Al surface was electro-polished in a solution of perchloric acid and ethyl alcohol and then in-situ cleaned for preparing a single crystal Al. Their experiment carried out in an ultrahigh vacuum (UHV) chamber equipped with LEED, Auger electroscopy and an automatic ellipsometer. Fig. 9 depicts their results of changes of ø and Δ as a function of oxygen exposure (L) for various Al crystallographic faces at 5x10~7 torr and 295 K. It seems that, the results of Hayden et al confirmed the decreasing behaviour of both ø and Δ. Their data reveals different oxidations rates for various Al single crystal faces, such as Al (111), Al (100) and Al(l 10).

Kotenev [5 ' 9 ' 10] sophisticated experimental set up revealed that the interaction between Al and oxygen at room temperature can be divided into 4 different stages. He claimed that the first stage corresponds to the formation and development of the first oxygen adsorption mono-layer, the second stage shows the growth of the second oxygen layer, the third stage belongs to formation of oxide nuclei and the fourth stages reveals the islet-like growth of a three dimensional amorphous oxide. These four stages are shown in Figure 6 and 7. Figure 6 shows the measured data of ellipsometry technique and Figure 7 the changes in augerpeak intensities and both are in good agreement in way of showing existing four stages of initial oxidation.

The decrease of ø and Ä as a function of oxidation and type of Al crystal face was also reported by Reichel et al [11], at 350 K and lxlO"4Pa. Figure 10 shows the kinetics of ultrathin (<1.5 nm) oxide-film growth on bare Al(100) and

The change in thickness of the oxide layer as a function of Langmuir value, L, has been measured by Kotenev et al

727

Figure 11 shows the semi-logarithmic plot of slowly increasing thickness of oxide film on an Al(l 11) substrate.

Al(llO) and Al(lll) substrates measured by Reichel et al [11] . The figure shows that the oxidation kinetics on Al(100) and Al(l 10) can be subdivided into an initial, very fast and a subsequent very slow oxidation stage. However, on Al(lll) at T< 450 K, no distinction between initial fast and a slow stage could be determined. Reichel et al [l 1] stated that this behaviour leads to an unexpected Al oxide growth at low temperature with respect to the rate of oxidation.

p*5Kr'Torr T » 295 K

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-0,1 -0,2-0,3-j £ <

Hayden et al [8] found that at lower temperatures (less than 600 K) the oxidation rate follows a logarithmic rate law and holds a proportional dependence to the square root of the pressure (p05). They also reported that at higher substrate temperature (600-800 K) the oxidation kinetics follows a parabolic relationship (Fig. 12). These measurements were carried out at a constant pressure of lxlO"4 Torr and calculated from changes of ellipsometric parameters.

%:

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exposure (Langmuir)

Figure 9. Changes of ø and Δ during the oxygen interaction wit various Al crystal faces at 295 K and 5xl0"7 Torr [8]

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^* K /*m*>

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Time (x10 _3 s)

Figure 12. Plot of slow oxidation stage of Al{ 111} at p= lx 10 4 Torr [8]

r=35QK

0.0

A«{110}

0-9

Benka and Steinbatz in 2003 [17] studied the oxidation of Al using secondary electron emission. The authors performed experiments in a UHV chamber with a base pressure of lxlO"10 mbar. Their Al sample was a 100 nm evaporated Al layer on a silicon wafer. In order to get a fresh non-oxidized Al surface, the samples were sputtered cleaned by Ar ions, but they had no means of measuring the structure of the Al surface. They found that electron emission yield depends on both oxygen pressure and exposure. They also claimed that at the beginning of oxygen exposure, the yield decreases and this behavior is due to chemisorption of oxygen, but when exposure continued the yield increased due to oxide formation (Results from their studies are shown in Figure 13).

AI{111> AK100}

r=450K Al{110}

> . ^ Ì a ^ ^ ^ À M H A»{100} Γ-500Κ ί(10'Λ8)

Thus, the minimum yield corresponds to the end of chemisorption and the onset of oxidation process. As the figure shows, for pressure greater than lxlO"7 mbar, the onset of oxidation process occurs at 25 L (1L= 10"6 torr.s). Other researchers have reported different values for the onset of oxidation from 100 L [10] up to 250 L [19]. A large discrepancy also exists for the reported sticking coefficient

Figure 10. Experimental and fitted oxide-film growth for the oxidation of Al crystal faces at 350 K, 450 K and 500 K and and po2=lxlO' 4 Pa[ll] Hayden et al [8] also reported the kinetics of the slow stages of Al(lll) oxidation at 525 K under different pressures.

728

(neutral density filter) is used for adjusting different beam intensities.

of oxygen adsorption on Al, with values from 0.005 to 0.015 reported [14, 20]. Ubi)

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049 0.48 0.47 0.46 0.45 0 44

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Figure 15 shows the live captured images of time evolution of the oxide layer in 10 ms steps just after laser ablation. In the center of each image, the dark area depicts ablated surface of Al which is assumed to be a fresh Al. Image (a) was taken prior to ablation process and (b) was taken at the end of the ablation. Images (b), (c) and (d) were taken at 10 ms time intervals. Comparison between these images (inside the circle) clearly shows that surface is turning gradually to the color of surface prior to ablation, which most probably relates to the formation of an oxide layer on the fresh Al surface.

r

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1

■1

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i

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1

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L.

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,

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,

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This result is promising and proves that this technique is a potentially suitable method to study the fast oxidation process in the millisecond time scale. Further studies on the image analysis techniques and measuring the rates of molten Al oxidation is in progress. For molten A oxidation studies, a high temperature stage is being incorporated into the apparatus described above and the result will be presented in subsequent publications.

exposure [LJ Figure 13. Electron emission yields of Al as a function of exposure time for different oxygen pressure from 2 x 10"9 to 3xl0"6[17] Introduction of a new technique Laser ablation is a technique for mass removal by coupling laser energy to the surface of a material. In recent years, femto-second laser ablation attracted a lot of attention because of small heat-affected-zone on the surface of target material [24].

tat tßmt _ g B 0 C p î l |

A new method for studying the initial stages of oxidation, where a fresh surface of Al is prepared by removing the oxide layer using an ultra-fast laser ablation technique, has been developed at Swinburne University of Technology. Following ablation of the surface, the kinetics of solid Al oxidation is quantified by means of a fast imaging.

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For the ablation of solid Al surface, the laser experimental set up was optimized. The sample was placed in a vacuum chamber with a silica window (50 mm diameter) for optical access. The chamber was connected to a rotary pump producing a vacuum of 2 mtorr (low vacuum range). Laser ablation and illumination light sources were focused through two different optical lenses (Fig. 14). A high speed camera was used to image the sample and was synchronised to a fast mechanical shutter in the ablating laser beam. In order to make shallow holes, the number of shots and laser intensity were optimised. The live images of ablated Al surface and the formation of sparks during laser ablation showed that 40 shots with 7-8 mW average power were enough to remove the oxide layer on the solid Al specimen. In order to capture images by high speed camera, a very strong light source was necessary. Thus, illumination of Al surface was a challenging issue. Focusing a very bright light source, using lenses and a fibre optic light, were optimized. The camera was adjusted to take 3000 frames per second in full resolution. Finally, the time evolution of the oxide layer growth on the ablated surface was recorded. Figure 14 shows a schematic set-up for the laser ablation and fast imaging technique. As the figure shows an ultra-fast laser beam (pulse duration in order of 100 femto-second =10"13 s) hits the Al surface through the objective lens. The BBO crystal (ί-barium borate) is a crystal used for frequency doubling and coverting a red laser beam (wavelength= 800 nm) to a blue laser beam (wavelength= 400 nm). The blue laser allows more accurate focusing and more efficient absorption of the laser energy bu the material surface. A beam splitter

ß UgM kam {Her optk o&h)

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Figure 14. Schematic set up of laser ablation and fast imaging technique for visualizing the oxide formation on solid Al specimen

Figure 15. Live images of Al surface taken by fast camera (3000 fps) a) Surface prior to ablation, b) surface just at the end of ablation c) 10 ms after ablation d) 20 ms after ablation

729

7. E.K. Lindmark, JJ. Nowak, M.T. Kief, Proceeding of SPIE Vol. 4099 (2000), pp 218-227

Conclusion The interaction of solid Al and oxygen is rapid and complicated and it occurs in a stepwise process that is controlled by several parameters such as surface defects, crystal orientations and the method of preparing the fresh surface. For those reasons, there are inconsistencies between various data reported by different researchers. Table (1) reports the values of the measured Longmuir parameter (L= 10"6 torr.s) at which oxidation of aluminium surfaces commences, by different authors.

8. B.E. Hayden, W. Wyrobisch, W. Oppermann, S. Hachicha, P. Hofmann and A.M. Bradshaw, Surface Science, Volume 109, No. 1, Aug (1981), pp. 207-220 9. V.A. Kotenev, Protection of metals, Vol. 36. No.4, (2000) pp. 366-374 10. V.A. Kotenev, Protection of metals, Vol. 36. No.5, (2000) pp. 409-418

Table 1. Reported L values for the onset of Al oxidation Author's Name Kotenev [5, 9,10] Hayden et al [8] Benka and Steinbatz [17] ReicheUifl/[ll] Zhukov et al [19]

11. F. Reichel, L.P.H., Jeurgens, E.J. Mittemeijer, Acta Materialia 56 (12), (2008), pp. 2897-2907

L value <100 for Al (111): 110 for Al (110): 30-50 for Al (100): 230-300 25 200-250 250

12. L. Lauderback, S. Larson, Surf. Sci. 233 (1990) 276 13. H. Brune, J. Wintterlin, R. Behm, G. Erti, Phys. Rev. Lett. 68 (1992) 624 14. H. Brune, J. Wintterlin, J. Trost, G. Erti, J. Chem. Phys. 99 (1993) 2128 15. J. Trost, H. Brune, J. Wintterlin, R. Behm, G. Erti, J. Chem. Phys. 108 (1998) 1740

Thus, there is considerable uncertainty in the literature about the point which Al oxidation commences. In general, all the experiments and varied techniques are carried out under ultra high vacuum conditions. There is no data available for oxidation process under low vacuum or normal pressure, nor is any kinetic data available for early stage molten Al oxidation. In this paper, a new technique has been introduced which may potentially provides better, faster surface treatment and also can be applied for measuring molten Al oxidation rate under lower vacuum or even normal pressure. This technique may allow observation of the initial stages of oxidation on very short-time scales, too.

16. M. Schmid, G. Leonardelli, R. Tschelieίnig, A. Biedermann, P. Varga, Surf. Sci. 478 (2001) L355 17. O. Benka, M. Steinbatz, Surface Science 525 (2003) 207-214 18. J. Crowell, J. Chen, J. Yates, Surf. Sci. 165 (1986) 86 19. V. Zhukov, I. Popova, V. Vomenko, J. Yates, Surf. Sci. 441 (1999) 240

References

20. V. Zhukov, I. Popova, J. Jates, Surf. Sci. 441 (1999) 251

1. A.S. Khanna; Introduction to High Temperature Oxidation and Corrosion; ASM International (2002), No. 06949G, ISBN: 0-87170-762-4

21. C. McConville, D. Seymour, D. Woodruff, S. Bao, Surf. Sci. 188 (1987) 1

2. H.L. Yu, M.C. Munoz and F. Soria, Surface Science 94(1980),L184-L190

22. D.O. Connor, E. Wouters, A. Denier van der Gon, J. Vrijmoeth, P. Zagwijn, W. Slijkerman, J. Frenken, J. van der Ven, Surf. Sci. 287/288 (1993) 438

3. E.K. Lindmark, JJ. Nowak, M.T. Kief, Proceeding of SPIE Vol. 4099 (2000), pp 218-227 4. P. Sarrazin, A. Galerie and J. Fouletier, Mechanism of High Temperature Corrosion: a kinetic approach, Volumes 36-37 of Material Science Foundations (2008), ISBN: 0-87849-484-7

23. P. G. Sheasby, R. Pinner, The Surface Treatment and Finishing of Aluminum and its Alloys, 2 (sixth ed.), Materials Park, Ohio & Stevenage, UK: ASM International & Finishing Publications, (2001) ISBN 0904477-23-1

5. V.A. Kotenev, Protection of metals, Vol. 39. No.5, (2003) pp. 415-423

24. S.H. Kim, I. Sohn, S. Jeong, Applied Surface Science 255 (2009) 9717-9720

6. F.P. Fehlner and N.F. Mott, Vol. 2, No. 1, (1970) pp. 59-67

Correspondent author's email: [email protected]

730

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

QUALITY ASSESMENT OF RECYCLED ALUMINIUM Derya Dispinar1, Anne Kvithyld1 and Arne Nordmark1 ^INTEF Materials and Chemistry, Trondheim, N-7465, Norway Keywords: Recycle, aluminium, metal quality, mechanical properties, remelting, bifilm index Melting procedures

Abstract One of the serious problems during remelting of aluminium is the presence of surface oxide including coating. In this work, a wrought and a cast alloy were selected and subjected to remelting experiments. 3000 wrought alloy sheets with three surfaces; (i) before anodising, (ii) between anodising and coating and (iii) after coating, was investigated. The bifilm index was measured as a measure of metal quality; 3-point bending and tensile testing samples were collected for mechanical testing. A good correlation between the mechanical properties and the bifilm index was found. For the surface treated sheet skimming reduces the bifilm index. After skimming the melt in (i), (ii) and (iii) have the same quality, that is a comparable bifilm index and mechanical properties. However, the bifilm index of the cast alloy decreased after remelting three times, thus decreasing the quality.

6 kg of each coil material was melted at 720°C in an induction furnace. 10 reduced pressure tests samples (Fig 1) were taken for metal quality check (i.e. bifilm index measurement). 10 cylindrical bars were cast for tensile testing (Fig 2) and 10 plates were cast for 3 point-bending tests. 65 kg of premium grade primary ingots were melted in a resistance furnace at 740°C. The melt was poured into plate-molds where the surface to volume ratio was designed to be high (80x20x700mm). All of these plates were then charged into the crucible and remelted in an induction furnace at 740°C. This procedure was repeated 3 times. Sampling was as described above. Reduced Pressure test (bifilm index)

Introduction

Molten aluminium is poured into a sand mould with dimensions given in Figure 1, leaving the metal to solidify under vacuum at 100 mbar, which enhances pore formation. The bifilm index [13] is the sum of the maximum length of the pores; giving a total oxide length for a given surface area Figure 2. A rule of thumb: is 10 mm; best quality, 1 0 - 5 0 mm: good and over 50 mm: bad metal quality.

Environmental concerns, particularly increased energy cost and consumption of natural resources have led to production by recycling. It is a well-known fact that recycling requires about 5% of the energy needed for primary production. Today recycled aluminium accounts for one-third of aluminium consumption world-wide. Nevertheless, there is a long going discussion about the quality issues of the secondary aluminium. One of the problems during remelting of the aluminium scrap is contamination from surfaces as well as the surface oxide of the scrap itself. After the melting process, approximately 10% of the charge is lost due to these oxides and removal of the slag [1-4]. It was also shown [5-7] that turbulent transfer and pouring of the melt increases the metal losses even more. During the melting of the charge in the crucible, the surface oxide of the material may thicken, becoming often micrometres or even millimetres thick [812]. Thus, the recycling/remelting of aluminium is not straight forward and simple, it requires extra attention. Since quality is of central importance for the final product, a series of remelting experiments were carried out in this work. The metal quality change was assessed by a reduced pressure test using the bifilm index [13] and employing mechanical tests.

Figure 1. Dimension of sand mould sample for the reduced pressure test

Experimental procedure Materials 0.5 mm thickness wrought 3000 coils (Table 1) were collected from different stages in the coil coating production: (i) before anodising (ii) between anodising and coating, (iii) after coating. Premium grade primary ingots of A356 alloy (Table II) were provided by Alcoa Norway. Figure 2. The maximum length of the pores are measured. Added values give the bifilm index.

731

Table I. The chemical composition of wrought 3000 coils Cu Mn Mg Cr Ni Zn Pb Ti 0.58 0.20 0.68 0.34 0.02 0.01 0.23 0.01 0.02

Si 0.53 Si 6.9

Fι

Table II. Chemical composition of cast alloy / ^356 Ti Sr Mn Fι Na 0.0012 0.002 0.116 0.11 0.0005

Mg 0.32

Al rem. P 0.0002

Al rem.

Effect of skimming Ten reduced pressure test samples were collected in each experiment. First two samples were collected before skimming. It was found that the bifilm index was high for the two samples. This indicated a high oxides content for (iii) anodized and coated (Fig 5). After the dross is skimmed off, the melt quality is observed to remain constant for the next four measurements as seen in Figure 5.

Gμ I L

«

J

Figure 3. Dimension of the tensile test pattern The oxide formed over the surface of the coil material was characterized by SEM and TEM Result The oxide layer thicknesses analyzed on the surface of the 500 μπι coil material are: (i) Untreated surface: 10 nm aluminum oxide layer (ii) Anodised surface: 200 nm aluminum oxide layer (iii) Coated surface: Primer : 2.2-3.0 μπι both sides Topcoat + 25 μπι front side Topcoat + 35 μιη back side

0

before skimming

after skimming

5 min

10 min

15 min

Figure 5: Effect of skimming observed in RPT Aluminium sheet The average bifilm index measurements of 3000 coils collected from the different section of the production line are given in Figure 6. As seen in figure, the bifilm index results for (i) 'untreated', (ii)4 anodized' and (iii)'anodized + coated' charges have values in the same range of 9 mm, 10 mm and 15 mm respectively.

Recovery 6 kg of 3000 charges ((i) untreated; (ii) anodized; (iii) coated were melted in each experiment, and after the melting, the dross was skimmed off. When the sample collection was complete, the dross and the remaining material were weighed separately, and the melting yield was calculated, shown in Figure 4.

25 100

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Figure 4: Melting yield of the 3000 coils with different surfaces (i), (ii) and (iii).

(Ȕ)

Figure 6: Bifilm index of remelted sheet materials.

732

Mechanical results for (i), (ii) and (iii) are given in Figure 7. Both the ultimate tensile results (a) and the elongation at fracture (b) values appear to be same within the uncertainty limits. For the max bend strength (c), the (iii) 'anodized + coated' material seems to have a slightly lower value.

Three times re melted

Figure 8: Bifilm index change remelted of A356

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1

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140

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(b) 320

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(c) Figure 7: Mechanical test results of the sheet materials (a) ultimate tensile strength, (b) elongation at fracture, (c) 3-point bend test Ingots The bifilm index change of the ingot material after three times remelting is given in Figure 8. It was not practical to skim the surface. Premium quality primary A356 had a bifilm index of 13 mm. After three times remelting and pouring into plate shapes with a high surface to volume ratio, the bifilm index significantly increased to 110 mm. A similar result was found with the mechanical tests (Fig 9). The ultimate tensile strength was 166 MPa and it dropped down to 150 MPa; with elongation at fracture dropping from 2.7% to 1.9%. Max bending strength was also decreased from 345 MPa to 319 MPa.

Three times remelted (C)

Figure 9: Mechanical test results of the ingot material (a) ultimate tensile strength, (b) elongation at fracture, (c) 3-point bend test

733

Discussion An unexpected result concerning metal quality of 3000 was the observation of similar bifilm index values for the melt from the three different surface finish samples. As seen in Figure 5, the bifilm index results for (i) 'untreated', (ii) 'anodized' and (iii) 'anodized + coated' charges were 9 mm, 10 mm and 15 mm, respectively. It was expected that especially the (iii) 'anodized + coated' material would have the worst quality due to the thick oxide layer and the coating. Not surprisingly, since they all had the same bifilm index, the tensile properties of these castings were also similar with ultimate tensile strength around 120 MPa and elongation at fracture around 5% (Fig 7).

(a)

These results can be explained as follows: the materials were charged into the crucible in stacks as shown in Figure 10 (a). When the temperature rises inside the crucible, the plates retain their form of stacking due to the rigid oxide structure. However, aluminium with its low melting point compared to the oxide that surrounds it will melt and settle at the bottom of the crucible. After all the material inside the plates is drained to the bottom of the crucible, the surface oxide of the original charge remains and is collected as dross on the surface (Fig 10b). This is schematically shown in Figure 11.

I (b)

(c) Figure 11. Schematic representation of melting procedure: (a) Charging of the materials into the crucible (b) start of the melting, (c) completion of melting and dross formation at the surface The bifilm index changes from good to bad for the ingot material after three times remelting as shown is given in Figure 8. The mechanical properties also decreased significantly. The reason is probably that removal of oxides by skimming was hardly possible as the remelting procedure was different than the wrought alloy. However, the correlation between bifilm index and mechanical properties of the ingots (cast alloy: A356) is also in good agreement (Figs 8-9).

(a) (b) Figure 10. (a) Charge of the materials in the crucible (b) Completion of melting and dross formation at the surface

It is important to note that fluxing or degassing was not carried out. Removal of surface oxide by skimming has a strong effect on melt quality. Since skimming is problematic, metal refining to remove oxides will be even more important for remelting than for primary production.

Tests from surfaces (i), (ii) and (iii) employed aluminiumfromthe same original coil. The metal quality proved to be the same for all melting experiments as seen in Figure 6 (i.e. same bifilm index). This result is also supported by the reduced pressure test results. As seen in Figure 5, RPT samples collected from the melt before skimming have a high bifilm index, which indicates a high content of oxides. When the dross is skimmed off, the melt quality is observed to be constant and good in the next four measurements.

Conclusions

The effect of surface treatments is observed to be critical for the melting yield. The reclaiming (mass) ratios were 96%, 90% and 70% for (i) untreated, (ii) anodized and (iii) anodized and coated materials, respectively (Fig 4). The metal loss was highest in (ii) anodized and (iii) coated materials, as expected.

1. Metal quality after skimming is the same for all three coil materials (i) untreated, (ii) anodized and (iii) anodized and coated.

Overall, there was a good correspondence between bifilm index and the mechanical properties of the coil material. This can be seen in Figures 6 and 7. Interestingly, the correspondence is most clear for the the max bend strength of the sheet material seen in Figure 7. Here a slight decrease in the max bend strength is observed for (iii) 'anodized + coated' material which has the highest bifilm index.

3. There is a good correlation between bifilm index and mechanical properties.

2. The loss of metal to dross is high for the (iii) anodized+ coated aluminium.

4. Bifilm index of primary cast alloy increased from good to bad after three times remelting, with no skimming. The mechanical properties decreased significantly.

734

Acknowledgment The authors acknowledge the financial support of SINTEF for this work. Also thanks are due to Hydro Aluminium Rolled Products AS, Holmestrand and Leif Ivar Koksvik for providing materials, and to Alcoa for supplying the cast material. Thorvald Abel Engh is acknowledged for assistance.

References 1.

2.

3. 4.

5.

6.

7.

8. 9.

10.

11. 12. 13.

Gronostajski, J. and A. Matuszak, The recycling of metals by plastic deformation: an example of recycling of aluminium and its alloys chips. Journal of Materials Processing Technology, 1999. 92-93: p. 35-41. Tenorio, J.A.S., M.C. Carboni, and D.C.R. Espinosa, Recycling of aluminum - effect of fluoride additions on the salt viscosity and on the alumina dissolution. Journal of Light Metals, 2001. 1(3): p. 195-198. Samuel, M., A new technique for recycling aluminium scrap. Journal of Materials Processing Technology, 2003. 135(1): p. 117-124. Amini Mashhadi, H., et al., Recycling of aluminium alloy turning scrap via cold pressing and melting with salt flux. Journal of Materials Processing Technology, 2009. 209(7): p. 3138-3142. Dispinar, D. and J. Campbell, Effect of casting conditions on aluminium metal quality. Journal of Materials Processing Technology, 2007. 182(1-3): p. 405-410. Verran, G.O. and U. Kurzawa, An experimental study of aluminum can recycling using fusion in induction furnace. Resources, Conservation and Recycling, 2008. 52(5): p. 731-736. Prakash, M., P. Cleary, and J. Grandfield, Modelling of metal flow and oxidation during furnace emptying using smoothed particle hydrodynamics. Journal of Materials Processing Technology, 2009. 209(7): p. 3396-3407. Sleppy, W.C., Oxidation of molten high-purity aluminum in dry oxygen. Journal of the Electrochemical Society, 1961. 108(12): p. 1097-1102. Impey, S.A., D.J. Stephenson, and J.R. Nicholls, Mechanism of scale growth on liquid aluminum. Materials Science and Technology, 1988. 4(December): p. 1126-1132. Seamans, G.M. and E.P. Butler, In situ observations of crystalline oxide formation during aluminum and aluminum alloy oxidation. Metallurgical and Materials Transactions A, 1975. 6A(November): p. 2055-2063. Fuoco, R., et al., Characterization of some types of oxide inclusions in aluminum alloy castings. AFS Transactions, 1999. 107: p. 287-294. Tenorio, J.A.S. and D.C.R. Espinosa, High-temperature oxidation of Al-Mg Alloys. Oxidation of Metals, 2000. 53(3/4): p. 361-373. Dispinar, D. and J. Campbell, Critical assessment of reduced pressure test. Part 2: Quantification. International Journal of Cast Metals Research, 2004. 17(5): p. 287-294.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

IN-LINE SALT-ACD1M: A CHLORINE-FREE TECHNOLOGY FOR METAL TREATMENT Patrice Robichaud1, Claude Dupuis1* Alain Mathis2, Pascal Cτtι3 and Bruno Maltais^ j

Rio Tinto Alcan, Arvida Research and Development Centre, P.O. Box 1250, Jonquiθre, QC, Canada G7S 4K8 2 Rio Tinto Alcan, Aluval, 725 Aristide Berges - B.P. 27 - 38341 Voreppe Cedex - France Sociιtι des Technologies de ΓAluminium du Saguenay Inc. (STAS), 1846, Outarde, Chicoutimi, QC, Canada G7K 1H1

3

Key Words: Aluminum Processing, In-Line Treatment, Salt Fluxing, ACD composed of a Flux Feeder for Degasser (FFD ) combined with the ACD™, is considered as the last technological step required to achieve a 100% chlorine-free casthouse. The same principle of in-line salt injection has been applied to other technologies with positive results [8].

Abstract A new generation of the Alcan Compact Degasser (ACD™), the Salt-ACD™, based on the utilization of salt fluxes in replacement of chlorine gas, was introduced to the aluminum industry [1]. This unique technology has been developed by Rio Tinto Alcan (RTA) since 2003 and in collaboration with La Sociιtι des Technologies de PAluminium du Saguenay (STAS). Its utilization in combination with the Rotary Flux Injector (RFI™) for furnace preparation and/or the Treatment of Aluminum in Crucible (TACm) for aluminum pre-treatment eliminates the use of chlorine in casthouses. The Salt-ACD™ technology has been successfully implemented and operated in RTA casthouses. It supports the objective of eliminating chlorine from the casthouse for health, safety and environmental reasons.

Since then, the deployment of the technology has been successfully completed in key Rio Tinto Alcan casthouses. From its original inception and prototype version, the FFD™ has now evolved to a mature industrial design well adapted for easy implementation and ensures reliable and precise operation. This paper presents recent advancements of the FFDtM technology, its application with the ACD™ and reports the metallurgical and operational performance.

This paper presents recent developments in terms of equipment, key components and retrofittability to existing ACD™ units. The operating experience and metallurgical performance are reviewed.

Key Technology Principles Optimized Treatment Zone Principles for optimizing the metallurgical performance of the Salt-ACD™ have been well established since the development of the first FFD™ prototype [1]. The key process characteristics are the interfacial contact area that is created between the metal and the dispersed salt droplets, the residence time of these salt droplets within the molten metal and the subsequent separation by flotation of the impurities to the metal surface [9].

Introduction In-line metal treatment of aluminum alloys is a key step to meet increasing customer's quality requirements for critical application products such as can stock, lithographic sheet, fine wire and micropore extruded tubes. Molten metal quality depends on the control of non-metallic inclusions, alkali/alkaline earth elements (Na, Ca, Li) and dissolved hydrogen. To maximize productivity, a casthouse must exploit every opportunity in the processing chain to treat the metal using the most efficient technology available. In a smelter casthouse, the three main time periods available for metal treatment are pre-treatment in the crucible for the removal of alkali elements and inclusions, treatment in the furnace after alloy preparation and in-line during casting of the product [2].

Figure 1 conceptually shows how the FFD™ is used in conjunction with the ACD™ to provide a precisely controlled flow rate of salt-gas mixture injected in the molten metal. The multi-stage design of the ACD m technology is a key feature to maximize the performance of the process. Injection of the saltgas mixture takes place through the first rotor of the ACD™ relative to the metal flow. The salt is dispersed into small droplets by the rotor and is further sheared by subsequent downstream rotors where the separation andflotationof impurities take place.

The evolution of metal treatment technologies over the years is a good reflection of the increasing demand in terms of quality, metal processing cycle time, investment and operational costs [37]. Another important priority of today's aluminum industry is the impact of our processes on the environment and worker safety. Chlorine is a reactive gas with a proven track record in the aluminum industry for molten metal fluxing, but it is also a toxic gas. Environmental pressure and increasingly severe legislation have pushed the industry and manufacturers to develop technologies based on the utilization of salt fluxes to replace chlorine.

The standard FFD™ comes as a volumetric controlled feeder system, which is sufficient to achieve the required dosing precision. As an option, it can be fitted with a weighing module to achieve Loss-in-Weight type control and positively record the amount of salt injected (in compliance with American EPA requirements, for example). With a controllable injection rate ranging from 1 to 30 g/min combined with the typical gas flow requirements of the ACD™ technology, the FFD capabilities cover a wide spectrum of process requirements in terms of aluminum treatment.

A new chlorine-free technology, the Salt-Alcan Compact Degasser (Salt-ACD™), developed at the Arvida Research and Development Centre of Rio Tinto Alcan since 2003 and La Sociιtι des Technologies de l'Aluminium du Saguenay (STAS), has been introduced to the aluminium industry. This technology,

737

significant difference in the alkali removal efficiency of the SaltACD™ obtained while using a salt flux mixture composed of 60% MgCls compared to 75% MgCl2.

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A pre-fused salt mixture of MgCl2 ^ d KCl is preferably used. Adjustment of the chemical composition and granulation to a specific size distribution results in salt particles that are most suited to maximize the chemical kinetics of the reactions taking place in the Salt-ACD™.

MgCl2 + 2Na — 2NaCl +Mg MgCl2 + 2 Li - 2LÌC1 + Mg MgCl2 + Ηa - CaCl2 + Mg

Production lots of salt are stringently controlled to ensure that residual moisture is always kept at minimum. The MgCl2 is a deliquescent salt, meaning that it will absorb moisturefromthe air. The deliquescence relative humidity of the MgCl2 under typical workplace conditions is approximately 33% [12,13]. At humidity levels above this, any contact with the salt will result in water contamination [14] and could potentially affect the performance of the salt-ACD™.

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Technology Presentation

Based on the KCl-MgCl2 phase diagram generated using FactSage™ presented in Figure 2, it can be seen that working with MgCl2/KCl mixtures close to one of the three eutectics at 36, 40 and 64% MgCl2 allows injection of a salt flux that transforms to the liquid state below the molten metal processing temperature.

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The FFD™ system comes as a stand-alone unit which includes the gas and electrical distribution panels as well as its own human machine interface. The gas distribution panel precisely doses the gas volume to efficiently convey the salt and gas mixture to the ACD™ rotor no. 1 relative to the molten metal flow. The salt dosing system, adapted with an optimized screw design, is secured in a hermetically sealed cabinet kept under dry atmosphere to prevent any humidity pickup by the salt. Figure 4 shows a perspective view of the unit.

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A considerable effort has succeeded in the development of a compact salt flux feeding system that can easily be added to a new ACD™ system or retrofitted to an existing unit. Economical aspects have also been considered in order to obtain an affordable package. Moreover, the retrofit was proven operational and efficient on all existing ACD™ configurations including sealed, non-sealed, and while employing a controlled atmosphere in the ACD™. The controlled atmosphere ACD™ consists in adding a small air flow in the ACD™ to produce a controlled atmosphere above the molten metal having about 2%-4% oxygen.

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It is well known that MgCl2 is the chemically reactive agent in the salt flux responsible for removing the alkali impurities according to reactions 1 to 3 while KCl is used to reduce the melting point of the salt mixture [10,11].

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738

The self-contained control panel includes all the necessary elements for reliable process control such as pressure monitoring, gas flow control, and a sophisticated motor controller for precise salt flow control in a very compact design. Industrial Experience The FFD™ has been implemented and operated for production in key Rio Tinto Alcan casthouses. The process is qualified for the production of numerous products with many customers. Operation of a FFD™ has proven to be stable with no problems related to salt accumulation (with sealed, non-sealed and controlled atmosphere ACD™). Figure 4. 3D View of the Flux Feeder for Degasser (FFD™)

The consistency of salt injection is monitored by measuring the pressure inside the flux feeder. To convey the salt mixture through the injection line, a nominal pressure is maintained within the unit during treatment. This pressure depends on many parameters such as the injection line dimensions, the salt flow rate, the rotor immersion depth in liquid metal, etc. Changes in pressure will indicate abnormal conditions such as line blockage. Figure 6 shows the typical pressure variation inside the injection line during the casting period while Figure 7 shows the initial pressure at cast start for 50 consecutive casts. As illustrated, the pressure of Cast no. 1 is very similar to that of Cast no. 50, indicating stable and reliable performance.

A system for continuous casting applications is also available. Integration of an airlock with the flux reservoir allows salt refilling without interruption of the process. A specific design of the ACD1M Rotor no. 1, where salt is injected, completes the system. A thermally insulated tube, placed inside the graphite shaft of the rotor maintains the salt mixture below its solidus until it enters a temperature zone above its liquidus thus preventing partial melting. The FFD™ can be installed in the vicinity of the ACD™, either on the ACD™ structure or floor-mounted on the shop floor, as shown in Figure 5. This arrangement allows for easy refilling of the unit at an ergonomie level without the use of platforms. It also provides direct access to the local controls of the feeder, as they are conveniently located directly on the unit.

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From a control point of view, benefits are achieved by the use of a dedicated local controller. The stand-alone configuration reduces the work required to link the FFD™ to the control panel of an existing ACD™. Specifically, hardware and software modifications are minimized in case of retrofits. However, expansion capabilities for connections to sophisticated level 2 systems are possible, including complete integration to the casting pit controls.

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Characterization of the metallurgical performance of the SaltACD™ was done by taking measurements and comparing the impurity levels at the inlet and outlet. Hydrogen concentration was measured using AÄSCAN™l. Alkali removal was measured using Optical Emission Spectroscopy measurements from horizontal disks, while inclusion concentration was measured using PoDFA.™.

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Hydrogen removal occurs as a result of diffusion mass transfer between the liquid metal and the dispersed gas bubbles that are generated by the rotary injectors [7]. The concentration gradient between the gas and liquid phases is the driving force. Parameters such as the interfacial gas-liquid contact surface area (the size and number of micro-gas bubbles) and residential time of the bubbles in the melt, affect the hydrogen removal rate. However, these are independent whether chlorine or salt is being injected with the gas.

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Alkali elements are removed by reaction with MgCl2 according to equations presented earlier. The kinetic factors affecting the alkali removal rate have been documented elsewhere [15]. It is generally accepted that the controlling mechanisms are directly dependent on the salt-melt interfacial area generated. Increasing the contact surface is achieved by optimizing salt droplet dispersal. Figure 10 shows the sodium removal performance when using stoichiometric ratios of 1 and ratios of 1.5 to 3 respectively. With a ratio of 1.0, the sodium removal efficiency is approximately 40%, while at stoichiometric ratios, the sodium removal performance increases to approximately 50%.

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When injecting an anhydrous salt, hydrogen removal performance remains essentially unchanged as compared to operation with chlorine. Figure 8 shows plant measurements which exemplify the transparency of converting from chlorine to salt use in the ACD™. As illustrated, H 2 levels (mL/100 g) and variation at ACD™ outlet are similar.

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Inclusion Removal The inclusion removal efficiency of the Salt-ACD™ varied between 69 and 88% over all industrial conditions tested as shown in Figure 11. This performance is the result of a good balance between the dispersal, dewetting and flotation mechanisms achieved in the multi-stage treatment approach of the Salt-ACD™ process [16,17].

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AÄSCAN is a registered trademark of Rio Tinto Alcan Inc.

740

Conclusions The Salt-ACD™ process has been successfully implemented in key Rio Tinto Alcan casthouses. A thorough understanding of the critical process parameters and a detailed process monitoring allowed reaching metallurgical performance equivalent to the utilization of chlorine. The FFD™ technology integrating key features for controlling the key parameters for ensuring a reliable injection of salt flux, has proven to be retrofitable to existing ACD units. Together with in-crucible treatment, as achieved with the TAC™ system, and the in-furnace treatment using the RFI™, the SaltACD1** technology completes the technology port-folio allowing for a complete elimination to the utilization of chlorine for metal treatment with no compromise on final metal quality for critical product applications.

Figure 11. Inclusion Removal Efficiency for Different Alloy Families Cumulated experience indicated that the inclusion removal efficiency achieved using the multi-stage Salt-ACD™ is equivalent or better than the ACD™ operated with chlorine gas.

Acknowledgment The authors wish to thank Rio Tinto Alcan for permission to publish the present paper. The authors also want to acknowledge the collaboration obtained with key Rio Tinto casthouses during the deployment of the technology.

Dross Generation No differences were observed regarding dross condition or quantity when comparing the Salt-ACDfM process with the ACD™ operated with chlorine. The dispersed salt is found to be a good dewetting agent to keep the dross dry and non-reactive.

References [1] Leboeuf, S. et al. "In-Line Salt Fluxing Process: The Solution to Chlorine Gas Utilization in Casthouses", Light Metals 2007, (The Minerals, Metals & Materials Society, 2007), 623-627.

Impact on Downstream Operations The in-line injection of salt using the Salt-ACD™ was observed to have no impact on the performances of downstream filtration technologies such as Ceramic Foam Filters and Deep Bed Filters. Figure 12 compares the metal cleanliness measured downstream of the filtration unit using LiMCA Π™ 2, before and after the retrofit of a FFD™ to an existing ACD™ . 1

[2] Le Brun ,P. "Melt Treatment - Evolution and Perspectives", Light Metals 2008, (The Minerals, Metals & Materials Society, 2008), 621-626. [3] Maltais, B. et al. "Metal Treatment Update", Light Metals 2008, (The Minerals, Metals & Materials Society, 2008), 547-552. [4] Gariιpy, B. et al."The TAC Process: A Proven Technology", Light Metals 1984, (The Minerals, Metals & Materials Society, 1984), 1267-1279.

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[5] Rasch, B. "Refining of Potroom Metal Using the Hydro RAM Crucible Fluxing Process", Light Metals 1998, (The Minerals, Metals & Materials Society, 1998), 851-854.

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[6] Dupuis, C. et al. "Rotary Flux Injection: Chlorine-Free Technique for Furnace Preparation", Light Metals 1998, (The Minerals, Metals & Materials Society, 1998), 843-847.

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[7] Waite, P.et al "The Alcan Compact Degasser: A TroughBased Aluminum Treatment Process Part I: Metallurgical Principles and Performance" Light Metal 1996, (The Minerals, Metals & Materials Society, 1996), 1001-1005.

Figure 12. Metal Cleanliness after Filtration

[8] Chesonis, C. et al "Chloride Salt Injection to Replace Chlorine in the Alcoa A622 Degassing Process" Light Metals 2008, (The Minerals, Metals & Materials Society, 2008), 569-574. [9] Waite, P. "A Technical Perspective on Molten Aluminum Processing", Light Metals 2002, (The Minerals, Metals & Materials Society, 2002), 841-848.

2

ACD, Salt-ACD, FFD, TAC, RFI, LiMCA II, PoDFA are trademarks of Rio Tinto Alcan Inc.

741

[10] D.H. DeYoung, "Salt Fluxes for Alkali and Alkaline Earth Element Removal from Molten Aluminum," Aluminium Cast House Technology (Proceedings of the 7th Australian Asian Pacific Conference, 2001), 99-113. [11] T.A. Utigard et al., "The Properties and Uses of Fluxes in Molten Aluminum Processing" JOM, 50 (11) (1998), 38-43. [12] Lietai Yang, et al, "Experimental Determination of the Deliquescense Relative Humidity and Conductivity of Multicomponents Salt Mixtures" Materials Research Society, 713. (2001), JJ11.4 [13] Greenspan L., "Humidity Fixed Points of Binary Saturated Aqueous Solutions", J. Res. Nat. Bur: Stand. (US.), 81A,1, (1977), 89-96 [14] G.J. Kipouros, and D.R Sadoway, "A Thermochemical Analysis of the Production of Anhydrous MgCl2", Light Metals (Elsevier Science Ltd, 1, 2001), 111-117. [15] Bilodeau, J-F. et al. "Modeling of Rotary Injection Process for Molten Aluminum Processing", Light Metals 2001, (The Minerals, Metals & Materials Society, 2001), 1009-1015. [16] R. R. Roy. et al. "Inclusion Removal during Chlorine Fluxing of Aluminum Alloys", Light Metals 1998, (The Minerals, Metals & Materials Society, 1998), 871-875. [17] R. R. Roy. et al. "Inclusion Removal Kinetics During Chlorine Fluxing of Molten Aluminum", Light Metals 2001, (The Minerals, Metals & Materials Society, 2001), 991-997.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S CAST SHOP for ALUMINUM PRODUCTION Melt Quality Control SESSION CHAIRS

Claude Dupuis Rio Tinto Alcan Saguenay, Quebec, Canada Steinar Benum Aloca Norway Mosjoen, Norway

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

THE EFFECT OF TIB2 GRANULES ON METAL QUALITY Maryam Mohamed Al-Jallaf1, Margaret Hyland2, Barry Welch 3, Ali Al Zarouni1, Fahimi Abdullah1 1

Dubai Aluminium Co, P.O.Box 3627, Dubai,UAE department of Chemical & Materials Eng., Univ. of Auckland N.Z 3 Welbank Consulting Ltd., N.Z. and S of Ch.Sci. & Eng. UNSW 2052 Aust. Keywords: TiB2, inclusions, metal cleanliness, granules Abstract and examined for presence of foreign material or inclusions by such methods. The characteristic of metal being free of inclusions is generally referred to as "metal cleanliness".

TiB2 granules were added to a fully graphitized electrolytic cell in a trial to provide a barrier coating on the carbon cathode to prolong cathode life. The consequential impact on metal cleanliness was evaluated by a detailed metallographic analysis using the PoDFA technique. Metal produced from the test cell was mixed with regular potline metal and cast into billets. Samples were taken from different locations in the process stream and also for three different types of metal charged into the furnace, namely regular potline metal, 25% of metal from test cell mixed with regular potline metal, and 50% of metal from test cell mixed with regular metal. The PoDFA analysis shows that samples containing metal from test cell had more grain refiner inclusions than regular potline metal but fewer carbide inclusions. However, there was no overall significant negative impact on the specified requirements of metal cleanliness.

Classification and effect of inclusions Different types of non-metallic inclusions and their effects on end product quality in aluminium have been described [2]. 1. Oxidefilmsclassified by length and thickness films. 2. Oxide flakes consisting of magnesium and aluminium oxides. 3. Spherical oxide particles. 4. TiB2 clusters. These are not considered very important. 5. Particles of salt inclusions soluble in water. 6. Other non -metallic inclusions such as carbide The effect of inclusions will depend on their nature [2].

Introduction

Since titanium boride is used as grain refiner in rod form, these agglomerations are considered less important. Titanium vanadium diboride (Ti,V)B2 is said to be hard in nature and small in size but the clusters they form are tolerated at levels much higher than oxides [9]. However, in the presence of oxide films and / or liquid chloride inclusions, they can result in complex agglomerates of 20 μ to a few mm and are extremely detrimental, especially in rolled products. Magnesium oxide inclusions are said to affect downstream process only if present in large patches. Spinel inclusions are also especially harmful to the process. Particles of salt contribute to the rejection rate of aluminium sheet and bar products. Carbide inclusions can degrade mechanical properties of thefinalproduct [10].

It has been shown in another paper published in these proceedings [1] that addition of TiB2 granules was effective in reducing cathode erosion significantly. But it was required also that addition of TiB2 should not affect hot metal quality in terms of it being inclusion free since significant deterioration would negatively affect the cast house products and consequently acceptability to the end user. To examine this aspect, the quality of hot metal tapped from the test cell was studied. Samples were taken from a control cell and from routine production for comparison. The samples were analysed using light microscopy combined with Porous Disc Filtration Apparatus (PoDFA) as well as Scanning Electron Microscope (SEM) / Energy Dispersive Xray analyser (EDX).

TiB2 inclusions / agglomerates

Inclusions and metal cleanliness - an overview

Since the trial involved addition of TiB2 granules to the test cell a more detailed discussion of its effect on metal quality is relevant. TiB2 particles when present in molten aluminium tend to form agglomerations. Since titanium boride is also used as grain refiner in casting process, this aspect is a concern for manufacturers and end users of products. Agglomerations of TiB2 can cause end product quality problems in certain applications since TiB2 particles are much harder than aluminium. Individually they are generally sub-micron in size, and when examined by microscope the grain refiner rods are rarely observed to be > 2 μ. The agglomerations formed can be an order of magnitude larger than the individual particles or even greater [11]. The affinity of TiB2 particles and oxide films to agglomerate is apparent from investigation of both the microstructure of grain refiners and autopsies of tears or pinholes in foil or bright trim products. The need to minimise the level of oxidefilmsin the casthouse has long

Any foreign phase occurring in aluminium is defined as a nonmetallic inclusion [2], based on the criteria that the phase already existed above the formation temperature of the first solid particle arising out of the melt. A recent thesis provides a good description of the various terms [3]. Many of these inclusions degrade the mechanical properties of the end product at aluminium foundries, extruders etc., [4,5] and also result in undesirable blemishes. In critical applications, any inclusion with diameter greater than 10 20 μ is considered to be deleterious although even smaller inclusions can cause problems if present in sufficient quantity. Inclusions present in fluxed and filtered aluminium metal typically range from 5 to 50 ppbv. Since this is extremely small, pre-concentration is required for metallographic study, and techniques like PoDFA are employed for this purpose [6,7,8]. Prior to the cast, from the launder, liquid metal can be sampled

745

casting table was also done to ensure minimisation of external contamination so that the effect of metal from test cell on billet quality could be detected easily.

been recognised [12]. In many applications avoidance of TiB2 agglomerations at the surface is important. Description of the cast house processes at the site

In Experiment 1, 25% cold metal from test cell was mixed with 75% molten metal from other cells in same potline. All metals were charged into 40 tonnes capacity furnace and alloyed. After achieving the desired alloy composition the furnace was kept on hold for 60 minutes prior to billet casting. During the casting some inline treatments were carried out. These included grain refiner addition [19] into the casting launder before degasser, degassing metal using Alcan Compact Degasser (ACD), and finally filtering the metal through 30 ppi Ceramic Foam Filter (CFF). Control 1 was for billet production done with routinely produced hot metal from the same potline but without any metal from test cell. In Experiment 2, 50% cold metal from test cell was mixed with 50% molten metal from other cells in same potline. Control 2 was for billet production done with hot metal from same potline but without any metal from test cell.

The process flow sheet is presented in Figure 1.

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Samples were taken from cells and at several points in the process using PoDFA unless specified otherwise (Figure 1).

Figure 1: Cast house process flow chart Typically, molten metal from an electrolysis cell is tapped in to a crucible of 5 -8 tonne capacity. The tapping frequency from each cell is once in 32 hours. Three to four cells are tapped to fill the crucible. Once the crucible is full, the molten metal is transported to the processing station for removing alkali metals, especially sodium by Removal of Alkali Metal (RAM) system [13,14] prior to reaching casthouse. Treated metal from several crucibles is poured in to the cast house furnace having capacity of about 50 tonnes. The molten metal is held in this furnace and mixed with alloying elements to manufacture special alloys.

Sampling from cells For the test cell, few days after the first addition of TiB2, hot metal sampling from the cell was started. Samples were taken every month for a period of one year. The control cell was sampled similarly. Pot metal elemental purity was monitored for both cells as a routine measure using in house methods [20]. Sampling from crucible -after RAM It was not possible to get 7.5 tonnes of metal exclusively from the test cell at one time. Hence 2.5 tonnes of metal from the test cell along with 5 tonnes of metal from neighbouring two cells had to be sent for RAM processing. Inclusions were measured before and after RAM treatment.

Metal is then passed through degasser to remove dissolved hydrogen gas and filtered [15,16] and finally cast in to extrusion billets or other products such as foundry ingots or sows. Prior to final cast, grain refiners [17,18] are occasionally added to improve the grain structure of thefinishedproduct.

Samples from casthouse for metal cleanliness Samples were taken from cast house furnaces while conducting Experiments 1 and 2. Furnace samples were also taken from routine production furnaces utilising hot metal from same potline for comparison. During casting, samples were taken using PoDFA at two time intervals. First set of samples was taken when cast length reached about 1.5 meter and another towards end of casting with another 1 meter left to cast. Samples were taken at outlet of furnace and CFF.

Adaptation of cast house process for experiments It was essential to study the metal cleanliness using metal from test cell. Since the cast house furnace required 40 to 70 tonnes per batch and the cell produced only 2.5 tonnes per tap, this study was challenge. To partly circumvent this problem, production from test cell was cast into sows of 500-650 kg. About 30 tonnes of metal from the test cell was thus converted to sows for further study in casthouse. Once the treated metal was cast into sow, RAM treatment had necessarily to be skipped. RAM treatment of the - 2.5 tonnes tapped metal from test cell was not possible due to the design limitations of the RAM process equipment.

Parameters measured Boron and titanium concentration These concentrations were measured for metal from furnace using arc spark emission spectroscopy.

These sows were used in two experiments at one of the furnaces in casthouse from which molten metal was sent for extrusion billet production. Metal cleanliness standard for extrusion billets is much more demanding than for most other products and hence it was decided to use the metal from the test cell in the billet production. Melting any other cold metal during the experiments was avoided to rule out interference. Prior to the experiments, the furnace was hot cleaned. In addition, general cleaning of other refractory such as launder, degasser's trough, filter box and

Concentration of inclusions by PoDFA Typical operation of PoDFA has been described in literature [4]. PoDFA disc of 25 mm diameter was polished using standard procedures [21,22,23] and viewed under light microscope at different magnifications to get an overall idea of the extent of inclusions in the sample. Photographs were taken.

746

Characterisation of inclusion types For every sample, in addition to total inclusions count, inclusions were also classified [24]. The inclusions were quantified by using light microscope and grid method. The counting for a particular field was done for every type of inclusion using Alcan catalogue of inclusions [9]. Identification of the inclusions was by light microscope [25] as well as SEM) / EDX. Characterisation was based on the interpretation by experienced metallurgists using information such as alloy type, sampling position, metallurgy process data, inclusion colour / shape, hardness and distribution.

Figure 2: Carbide inclusions-control cell (left) and test cell (right) Before and after RAM treatment Result presented in Table 2 is for metal from one crucible based on three cells including the test cell. The total inclusion counts as well as classification according to inclusion type is shown. The reduction seen in total inclusions is in line with findings of 43 % reduction in total inclusions after RAM treatment [27].

Results and discussion Metal sampled directly from cells The comparison of the results based on 12 samples is presented in Table 1 . Detailed classification data for inclusions was not available for all the 19 routine production samples, but based on available data it was seen that this was similar to control cell with about 89-96 % of carbides and about 5-10 % of alumina and other inclusions with no presence of (Ti,V)B2. After RAM treatment inclusions were reduced even further, so the levels seen in test cell did not pose any concern. Values for control cell are based on a single sample only.

Table 2: Metal cleanliness of crucible sample for test cell before and after RAM Unit Total inclusion in metal (Ti,V)B2

mm /kg

AI4C3

Table 1: Comparison of metal cleanliness at potlines

Number of samples Total inclusion in metal (Ti,V)B 2 | Oxide AI4C3 AI4C4B | Alumina Others

Control

No"!

Average(iSP) 12

cell 1

Prodn, 19

0.901 (0.513)

0.885

0.908 (0.308)

64(12)

0

2(2) 22(10)

0 90

4(5)

0

KD

5 5

-

7(3)

Alumina dispersed Others

Routine!

Test cell

% % % % % %

RAM

1.287

0.578

75

53

AI2O3

Unit

mm 2 /kg

Before RAM

31 G-

13

There is only 7 % (Ti,V)B2 for crucibles which were not RAM treated but 75 % aluminium carbides are seen. Since there is practically no (Ti,V)B2 in regular pot metal amount present in metal from test cell gets diluted since 2.5 tonnes of this gets mixed with 5 tonnes tapped from adjacent cells. Quantities of (Ti,V)B2 work out to 0.023 and 0.090 mm2/kg in metal from test cell and diluted cell respectively and this is approximately matches with the expected figure from dilution. After RAM it is seen that aluminium carbides as well as (Ti,V)B2 dropped and in aluminium oxides increased. It is seen that RAM system also reduced inclusions. At Cast house

This is further described in Figure 2. All the pinkish particles are boride inclusions and specifically (Ti,V)B2. The grey particles are carbide inclusions. For control cell it is seen that 90 % of total inclusions are carbide inclusion but for the test cell the aluminium carbide content dropped to 22 %. It is hypothesized that the TiB2 granules, which were added into test cell, settle down and cover the cathode in the bottom of cell and prevent aluminium carbides to be released into the bath. This suggests that the carbide transfer to bath is retarded. This is a positive sign that settled TiB2 may be reducing carbide corrosion. And at the same time based on much higher value of (Ti,V)B2 for test cell compared to control cell, it is postulated that TiB2 dissolved in metal has reacted with vanadium in pot metal and produced (Ti,V)B2. It is known that many transition metal borides including vanadium are insoluble and precipitate out when boron is added to aluminium metal[26]. The free energies of formation of VB2 as well as TiB2 are both negative at 970 ° C.

Experiment 1 : cast house results - inclusions The average measured data at the specific locations for Experiment 1 and is presented along with Control 1 in Table 3. Based on this data it was concluded that the addition of TiB2 to test cell did not cause any significant increase in total inclusions as compared to control. The amounts observed in Experiment 1 were considered quite acceptable by cast house. The specifications for inclusions vary a great deal and are customer specific. In casthouse for the total inclusions count, international norms are 0.1-0.3 mm2/kg [28]. Experiment 2: Cast house results - inclusions The average measured data for the same individual location for Experiment 2 is presented in Table 4 along with control sample. This experiment provided further support to concluding that the addition of TiB2 to test cell did not cause any significant increase in total inclusions as compared to control. The differences in the results of Experiments 1 and 2 are due to other factors that affect

747

metal cleanliness in casthouse such as furnace cleanliness, casting parameters etc. Therefore it is concluded that using 25- 50% of metal from treated cell in one batch of about 40 tonnes will not affect the total amount of inclusions significantly as this value is within the typical range mentioned earlier. The amounts observed in Experiment 2 were considered quite acceptable by cast house. Normally, PoDFA sampling is done in cast house at furnace outlet after titanium boride grain refiner addition. For Experiment 2, in addition to this normal sampling, samples were taken before titanium boride grain refiner addition as well. This was to see the differences between borides that are coming from treated cell from those are picked up in casthouse after addition of titanium boride grain refiner (A1TÌ5B1). This is summarized in Table 5. Table 3: Metal cleanliness - Experiment 1 (Average of 2 samples) & Control 1 Unit

Total inclusion TiB2 (Ti,V)B2 Thin & thick oxide films A14C3 MgO, dispersed MgO, patch Spinel Others

Furnace outlet (after A1T5ÌB1)

CFF outlet

mm2/kg

Expt. 1 0.197

Ctrl. 1 0.202

Expt. 1 0.056

Ctrl. 1 1 0.060

% % %

50 13 2

37 2 6

60 10 4

61 3 11

% %

10 19

14 0

9 7

9 0

% % %

0 1 4

25 3 13

3 2 5

4 0 10

Table 5: Borides before and after AITÌ5B1 addition

The TiB2 is generally not considered as a detrimental inclusion as it is deliberately added for grain refinement. But it was included in the total inclusion counts in order to target high standards of metal cleanliness. The oxide films were also included in the PoDFA count. The variation in the inclusions from sample to sample is due to non uniform distribution of inclusions in the bulk. Trace elements pick up in casting furnace It is well-known that chemical trace elements such as Ti, V, Na, etc., influence the properties of thefinalproducts at the cast house and for the end user [29]. Therefore it was needed to ensure that TiB2 treatment in test cell should not affect metal quality. This aspect was examined for Experiments 1 and 2, compared with controls. For each experiment, three samples were cut from the sows produced exclusively from test cell. Two samples were also taken from casting furnace after re-melting these sows. These samples were analyzed by Arc/ Spark Optical Emission Spectrometry. It is seen from average results described in Table 6 and Table 7 that B and Ti content in cast house metal are similar to the metal from other cells where no TiB2 was added. Hence it can be concluded that there is no significant negative impact on metal chemistry by using TiB2.

Table 4: Metal cleanliness - Experiment 2 (Average of 2 samples) & Control 2 Unit

Total inclusion TiB2 (Ti,V)B2 Thin & thick oxide films A14C3 MgO, dispersed MgO, patch | Spinel Others Graphite particles

Furnace outlet (after AITÌ5B1)

CFF outlet

Expt. 2

Ctrl. 2

Expt 2

Ctrl. I 2

0.537 46 12

0.559 27 3

0.295 59 7

0.263 66 3

7 13

9 31

5 7

6 12

0 9 0 13

0 13 6 11

1 9 0 12

1 4 0 7

0

0

0

1

mm /kg

% % % % % % % % %

Table 6: Metal purity at cast house - Experiment 1

f

B Ti Mg Si Fι

% 0.0030-0.0200 0.4700-0.5300 0.4200-0.4800 0.1500-0.1900

% 0.0001 0.0051 0.5062 0.4496 0.1586

Controll (average), 0 0.0111 0.5045 0.5945 0.1729

Table 7: Metal purity at cast house - Experiment 2

B Ti Mg Si Fι

Alloy specifications,

Experiment 2 (average ),

Control 2 (average),

0.0000-0.0200 0.4500-0.5100 0.4000-0.4600 0.1500-0.1900

0.0001 0.0065 0.4776 0.4182 0.1755

0 0.0059 0.483 0.4267 0.1795

% -

Electrical conductivity Electrical Conductivity of billet was measured by conductivity meter model Auto Sigma 3000 manufactured by GE Inspection Technologies [30].There was no difference in conductivity observed between the billets containing metal from the test cell as compared to that produced with metal from regular production and the reading was 62IACS units for both. Grain size of homogenized billet A very important property of the structure of billets is the size and shape of the grain. The morphology of grain is also important and should not be feathery or columnar in shape. Grain size of homogenized billet was measured by light microscopy and CLEMEX Image Analyser using Feret method. The main difference here was that the samples are cut from a cast billet. The results are presented in Table 8.

Edge

50X

Mid radius

50X

Centre

50X

Figure 4: Grain configuration and morphology for Experiment 2

Table 8: Grain size of billet from Experiments 1 & 2 Aiverage (SD)

Ed,

ill

e rtioi1

135( 15)

Mid

Center

137(17)

149(15)

Ac(lept ance C rite ria ain size

ink 'S

Experiment 1

115

110

128

Experiment 2

147

159

188

Edge

50X

Centre

Centre

50X

Conclusions a.

b. c.

Mid radius

50X

Confirmation of analytical technique To confirm the metal cleanliness results, 11 PoDFA samples from above were sent to an accredited laboratory for inclusions count, & break-up, and metallographic analysis. Results indicated that inclusions measured (mm2/kg) were within the range of reproducibility allowance specified by the PoDFA manual.

The grain size and the distribution were found to be similar to those from routine production. From this study it was concluded that addition of TiB2 to the test cell did not affect the grain structure of end product.

50X

Mid radius

Figure 5: Grain configuration and morphology for routine production

n

The above results were found to be within acceptance limit and similar to routine production samples for same type of alloys. The micrographs for the two experiments and for routine production are presented in Figure 3 onwards.

Edge

50X

d.

50X

Figure 3: Grain configuration and morphology for Experiment 1

749

The total amount of inclusions for pot metal from test cell remained more or less same as control cell where no TiB2 was added. It was also within typical range of 0.3-1.5 mm2/kg. Aluminium carbide was much less at 22 % for test cell than control cell where it was the main inclusion at 90 %. Titanium vanadium diboride (64 %) was the main type of inclusion observed in metal from test cell, which was not observed in metal from control cell at all. It is postulated that TiB2 granules, added into test cell settled on top of the cathode and prevented further reaction of cathode with aluminium thus hindering aluminium carbide formation. Some of the TiB2 may have reacted with vanadium in the metal to form (Ti,V)B2. Based on the two trials in cast house it is seen that the total amount of inclusions remained more or less same irrespective of metal from test cell being used or not. Titanium diboride (about 60 %) was the main type of inclusion observed at CFF outlet for the two trials in casthouse. This is quiet normal as titanium boride is added as a grain refiner into the launder in the liquid metal at furnace outlet. It is also seen that the CFF is able to remove the inclusions quite effectively. Higher quality grades of CFF may be able to counter act higher inclusion levels

e.

f. g. h.

i.

As far as metal impurity levels were concerned, no negative impact was observed in casthouse based on the two trials carried out. No significant B and Ti pick up was seen in cast house metal. No impact was seen on grain size of the billets for billets produced from the two trials and hence the test did not affect this property. The conductivity data also remained similar for the metal from test cell and the control cell implying that the test did not affect the conductivity. Due to insufficient quantity, the effect of inclusions in test metal could not be evaluated by using 100 % of this material in cast house, and this would have been a more rigorous test. But on the other hand there was no RAM treatment of the metal prior to the results and the metal did not have any flux addition during processing and was processed in one of the casting stations which historically produced more non conforming product. These aspects may have counter balanced each other to some extent. Inclusions in metal are influenced by other factors such as casting station, quality of CFF used, type of degassing facility used, vibration in the metal filtration area etc., and hence the cast house can control inclusions to some extent by fine tuning the process.

9 M.Ryvola, "Catalogue of Inclusions in Aluminium and its Alloys", Alcan International Limited, Revised Version November 1991, 78 pages. 10 J.R0dseth, B.Rasch, O.Lund and J.Thonstad, Light Metals 2002, 883-887. I I P . Cooper and A. Barber, "Review of the Latest Developments and Best use of Grain Refiners", 2nd International Melt Quality Workshop, Prague, Czech Republic, 16-17th October 2003, 10 pages. 12 Q.G.Wang, D.Apelian and D.A.Lados, J. Light Metals, 2001, 73-84. 13 B. Rasch, E.Myrbostad and K.Haf, Light Metals 1998, 851854. 14 "Hycast RAM System User Manual", Hycast AS, Hydro Aluminium, Revision 0, 2007. 15 C.E.Eckert, R.E.Miller, O.Apelian and R.Mutharasan, Light Metals 1984, 1281-1304. 16 Y.Ohno, Journal of Japan Institute of Light Metals, No.51, 2001, 134-137. offprint from Pyrotek company website, 1-10-09 (English translation of Japanese paper). 17 M.Bryant and P.Fisher, "Grain Refining And The Aluminium Industry - Past. Present And Future", Proc. 3rd Austr. Alum. Conf, editor Nilamani, TMS 1993, 281-291. 18 A..M.Detomi, A.J.Messias, S.Majer and P.S.Cooper, Light Metals 2001, 919-926. 19 P.Cooper and A.Barber, 2nd International Melt Quality Workshop, Prague, Czech Republic, 16-17th October 2003, 10 pages. 20 DUB AL Internal procedure PR/TSK.14. 21 C.Kammer, "Aluminium Handbook", vol. 1, 1st. ed., Aluminium Verlag GmbH,1999, 511-520. 22 M.Warmuzek, "Metallographic Techniques for Aluminum and Its Alloys, Metallography and Microstructures, vol. 9, ASM Handbook, ASM International, 2004, p. 711-751. 23 J.E.Hatch, "Aluminium Properties and Physical Metallurgy, Chapter 3 : Microstructure Of Alloys", American Society for Metals, 1984, 58-104. 24 "PoDFA Inclusion Catalog", ARVIDA Research and Development Centre, ALCAN International Limited, 07-10-1997, 27-42 & 47. 25 "Metallography and Microstructures", Metals Handbook, Ninth Edition, vol.9, American Society for Metals, United States of America, 1985. 26 S.Karabay, Materials & Design, vol. 29, 2008, 1364-1375. 27 Internal plant data at DUB AL. 28 M.V.Canullo and R.A.Laje, "Metallurgical Quality of Aluar Billets , Ten Years of Increasing Quality", ET08 -9th International Aluminium Extrusion Conference, May 2008, 15 pages. 29 C.J.Simensen, "Sources of Impurities in Aluminium Melts and Their Control", Aluminium Melt Refining and Alloying - Theory and Practice, Melbourne, July 10-12, 1989, Cl -C19. 30 "Auto Sigma 30000 ;Technical and Operations Reference Manual", GE Inspection Technologies, LP, USA, Issue 04, 05/2005.

Acknowledgements The authors would like to thank the management of DUB AL for sponsoring and supporting this project. The authors would also like to acknowledge the contribution of various departments especially from the smelter and casthouse operations. Special thanks to G.Meintjes for coordinating the trial in line 7, N.A1 Jabri for supporting this trial whole heartedly, N. Rana for the statistical analysis, the metallurgy team for coordinating various metallographic tests, S.K.Howaireb and her team for chemical analysis of numerous samples, and Dr.K.Venkatasubramaniam for reviewing the paper. Support provided by all the other departments is also gratefully acknowledged. Thanks to J.Proulx of ABB Metallurgical services for sharing expertise regarding inclusions. References 1 Part 1 of this paper, "Simplifying Protection System To Prolong Cell Life" by M.M.Al-Jallaf, M.Hyland, B.Welch, A. Al Zarouni, and A. Fahimi being published in TMS 2010. 2 J.Langerweger, Light Metals 1981, 688- 705. 3 N.Habibi, Ph.D Thesis, Universitι du Quιbec à Chicoutimi (UQAC), Canada, August 2002,15-23. 4 D.Doutre, B.Gariepy, J.P.Martin and G. Dube, Light Metals 1985,1179-1189. 5 "The Influence and Control of Porosity and Inclusions in Aluminium Castings", Downloaded from ASM website on 15-1109. http://www.asminternational.org/pdf/spotlights/5114alum castc5. pdf 6 L.Liu and F.H.Samuel, Journal Of Materials Science, 32 (1997), 5901-5925. 7 L.Liu and F.H.Samuel, Journal Of Materials Science 32 (1997) 5927-5944. 8 "Determination of PoDFA Inclusion Concentrates", Alcan Report AR-90/0032, Dec 1990 (Confidential).

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

THERMODYNAMIC ANALYSIS OF Ti, Zr, V AND Cr IMPURITIES IN ALUMINUM MELT A. Khaliq, M.A. Rhamdhani, G.A. Brooks. J. Grandfield Swinburne University of Technology, Faculty of Engineering and Industrial Sciences, Melbourne, Australia CAST Cooperative Research Centre (CAST CRC), Australia Grandfield Technology Pty, Ltd, Victoria, Australia Keywords: transition metals, boride formation in Al melts, boron treatment their application in the petroleum industry is limited. It was also reported that there was an increasing trend in the concentration of impurities (including V) in the coke which later ends up in aluminum [12]. In anticipation of this, it is imperative that the current methods available for removing these impurities from aluminum to be optimized; or for new methods to be developed.

Abstract Aluminum is widely used as the main material for overhead power cable because of its good electrical conductivity and light weight. Metal impurities, in particular Ti, Zr, V and Cr in the solution, affect the electrical conductivity of aluminum significantly. Industrially, boron treatment has been used to remove these impurities through the formation of borides. However, studies have shown that solution thermodynamics and the detailed reaction mechanisms of the borides formed in aluminum melts are not well understood. In the present work, thermodynamic analysis has been carried out to investigate the relative stability and to elaborate on the preferential formation of various borides in aluminum melt. It is shown that diborides (MB2) are the most thermodynamically stable boride compounds of these impurities in the given working conditions. The ZrB2, TiB2 and VB2 phases are more stable compared to A1B2 and CrB2 hence do not dissolve readily. It is also shown that the relative stability of the boride phases is affected by the presence of other metal diborides.

Commercially, boron treatment has been used to remove transition metal elements through the formation of borides. This is carried out by adding boron or aluminum-boron master alloys to the melt. [2, 12, 13] The formed borides or diborides are separated by gravity settling or by flux addition combined with mechanically mixing in the melt [14]. The reactions depend upon diffusion of the impurities in the aluminum melt and the availability of boron/aluminum borides. It has been reported that this reactions of boron with transition metals are fast enough that 70% of increase in electrical conductivity is achieved in the first couple of minutest 15]. Further increase in electrical conductivity is quite slow and takes more than two hours, i.e. after the heavy boride particles settle at the bottom of furnace [15-19]. It has been established by previous works that Ti, Zr and V can be removed from aluminum melt by boron treatment [13, 15, 20]. Laun [21] reported the removal of Cr and Mn during boron treatment along with Ti and V. Setzer[20] and Cooper [15] studied on the removal of Ti, Zr, V and Cr , Mn was not mentioned. Wang et al. [20] reported in their study on boron reactions with transition metals that there was no evidence of borides of Cr and Mn in the reaction products. However, there are different statements about removal of Cr and Mn through formation of borides[22].

Introduction Aluminum can be used as an electrical conductor if the level of impurities is controlled precisely, in particular, the concentration of transitional metals Ti, Zr, V, and Cr. Their effect is minimized when they are in a combined form rather than in solution with aluminum. The detailed discussion on the effect of solute impurities to aluminum properties has been described in the literature [1-3]. The maximum solubility and the effect of transition metal impurities on electrical resistivity in and out of solution are given in Table 1. It can be seen that vanadium and chromium has the highest effect on the resistivity; and the presence of these elements in solution increases the resistivity by a factor 10 to 20.

It is obvious from previous research findings, that the detailed mechanisms of reactions between boron and transition metal impurities in liquid aluminum are not fully understood. Moreover, the exact nature of the borides type, their morphology and composition need to be evaluated and analyzed to get a full understanding of borides formation mechanism for effective boron treatment of aluminum.

Table 1: Transition metal impurities, their maximum solubility and effect on resistivity of aluminum [3] Avg increase in resistivity Max. Solubility in Al per wt% μΩχιη Elements (wt %) Out of In Solution solution Titanium 0.12 1 2.88 Zirconium 0.28 1.74 0.044 Vanadium 0.5 3.58 0.28 Chromium 0.77 4 0.18

In this article, a literature study on thermodynamic information and previous works on the transitional metals (i.e. Ti, Zr, V, and Cr) and their borides in Al melt will be presented. In addition, investigations on the possible stable phases; their relative stability; and their preferential formation in the presence of other transition metals borides have also been carried out using thermodynamic packages. Transition Metal Borides in Aluminum

These impurities mainly come from carbon anodes (from petroleum coke) and bauxite used in the primary production of aluminum [4]. Although there are many methods to remove transition metals (in particular V) from the petroleum coke [5-11],

Previous investigators have indicated that there was some solubility of aluminum in transition metal diborides; vice versa some solubility of transition metals in aluminum diborides formed

751

in aluminum melts [19, 23, 24]. Setzer and Boone [20] reported in their study that transition metal borides formed in Al melt were not pure and they formed cluster of diborides (containing V, Ti, Zr, and Cr). They also investigated the morphology of the borides and observed two distinct morphologies, i.e. equiaxed hexagonal (occasionally platelet) and clustered platelet. They attributed these to the different initial boride particles used for precipitation, i.e. A1B2 and A1B12, respectively. Although the morphology of the borides were presented, it was not clear whether these borides were in some form of solid solution or clusters of separated boride phases. Wang et al. [22] studied the reactions of boron with transition metal and observed similar cluster of borides. Energydispersive x-ray spectroscopy results suggested the presence of Ti, V, and Zr. They speculated that A1B2 formed a continuous solid solution of (Al, Ti, V, Fe)B 2 but no further evidence was presented.

assumptions was used. In the sub-lattice model, it is considered that crystalline species are formed in two or more different lattice structures [32]. The thermodynamic analyses carried out include: Evaluation of Gibbs free energy formation of aluminum and transitional metal borides. Evaluation of binary and ternary phase diagrams (Al-B, AlTi-B, Al-Zr-B and Al-Cr-B). Equilibrium calculations of different compositional systems (Al-Ti-B, Al-Ti-Zr-B, Al-Ti-Zr-V-B, and Al-Ti-Zr-V-Cr-B) with stoichiometric and excess boron concentration for transition elements (assuming the formation of diborides). Borides Gibbs Free Energy Evaluation The Gibbs free energy formation of various metal borides phases has been evaluated using HSC Chemistry 7.0 package in the temperature range of 650°C to 900°C. The results obtained are presented in Figure 1. During the evaluation, it was found that diborides phases (A1B2, TiB 2 , ZrB 2 , VB 2 , and CrB2) of transitional metals were the most stable in the given conditions as compared with their other possible phases (A1B12, TiB, VB, V 3 B 4 , V 5 B 6 , CrB, Cr 3 B 4 , Cr 5 B 3 ). Figure 1 suggest the order of stability of pure diborides in aluminum melt from ZrB 2 , TiB 2 , VB 2 , A1B2, CrB 2 to A1B12 in the given temperature range (650°C to 950°C). This order of stability also suggests that the elements Zr, Ti, V can be removed by the addition of Al-B master alloys through the formation of borides (as has been demonstrated experimentally by previous investigators). Chromium diborides (CrB2), however, shows a lower stability compared to A1B2, thus boron treatment may not be the appropriate method for removing chromium. This agrees with the findings of Wang et al. [20] who reported that there was no evidence of borides of Cr and Mn in the reaction products during reactions between boron and transition metals.

Fjellstedt [19] used optical microscopy, energy-dispersive x-ray spectroscopy and x-ray diffraction to investigate the mutual solubility in Al-Ti-B system. It was reported that the A1B2 and TiB 2 particles were stable and found as separate phase having limited solubility with each other in the aluminum melt[19]. In some case, A1B2 was observed to be surrounded by ring of TiB 2 phase. It was further reported that complex compounds of (Ti,Al)B2 were unstable and that A1B12 showed very low solubility of Ti [19, 23, 24]. During research of grain refinement of aluminum and its alloys using inoculants, McCartney [25] reported that TiB 2 , A1B2 were stable but (Ti,Al)B2 phases were unknown in stability so far. Many other researchers have also discussed the metastability and presence of (Ti,Al)B2 but none have provided clear and complete explanations [25-29]. Higashi et al. [23] studied the solubility of aluminum in various metal borides including TiB 2 , ZrB 2 , HfB2, VB, V 3 B 4 , NbB 2 , TaB and W 2 B 5 by growing these crystals from aluminum melt at 1300°C to 1550°C. They reported that the solubility of aluminum in these borides was 0.1 wt% maximum. This was interpreted as A1B2 did not form mixed crystals with the above borides. Recently, Otani et al. [30] investigated the solid solution ranges of ZrB 2 with refractory diborides (HfB2, TiB 2 , TaB 2 , NbB 2 , VB 2 and CrB2). Mixed-boride samples were heated and melted using arc melting furnace under Ar atmosphere at 1600°C. The samples were analyzed using x-ray diffraction and inductively-coupled plasma techniques. Their study showed that ZrB 2 formed a perfect solution with NbB 2 , HfB2 and TaB 2 . In the case of ZrB 2 -TiB 2 system, they observed two phases, i.e. ZrB2-rich (Zro.76Tio.24B2) and TiB2-rich (Zr0.nTi0.89B2) phases. It was also reported that ZrB 2 also made a solid solution with VB 2 and CrB 2 but the solubility was very low i.e. < 3 mol% and < 1 mol %, respectively at 1600°C.

!

-20

1«eo

1

r— ^

, — ,

1—

.., ........ —

-40

w—

Jjn

~—A·

-80

r

»

« 0.08AIBJ ~#-0.5CrB2 I

Jfc—

<

0.5TiB2 1

-A ->-0.5Zr8, j

-100 - ►"

1

J

-120 -140 O

Thermodynamic Assessment of Borides Formation

ARf\

-4—

f-,

650

<| -—! —t 1

700

4

^

^

._ 4 [ —

lb 1

S

t

p. i

750 800 850 TemperaturefC)

4

__*>

,

Τ00

J ·

1

950

t

1Τ0Τ

Figure 1 : Gibbs free energy evaluation of pure metal borides

In the current study, thermodynamic assessment of different compositions of transition metals and boron in aluminum have been carried out in the temperature range of 675°C to 950°C using the HSC Chemistry 7.0 and FactSage 6.1 thermodynamic packages. The light metal database (FTlite) has been used in FactSage package where modified quasichemical model has been employed for the assessment of liquid phases. This model is based on the short range ordering tendency of atoms or molecules in liquid solutions [31]. For solid boride phases (MB, and MB 2 ), the sub-lattice model which based on the substitutional solution

Phase Diagrams and Equilibrium Calculations Equilibrium calculations of different Al-M-B (M = Ti, Zr, V, Cr) systems in Al melt were carried out using FactSage 6.1 thermodynamic package in the temperature range of 675°C to 900°C. The "Equillib" module of FactSage was used for the calculations which incorporate the Gibbs free energy minimization technique to predict the thermodynamically stable

752

phases in different conditions. Due to lack of solution thermodynamic information, in these calculations it was assumed for the boride systems that:

In case of Al-Ti-Zr-B system, mixed borides are composed of almost equal proportion of TiB 2 and ZrB 2 . In the alloy systems containing V and Cr with B and Al, VB2(s) and CrB2(s) phases have been predicted along with A1B2 borides. It is assumed that there is some solubility of TiB 2 and ZrB 2 in A1B2 and vice versa. According to the given analysis there is no solubility of Ti, Zr or Al in CrB 2 and VB 2 as they have been predicted as separate solid phases. The stability of each and every phase depends upon the temperature and composition of solute in the aluminum melt.

ZrB 2 formed an ideal solution with TiB 2 , i.e. (Zr,Ti)B2 There was only a very small solubility of Al in (Zr,Ti)B2 and vice versa A1B12, VB 2 , CrB 2 (and other borides) were pure The total concentration of the transition metals varied from 1.1 1.5 wt% in the Al alloy-boride systems under discussion. Stoichiometric and 75wt% excess amounts of boron (assuming diborides formation) were added to the systems to investigate the boride formation preferences. It has been suggested by previous investigators [13, 15] that excess boron may be required for efficient formation of the metal transition borides.

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When considering these diborides systems, the following reactions may occur: [M] + A1B2(S) = MB2(S) + A1(1) (1) [M] + [B] = MB(S) (2) [M] + 2[B]=MB 2 ( S ) (3) A1(1) + 2[B]=A1B 2(S) (4) [M] + A1B2(S) = (M,A1)B2(S) (5) [MJ + [M 2 ] + A1B2(S) = (M^M^Bac,: + Aln (6) (7) [Md + [MJ + A1B2(S) = (M1,M2,A1)B2(8)

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The general features of aluminum interaction with transition metals include the formation of supersaturated solid solution, quasi-crystalline and amorphous compounds at some particular composition and cooling rate during solidification[33]. The current equilibrium calculations showed that at 1.1 - 1.5 wt%, boron, aluminum and transition metals (Ti, Zr, V and Cr) may possibly form A1B2, TiB 2 , (Ti,Zr)B2, VB2(s), CrB2(s) in the temperature range of 675°C to 900°C. The possible phases in aluminum binary and ternary alloys with boron and some transition metals are summarized in Table 2. Table2: Thermodynamically possible stable phases in Al binary and ternary systems in temperature range of 675°C to 900°C. Alloy Systems Possible Stable Solid Phases Al-B A1B2 Al-Ti-B TiB 2 * Al-Zr-B ZrB 2 * Al-V-B VB 2is) ,Al 7 V is) Al-Cr-B CrB 2(s) *very small solubility of Al

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Under equilibrium solidification Al3Ti, A1B2 and TiB 2 were predicted in the Al-Ti-B ternary systems as proposed by Abdelhamid et α/.[35]. The Al3Ti phase possesses a tetragonal crystal structure but A1B2 and TiB 2 phases are composed of hexagonal crystal structures with only a minor difference in their lattice parameters. Hence, there are possibilities for the stability of a continuous range of (Al, Ti) B 2 solid solution [33]. In the thermodynamic analysis complete solid solutions of ZrB 2 and TiB 2 was predicted. On the other hand solid solution of A1B2 with TiB 2 and ZrB 2 does exist having limited solubility. The previous research by Higashi[23], Jones[24] and Fjellstedt[19] found that there was some solubility of Ti in A1B2 and some Al in TiB 2 but there was no evidence of any solubility among these borides particles. Moreover, Fjellstedt[19] explained that (Ti,Al)B2 type compounds are unstable and cannot exist at room temperature. From the FactSage assessment and previous studies, it may be assumed that TiB 2 , ZrB 2 and A1B2 are mixed with each other and making a metal boride solid solution (MB2) phases.

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Figure 6 shows the associated equilibrium solute concentration changes in the Al melt. In the temperature range studied, no solute Zr and Ti were predicted in the Al melt as they were present in the stable (Zr,Ti)B2 phase. V was predicted above 825°C when VB2 started to dissolve. The equilibrium concentrations of Cr and B were 0.16wt% and 0.04wt% at 675°C, respectively. CrB2 was the least stable phase compared to other diboride phases. It was assumed that CrB2 will start dissolving first in the melt which may result in the early increase of Cr and B concentrations with increase in temperature. At above 750°C no further increase in Cr and B concentrations were predicted, as all CrB2 had already dissolved. A slight increase in B solute concentration was predicted above 825°C, likely associated with dissolution of VB2 to the melt.

Equilibrium Composition Analysis In the first case, equilibrium calculations of liquid aluminum with lwt % V and stoichiometric (and 75 wt% excess) amounts of boron were conducted. During the thermodynamic assessment of Al-V-B system, separate solid VB2(S) particles and A1B2 were predicted (See Table 2). In the Al-lwt%V-B system, vanadium diboride (VB2) was predicted to be little stable in the given temperature range but dissolve slowly at higher temperatures(900°C).The changes in the vanadium, VB2 and A1B2 equilibrium concentration in Al melt with temperature are given in figure 4. It can be seen from figure 4 that in the case of stoichiometric addition of boron, the concentration of V in the metal at 675°C was 8 to 10 ppm. The concentration of V increased with increasing temperature which suggested that the VB2 was dissolving back into melt with increasing temperature. With excess boron, both VB2 and A1B2 were predicted. It appeared that the excess boron combined with Al to form A1B2. By adding excess boron, the VB2 also appeared to be more stable in the temperature range studied. No dissolution of VB2 was predicted in the temperature range of 675°C to 900°C, as shown in figure 4, i.e. no V was predicted in the Al melt as a solute. 15000 Jfc.13500

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Figure 4: V, VB2 and A1B2 equilibrium concentrations in Al melt with stoichiometric and excess B addition having 1.2 to 1.4 wt% of Al. The second case studied was the equilibrium between Al-lwt%M (where M = Zr, Ti, V, Cr) with stoichiometric and 75% excess boron additions. In the case of stoichiometric addition of boron, the following diborides were predicted: (Zr,Ti)B2 with very little solubility of A1B2 (i.e. 0.001 wt% at 750°C), VB2 and CrB2. Figure 5 shows the equilibrium concentrations of metal diborides predicted in the system Al-lwt%M (M = Zr, Ti, V, Cr each 0.25wt%)-0.357wt%B. The figure shows (Zr,Ti)B2 was stable and the concentration remained unchanged in the temperature range studied. On the other hand, the diborides of Cr and V were relatively unstable and predicted to be dissolving back into the melt with increasing temperature. VB2 was predicted to be stable only up to 825°C, where above this it started to dissolve. The equilibrium concentration of CrB2 was predicted to be 12.5wt% at 675°C and decreased with increasing temperature; at 750°C CrB2 was completely dissolved in Al melt over the temperature range studied.

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and Zr in the A1B2 phase; and Al in (Zr,Ti)B2, i.e. 0.001 wt% at 750°C.

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It can be seen from Figure 7 that (Zr,Ti)B2 and VB2 are more stable compared to A1B2 and CrB2 under the given conditions. The equilibrium concentration of CrB2 was predicted to be 9.2 wt% of total metal borides at 675°C. As the temperature was increased, CrB2 dissolved first, hence the decrease in CrB2 concentration up to 750°C where all CrB2 was completely dissolved. Afterward A1B2 started to dissolve. The concentration of (Zr,Ti)B2 and VB2 remained unchanged in the melt in temperature range 675°C to 900°C.

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A thermodynamic assessment of various aluminum systems with Ti, Zr, V, Cr and B have been carried out at temperature range between 675°C to 900°C. Two main cases were analyzed one stoichiometric composition and the other with 75 wt% excess of boron addition to aluminum melt composed of 1.4 wt% of total transition metal (Ti, Zr, V and Cr). According to the results, the Gibbs free energy for formation of the pure diboride compounds, ZrB2 was found to be most stable under the given conditions as compared with other transition metal diborides. The order of stability (from most stable to least stable) of these boride phases was ZrB2, TiB2, VB2, A1B2, CrB2 within the given temperature range.

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Also of note is the relative stability of phases depends upon the presence of other diboride phases. For example, in the first case (stoichiometric addition of boron), where no A1B2 was present, the VB2 phase remained thermodynamically stable only up to 875°C (as shown in Figure 5). In the presence of A1B2 (excess boron added), VB2 was predicted to be stable up to 900°C, as shown in Figure 7. 6000

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The equilibrium calculations predict that Cr could not be effectively removed from aluminum above 750°C through boron treatment as aluminum diboride was more stable compared to chromium diboride. When boron was added to the aluminum melt containing Zr, Ti, V and Cr, the results suggested that boron would be tied up by Zr and Ti, then V, Al, and Cr, forming (Zr,Ti)B2, VB2, A1B2 and CrB2, respectively. In the case of excess boron, with increasing temperature the least stable CrB2 dissociated first, followed by A1B2 at 750°C leaving behind the Al melt with stable borides (Zr,Ti)B2, and VB2. It may be suggested (from thermodynamic perspective only) that boron treatment would be most effective for removing Zr and Ti.

Figure 7: Equilibrium concentration of metal diborides in Al-MB system with 75% excess B addition having total 1.4 wt% of Al. The ZrB2 and TiB2 are in the form of (Zr,Ti)B2. Figure 8 shows the solute concentrations in Al melt in the case of excess boron addition. The solute concentrations in the case of stoichiometric boron addition are also shown for comparison. It can be seen from the figure 8 that the addition of stoichiometric or excess boron may be carried out to remove Zr, Ti and V from the melt. However, the model predicts that Cr would remain in the melt; as the excess boron appeared to be reacting with Al to form A1B2 rather than forming CrB2.

It should again be emphasized that the current study took into account thermodynamic factors only which have provided the theoretical limits. The equilibrium calculations were also carried out with various assumptions. These limitations aside, the results provided an insight on the relative stability and preference of transition metal borides which may form in aluminum melt. Further research including experimental work is required for a complete understanding of boride formation in liquid aluminum. Acknowledgement The authors thank CRC CAST and Swinburne University of Technology for funding of this project.

755

[21]

References: [I] [2] [3] [4]

[5] [6] [7] [8] [9] [10] [II] [12] [13]

[14] [15]

[16]

[17]

[18]

[19]

[20]

T. Engh, Principles of Metal Refining: Oxford University Press, 1992. G. Gauthier, J. Inst. Met, vol. 59, pp. 129-150, 1936. L. Willey, "Effects of Alloying Elements and Impurities on Properties," Aluminum, vol. 1, p. 174, 1967. K. Grjotheim, M. Malinovsky, K.Matiasovsky, J. Thonstad, Aluminium Electrolysis: Fundamentals of the Hall Heroult Process, 2nd ed.: Aluminium-Verlag GmbH, D-4000 Dusseldorf, Germany, 1982 L. McCorriston, "Process using Sulphate Reagent for Recoveing vanadium from cokes drived from heavy oils," 1983, US 4 389 378. Queneau et al., "Recovery of V 2 0 5 and Nickel values from petroleum coke," 1984, US 4 443 415. Schemel et al.,"Method for leaching and recovering vanadium from vanadium bearing by product materials,", 1985, US 4 539 186. L. McCorriston, "Process Using Carbonate Reagent for Recovering Vandadium From Cokes and Ashes Drived From Heavy Oils," 1985, US 4 536 374 . Thornhill et al., "Process and Apparatus for Heavy Metal From Carbonaceous Materials," 1994, US 5 277 795. Malone et al., "Recovering Vandadium from Petroleum Coke as Dust," 2001, US 6 241 806 Bl. W. Zhang, et al., "Modelling of impurity balance for an aluminium smelter," 1996, pp. 405-411. W. Dean, "Effects of Alloying Elements and Impurities on Properties," Aluminum, vol. 1, p. 174, 1967. W. Stiller and T. Ingenlath, "Industrial Boron Treatment of Aluminium Conductor Alloys and Its Influence on Grain Refinement and Electrical Conductivity," Aluminium (English Edition), vol. 60, 1984. G. Dube, "Removal of Impurities from molten aluminium," 1983, US 4 470 846. P. Cooper and M. Kearns, "Removal of transition metal impurities in aluminium melts by boron additives," Aluminium Alloys: Their Physical and Mechanical Properties, Pts 1-3, vol. 217, pp. 141-146, 1996. R. Cook, M. Kearns P. Cooper, "Effects of residual transition metal impurities on electrical conductivity and grain refinement of EC grade Al," Light Metals, pp. 809-814, 1997. S. Karabay and I. Uzman, "Inoculation of transition elements by addition of A1B2 and A1B12 to decrease detrimental effect on the conductivity of 99.6% aluminium in CCL for manufacturing of conductor," Journal of Materials Processing Technology, vol. 160, pp. 174-182, 2005. S. Karabay and I. Uzman, "A study on the possible usage of continuously cast aluminium 99.6% containing high Ti, V, and Cr impurities as feedstock for the manufacturing of electrical conductors," Materials and Manufacturing Processes, vol. 20, pp. 231-243, 2005. J. Fjellstedt, et al., "Experimental analysis of the intermediary phases A1B2, A1B12 and TiB2 in the Al-B and Al-Ti-B systems," Journal of Alloys and Compounds, vol. 283, pp. 192-197, 1999. W. Setzer and G. Boone, "Use of aluminum/boron master alloys to improve electrical conductivity," Light Metals 1992, pp. 837-844, 1991.

[22]

[23]

[24]

[25] [26]

[27] [28]

[29] [30]

[31]

[32]

[33] [34] [35]

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T. Luan Bao-gui, Haung Chong-qi, "Increasing the electrical conductivity of aluminum conductor by treating the melt with boron," Electrical Wire and Cables, ( in Chinese), pp. 36-40, 1984. G. Wang, et al., "Reaction of boron to transition metal impurities and its effect on conductivity of aluminum," Transactions of Nonferrous Metals Society of China, vol. 12, pp. 1112-1116, Dec 2002. I. Higashi, et al., "crystal-growth of borides and carbides of transition-metals from molten aluminum solutions," Journal of Crystal Growth, vol. 33, pp. 207211,1976. G. Jones and J. Pearson, "Factors affecting the grainrefinement of aluminum using titanium and boron additives," Metallurgical Transactions B, vol. 7, pp. 223-234, 1976. D. McCartney, "Grain refining of aluminium and its alloys using inoculants," International Materials Reviews, vol. 34, pp. 247-260, 1989. G. Sigworth, "Grain refining of aluminum and phase relationships in the al-ti-b system," Metallurgical transactions. A, Physical metallurgy and materials science, vol. 15 A, pp. 277-282, 1984. M. Guzowski, et al., "The role of boron in the grain," Metallurgical Transactions A, vol. 18, pp. 603-619, 1987. M. Guzowski, et al., "Role of boron in the grain refinement of aluminum with titanium," Metallurgical transactions. A, Physical metallurgy and materials science, vol. 18 A, pp. 603-619, 1987. M. Easton, D.StJohn , "The Effect of Alloy Contents on the Grain Refinement of Aluminium Alloys. "Report, CRC Cast Metals Manufacturing (CAST), Australia S. Otani, et al., "Solid solution ranges of zirconium diboride with other refractory diborides: HfB2, TiB2, TaB2, NbB2, VB2 and CrB2," Journal of Alloys and Compounds, vol. 475, pp. 273-275, 2009. A. Pelton, et al., "The modified quasichemical model I Binary solutions," Metallurgical and Materials Transactions B: Process Metallurgy and Materials Processing Science, vol. 31, pp. 651-659, 2000. M. Hillert, et al., "A two-sublattice model for molten solutions with different tendency for ionization," Metallurgical Transactions A, vol. 16, pp. 261-266, 1985. N. Belov, Multicomponent phase diagrams : applications for commercial aluminum alloys Amsterdam ; Oxford : Elsevier 2005. D. Kammer, Aluminium Handbookl, First Edition ed.: Aluminium-Zentrale e.V., Am Bonneshof 5, 40474 Dusselddorf, Germany, 1999. A. Abdelhamid and F. Durand, "Liquid-solid equilibria of Al-rich Al-Ti-B alloys .1. nature of the 4-phase and 3-phase reactions," Zeitschrift Für Metallkunde, vol. 76, pp. 739-743, 1985.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CURRENT TECHNOLOGIES FOR THE REMOVAL OF IRON FROM ALUMINUM ALLOYS Lifeng Zhang, Lucas N. Damoah Department of Materials Science & Engineering, Missouri University of Science and Technology (Missouri S&T) 223 McNutt Hall, Rolla, MO 65409-0330, U.S.A. Email: [email protected] KEY WORDS: Aluminum, Iron-Rich Phases ί-AlFeSi [3]. Since the solid solubility of iron in aluminum is less than 0.05 % at equilibrium, almost all iron forms second phases in aluminum [4]. The binary Al-Fe and ternary Al-Fe-Si phases are the main Fe-rich phases in aluminum alloys [3]. Three dimensional morphology of the Fe-rich intermetallic compounds observed by current authors (Figure 1 [5]) suggests that Fe-rich intermetallic phases have much more complex morphologies, with fragile and brittle appearance than what is shown in two dimensional observation. These morphologies imply why they are detrimental to the mechanical properties of aluminum.

Abstract In the current paper, the Fe-rich phases in and their detrimental effect on aluminum alloys are summarized. The existence of brittle platelet ί-Fe-rich phases lowers the mechanical properties of aluminum alloys. The methods to neutralize the detrimental effect of iron are discussed. The use of high cooling rate, solution heat treatment and addition of elements such as Mn, Cr, Be, Co, Mo, Ni, V, W, Cu, Sr, or the rare earth elements Y, Nd, La and Ce are reported to modify the platelet Fe-rich phases in aluminum alloys. The mechanism of the modification is briefly described. Technologies to remove iron from aluminum are extensively reviewed. The precipitation and removal of Fe-rich phases (sludge) are discussed. The dense phases can be removed by methods such as gravitational separation, electromagnetic separation, and centrifuge. Other methods include electrolysis, electro-slag refining, fractional solidification, and fluxing refining. The expensive three-layer cell electrolysis process is the most successful technique to remove iron from aluminum so far. Introduction

Figure 1. 3-D morphologies of Fe-rich intermetallic phases ofί-Al(FeMn)3Si[5].

During refining and recycling of aluminum alloy scraps, iron gradually accumulates [1] and is of more difficult to be removed with decreasing Fe content [2]. Most aluminum alloy production requires tight composition controls on iron. For example, iron content level above 0.15 wt% is unacceptable in premium aerospace alloys such as 7050. Iron is the most pervasive impurity element in aluminum alloys, which stems from the bauxite and steel tools used during both primary and secondary production. Iron usually forms second phases in the aluminum alloys owing to its low equilibrium solid solubility in the aluminum (max. 0.05%), such as Al3Fe, a-AlFeSi and

A number of Fe-rich phases in Al-Fe-Si ternary system have been identified as shown in Table 1. Al3Fe (also reported as 0-Al3Fe or 0-Ali3Fe4 [4, 6] ), a common equilibrium phase, forms a eutectic with aluminum at about 655 °C [3]. The most important Fe-rich phases in aluminum alloys containing silicon are ί-phase and a-phase. Among all the Fe-rich phases, ί-AlFeSi is thought to be the most deleterious, and most efforts have been devoted on how to avoid the formation of ί-AlFeSi, which is brittle and generally assumed to act as stress raisers and points of weak coherence.

757

There are no effective practical methods to directly remove iron from aluminum alloys by conventional refining. The techniques or theories on direct iron Fe-rich intermetallics AlmFe Al6Fe AlxFe AlDFe

removal from aluminum have made no satisfactory progress so far.

Table 1. Identified Fe-rich phases in aluminum alloys Crystal structure Bet Orthorhombic C-centred orthorhombic

0-Al3Feore-Al13Fe4 a-Al8Fe2Si or a-Al12Fe3Si2 a-Al15Fe3Si2 ί-Al5FeSi Al9Fe2Si2 τ-Al4FeSi2 τ-Al3FeSi2 qrAlFeSi q2-AlFeSi Y-Al8FeSi p-Al8Mg3FeSi6 7C-Al8Mg3FeSi6

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has the same Bravais lattice with Fe. Thus Cr and Co can also be used to precipitate Fe-rich phases. Mathta [18] reported that the optimum ratio of Co/Fe was -1.0 for A413 Al-llSi alloys and the Co-Fe phases were identified approximately as Ali5(Fe,Co)4Si2, while Murali [22] stated the Co-Fe phase was Ali4Co2(Fe,Si).

Precipitation of High Fe-Rich Phases The method is mainly applied to purify Al-Si cast alloys. Iron is removed by the formation of primary Fe-rich intermetallics, generally primary a-Ali5(Fe, Mn)3Si2 or a-Ali5(Fe, Mn, Cr)3Si2 (called "sludge"). Manganese

0.3 r

Although Mn is harmful to the mechanical properties of aluminum alloys, it is widely used to neutralize Fe in Al-Si cast alloys. Mn has an atomic radius and crystal structure close to Fe [4, 18]. Yoo reported that the crystal structure of precipitated a-Ali5(FeMn)3Si2 depended on the Mn/Fe ratio. The crystallization of the sludge has proved to be an effective method to remove Fe from Al-Si alloys, as shown in Figure 2 [19]. After the formation of the sludge, further steps including gravity separation, filtration, electromagnetic separation or centrifuge separation are conducted to remove the sludge phase from the molten aluminum. The iron concentration generally decreases from 1 - 2 wt% to at most 0.4 wt% after the treatment [20-21].

2. 0.0n to

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Time (minutes) < Figure 2. Effect of holding time and initial Fe, Mn content on the formation of sludge at 605°C [19] Chromium Mondolfo [4] claimed that Cr was a possible additive to neutralize Fe phases in aluminum. Cr was believed more effective than Co, and a ratio of Cr/Fe= 0.33 can prevent the formation of ί-AlFeSi [18]. The function

Cobalt Fe, Cr and Co have similar atomic radii and Cr also

758

of Cr is similar to that of Co. The calculated isothermal section of Al-Fe-Cr at 700°C was shown in Figure 3. [23] The formation of AlnCr 2 and Ali3Cr2at high Cr content (-10 at%) above the liquidus surface may remove iron from the Al melt. However, the iron solubility of AlnCr2 and Ali3Cr2 is too small to effectively remove iron, and a lot of Cr should be added into the Al melt, which is impossible for the real industrial practice.

sedimentation. Most of the sludge settled to the bottom and thus the upper alloy was purified. Filtration Primary Fe-rich inter-metallic particles also can be removed by porous filters similar to the removal of nonmetallic inclusions by filtration. Figure 5 [20] shows the schematic steps of the filtration operation. After a short time (10-20min) holding at sludge formation temperature, the melt is decanted through a preheated filter. Small amount of sludge also precipitates at the bottom of the melt during the holding time. The holding time used for filtration is much shorter than holding time used for gravity separation and so filtration is suitable for the continuous treatment. Donk [33] concluded that finerpore foam filters can remove small size sludge particles but actually only slightly increased the removal efficiency because the captured small particles easily blocked the filter pores. The efficiency increases with increasing Mn/Fe (>1) or initial Fe and Mn content (Table 2) [36-37].

**#»* *mo*m * 'm#m i?m+wifmf.tì

Figure 3. Al-Fe-Cr phase diagram (700°C and 1 atm) [23] Beryllium Murali considered beryllium was a more effective neutralization element than Mn, Co and Cr [24]. Crepeau reported that the addition of Be > 0.4 wt% was required [2], while other references showed that the trace addition of Be 0.06-0.27 wt% was enough [24-27]. Strontium In aluminum wrought alloys, Sr was also applied to transform the platelet Fe-rich phases to a-AlFeSi (Al8Fe2Si) [28-31]. It was reported that the 0.01-0.10 wt% addition of Sr to Al-Cu-Mg-Zn wrought alloys refined the intermetallic phases [32].

f ""'^^

<**3H^v-

P\ ,

*

*

>

m

4 'M ^ifl

k

Figure 4. Fe-rich phase precipitation and subsequent gravity separation in an aluminum alloy [35] Stepl „

^ C

*«P2

^

^CoolingtoT2 T1>T2

Al, Si, Fe High Temp.TI

Primary o-AI(Fτ,Mn)Si

Gravity Sedimentation Step 3 Holding for 10*20mta

Gravity sedimentation is a method to remove the heavier phase, like Fe-rich phase, from aluminum. Donk [33] found that sludge formed and segregated immediately during the cooling from 840 °C to 600 °C. Reported removal efficiency data was summarized in Table 2. Flores and Cao reported an Fe removal fraction over >70 % with Mn/Fe>l and relative high Fe content [19, 21, 34]. Figure 4 [35] shows a microstructure of an aluminum alloy after gravity

Step 4

Filter

Prìmàty a~AA(F*M
Figure 5. Schematic of the filtration process (Tl: melting temperature, T2: holding temperature) [20]

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Table 2. Reported removal efficiency of iron fronl aluminurn alloys Holding After Purification Fe removal Authors and Composition (wt.%) fraction Mn/Fe Temp Time reference (wt.%, Al balance) (min) (°C) Mn Fe (%) Kim [38] NA 6Si-l.64Fe-l.66Mn 1 0.45 72.6 690 20 Li [35, 39-40] 12Si-l.13Fe-l.22Mn 0.41 63.7 1.1 640 30 0.26 Jiao [41] 70 10Si-lFe-l.lMn NA NA 0.39 0.30 1.1 Xu [421 0.41 11.7Si-1.2Fe-1.8Mn NA NA 65.8 1.5 0.26 0.52 57.7 Cao[21,43] 12Si-l.23Fe-l.llMn 0.9 600 240 0.36 Cao[21] 12Si-l.llFe-l.03Mn 0.57 48.6 0.9 600 240 0.29 Cao [21] 12Si-l.22Fe-2.15Mn 0.35 71.3 600 240 0.30 1.8 Flores [19] 9.5Si-1.6Fe-2.2Mn 1.4 0.5 0.37 77 605 380 Flores [34] NA 0.4 9.2Si-1.5Fe-2.2Mn 73 1.5 640 180 0.42 0.64 42 12Si-l.lFe-lMn 20 0.9 605 Nijhof[36] 8Si-l.22Fe-l.12Mn 0.82 32.8 0.9 630 20 0.67 57.2 12Si-l.lFe-l.8Mn 605 20 0.57 0.47 1.6 35.2 8Si-l.22Fe-l.12Mn 0.79 0.9 640 30 0.63 0.42 0.64 42.9 11.5Si-1.12Fe-0.99Mn 0.9 605 15 VanDerDonk 56.9 11.4Si-1.09Fe-1.86Mn 0.58 0.47 1.7 605 15 [37] 61.4 11.2Si-1.58Fe-1.9Mn 1.2 0.61 605 15 0.50 58.9 12Si-2.07Fe-l.91Mn 0.53 0.85 0.9 605 15 NA 0.4 69.7 12Si-1.32Fe-lMn 30 0.75 605 12Si-1.30Fe-lMn NA 0.48 63.1 0.77 625 30 DeMoraes[20] 9.5Si-1.34Fe-1.5Mn NA 0.25 81.3 1.1 605 30 0.41 9.5Si-1.14Fe-1.5Mn NA 64.0 1.4 610 30 82 9.5Si-0.98Fe-1.5Mn NA 1.5 605 30 0.18 HSi-2.07Fe-2.03Mn 1 0.36 82.6 0.30 HSi-2.07Fe-2.53Mn 1.22 0.23 88.9 0.20 Matsubara [44] NA NA HSi-2.07Fe-3.00Mn 0.12 92.8 0.15 1.45 93.7 HSi-2.07Fe-4.15Mn 2 0.1 0.13

Filter (ppi)

Sludge separation method

NA

EM separation

NA NA NA NA 30 10 10 10 10 20 20 30 20 20 NA

Gravitation al separation

Filtration

Centrifuge

Centrifugal Separation

Electromagnetic Separation

The centrifugal separation technique was applied to directly remove iron-rich phases from the partially solidified aluminum alloy melts without any other elements addition. Matsubara et al [44] studied the sludge removal from Al-Il Si alloys by the centrifugal separation technique. The iron-rich phases moved to the edge side of the melt and the central part was purified, as shown in Figure 6 [44]. The rotational speed has a great influence on the purification efficiency [44].

The principle of Electromagnetic (EM) separation of particles from liquids was first proposed by Kolin in 1953 [45], and by Leenov and Kolin in 1954 [46]. When a uniform electromagnetic force is applied to a liquid metal, the metal is compressed by the electromagnetic force (Lorentz force) and a pressure gradient is generated in the metal. The non- or lessconductive particle suspended in the liquid metal receives only the pressure force because it does not experience the electromagnetic body force. As a result, the particle is forced to move in the opposite direction of electromagnetic force. The conductivity of Fe-rich phases is less than that of molten aluminum Researchers [35, 38-42, 47-51] have applied electromagnetic separation method to remove iron (in the form of sludge) from the melt. Kim et al [38] successfully purified Fe from 1.64 wt% to 0.45 wt% in A1-7SÌ scraps. Figure 7 [38] shows that the angular Fe-rich sludge settled on the side and bottom of the tube.

Figure 6. Microstructure of the transverse cross section of the centrifugal separated melt [44]

760

extreme purity aluminum (99.999 wt% Al) [64-67] It is reported that iron content decreased from 747 ppm to 24 ppm after 2 hours processing in a fractional solidification apparatus [68]. Electroslag Refining Electroslag refining process (ESR) is a secondary refining process in which the slag or flux is used both as a heating source and as a refining medium [69]. The process is already well established for ferrous metals but has not been used for aluminum refining. Mohanty's investigation reported 26% iron removal (from 0.22wt.% to 0.16wt.%) from commercial aluminum by ESR with a flux containing aluminum phosphide. [70]

Figure 7. Distribution of sludge with induction (left) 30 A and (right) 40 A [38] Table 2 shows reported iron removal efficiency by different sludge separation methods. Comparing with gravity and filtration, EM separation appears more efficient under similar experimental conditions. Furthermore, EM separation can also remove nonmetallic inclusions, which has been proven by many researchers [52-55]. Gravitational separation can achieve a high iron removal but needs long settling time. The combination of gravitational separation and filtration gives impressive iron removal results but the disadvantages during gravitational separation step are still there, in addition to the problem of clogging of filter pores. Although the centrifuge has relatively high removal efficiency, it is also unpractical for the large scale processing due to challenges with temperature control. [56]

Fluxing Refining Few references have reported that significant iron removal can be achieved by using flux. The studies of the current author showed that the addition of Na 2 B 4 0 7 flux with NaCl and KCl significantly lowered the iron content from 0.33 wt% to 0.18 wt% in laboratory experiments, as shown in Figure 8. [71] However, industrial scale experiments showed little removal of iron. Nijhof et al reported iron removal with a mixed flux of NaCl, KCl, NaF, an iron removal from -0.9 wt% to -0.7 wt% was obtained. [36]

Electrolysis The expensive process of three-layer-cell electrolysis is the most successful technique for the removal of iron and silicon from the molten aluminum so far [33]. Because the pure molten aluminum is the lightest, it will stay on the top of the three layers. Thus, the purified aluminum is obtained. However, the energy consumption for this process is relatively high, ~ 1314 kWh per kg so far [57-60]. This technique is also applied to purify commercial aluminum with low level of initial impurities into high purity [61]. C

Fractional Solidification

Na284Ö7 · * * · *

Figure 8. Relations between Fe concentration and Na2B407 addition at different holding time [71]

The technique is based on the distribution coefficient (k) of impurities. For k
Summary Different technologies to remove iron from aluminum have been summarized in the current paper. Manganese is the most common element used for neutralization. So far, three-layer-cell electrolysis is the most successful technique to remove iron from the aluminum but it is expensive and only suitable for

761

high purity aluminum production. The technology of EM separation can efficiently remove Fe-rich phases and could be a continuous process. EM separation is faster than gravitational separation, and avoids the problem of pore blockage during the filtration process. Acknowledgements This research is supported by the Research Board Grant, Laboratory of Green Process Metallurgy and Modeling, Material Research Center, and Intelligent Systems Center at Missouri University of Science and Technology (Missouri S&T). References 1. Green, J., Aluminum Recycling and Processing for Energy Conservation and Sustainability.Vol. 274. 2007. 2. Crepeau, P.N., Transactions of the American Foundrymen's Society, 1995.103: 361-366. 3. Khalifa, W., et al, Metallurgical and Materials Transactions A, 2003. 34(3): 807-825. 4. Mondolfo, L.F., Aluminum alloys: Structure and properties. 1976, London: Butterworths. 5. Damoah, L.N.W. and L. Zhang, Metallurgical and Materials Transactions B, No.4, 2010. 6. Stefaniay, V, et al, J. Mater. Sci, 1987. 22(2): 539-546. 7. Liu, P., et al. Mater. Sci. Tech., 1986. 2(10): 1009-1018. 8. Young, R., et al, Scr. Metall, 1981.15(11): 1211-1216. 9. Shabestari, Mater. Sci. & Engr A, 2004. 383(2): 289. 10. Skjerpe, P., Metall. Trans. A, 1987.18(2): 189-200. 11. Cooper, M., Acta Crystallogr, 1967.23(6): 1106-1107. 12. Kral, M., et al. Scripta Materialia, 2004. 51(3): 215. 13. Lu, L., Metall. & Mater. Trans. A, 2005. 36(3): 819. 14. Carpenter, G., et al, Scripta Metallurgica et Materialia, 1993. 28(6): 733-736. 15. Zheng, J.G., et al, Phil. Magazine A, 2000. 80(2):493. 16. Kral, et al. Metall. Mater. Trans. A, 2006. 37(6): 1997. 17. Barresi, J.G., et al. 1993. 18. Mahta, et al., Int. J. Cast Metals Res., 2005.18(2): 73. 19. Flores, A., et al., Intermetallics, 1998.6(3): 217-227. 20. de Moraes, et al. Mater. Trans., 2006. 47(7): 1731. 21. Cao, X., et al. Metall. Mater. A, 2004. 35A(5): 1425. 22. Murali, S., et al, Materials Chara., 1994. 33(2): 99. 23. Zhang, L., Removal of ironfromaluminum 2009. 24. Murali, S., Mater. Sci. Engr. A,1995. A190(l-2): 165. 25. Murali, S., et al., Mater. Sci. Tech, 1997.13(4): 337. 26. Wang, et al,Mater. Sci. Engr. A„ 2000. 280(1): 124. 27. Yie, S., et al., Mater. Trans., JIM, 1999. 40(4): 294-300. 28. Mulazimoglu, et al., Light Metals, 1994. 27(3): 1047. 29. Morris, L.R. and F.B. Miners, aluminum alloys. 1975 30. Mulazimoglu, et al. Aluminium, 1992. 68(6): 489-493. 31. Closset, B., et al., Light Metals, 1996: 737-744. 32. William, D.V. and W. Bernard, Aluminum base alloys oftheAl-Cu-Mg-Zntype. 1987. 33. Van der Donk, 3rd Inter. Sym. Recycling of Metals and Engineered Materials. 1995, Alabama; US A : 651 -661.

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34. Flores, A., et al., Light Metals, 1992: 845-850. 35. Li, T.X., Application of Electromagnetic Seperation in Aluminum Melt, 2004, Shanghai Jiao Tong Univ. 36. Nijhof, G.H., et al., Light Metals, 1996: 1065-1069. 37. Van der Donk, H.M., Method for Refining an Aluminum Scrap Smelt. 1998. 38. Kim, J.H., et al, Journal of Materials Science Letters, 2000.19(3): p. 253-255. 39. Li, TX., et al., Transactions of the Nonferrous Metals Society of China, 2003.13(1): p. 121-125. 40. Li, et al., J. Shanghai JiaoTong U., 2001.35(5): 664. 41. Jiao, W.-L., et al. Foundry, 2006. 55(2): 179-181. 42. Xu, Z., et al J. Mater. Sci., 2003. 38(22): 4557-4565. 43. Cao, X., N. Saunders, and J. Campbell, Journal of Materials Science, 2004. 39(7): 2303-2314. 44. Matsubara, J. Japan Ins. Light Metals, 1998. 48(2): 93. 45. Kolin, A., Science, 1953.117(2): 134-137. 46. Leenov, D. et al, J. Chem. Phys., 1954. 22(4): 683-688. 47. Yao, et al. J. of Northeastern U.,, 2001. 22(2): 127-129. 48. Xu, Z., et al., Cailiao Kexue Yu Jishu, 2001.17:306. 49. Zhang, L., et al., Special Casting & Nonferrous Alloys, 2005. 25(3): 131-132. 50. Jiao, Zhang, Foundry Technology, 2006. 27(3): 269. 51. Zhang, et al., Mater. Sci. & Tech., 2006.14(5): 524. 52. Shu, et al., Metall. Mater. Trans. A, 1999. 30(11): 2979. 53. Li, T X , et al, Acta Metall. Sinica, 2000.13(5): 1068. 54. Shu, et al, Metall. Mater. Trans. B, 2000. 31(6): 1535. 55. Shu, et al, Metall. Mater.Trans. B, 2000. 31(6): 1527. 56. Singleton, et al, J. Ins. of Metals, 1971. 99: 155-159. 57. Zhao, et al, Light Metals, 2008. 58. Lu, H, et al. Light Metals, 2004: 303-305. 59. Benkahla, B , et al. Light Metals, 2008: 451-455. 60. Martin, O, et al. Light Metals, 2008: 255-260. 61. Vire;, S. and L. Gauckler, Cell for the refining of aluminum. 1985. 62. Lux, A.L, et al, Metal. Trans. B, 1979.10(1): 71-78. 63. Lux, A.L. et al, Metal. Trans. B, 1979.10(1): 79-84. 64. Dawless, et al, J. Cryst. Growth, 1988. 89(1): 68-74. 65. Zhang, J, et al. Transactions of the Nonferrous Metals Society of China, 2006.16(1): 1-7. 66. Zhang, J, et al, Shanghai Jiaotong Daxue Xuebao, 2005.39(11): 1787-1791. 67. Zhang, J, et al, Chinese Journal Of Mechanical Engineering, 2006. 42(4): 64-68. 68. Toshiaki Iuchi, S, et al, process and apparatus for refining aluminum. 1988. 69. Benz, M.G. and B. Hills, Direct Processing of Electroslag Refined Metal. 1992.. 70. Mohanty, B.P, et al. Trans. Indian Inst. Met, 1986. 39(6): 646-651. 71. Gao, J.W, et al. Scripta Materialia, 2007. 57(3): 197200.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ELECTROMAGNETICALLY ENHANCED FILTRATION OF ALUMINUM MELTS Mark William Kennedy1, Shahid Akhtar1, Jon Ame Bakken1, Ragnhild E. Aune1'2 department of Materials Science and Engineering, Norwegian University of Science and Technology, N-7491 Trondheim NORWAY department of Materials Science and Engineering, Royal Institute of Technology, 100 44 Stockholm, SWEDEN Communicating author: [email protected] Keywords: Aluminum, Ceramic Foam Filters, Magnetic Field, Meniscus, Coil Abstract The major drawback of the use of Ceramic Foam Filters (CFF) for purification of aluminum is their low efficiency for particles in the range of 10-30 μπι. The application of electromagnetic force from an induction coil in combination with a filter can cause back mixing and recirculation through the filter media. In the present work an experimental set-up has been designed, built and verified by studying the meniscus behavior of molten aluminum under varying magnetic field strength. Batch typefiltrationexperiments with 30 ppi CFF were also conducted with and without a magnetic field using an A356 aluminum alloy containing 20% anodized and lacquered plates, as well as 20% composite material (A356 base and 15% SiC particles with size range 10-50 μπι). The presence of a magnetic field has proven to have both an effect on the build up of the filter cake, as well as on the re-distribution of particles within the filter.

Theory The use of electromagnetic fields is an emerging technology for the production of high-quality aluminum alloys with increased melt cleanliness[5]. Electromagnetic fields provide a means of influencing separation processes, without physical contact and added risk of contamination. The liquid metal will be acted upon by the electromagnetic Lorentz force F (N/m ) inducing motion: (1)

F = JxB

where J is the induced current density (A/m ), andί the magnetic flux density (also simply referred to as the magnetic field (T)). In addition to generating motion the electromagnetic field will also heat the metal.

Introduction

The convective flow induced in the liquid metal by the axial variation of the electromagnetic field of a short coil, may adversely affect the separation process[6]. In Figure 1, the variation of the axial magnetic field along the coil length for various coil diameters to length ratios, evaluated using the BiotSavart Law for an empty coil, are presented. It can be seen in Figure 1, that longer coils give a more powerful magnetic field, which can be very homogenous over much of their length, but with significant end effects. Shorter coils give less axial variation and overall weaker fields, as well as more radial variation of the magnetic field strength (not shown). The possible impact this has on the separation process is not clear; however, it could be expect that a very short coil behaves differently from a very long coil.

Due to increased demand for clean high-performance aluminum products, it has become an increasingly important task to reduce the level of impurities, and especially inclusions, in aluminum melts. It is a well known fact that non-metallic inclusions not only form porosity, but also result in stress concentration, which in turn may affect the static and dynamic properties of aluminum alloy products[l-2]. Traditional processes such as gravity sedimentation/flotation, degassing, flux refining and filtration, have difficulty meeting the cleanliness levels demanded in many applications, due to their low efficiency in removing micrometersized inclusions[3]. El-Kaddah et.al[A] presented the concept of combining electromagnetic separation withfiltrationto enhance the inclusion removal efficiency. They did, however, not provide sufficient experimental verification of the idea. In the present work an experimental set-up including several different induction coils has been designed and verified with respect to the generation of magnetic pressure and hence flow using a 50 Hz AC field. The molten metal meniscus behavior has also been studied as a way of directly evaluating the influence of the magnetic field on the molten metal sample. Different coil designs and a number of different currents were used. It was hoped to qualitatively observe mixing, the size of the meniscus and how great a magnetic pressure that could be developed, without damaging the apparatus or excessive overheating the molten aluminum. In order to demonstrate the effect of back mixing of the fluid through the filter media, batch filtration tests have been conducted with and without a magnetic field.

0

0.1

0.2

| - « - 0.1

0.3

0.4

0.5 0.6 Relative Length

0.7

Diameter over length Ratio for Coil * 0.3 * 0.6 — - i -» *—1.5

0.8

0.9

1

«- 2 ~ » ~ 3 - — 3J

Figure 1. Relative (short coil/equivalent length of infinite coil) axial magnetic flux density for an empty coil is plotted against the shape at constant current and turns per unit length.

763

1000.0

In a short coil, the presence of the work piece alters the strength of the axial component of the magnetic field in the air gap, due to the reduced coil aperture. Kennedy et. al.[l] have successfully used Equation (2) to estimate the strength of the axial magnetic flux density in a short coil, accounting for the length of the coil and the presence of the work piece: 5 0 * = kN* B^ = kN*ί0 Hœ = kN*M0 Nc Ic/Cc

(2)

where B0* is the axial magnetic flux density in the air gap of a short coil containing a work piece (T), Bœ the magnetic flux density of an equivalent length of infinite coil without a work peice (T), kN* a modified shortness correction factor for a loaded coil, ì0 is the magnetic permiability of free space = 4 π IO"7 H/m, Nc is the number of coil turns, Ic is the coil current (A) and ( is the length of the coil (m).

01

_ _

J Frequency, Hz [- - Solid

Liquid |

Figure 2. The electromagnetic penetration depth is plotted against the frequency for solid and liquid aluminum. The use of high frequency does, however, improve the efficiency of induction heating by increasing the ratio of (Radius/δ) for a fixed work piece radius. Overheating of the work piece will then limit the amount of current and hence the magnitude of the magnetic field that can be effectively used.

Vaughan and Williamson[8] developed an equation for a modified short coil correction factor, to account for the presence of the work piece: kN* = kN(l- [Dw/Dc]2) + [Dw/Dcf

I

(3)

where kN is the Nagaoka coefficient, Dw the diameter of the work piece (m), and Dc the diameter of the coil (m). The Nagaoka coefficient is a theoretical factor, which accounts for the reduced strength of the magnetic field in an empty short coil and it has been tabulated elsewhere [9-10].

Conducting experiments at mains frequencies, i.e. 50 or 60 Hz, have therefore the following significant advantages: • deep electromagnetic penetration, and • powerful magnetic fields without overheating of the work piece.

For the dimensions of typical induction coils, the Nagaoka coefficient can be conveniently estimated by the Wheeler formula[ll] as reformulated by Knight[12]:

Experimental

kN = 1/[1 + 0.4502 (Dc + Sc) / Q]

SetUp Two separate coil designs were produced for use with the meniscus and batch experiments. One coil was a traditional "long" coil (Coil 1), and the other an unusual multi-layer "short" coil (Coil 2), as pictured in Figures 3 a) and b). The multi-layer coil was specifically designed to produce the most powerful magnetic field possible over the desired length of the sample, using the available power supply. The specifications for the coils are given in Table I.

(4)

where τc is the electromagnetic penetration depth in the coil (m). The magnetic flux density in the air gap is extremely important as it directly relates to the magnitude of the electromagnetic pressure Pm (Pascals) that can be developed as follows: Ρ„ = Â0*2/(2ìïìΓ)

(5)

One problem associated with the electromagnetic separation techniques is the difficulty in generating large Lorentz forces deep within a highly conductive metal. This is especially true for AC fields at high frequencies, due to the skin effect: δ = (ñ/[π/ì0ìΓ])α5

Table I. Specifications for the long (Coil 1) and multi-layer coil (Coil 2).

CoiM Coils: Expt2 Inside diameter, mm 118 124 Average diameter, mm 300 Heiqht, mm Coil copper tube diameter, mm 6.35 Coil copper tube thickness, mm 0.762 Number of turns 41.0 71.7 Measured inductance of empty coil, μΗ Measured resistance at maximum load, ohms 0.0319 62 Avg. Coil temperature at maximum load, Deg C:

(6)

where p is the electrical resistivity (ohm m),/the frequency (Hz) and ìΓ is the relative permeability of the metal (1 for aluminum). The electromagnetic penetration depth (S) is reduced in proportion to the square root of the frequency (lNf), as presented in Figure 2 and Equation (6). As can be seen from Figure 2, the penetration depth becomes very small at high frequency, e.g. ~1 mm for pure liquid aluminum at 100 kHz, while at mains frequencies of 50 or 60 Hz, it is -35 mm. For magnetically thick samples (Radius/τ>3) the magnetic flux density declines nearly exponentially with each penetration depth, resulting in that only a negligible flux remains after about 3 penetration depths. As a result of this, it is therefore not possible to produce significant Lorentz forces deep within a thick work piece of a high conductivity metal at high frequency.

Coil 2 2-1 126 132 104 6 1 15.5 25.1 0.00928 23

Coil 2 2-2 140 146 105 6 1 15.5 29.0 0.01018 23

Coil 2 2-3 153 159 105 6 1 15.5 33.2 0.01126 25

Coil 2 2-1 + 2-2 126 132 107 6 1 31.0 103.3 0.01962 23

Coil 2 I 2-2 + 2-3 140 146 108 6 1 31.0 119.9 0.02193 27 |

The various sub-coils in Coil 2 could be placed in series or parallel to achieve the most advantageous configuration. The strongest magnetic fields (-0.2 T), as well as the most stable thermal operation, were achieved when the two inner coils were electrically connected in series and operated at the maximum available voltage (-28 V) with the water cooling in parallel and flowing counter currently. A cooling water flow of approximately 3-4 m/s was obtained at 4-6 Bar line pressure.

764

Alloy A standard non-grain refined A356 alloy was used for the study of the meniscus behavior. The chemical composition of the alloy is given in Table II. :i

mMM&:5& ,

pr

Table II. The chemical iomposil ion of the A356 ίdloy in wt. %. Mn Zn Alloy Si Mg Fe Ti Al type A356 7.03 0.41 0.091 0.008 0.005 0.11 Bal

a For the batch filtration tests a feed recipe was prepared containing: 60% A356 alloy, 20% anodized and lacquered plates, and 20% A356 composite material containing SiC particles with a size range 10-50 μιη. For the meniscus tests, as well as the filtration tests, the alloy was melted in an induction oven at 750 °C.

Figure 3. Photographs of a) the traditional long coil (Coil 1), and b) the multi-layer coil (Coil 2). The transformer used as the power supply was limited to a nominal voltage of 30 V and a maximum of 1500 A. Due to the high resistance of the coils, it was not necessary to compensate the power factor on the secondary side using a capacitor bank. Electrical measurements were conducted using a Fluke 43B Power Quality Analyzer. In Experiments 1 and 2, a 500 A Fluke current probe was used. This probe was however, replaced with a Fluke il000s AC current probe, with an accuracy of 1% and a precision of 1 A, on all subsequent experiments. The amount of energy entering the liquid metal was determined by subtracting the power lost in the coil (measured empty) from the total power of the system while operating with liquid metal. Electrical values were qualitatively verified using batch calorimetrie measurements, which agreed within the accuracy of ~ 5%. For Experiment 2, the obtained values for the current, as well as the power are only indicative as the current probe was used outside of its linear range (readings were later adjusted by calibration against the i 1000s probe).

Two filtration tests were conducted with 30 ppi CFF, i.e. one test with and one without a magnetic field, in the set-up shown in Figures 4a) and b). Maximum voltage was applied to the coil i.e. -28 V, for ten minutes after pouring the molten metal. The power was then stopped and the sample solidified. Results and Discussion Meniscus tests Four meniscus experiments and one batch filter test using magnetic field were conducted as summarized in Table III. Table III. Summary of experimental results Voltage Currenl Coil (V) (A) 1 1 14.21 364 2 1 thicker insulatioi 27.96 684 3 2 (1+2 in series) 28.15 730 4 meniscus 2 (1+2 in series) 14.40 373

Experiment

The experimental apparatus were constructed using two lengths of BIMEX 400 fiber riser supplied by Intermet Refractory Products Ltd., with a nominal inside diameter of 102 mm, an outside diameter of 120 mm, and a length of 150 mm. The two lengths were joined together using Fibrefrax moldable cement. The bottom portion of the apparatus was imbedded in a sand mold to provide a leakfreebottom. Bimex 400 fibre crucible, nominal 102mm inside dia., 300 mm high 2 inside coils (2-1 and 2-2) with 16.5 turns each, 125 mm inside dia., -108 mm high.

[* batch filte 2 (1+2 in series) 28.23

723

Total Load (W) 4140 16536 11900 3155

(W) 260 680 1270 460

12200 1930

Magnetic Theoretical Flux Magnetic Density Pressure (Bar) (T) 0.06 0.016 0.11 0.047 0.20 0.16 0.11 0.049 0.20

0.15

Liquid 1 Average Metal MerriscusJTemperatue Depth (mm) (oC) 1 (mm) 105 10 N/A 1 125 58 808 108 70 774 114 44 650 100 over and 100 709 under the filter N/A

The aluminum was, as previously mentioned, heated to 750°C and added to the unheated crucible where it cooled to ~710°C. Coil 1, the traditional long coil, was used initially. In the first experiment the lowest available voltage (14 V) was used, and in the second experiment the highest (28 V). It was not known at the time if sufficient heating could be produced to hold the sample temperature at either voltage. The meniscuses produced in the first and second experiments are presented in Figure 5a) and b).

Refractory blocks. Approximately 8.6 kg to balance possible magnetic over pressure (maximum 0.16 Bar). 30 PPI Ceramic Foam filter, 50mm thick 105 mm dia.

Casting sand -25mm up into crucible

Figure 4. a) A schematic diagram of the batch test apparatus, and b) a photograph of the same. In the meniscus experiments the coil was positioned in such a way that the bottom of the coil was in level with the top of the sand inside the fiber risers. In the batch filtration experiments, a CFF was cemented in at the junction of the two risers, as shown in Figure 4a). The coil was positioned in such a way that the midline of the coil (and the strongest point in the magnetic field) was in level with the bottom of the filter.

Figure 5. The meniscus measurement with long coil (Coil 1) with a magnetic flux density equal to a) -0.06 T, and b) -0.11 T. It was established that the long coil (Coil 1) was just capable of compensating for heat losses at the lowest voltage, i.e. 14 V, with

765

260 W delivered for heating the sample. No temperature control was, however, possible. At the highest voltage, i.e. 28 V, 680 W was produced heating the sample to over 850°C within about 10 minutes. The power was periodically interrupted during the experiment to avoid excessive overheating.

With the multi-layer coil (Coil 2), the aluminum was pushed up and partly out of the coil by the magnetic pressure. The observed meniscus height with the multi-layer coil, at the maximum transformer voltage, was sufficient to create back mixing through a filter depth of ~ 50 mm thickness.

Between experiments 1 and 2, it was found necessary to improve the electrical insulation of the coil, by the addition of a glass fiber sleeve. This caused the length of the coil to change slightly from 275 to 300 mm, and the inside diameter from 120 to 118 mm.

Most of the Lorentz forces generated due to Equation (1) are dissipated due to viscous forces (flow) and this results in much less than the theoretically possible height increase in the liquid metal, as indicated in Table ΙΠ (-10-30% of theoretical was observed).

Based on the first experiments, it was realized that a more powerful magnetic field could be produced if the coil was shortened, and if a multi-layer coil was used (giving similar total conductor length and impedance). Fabricating a coil in segments gave more electrical flexibility, and better cooling could be achieved as the length of each segment was shorter and more cooling water could be supplied at the available line pressure.

Batch tests In order to observe the largest possible difference between the gravity and the magnetic batch tests, the highest available voltage (-28 V) was applied to the coil to create the most powerful magnetic field possible. By adopting Equation (2), the magnetic flux density in the air gap between the coil and the work piece was estimated to be -0.2 T.

The inner two layers of the multi-layer coil arrangement (Coil 2) were used in Experiments 3 and 4, first at the highest voltage and subsequently at the lowest. The meniscus formed at the highest magnetic field strength was reduced in diameter and increased in height as presented in Figures 6a) and b). Its motion was extremely dynamic, and it had a tendency to fall onto the side of the coil missing Vi of a turn.

The filter element was positioned in such a way that the bottom of the filter was located in the position of the highest magnetic flux density, as shown in Figure 1. It was believed that this would cause a maximum upwards flow through the filter element. In the batch filtration tests it was found necessary to pre-heat the filter before starting the experiment to avoid an initial blockage under gravity conditions. Under the influence of the magnetic field preheating was, however, not required due to the resulting conditions caused by the field. The experimental data collected during the magnetic filtration experiment are also summarized in Table III. During the batch test it was established, that with a sufficient metal height over the filter it was possible to avoid the formation of an excessive meniscus when under the influence of a stronger magnetic field. This experimental condition could be used to reduce the degree of oxidation and dross formation caused by the melt circulation. The degree of melt circulation was established to be so intense, that some of the sand from the bottom of the crucible was drawn from the sand base of the experimental set-up and dispersed into the liquid metal phase located beneath the filter, see Figure 4. Sand particulates were in fact lifted and pushed into the bottom of the filter element giving proof that intense back mixing was taking place during the experiment. The sand base at the bottom of the crucible remained, however, intact during the gravity filtration experiment. The cross section of the bottom part of the ingot with and without the influence of a magnetic field is presented in Figures 8a) and b).

Figure 6. The multi-layer coil (Coil 2) with a magnetic flux density of a) -0.20 T, and b) -0.12 T. The experimental data obtained from Experiment 3 and 4 are summarized in Table III. The meniscus heights as a function of the estimated magnetic flux density in the air gap between the coil and the work piece are plotted in Figure 7.

■

0.10

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Estimated air gap magnetic flux density, T

Figure 7. The obtained meniscus height (+/- 5mm) is plotted against the magnetic flux density estimated using Equation (2) for the long coil (Coil 1) and the multi-layer coil (Coil 2).

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Figure 8. The cross section of the bottom part of the ingot obtained from the batchfiltrationtests a) with a magnetic field, and b) without a magnetic field.

For the long coil (Coil 1) a greater meniscus height was created for a given magnetic flux density as the coil, and therefore the axial magnetic field, extended well beyond the aluminum pool.

766

The cross section of the filter media obtained from the experiments with and without a magnetic field is presented in Figures 9a) and b). In the sample influenced by the magnetic field it was established that the center part of the filter media at the top had been worn down as a result of the strong upward melt circulation through the filter. The filter media with the gravity pouring conditions showed, however no sign of degradation.

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Representative Scanning Electron Microscope (SEM) micrographs of the filter cake with and without the influence of a magnetic field are presented in Figures 10 a) and b). It can be seen from the figure, that there are more SiC particles present in the filter cake in the sample poured under gravity conditions, while the SiC particle density is less in the sample poured and stirred with the magnetic field. It is assumed that the SiC particles, in this case, were more evenly dispersed due to the upward flow of the aluminum caused by the magnetic back mixing.

Figure l l . A simulation of theflowvelocity and magnetic head produced in cold crucible experiments [14]. Acknowledgements The authors wish to express their gratitude to Egil Torsetnes at NTNU, Trondheim, Norway, for helping with the design and construction of the experimental apparatus. Deepest gratitude is also due to Kurt Sandaunet at Sintef, Trondheim, Norway, for the use of the Sintef laboratory and his contribution in the execution of the experiments. Special thanks to Liss Pedersen at Alcoa, Lista, Norway, for the supply offiltermaterials. The author would also like to acknowledge the funding from the Norwegian Research Council through the RIRA project.

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Figure 10. SEM micrographs of thefiltercake a) with a magnetic field, and b) without a magnetic field (magnification X 100).

1.

Conclusions

2.

An experimental set-up has been designed and verified for the fluid re-circulation under different transformer voltages and coil designs. The largest meniscus height was observed using a multilayer coil arrangement and a maximum transformer voltage of 28 V. Melt circulation under the influence of a magneticfieldhas proven to be effective in distribution of SiC particles.

3.

4.

Future Work

5.

In industrial casting operations, a bulk flow field is produced in the filter by the gravity flow of metal, passing though thefilterat a velocity of ~1 cm/s[13]. An example from the literature[14] is presented in Figure 11, indicating that a magnetic field should be capable of generating an internal flow velocity an order of magnitude greater than the normal industrial flow velocity range. Additional parametric studies will be conducted to explore the influence of the magnetic field on aluminum and inclusions flowing through a filter element.

6.

7.

767

O. Majidi, S.G. Shabestari, and M.R. Aboutalebi, "Study of Fluxing Temperature in Molten Aluminum Refining Process", Journal of Mat Pro Tech, 182 (2007), 450-455. B.W. Zhang, Z.M. Ren, and J.X. Wu, "Continuous Electromagnetic Separation of Inclusion from Aluminum Melt Using Alternating Current", Trans Nonferrous Met SOC China, 16(2006), 33-38. K. Li, J. Wang, D. Shu, T.X. Li, B.D. Sun, and Y.H. Zhou, "Theoretical and Experimental Investigation of Aluminum Melt Cleaning Using Alternating Electromagnetic Field", Materials Letters, 56 (2002), 215- 220. N. El-Kaddah, A.D. Patel, and T.T Natarajan, "The Electromagnetic Filtration of Molten Aluminum Using an Induced-Current Separator", Journal of Materials, 46 (1995). Z.T. Zhang, Q.T. Guo, F.Y. Yu, J. Li, J. Zhang and T.J. Li, "Motion Behavior of Non-Metallic Particles Under High Frequency Magnetic Field", Trans. Nonferrous Met SOC China, 19 (2009), 674-680. D. Shu , B. Sun, K. Li, and Y. Zhou, "Particle Trajectories in Aluminium Melt Flowing in a Square Channel Under an Alternating Magnetic Field Generated by a Solenoid", Scripta Materialia, 48 (2003), 1385-1390. M.W. Kennedy, S. Akhtar, J.A. Bakken, and R.E. Aune, "Review of Classical Design Methods as Applied to

8.

9. 10. 11. 12. 13.

14.

Aluminum Billet Heating with Induction Coils", submitted to EPD Congress 2011, TMS. J. Vaughan and J. Williamson, "Design of Induction-Heating Coils for Cylindrical Nonmagnetic Loads," American Institute of Electrical Engineers, Transactions of the, 64, (1945), 587-592. H. Nagaoka, "The Inductance Coefficients of Solenoids," Journal of the College of Science, 27, (1909), 18-33. E. B. Rosa and F. Grover, "Formulas and Tables for the Calculation of Mutual and Self Induction," Scientific Papers of the Bureau of Standards, No. 169., (1916), 5-231. H. Wheeler, "Simple Inductance Formulas for Radio Coils," Proceedings of the IRE, 16,(1928), 1398-1400. D. Knight. (2010, August 25). 3.1. Solenoids: Part 1. http://www.g3ynh.info/zdocs/magnetics/part_l.html C. Dupuis, G. Bιland, J.P., "Filtration Efficiency of Ceramic Foam Filters for Production of High Quality Molten Aluminum Alloys," (Paper presented at the 32nd Annual Canadian Conference of Metallurgists, 29* August - 2nd September), 1993. E. Baake, A. Umbrashko, B. Nacke, A. Jakovics, and A. Bojarevics, "Experimental Investigations and LES Modeling of the Turbulent Melt Flow and Temperature Distribution in the Cold Crucible Induction Furnace", (Paper presented at the 4 th International Conference on Electromagnetic Processing of Materials, Lyon, France, 14-17 October, 2003).

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

A REVIEW OF THE DEVELOPMENT OF NEW FILTER TECHNOLOGIES BASED ON THE PRINCIPLE OF MULTI STAGE FILTRATION WITH GRAIN REFINER ADDED IN THE INTERMEDIATE STAGE John H Courtenay (1) Frank Reusch (2) Stephen Instone (3) (1) MQP Limited, 6, Hallcroft Way, Knowle, Solihull West Midlands, B93 9 EW (2) Drache Umwelttechnik GmbH, D-65582-Diez, Germany (3) Hydro Aluminium Deutschland GmbH, Research and Development Centre.P.O.Box 2468, D-53014 Bonn, Germany Keywords: Filtration, Cyclone

grain refiner addition, via a suspected agglomeration behavior and alteration of filtration mechanism, prevented bridge formation in finer pore ceramic foam filters thus reducing the hitherto observed very highfiltrationefficiencies reported in earlier work (5).

Abstract Recent developments in filtration technology based on the principle of using a three stage process where a ceramic foam filter is operated in cake mode in the first stage; grain refiner is added in a second chamber and a further filtration means is used in the third stage to remove oxide inclusions or agglomerates originating from the grain refiner addition are reviewed. The first development - the XC Filter, was presented by Instone et al in 2005 and described a system where a small deep bed filter (DBF) was successfully applied in the third chamber. A second prototype multi stage filter was described at TMS 2008 based on the same principle but with a cyclone deployed in the final chamber. An industrial prototype was constructed based on water modeling work and plant trials were undertaken. The current stage of development of each system and their relative merits are evaluated.

A more recent study by Lae et al (5) using a filtration pilot showed that from the point when standard AT5B grain refiner was added to alloy 5182 at a casting speed of 1.8cms/s the post filter Limca count increased from 9k/kg up to 20k/kg and the filtration efficiency decreased from 71% to 31%. It was concluded that this was due to the interaction of grain refiner particles with the bridge formation mechanism observed in non grain refined melts. The XC Filter - three stage filter with mini bed In 2005 Instone et al (6, 7) described a new design of filter unit named the XC filter which gave superior filtration efficiency achieved by the combination of ceramic foam filtration and deep bed filtration. Importantly this design comprised a three chamber unit with a ceramic foam filter in the first chamber, grain refiner addition in the second chamber and a small bed filter in the third chamber.

Introduction Peter Waite (1) stated that there was a definite need to develop an efficient, low hold up volume, filtration process capable of treating high flow metal rates. The object of the work reviewed is in each case to develop a filter that could deliver the high efficiency performance of a deep bed filter but with low hold up volume, low floor space requirement and the ability to be used economically in conjunction with frequent alloy changes.

The design concept is shown schematically in Figure 1.

The phenomena of enhanced filtration efficiency in ceramic foam filters that could be achieved by adding the grain post filter reported by Towsey et al (2) provided the starting point. This phenomenon was first reported by Kakimoto et al (3) in 1996 in relation to the operation of porous tube filters. Kakimoto concluded that bridges of CaO particles that tended to form at the pores at the surface of a tube filter were "suppressed" by the addition of boron containing grain refiners. That is the addition of titanium diboride particles prevented the formation of a stable filter cake which is initiated by the formation of bridges as a first stage to support the subsequent cake formation. This conclusion was reached on the basis of metallographic examination of spent tube filters. In 2002 Towsey et al (4) reported the results of an extensive study on the effect of addition of various grain refiner compositions on the performance of ceramic foam filters with the conclusion that

Figure 1. XC Filter design, Instone et al, (6, 7) Several prototypes of this filter were built and tested over the period 2000-2005 at the pilot DC casting center at the Rheinwerk smelter in Neuss Germany. More than 80 evaluation casts using

769

this technology were conducted in this period. This evaluation program extended through to a three week pre-production trial. The results of the pilot testing compiled using LiMCA and PoDFA measurement techniques showed that excellent filtration efficiency could be achieved. These results have been reported previously by Instone et al (6, 7). In this paper the results achieved during the pre-production trials will be discussed.

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LiMCA 1 & PoDFA Furnace Exit ("before")

Casting Furnace

XC-Filter Grain Refiner Addition Point

This extended casting campaign was performed with the XCFilter prototype shown in Figure 2 to demonstrate the long-term stability of the technology under production conditions. This work also helped to finalise the various scale-up parameters for a production unit. Unlike previous trials which were conducted using potroom metal, these trials were performed using metal prepared from recycled foil scrap or processed dross and delivered to the furnace as molten pre-alloyed metal.

LiMCA 2 & PoDFA After XC-Filter ("after") Mould Table

Figure 3. Plan view schematic of centre 40 at NESIC, showing positions for melt quality measurement by LiMCA and PoDFA and the position for the addition of the grain refiner Melt treatment: The furnace was de-drossed after filling with the liquid metal. No alloying or gas injection was performed. On occasion, in order to increase the inclusion loading during casting, the melt in the furnace was stirred manually with a paddle. A typical LiMCA curve showing the effects of this stirring operation is shown in Figure 4. Inclusion removal efficiency was found to be higher for the stirred melt with higher inclusion levels. This effect is shown for several casts in Figure 5. Figure 6 shows the increased filtration efficiency for different particle sizes. This figure also shows that the level of inclusions at the filter exit did not increase as a result of the stirring operation, which demonstrates the robustness of the XC filter. Figure 2. Industrial prototype XCfilterused for these trials An overview of the casting program and a summary of several important parameters is presented here: • • •

Total of 37 casting trials performed in this campaign Changed bed filter filling after 13 casts due to small leak at tap hole. 7 CFFs used - changed out after on average 5 casts

The following parameters were used for all casts: • • • • • • •

1 mould of 1750mm x 600mm size Flow rate of =10t/h Casting speed of 62 mm/min 30ppi CFF - 3 suppliers 5-layer DBF- bed filter construction as used in the previous work Grain refinement after CFF and before DBF (3:1 TiB rodfromLSM at 0.6kg/t) Layout of casting and measurement equipment shown schematically in Figure 3.

Time (min)

Figure 4. LiMCA N20 run chart and average N20 counts for charge 42997

770

9 before filtration, before stirring ■ after filtration, before stirring H before filtration, after stirring 3 after filtration, after stirring i- Efficiency before stirring - Efficiency after stirring

Trial 1 2 3 4 5 Figure 5. N20 values measured at the two LiMCA positions before and after stirring of the melt in the furnace 5 i

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Figure 7. Effect of inclusion loading on filtration performance Capacity of CFF and bedfilterin the XC-Filter An important factor of the performance of the CFF in the XCFilter is the capacity of the filter, that is, the amount of metal that can be passed through the filter before it has to be replaced. None of the CFFs was replaced due to blockage of the filter or deteriorations in the filtration efficiency. Never the less, the filter capacities achieved in these trials were significantly higher compared to the results gained in previous casting campaigns. The main difference between these two campaigns is the metal source. In the past only pot room metal was used for the trials while remelted foil scrap or metal won from processed dross was used in the current campaign. This recycled metal had of course already been inoculated with TiB grain refiner.

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Metallographic inspections of the used filters supported this argument. Al-Oxides, predominantly present as oxide-skins were mainly observed in the samples. No bridging in the uppermost cells of the CFFs, such as previously documented, was observed in any of the used filters examined. As has been previously suggested, Titanium Boride particles present in the recycled metal may have prevented the formation of bridges in these filters.

Particle Size (pm)

Figure 6. Inclusion size distribution for stirred and settled melts measured at the two LiMCA positions. Performance of XC-filter at different levels of inclusion loading Figure 7 shows the average LiMCA N20 values for all charges before and after filtration. The filtration efficiency is sorted by the inclusion loading in the metal coming from the furnace. As seen in the previous examples, the filtration efficiency results show a correlation between filtration efficiency and inclusion loading. Four ranges of inclusion loading were defined: • • • •

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A reduced efficiency of the CFF may have resulted due to the presence of these grain refiner particles in the melt. This in turn would have increased the number of inclusions entering the bed filter. A higher total filtration efficiency and bed filter capacity under production conditions could be expected to when using exclusively potroom metal.

ultra low loading below 1. Ok/kg, low loading between 1 .Ok/kg and 2.0k/kg, medium loading between 2.0k/kg and 5.Ok/kg, high loading above 5.Ok/kg.

The overall filtration efficiency when using recycled metal was also considerably higher than that of a standard CFF when used in conjunction with TiB grain refiners.

For these four ranges the average inclusion loading before and after filtration and filtration efficiency were calculated. For inclusion loading >lkcounts/kg (N20) filtration efficiency was greater than 85%.

PoDFA samples, before and after filtration, were taken at a casting length of about 2m for each cast. Selected samples were chosen for further metallographic investigation. The same methodology as used for the evaluation of the LiMCA data, was applied to the PoDFA data and confirmed at least semiquantitatively the good filtration performance of the XC filter. The metal exiting the furnace contained a variety of particles such as Al-oxides, Mg-oxides and Al-nitrides. After passing through the filter the amount of inclusions present in the melt was significantly reduced with the predominant inclusion type being Al-oxides.

771

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Optifîlter - three stage filter with Cyclone

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In 2005 Katgerman (8) described work on water modeling and computer modeling of flow control devices. It is well known that dams and weirs placed in a launder section or a chamber, such as that forming part of a degassing apparatus, contribute to the removal of inclusions. However, Katgerman concluded that, although this could be effective for small concentrations of particles, this technique suffered from the drawback that small fluctuations in the flow behavior may reintroduce the sunken particles (collected at the base of the dams due to settling out by virtue of their higher density relative to liquid aluminium) into the metal flow.

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Instead, based on flow calculations, the concept of a cyclone was developed and subsequently proven in terms of effectiveness by further numerical and water modeling experiments.

Figure 8. Predicted efficiency for particle removal by cyclone The model shows an expected removal efficiency of approximately 50% for particles >60micron with a flow velocity of 0.5 m/s rising to 80% for particles > 100 microns. However, the true removal efficiency can only be determined by actual operation of the unit in practice.

Design of the prototype A design concept was determined based on the above considerations and comprised three chambers: • • •

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Afirstchamber containing a ceramic foam filter A second chamber for the addition of grain refiner A third chamber containing a cyclone.

A key requirement of the design was that it should be able to work effectively with a maximum available head height at the casting pit of 1000 mm. The cyclone itself was required to fit into maximum external space of 1000mm x 1000mm x 1000mm which meant in practice, after allowing for the metalwork and refractories, that the maximum internal height for the cyclone would be 740 mm. The efficiency predicted is shown below. The figure depicts the % particle removal efficiency for different flow velocities and two types of cyclone design, type 1 without a container for inclusion capture at its base and type 2 with a container for inclusion capture. The results of the flow modeling and design of the cyclone have been presented in detail separately by Turchin et al (9)

Figure 9. Final design of three stage "OptiFilter" filtration unit. First casting trials Casting trials were carried out at Trimet Aluminium under the following conditions: • • • • • •

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Alloy: 1000 series 790 Number of billets: 32 Billet length: 2000 mm Cast size: 13,000kg Casting speed: 200kg/min Casting temperature: 790° C

Some difficulties were experienced with preheating the first chamber containing the ceramic foam filter and the cyclone chamber

Type 2

772

Despite these initial problems a heat was successfully started at the second attempt and metal flowed through the filter for some 20 minutes before freezing off on the casting table.

It was considered that restriction to the outlet flow must be due to either insufficient head height to overcome the resistance to flow of the cyclone and or an insufficient cross sectional area at the cyclone inlet slot to allow adequate flow through the cyclone. Further water modeling It was decided to re validate the flow modeling and conduct further water model tests at Delft University to verify the model. The water model was set up with a facility to use a variable inlet and head height. After extensive testing it was concluded that both an enlarged inlet slot cross sectional area and additional head height of + 150mm would be required to ensure that the outlet pipe was completely filled. Modifications were subsequently made to the prototype to provide an additional 150mm of head height and further liquid metal trials were undertaken on a sow caster at Trimet. The result was very satisfactory with the outlet pipe being completely filled with a 2.4 tonne crucible being cast through the Optifilter in less than 5 minutes.

Figure 10. The prototype "Optifilter" filtration unit under pre heat at Trimet Aluminium.

The next stage will involve setting the system up on the research casting pit with a 150mm available head height difference and casting two billets over the period of one hour. Podfa samples and billet slices will be taken to enable an assessment of the effect on cleanliness of the Optifilter to be undertaken.

Second casting trial Following modifications to the prototype to improve insulation and pre heating a second casting trial was attempted using the same conditions as in the previous test. However, although liquid metal passed successfully through the cyclone it was observed that the flow rate was very low - the exit tube only filled to approximately 10% of the exit cross section and it was estimated that the flow rate being achieved was of the order of 3t/h instead of the 15 t/h target.

Discussion The two technologies reviewed stem from a common starting point: the observation that addition of grain refiner in front of a CFF reduces filtration efficiency by preventing the formation of bridges at the top surface of the filter which act to form a filter cake which changes the filtration mode from being substantially depth filtration to cake mode filtration. To use this phenomena in a practical way it is necessary to employ a third chamber to ensure that no oxide stringers or boride agglomerates from the grain refiner can pass from the filter to the casting table. Two approaches have been developed to the prototype stage, in one case using a mini bed in the third chamber, in the other applying a cyclone. The XCfilteron the one hand has progressed successfully to a pre production stage trial were 37 heats were cast over a period of three weeks indicating the robustness of the technology and its suitability for casthouse operation. The results in terms of filtration efficiency were excellent with efficiencies of 97% for inclusion sizes over 50 microns together with average post filter N-20 Limca counts of 400/kg.

Figure 11. View of the exit flow from the cyclone showing that the exit flow tube is only partially filled. From these second series of trials it was concluded that:

The Optifilter initially suffered from difficulties in achieving the desired flow rate through the cyclone. At first this was thought to be due to problems with pre heat or insufficient inlet slot cross sectional area. Both of these were corrected but the flow rate was still inadequate and subsequent further water modeling pinpointed insufficient head height. When the head height difference was increased to 150 mm the outlet pipe was completely filled and a flow rate of >20t/hr was achieved on a sow casting station.

The measures to improve insulation and increase preheat temperature had been successful. Nonetheless despite this the unit still froze off and this was due to the outlet flow rate having been restricted to only 3t/h.

773

Ultimately it is anticipated that with sufficient head height that the cyclone will function satisfactorily however the requirement for additional head height of the order of 150mm may make it difficult to retro fit the Optifilter into existing casting lines without significant modification. Nonetheless should it be successful the cyclone would provide a low maintenance third chamber which would make it possible to use the system for frequent alloy change or low volume operations were it is necessary to drain the box completely between charges.

demonstrated in casting trials that a head height difference of approximately 150mm will be required between the inlet and outlet levels to drive metal through the cyclone at a sufficient rate to achieve the required flow rate for casting. 2. Further trials are planned with a modified prototype with a 150mm head height difference during which filtration efficiency measurements will be conducted. 3. Providing the above are successful the system would potentially provide high efficiency filtration solution for casthouses where frequent alloy changes or low volume operations necessitate fully draining the filtration unit between casts.

On balance the ability of the XC filter to deliver ultra high efficiency is already proven under production conditions and the system will be suitable for application were either long runs of the same alloy are made or were it is possible to flush through the relatively small hold up volume to allow a number of alloy changes to be made in sequence without the need to drain the bed. On the other hand if successful the Optifilter will provide a similar high efficiency filtration solution for frequent alloy change or low volume operations where it is necessary to fully drain the filtration unit.

Acknowledgments The Authors would like to express their sincere thanks to the management teams and casting personnel of: Hydro Aluminium Deutschland GmbH Research and development centre Bonn; and Trimet Aluminium AG casthouse at Essen. References P.Waite, "A Technical Perspective on Molten Aluminium Processing" Light Metals. 2002. 841-848, N.Towsey, W. Schneider, H-P.Krug, A.Hardman and N.J.Keegan, "The Influence of Grain Refiners on the Efficiency of Ceramic Foam Filters", Light Metals. 2001. 973 ^977 T K.Kakimoto, T. Yoshida, K.Hoshino and T.Nishizaka, "The Filtration of Molten lxxx Series alloys with Rigid Media Filter", Light Metals. 1996.833-838 « N.Towsey, W. Schneider and H-P. Krug, "The Effects of Rod Grain Refiners with Differing Ti/B Ratio on Ceramic Foam Filtration", Light Metals. 2002. 931-935 « E.Lae, H. Duval, C.Riviere, P. Le Brun and J.-B.Guillot, "Experimental and Numerical Study of Ceramic Foam Filtration", Light Metals 2006. 753-758 · S.Instone, M.Badowski and W.Schneider, "XC Filter - A Filter for Increased Filtration Performance", 9 th Australasian Conference and Exhibition Aluminium Casthouse Technology, 2005, 259 267. ^ S.Instone, M.Badowski and W.Schneider, "Development of Molten Metal Filtration Technology for Aluminium", Light Metals 2005. 933 - 938 L.Katgerman and J. Zuideman,"Upstream Fluid Flow Particle Removal", Light Metals. 2005. 927 - 931 A.N.Turchin, D.G.Erskin, J.H.Courtenay and L.Katergerman, 10th Australasian Conference and Exhibition Aluminium Casthouse Technology 2007, 225-230.

Conclusions In summary the following conclusions can be drawn regarding the stage of development and relative merit of each of the two technologies: XC Filter - three stage filter with a mini bed 1.During a production trial campaign involving casting 37 heats in sequence over a three week period a consistently high metal quality for the filtered metal was achieved through out and the prototype system functioned satisfactorily. The metal quality (LiMCA N20) after filtration averaged 0.4k counts/kg and consistently bettered the target value of
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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

WETTABILITY OF ALUMINIUM WITH SIC AND GRAPHITE IN ALUMINIUM FILTRATION Sarina Bao 1 , Anne Kvithyld2, Thorvald Abel Engh1, Merete Tangstad1 Norwegian University of Science and Technology, Trondheim, NO-7491, Norway 2 SINTEF Materials and Chemistry, Trondheim, N-7465, Norway Keywords: Wettability, Contact angle, Aluminium, Graphite, SiC, Time dependent property Abstract The aim of aluminium filtration is to remove inclusions such as AI3C4. For inclusions to be removed they have to come in close contact with the filter walls composed of A1 2 0 3 or SiC. It is therefore important that the molten aluminium has close contact with the filter wall. In addition to wetting properties between inclusions (AI4C3) and molten aluminium, the wettability of the filter (SiC) by aluminium is determined in sessile drop studies in the temperature range 1000-1300°C. Wettability changes with time in three successive steps and improves with time. To describe wettability at filtration temperatures employed in the industry of around 700°C, the results will be extrapolated to this temperature in future work. Introduction In filtration it is important that particles to be removed contact, or come very close to the filter walls. Therefore the molten metal carrying the inclusions must come into close contact, i.e. wet the filter material. In filtration of aluminium, alumina is the most common filter material, even though alumina is not wetted very well by aluminium [1-3]. Therefore one should investigate the use of alternative filter materials with improved wetting. In the laboratory, SiC and graphite demonstrate good wetting by molten aluminium. Problems with these materials exist, as SiC is easily oxidized to Si0 2 and graphite reacts with Al to give AI4C3. Once metal has entered the filter, oxidation is not a problem because the solubility of oxygen in aluminium is low, around 1.43xlO- 4 at.%at700°C[4]. In the molten aluminium/filter environment, the oxygen potential is very low. To study wetting of SiC and graphite in such a system, a high vacuum laboratory furnace containing only a minute amount of oxygen has been chosen. 1.1 Al-SiC System Wettability of aluminium has been reported for single crystal SiC [5], reaction bonded SiC [6], and sintered SiC [6, 7] (Figure 1). Wettability may change with the preparation process and the sintering aids for SiC. Aluminium has a decreasing contact angle on SiC with increasing temperature. Reaction bonded SiC has better wettability with aluminium than single crystal SiC and sintered SiC. For example, Figure 1 shows that at 830 °C, reaction bonded SiC, single crystal SiC, and sintered SiC have contact angles of 37°, 60°, and 107°, respectively.

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600 700 800 900 1000 1100 1200 1300 Temperature/°C Figure 1 The equilibrium contact angle vs. temperature for aluminium on SiC from literature [5-10] SSiC- sintered SiC; RBSiC- reaction bonded SiC; SCSiC- single crystal SiC In aluminium filtration, filters are primed to allow metal to flow through the filter without freezing. The pre-heating temperature is probably up to 1000°C and SiC filter materials will oxidize at this temperature. The oxidation of SiC is very slow at room temperature [11], but SiC will react with air to form a silica-rich surface layer at temperatures above 700 °C in air. SiC oxidation is controlled by the diffusion of oxygen molecules (or oxygen ions) through the thin oxide film [12]. The oxidation behaviour of SiC is also influenced by factors such as moisture, SiC particle size, and metal impurities [12]. In addition to preheating temperature of 1000°C, one should also take into account that the ceramic foam filters (CFF) are "baked" at temperatures above 1000°C. Thus the wettability of oxidized SiC by molten aluminium is important to study in aluminium filtration. V. Laurent et.al.[8] have shown that silica acts as an oxygen donor to extend the life time of the AI2O3 layer on an aluminium drop by the reaction 3Si0 2 +4 Al(l)->2 Al 2 0 3 +3 [Si]

(1)

The reaction between aluminium and Si0 2 does not improve the wetting of aluminium on SiC, shown as the circle marker in Figure 1. The silica layer is changed into alumina and the equilibrium contact angle is the same as aluminium on alumina. With time in high vacuum, a thin initial layer of silica can be removed by reactions (1) and (2) successively. 4Al(l)+Al 2 0 3 -+3Al 2 0(g)

(2)

Aluminium is then in direct contact with SiC and wetting is determined by the Al-SiC system. The role of the silica here is to delay the wetting of aluminium on SiC [8].

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Addition of Mg, Ca, Ti and Pb to aluminium enhances the wetting of Al-SiC [10, 13]. For example, the addition of 5 wt% Mg in aluminium results in wetting (θ<90°) at 700-960°C [6] since Mg reacts with A1203 and consumes the oxide layer on Al. The addition of Si will not influence wetting [13].

graphite radiation shields, located in a water-cooled vacuum chamber. The chamber was fitted with windows to allow a digital video camera (Sony XCD-SX910CR) to record the shape of the droplet. The contact angles and linear dimensions of the images were measured directly from the image of the drop using the Video Drop Shape Analysis software.

When the liquid composition reaches the peritectic point at 650°C, the following ternary quasi-peritectic reaction [9] occurs isothermally (See Figure 20). 3SiC+4Al(l) -> Al4C3+3[Si]

The experiments were carried out using 99.999% pure aluminium and substrates produced from 98.9% single crystal SiC (Washington Mills, Norway) and ISO-88 Graphite substrates. The average roughness of SiC and graphite was 51.25nm and 179.76nm, respectively. The aluminium rod with a diameter of 2mm was cut into small pieces around 2 mm in length, then polished by 500 mesh SiC paper and cleaned with ethanol in order to prevent further oxidation. The SiC single crystal was cut and ground with 1200 mesh diamond paper to get a flat surface, then dried in a closed furnace at 100°C until the experiment was performed in order to minimize the oxidation of the SiC surface. When the wetting furnace attained the high vacuum of 10"5 mbar, the sample was quickly heated to 950°C in about 80s, then heated to 1000, 1100, 1200, and 1300°C at the rate of 50°C/min, as shown in Figure 2. Although the furnace temperature overshoots to 1100°C at the first 80s, it does not affect the wettability at a lower temperature such as 1000°C, since the oxide skin holds the liquid metal at the beginning. In all of the experiments, the contact angle and dimensions of the drop were recorded during the isothermal period at 1000, 1100, 1200, and 1300°C. For the dimension reading, time=0 was taken to be the beginning of the isothermal period.

(3)

Thus the addition of Si to aluminium prevents the formation of AI4C3 and does not affect wettability [9, 14]. The free Si in reaction bonded SiC also prevents the formation of A14C3 [15] and is reported to be effective in promoting wetting by liquid aluminium in the temperature range of 700-1040°C [6]. This may be due to the additional Si-Al bond on the Al-SiC interface and the aluminium penetration into the SiC along the intergranular free silicon resulting from the reaction bonding process in the extensive reaction zone [16]. 1.2 Al-Graphite System Carbon filters have been industrialized as petrol coke filter-DUFI during 1970-1985, known for removing hydrogen, alkaline metals, and non-metallic inclusions in plants operated by ALUSUISEE Group and elsewhere [17]. However AI4C3 has a negative effect on the possible use of the graphite as a filter. The AI4C3 crystals formed as a reaction product are brittle and highly sensitive to moisture and promote accelerated fatigue crack growth rates due to their hydrophilic nature [18]. The wetting behaviour of the Al-graphite (or AI-AI4C3) system will be investigated from the view point of inclusion-metal wettability in filtration. During the last few decades, several studies had been devoted to the wetting behaviour of the Al-graphite system. As a reactive wetting system, it is agreed that 1) the final or steady contact angle is equal or close to the equilibrium contact angle of the liquid on the reaction product, A14C3 [19]; According to Tomsia et al. [6] in [19], the wetting behaviour of the reaction product is governed by the formation of adsorption layers at the interface, rather than by the subsequent nucleation and growth of the reaction product; 2) wettability would not be improved by the chemical reaction itself. The interatomic force is not correlated to the Gibbs free energy as an exchange of atoms is involved in a chemical reaction [20]. In the Al-graphite system, the final contact angle, θ2 on the reaction product AI4C3 is much lower than the initial contact angle, θ0 on the original substrate graphite.

Time/min Figure 2 Example of the registered temperature for the experiment holding the sample at 1200 °C Results

Experimental Procedure

3.1 Al-SiC System

The study of the wettability of Al-graphite and Al-SiC by the sessile drop technique gives rise to two main points. The first point is that liquid aluminium reacts with substrates forming AI4C3 at and close to the interface, after which the wetting of Alsubstrate transforms to the wetting of an AI-AI4C3 system. The second point is that liquid aluminium is covered by an oxide layer which inhibits wetting. Contact angles of liquid aluminium on the substrate were measured using the sessile drop method. The experimental apparatus is schematically shown in [21]. The apparatus essentially consists of a horizontal graphite heater surrounded by

Figure 3 and Figure 4 show the time dependent variations of wetting properties for liquid aluminium on SiC at 1000°C. Three kinetic steps can be distinguished: the first step, where the contact angle decreases rapidly (in 45min) from the initial contact angle θ0~120° to θ!~81°; the second step where the contact angle continues to decrease to a relatively low value of θ2~65°, but at a slower rate; and the third step where the contact angle stabilized at θ2 after approximately 150min. The stable base diameter and the sessile volume allow the stable contact angle in the third stage to be measured.

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Similar three kinetic steps: (1) rapid decrease, (2) slow decrease and (3) stable contact angle at 1100°C are shown in Figure 5 and Figure 6. The stable base diameter and the sessile volume guarantee the stable contact angle in the third stage. A lower stable contact angle is obtained in a shorter time at 1100°C than at 1000°C (See Table 1). Efforts to obtain the Al-SiC contact angle at even higher temperatures, for example 1200°C, were made, as shown in Figure 7 and Figure 8. Unfortunately, the evaporation of aluminium at higher temperatures is so high that the aluminium droplet disappeared quickly. The experiment was repeated at a later date

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Figure 8 Sessile volume and base diameter vs. time for aluminium on SiC at 1200°C Measurement of the contact angle, sessile drop volume, and base diameter has also been carried out at 1200°C, shown in Figure 13 and Figure 14. Unfortunately, the evaporation of aluminium is high and no stable contact angle was detected.

3.2 Al- Graphite System Figure 9 and Figure 10 give the wettability between aluminium and graphite at 1000°C as a function of time. The linear decreasing contact angle and the linear increasing base diameter indicate that further holding is required to attain the equilibrium contact angle. The almost constant sessile volume shown in Figure 10 is explained by low evaporation from the aluminium droplet because of its oxide layer. The wetting properties between aluminium and graphite at 1100°C are shown in Figure 11 and Figure 12. The similar three kinetic steps can be distinguished: the first step, where the contact angle decreases rapidly (in 30min) from the initial contact angle θ0~157° to θ ι ^ Ι 0 angle; the second step where the contact angle continues to decrease down to Θ2 but with a slower rate; and the third step corresponding to a nearly constant angle θ 2 ~ 65° at time 200min. The decreasing sessile volume indicates that de-oxidation of the oxide layer and evaporation from the aluminium drop take place. The stable base diameter and the sessile volume allow the measurement of the stable contact angle in the third stage.

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The initial sessile volumes, 5.6 μΐ at 1000°C, 5.7μ1 at 1100°C, 4.9ul at 1200°C, and 3.9 μΐ at 1300°C indicate that evaporation (from the initial heating to the time=0) is almost the same at 1000°C and 1100°C, and evaporation is greater at higher temperatures. The initial base diameters of 0.86mm at 1100°C, 2.2mm at 1200°C, and 3.1mm at 1300°C, as well as decreasing initial contact angles with temperature indicate that better wetting already occurs at the higher temperature before defined time zero.

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On the Al-SiC interface no AI4C3 layer was found on the sample cooled in air after 335 min at 1000°C (See Figure 16). However, Figure 17 shows the presence of a continuous layer of a reaction product, AI4C3 at the Al-graphite interface with a thickness of 130μιη. The graphite-A14C3 interface is rougher than the graphite substrate before the experiments and pores are present around some particles. Extra aluminium in the AI4C3 layer and the discrete A14C3 particles indicates that the reaction proceeds by dissolution of carbon into aluminium. The final contact angle is determined by the AI-AI4C3 system.

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Evaporation in both systems, Al-graphite and Al-SiC, depends only on reactions (2) and (4), independent of the interfacial reaction and the oxidation of the substrate. The same decreasing

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Discussion The three successive steps of wetting kinetics in Figures 3-14 are discussed successively.

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The sharp decrease of the contact angle from θ 0 to Q\ on both AlSiC and Al-graphite is similar to the contact angle curve observed previously for AI-AI2O3 for the same step in the same wetting furnace. This reduction is due to de-oxidation of the oxide layer according to reaction (2). This is effective when the outgoing flow of oxygen in A120 (g) is higher than the incoming flow of oxygen. The PjotaplO 3 Pa (10 5 mbar) in the furnace verifies that PMIO is lower than the equilibrium P A12O =3.7xl0" 3 Pa at 1000°C [22]. Thus the wetting behaviour of the first step is controlled by deoxidation of the oxide layer on the aluminium drop.

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There are three possible explanations concerning the second step from 0! to θ 2 in Al-SiC system: a) the dissolution of SiC into aluminium according to reaction (5), b) coverage of the interface by AI4C3 according to reaction (3), and c) de-oxidation of silica on the interface according to reactions (1) and (2).

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(5)

The maximum solubility of carbon in liquid aluminium, on the order of 30 ppm at temperatures close to 1000°C [23] is too low to support the dissolution mechanism in a). The kinetics of the reaction (3) is slow and leads to the limited formation of discrete particles of AI4C3 at the interface. Thus the spreading of de-oxidized aluminium is controlled by de-oxidation of SiC with limited amounts of discrete particles of A14C3. This is supported by the delayed equilibrium contact angle obtained in the Al-oxidized SiC system [8] at 660-900°C. 4Al(l)+3C(s)-^Al4C3

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(6)

In the second step from time ti to t2 in Al-graphite system, the spreading velocity is lower than the previous step and the base diameter is a linear function with respect to time. This second step does not exist in the non-reactive Α1-Α12θ3 system [24]. The interfacial reaction (6) has Gibbs energy of -136kJ/mol to 102kJ/mol [25] at temperatures of 660 to 1300°C. Figure 17 also supports that the second step in Al- graphite system is controlled by the formation of A14C3. The kinetics of reaction (3) forming A13C4 from SiC is slower than that of reaction (6) forming AI3C4 directly, so the stable

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To summarize the discussion, schematics of the wettability for the Al-SiC and Al-graphite systems are shown in Figure 18 and Figure 19. The contact angle initially changes with time and finally approaches an "equilibrium value". Both systems produce AI4C3 at the end and the interface reaction product promotes the wetting. However, extra silicon in aluminium could prevent the AI4C3 formation in Al-SiC system. Wetting in Filtration AI4C3 inclusions have relatively good wetting with aluminium. AI4C3 leads to fatigue cracking in casting and is not easily removed by settling because of its density of 2.36 g/cm3. The Al-SiC interface properties depend on filtration time. Priming before filtration leads to oxidation of the SiC filter. The silica changes into alumina by reaction (1). Alumina formed by the

reduction of the silica surface likely inhibits nucleation of AI4C3. An aluminium alloy with more than 10% silicon would prevent the AI4C3 formation according to reaction (3) (See Figure 20).

References 1.

2.

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Figure 20 The theoretical silicon level required to prevent AI4C3 formation [26] Conclusions

7.

The wettability of Al-SiC and Al-graphite with respect to time can be divided into three kinetic steps: (i) de-oxidation of AI2O3 layer, (ii) de-oxidation of silica for Al-SiC system with a small amount of AI4C3 (or AI4C3 formation for Algraphite system), and (iii) the stable contact angle. The equilibrium contact angles of the Al-SiC and Al-graphite system under 10"8 bar vacuum are shown in Table 3.

8.

9.

Table 3 The equilibrium contact angle Temperature [°C] 1000 1100 3. 4. 5.

The Al-SiC system [°] 65 57

[min] 150 125

10.

The Al-graphite system [min] [°] <97 >500 65 200

11. 12.

Both the SiC filter and AI4C3 inclusions have relatively good wetting with aluminium at temperatures greater than 1000°C. SiC filters are of special interest for the filtration of high Si aluminium alloys. At 700°C initially SiC and graphite are poorly wetted by aluminium. Times for wetting to take place may be hours. This will be discussed in future work.

13.

14.

Future Work 15.

The time dependent wetting properties of Al-graphite and SiC at 700°C will be extrapolated from the high temperature data. Pilot trials will be carried out in industry to investigate the effect of wettability on filtration efficiency.

16.

Acknowledgment 17.

The authors acknowledge the financial support from RIRA project by the research council of Norway. Thanks are also given to Don Doutre for a fruitful discussion and for providing Figure 20.

18.

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Laurent, V., et al., Wettability of Monocrystalline Alumina by Aluminium between Its Melting Point and 1273K. Acta Metallurgica et Materialia, 1988. 36(7): p. 1797-1803. Klinter, A.J., G. Mendoza-Suarez, and R. A.L.Drew, Wetting of pure aluminum and selected alloys on polycrystalline alumina and sapphire. Materials Science and Engineering A, 2008. M.Ksiazek, et al., Wetting and Bonding Strength in AUAI2O3 System. Materials Science and Engineering, 2002. A324: p. 162-167. Tang, K., Wettability of Al on Al203. SINTEF report, 2009: p. 6. Laurent, V., C. Rado, and N. Eustathopoulos, Wetting Kinetics and Bonding of Al and Al Alloys on a-SiC. Materials Science & Engineering A, 1996. 205: p. 1-8. Han, D.S., H. Jones, and H.V. Atkinson, The wettability of silicon carbide by liquid aluminium: the effect of free silicon in the carbide and of magnesium, silicon and copper alloy additions to the aluminium. Journal of Materials Science, 1993. 28(10): p. 2654-2658. Shimbo, M.N., M. and I. Okamoto, Wettability of silicon carbide by aluminum, copper and silver. Journal of Materials Science Letters, 1989. 8(6): p. 663-666. Laurent, V., D. Chβtain, and N. Eustathopoulos, Wettability of Si02 and oxidized SiC by aluminium. Materials Science & Engineering A: Structural Materials: Properties, Micro structure and Processing, 1991.135(1-2): p. 89-94. Laurent, V., D. Chβtain, and N. Eustathopoulos, Wettability of SiC by aluminium andAl-Si alloys Journal of Materials Science, 1987. 22(1): p. 244-250. CANDAN, E., Effect of Alloying Elements to Aluminium on the Wettability of Al/SiC system. Turkish J. Eng. Env. Sci., 2002. 26: p. 1-5. Guy Ervin, J., Oxidation Behavior of Silicon Carbide. Journal of the American Ceramic Society, 1958. 41(9): p. 347-352. Quanli, J., et al., Effect of particle size on oxidation of silicon carbide powders. Ceramics International, 2007. 33 p. 309-313. Moraes, E.E.S., M.L.A. Graηa, and C.A.A. Cairo, Study of Aluminium Alloys Wettability on SiC Preform. Congresso Brasileiro de Engenharia e Ciência dos Materials, 2006.15(19): p. 4217-4224. Li, J.G., Wetting of ceramic materials by liquid silicon, aluminium and metallic melts containing titanium and other reactive elements: a review. Ceramics international, 1994. 20: p. 391-412. Iseki, T., T. Kameda, and T. Maruyama, Interfacial reactions between SiC and aluminium during joinng. Journal of Materials Science, 1984.19: p. 1692-1698. Foister, S.A.M., M.W. Johnston, and J.A. Little, The interaction of liquid aluminium with silicon carbide and nitride-based ceramics. Journal of Materials Science Letters, 1993.12(4): p. 209-211. Bornand, J.-D. and K. Buxmann, DUFI: A Concept of Metal Filtration. Light Metals, 1985: p. 1249-1260. Etter, T., et al., Aluminium carbide formation in interpenetrating graphite/aluminium composites. Materials Science and Engineering A, 2007. 448(1-6).

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20. 21.

22.

23.

Kalogeropoulou, S., C. Rado, and N. Eustathopoulos, Mechanisms of reactive wetting: the wetting to nonwetting case. Scripta Materialia, 1999. 41(7): p. 723728. Zhou, X.B. and J.T.M.D. Hosson, Reactive wetting of liquid metals on ceramic substrates. Acta Materialia, 1996. 44(2): p. 421-426. Bao, S., et al., Wetting of Pure Aluminium on Filter Materials Graphite, AIF3 and A1203. Light Metals: 138th TMS Annual Meeting and Exhibition, Moscone West Convention Center, Feb., San Francisco, CA, USA, 2009: p. 767-773. Standards, N.B.o., JANAF Termochemical Tables, 2nd edition. 1971.37.

24. 25. 26.

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Yokokawa, H., et al., Phase relations associated with the aluminium blast furnace. Aluminium oxycarbide metals and Al-C-X (X=Fe, Si) liquid alloys. Metallurgical Transactions B, 1987.18B: p. 433-444. Bao, S., et al., Wettability of Aluminium on Alumina. To be published. Isaikin, A.S., et al., Compatibility of Carbon Filaments with A Carbide Coating and an Aluminium Matrix. Sci. Heat Treatment, 1980. 22: p. 815-817. Lloyd, D.J., The solidification microstructure of paniculate reinforced aluminium/SiC composites. Composites Science and Technology, 1989. 35: p. 159179.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Study of Microporosity Formation under Different Pouring Conditions in A356 Aluminum Alloy Castings Lu Yao, Steve Cockcroft, Daan Maijer, Jindong Zhu and Carl Reilly Department of materials engineering, University of British Columbia, 309-6350 Stores Road, Vancouver, B.C. Canada V6T1Z4 Keywords: Microporosity, Nucleation site distribution, Oxide film, X-ray micro-tomography castings. The focus is on oxide films formed during the pouring process, as studies have shown that pouring generates the largest free surface area during a gravity casting process [12]. The approach involves estimating the pore nucleation population by fitting a pore size distribution, which is predicted by a comprehensive mathematical model, to data obtained from a series of experimental castings in which the pouring conditions were varied. The results are discussed in terms of nucleation number density and nucleation potency (the latter defined as the supersaturation necessary for pore nucleation). In addition to the role of oxide film inclusions, the effect of hydrogen content and Sr modification on microporosity nucleation is also discussed.

Abstract In this work, the formation of microporosity has been examined under different casting conditions aimed at manipulating the tendency to form and entrain oxide films in small directionally cast A356 samples. Porous disc filtration analysis (PoDFA) was used to assess the melt cleanliness and identify the inclusions in the castings. The porosity volume fraction and size distribution were measured using X-ray micro-tomography (XMT) analysis. By fitting a pore formation model to the experimental results, an estimate of the pore nucleation population has been made. The results from the model predictions indicate that increasing the tendency to form and entrain oxide films not only increases the number of nucleation sites but also reduces the supersaturation necessary for pore nucleation in A356 castings.

Experimental method Casting procedures

Introduction

A series of experimental castings of A356 alloy were directionally solidified in a tapered, cylindrical refractory ceramic mold against a water-cooled copper chill under different pouring conditions conducive to varying the entrained oxide film content. To quantify the cooling conditions within the casting the evolution of temperature with time was measured with three K-type thermocouples positioned within the mold cavity at 10, 30 and 50 mm from the chill, as shown in Figure 1. To determine the hydrogen content, Ransley samples were cast, machined and analyzed using the LECO vacuum fusion technique, which is accurate to within 0.01 ppm. In order to vary the tendency to form and entrain oxide films within the casting, three different pouring methods were adopted: 1) normal pouring (intermediate tendency), Figure 1 (a); 2) Ar-shielded pouring (low tendency), Figure 1 (b); and 3) high-surface area pouring (high tendency), Figure 1 (c). In the first and second methods, the pour surface area was held constant, but shielding with argon in the second method decreased the ambient oxygen content. In the third method, the pour surface area was increased relative to the base or

The limited understanding of the nucleation mechanism is one of the main challenges still to be overcome in simulating microporosity formation in aluminum alloy castings. Previous studies have shown that porosity formation can be significantly influenced by inclusions, especially oxide films in the melt [1-6]. Conventional theories suggest that solid inclusions in the melt provide heterogeneous nucleation sites, which reduce the nucleation energy barrier required for pore nucleation [3, 7, 8]. In this theory, pore nucleation is promoted by a smaller contact angle between the liquid/pore interface and the solid substrate. This theory has been questioned by Campbell, who pointed out heterogeneous nucleation seems highly improbable due to the restriction on the minimum contact angle attainable, which is observed as close to 20 degrees in practice [9]. He proposed a nucleation-free mechanism for porosity formation based on a continuous solid oxide film present on the free surface of the liquid in aluminum alloys [9-11]. Due to free-surface turbulence (such as occurs during pouring), the oxide film becomes entrained in the melt and are often broken and folded as double-sided films. Within these oxide films it is proposed that air is entrapped and becomes a gas cavity. These gas cavities can be considered as preexisting pores in the melt. Pore growth can occur by hydrogen diffusion to these gas cavities and/or the unfurling of the bi-films due to a drop in the surrounding pressure associated with high fraction solid restricted feeding. This theory has gained considerable attention in the field of aluminum alloy casting research; however, no direct experimental evidence has been provided because of the nanometer scale thickness and transparent nature of the bi-films. The identification of the pore/oxide film interaction mechanism is still a subject of active research.

Figure 1. Schematic of the experiment setup for (a) normal pouring (NP), (b) Ar shielded pouring (ArP) and (c) high surface area pouring (HSAP) conditions.

The main objective of this work is to study the effect of oxide film content on microporosity formation in A356 aluminum alloy

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Table I. Summary of pouring conditions ^^^^^^Sto/w/?/e Label CAPTN-NP Casting CondïtïorT—-^^__ Metal Source Alcoa Ar Shielding No Pouring Surface Area Normal Hydrogen Content (ppm) 0.30±0.01

CAPTIN-NPL

CAPTIN-ArP

UBC-NP

UBC-HSAP

Alcoa No Normal 0.17±0.01

Alcoa Yes Normal 0.35±0.01

Rio Tinto Alcan No Normal 0.28±0.01

Rio Tinto Alcan No High 0.30±0.01

Table II. Chemical composition analysis for A356 aluminum alloy samples cast at CAPTIN and UBC (weight percentage) CAPTIN UBC

Si 6.89 7.49

Mg 0.34 0.38

Fe 0.15 0.10

Ti 0.13 0.10

Sr 0.0092 0.0030

Al Bal. Bal.

concentrated impurities and particulates from the melt are accumulated above the filter in the so-called 'cake'. The PoDFA sample included the residual metal, the inclusion cake and the filter, which was analyzed using standard metallographic procedures by Rio Tinto Alcan Limited, Arvida Research and Development Centre (ARDC), Jonquiere, Quebec, Canada. The types of inclusions present were identified under an optical microscope by comparison with documented images of the inclusions based on their morphological and chemical characteristics. The total area of each type of inclusions was measured using the grid method and divided by the total mass of the filtered melt to calculate the inclusion content. The results were given in the unit of mm2/kg.

normal case by pouring the metal into a course 20 ppi reticulated foam filter installed above the mould (unsaturated filtration). The filter acted to separate the liquid metal into several pour streams thereby increasing the surface area in contact with ambient air. Effect of Ar shielding. This component of the experimental work was undertaken at Canadian Auto Parts Toyota Inc. (CAPTIN) at their wheel casting facility located in Delta, BC, Canada. The A356 alloy was comprised of ingot from Alcoa and some revert from the plant. The metal was melted in a commercial reverberatory furnace, poured into a ladle and Sr was added for modification. The metal was not degassed. A boron-nitride coated, steel, pouring cup was used to transfer the metal from the ladle and pour it into the mould. The melt temperature was -710 °C at the time it was poured.

X-ray micro-tomography (XMT) analysis The cast porosity volume fraction and size distribution were measured using 3D X-ray micro-tomography (XMT). A 3mm diameter cylinder of material was machined from the centre of each casting. Three scans were performed on each cylinder at 10, 30 and 50mm from the chill, corresponding to where the thermocouples were positioned. Each scan was comprised of 500 slices with a diameter of 3 mm and a volume of 13.1 mm3. The three-dimensional images of the pores were generated from these 500 slices (2-D X-ray images) using the software ImageJ and AMIRA. The scans were conducted using the synchrotron X-ray microscope located in the TOMCAT lab at the Swiss Light Source (SLS), Paul Scherrer-Institute, Villigen, Switzerland. The tomography data were acquired by applying a beam energy of 27 keV and a resolution of 3.7 urn.

In addition to assessing the effect of Ar shielding at CAPTIN, there was one additional test completed with no Ar shielding in which the metal contained a lower hydrogen concentration. This difference arose due to differences in ambient humidity or holding time in the melt furnace, not degassing. Effect of Pour Surface Area. The second casting campaign was undertaken in the Advanced Materials and Process Engineering Laboratory (AMPEL) at University of British Columbia. Ingots sourced from Rio Tinto Alcan were melted in a graphite crucible using a lab-scale resistance furnace and held at 800°C for approximately 3hrs. In preparation for casting, 0.5 g of Al-15 wt% Sr alloy was added, the melt was held for a further 10 minutes for homogenization and the surface was skimmed to remove the accumulated dross. The melt was then finally poured at -750 °C. A higher pouring temperature was used to increase fluidity of the melt and enable it to pass through the filter.

Numerical model The readers are referred to reference [15] for details of the pore nucleation and growth model developed previously to predict the microporosity formation in A356 castings. Here only some key concepts in the model are presented for the sake of brevity.

The casting conditions and labels assigned to each of the cases for the purpose of discussion are summarized in Table I. The chemical compositions of the cast samples from the two sets of experiments are shown in Table II.

The model accounts for pore nucleation assuming a relationship between nucleation site distribution and the supersaturation of the liquid. The nucleation site distribution is defined as the change of the number of nucleation sites due to hydrogen supersaturation. The numerical form of the microporosity nucleation site distribution is a Gaussian function of hydrogen supersaturation, ss, as presented in Equation (1):

Inclusion analysis Concurrent with the casting operation, porous disc filtration analysis (PoDFA) samples were taken under each casting condition to measure the type and content of inclusions. The PoDFA was undertaken following the instructions [14] established by BOMEM Inc., Quebec, Canada. The procedure entailed pouring 1 kg of melt into a crucible where it was forced to flow through a fine filter under pressure. During filtration,

dNm dss

784

4ΐπó

exp(-

(ss-ss0) 2ó2

(1)

Table III. ResultsfromPoDFA analysis showing major inclusions found in samples cast at CAPTIN and UBC ^ ^ ^ ^ ^ ^ Sample Inclusion 7 ^ e ^^ Α1403(<3μπι) Α1Λ(>3μπι) MgAl 2 0 4 (TiV)B2 | TiB2/TiC

CAPTN-NP (mm2/kg) 0.47 0.86

-

CAPTIN-NPL (mm2/kg) 0.18 0.72

CAPTIN-ArP (mm2/kg) 0.49 0.49

-

-

where A, ss0 and σ are parameters describing pore nucleation kinetics. ss0 (mol/m3) is the mean hydrogen super-saturation required for pore nucleation, which may be equated with the potency of the pore nucleation population in the melt, σ (mol/m3) is the width of the nucleation site distribution, which may be equated with the variability in potency of the population. In this work, A is determined by the number of pores near the chill bottom measured from XMT experimental data. This assumes that the supersaturation achieved near the chill exceeds ss0 +3ó - i.e. full integration of Equation (1) gives A. The highest supersaturations are achieved at the chill because of the high cooling rates present there.

UBC-HSAP (mm2/kg) 0.03

0.22 0.06 1.46

Trace Trace 0.79

-

-

analysis, an analysis of the filter cake was undertaken using an SEM. The SEM analysis was able to identify structures that appeared to be film like. Furthermore and more relevant, EDX mapping of pores found in the PoDFA filter residue using the SEM have showed the presence of oxygen associated with pores. An example image is shown in Figure 2. This observation appears consistent with microporosity nucleation on oxide film inclusions. Unfortunately, it proved impossible to quantify the area fraction or number of oxide films present in the various samples of filter cake using the SEM.

Pore growth is simulated as a hydrogen diffusion controlled process. The mass transfer rate of hydrogen is evaluated according to Equation (2). dmh (2) ^ = Dr47rt- l t lp)/l diff dt where mH (mol) is the number of moles of hydrogen transported to the pore, Z), (m2/s) is diffusivity of hydrogen in the liquid A356 alloy, rp(m) is the radius of the growing pore, ldiff (m) is the thickness of the diffusion boundary layer, approximated as the pore radius, and ö is factor used to account for the reduction in pore surface area in contact with liquid metal as the solid fraction increases during solidification. The impingement factor is assumed to be equal to (l-fs)m, where fs is the solid fraction and m is a parameter to be determined byfittingwith experimental data.

Figure 2. EDX mapping analysis that shows presence of oxygen associated with pores. Microporosity analysis from XMT measurement The XMT data from the samples cast under the different conditions was analyzed and the volume fraction, number density and size distribution of the pores as a function of distance from the chill was determined. The data is presented as: 1) plots of volume fraction and number density vs. distance from the chill see Figures 3, 5 and 7; 2) a 2D projection of the 3D rendering of pores at 10, 30 and 50mm from the chill - see Figures 4, 6 and 8; and 3) a size distribution plot (number density vs. pore radius) see Figures 4, 6 and 8. In the line graphs, the experimental results appear as the symbols, where as in the size distribution plot (bar chart) the experimental data appears as the shaded bars. The error bars for volume fraction indicate the potential deviation in the measurement related to choice of gray scale threshold used to distinguish the pores from matrix in the XMT analysis [15].

Experimental results Inclusion analysis based on PoDFA measurement Originally, the PoDFA inclusion analysis was undertaken in an attempt to quantify the oxide film content in the melt. Table III shows the types and quantity (expressed in units of specific area) of the main inclusions identified in the analysis for both the Alcoa and Rio Tinto Alcan alloys under different pouring conditions at CAPTIN and UBC. In addition to the inclusions shown in Table III, traces of chlorides, MgO, graphite and refractory materials were also observed.

Effect of hydrogen content and Cooling Rate. The effect of hydrogen content on porosity is shown in Figures 3 and 4, in which the CAPTIN-NP and CAPTIN-NPL samples are compared. The CAPTIN-NP sample contained 0.30±0.01 ppm hydrogen and the CAPTIN-NPL sample contained 0.17±0.01 ppm hydrogen. As expected the higher hydrogen content yields a higher volume fraction of pores, which is a consequence of the higher pore number density in Figure 3(a) and trend to larger pore sizes in Figure 4. The high degree of sensitivity to hydrogen content is well established [4,5]. In addition, the data also shows a strong dependence on cooling rate with the trend to fewer (Figure 3(a)) larger pores (Figure 4) with increasing distance from the chill. The trend to fewer pores with increasing distance from the chill

There are a couple of observations that can be made from the results of the PoDFA analysis. Firstly, quantification of oxide film content proved to be difficult as only a limited number of the oxide films were detected. Secondly, based on the inclusions that could be identified, the coarse filter used to produce the UBCHSAP sample was able to "clean the melt", as evidenced by a reduction in the surface area of inclusions in thefilterresidue. To follow up on the lack of oxide films observed in the PoDFA™1 1

UBC-NP (mm2/kg) 0.26

PoDFA is a trademark of Rio Tinto Alcan Inc.

785

would have normally resulted in an increase volume fraction and trend to larger pores as previously discussed.

(decreasing cooling rate) relates to the nucleation process and points to the fact that there is some mechanism by which the number of nuclei activated is sensitive to cooling rate. The mechanism is discussed in more detail with the aid of the numerical analysis.

A

CAPTIN-NP, measured CAPTIN-ArP, measured CAPTIN-NP, predicted CAPTIN-ArP, predicted

•35

I

CAPTIN-NP, measured * CAPTIN-NPL, measured CAPTIN-NP, predicted CAPTIN-NPL, predicted

E —|

1

1

.

r—

10 30 50 distance from chill (mm)

! 2.0-1 E

(a) (b) Figure 5.Comparison between measured (symbol) and predicted (dashed line) pore volumefractionand number density in samples resulted from normal pouring (CAPTIN-NP) and Ar shielded pouring (CAPTIN-ArP) conditions.

-i > 1 ■ r 10 30 50 distance from chill (mm)

—| ■ 1 « r— 10 30 50 distance from chill (mm)

10 30 50 distance from chill (mm)

(a) (b) Figure 3.Comparison between measured (symbol) and predicted (dashed line) pore volume fraction and number density in samples resulted from medium and high hydrogen levels.

40

80

120

160

200

240

pore radius (/im) 80

(a)

120

160

200

240

280

240

280

(a)

pore radius (/-im)

CAPTIN-NP, [He)=0.30 ppm

■ ■ ■ f ^ i ^ F ^ ^ l EU

0

40

80

120

160

200

80 120 160 200 pore radius (μηι)

pore radius (μηι)

(b) Figure 4. Comparison between measured (solid columns) and predicted (transparent columns) pore size distribution at different distances from the chill in samples resulted from (a) medium and (b) high hydrogen levels.

(b) Figure 6.Comparison between measured(gray columns) and predicted(transparent columns) pore size distribution at different distances from the chill in samples resulted from (a) normal pouring (CAPTIN-NP) and (b)Ar shielded pouring (CAPTINArP) conditions.

Effect of Ar shielded pouring. Figures 5 and 6 compare the volume fraction and pore size distributions resulting from the normal and Ar shielded pouring (CAPTIN-NP and CAPTIN-ArP). From the various plots it is clear that the volume fraction, pore number and pore size all decrease under Ar shielded pouring. As there is higher hydrogen content in the CAPTIN-ArP sample (0.35 ±0.01 ppm) compared with the CAPTIN-NP sample (0.30±0.01ppm), the decreased in volumefractionand tendency to smaller pore sizes under Ar shielded pouring must relate to difference in the nucleation population - i.e. a shift to higher supersaturations and higher fraction solid. In the absence of a change in the nucleation processes, the higher hydrogen content

Effect of high-surface area pouring. The results from XMT measurement of the pore volume fraction, number density and size distribution under normal and high-surface area pouring conditions are shown in Figure 7 and 8. The results indicate a significant increase in the volume fraction and pore size in the high surface area casting (UBC-HSAP) relative to the normal pouring condition (UBC-NP). Interestingly, there is no significant difference in the number density of pores. It is also important to point out that the UBC-HSAP casting contained slightly more hydrogen than the UBC-NP casting (0.30 vs 0.28ppm) and thus it is difficult to conclude that the difference observed is due to a difference in nucleation population resulting from a change in tendency to form oxide films or is due to the difference in

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accurately reproduce the variation in volume fraction and number density of pores with distance from the chill quantitatively at most of the comparison points. Furthermore, the model is also able to reproduce the pore size distribution observed in the sample to within a reasonable degree of accuracy.

hydrogen content. The final conclusion will need to be deferred until the model-based analysis of the data has been completed. 3.0η 7 0-

UBC-NP, measured A UBC-HSAP, measured UBC-NP, predicted UBC-HSAP, predicted

2

§ 1.0-

•fs

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1'

1 0.5-

1.0-

0.0-

1.5ç

1 1 1 1 1 10 30 50 distance from chill (mm)

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.

1

J

Beginning first with the impingement factor, as the impingement factor is linked to the solidification morphology, it should only be dependent on cooling rate and alloy composition, not the pouring condition. Figure 9 shows the impingement factor m as a function of cooling rate determined by fitting with the experimental measurements. The variation of m with cooling rate is consistent with a reduction in impingement with decreasing cooling rate (increasing grain size). It can also be seen that a smaller m was applied for the CAPTIN samples, corresponding to a lower impingement between pores and solid grains at a given fraction solid. This can be explained by comparing the microstructures of the CAPTIN and UBC metal, shown in Figure 10. As can seen the CAPTIN metal (Figure 10 (a)) solidifies into a structure that is more globular dendritic than the UBC material. As a result, one would expect a smaller pore/solid impingement in the CAPTIN structure at a given solid fraction. Within a given alloy group i.e. CAPTIN vs UBC metal - it was possible to use a single cooling rate dependent m, for evaluation of the various pouring conditions.

.-·' .

,

1

10 30 50 distance from chill (mm)

(a) (b) Figure 7.comparison between measured (symbol) and predicted (dashed line) pore volumefractionand number density in samples resulted from normal pouring (UBC-NP) and high-surface-area pouring (UBC-HSAP) conditions.

0

40

80

(a)

120

160

200

240

pore radius (/ / m ) β «2

3

UBC

1»

—

φ — ""

.S

1 ·— 0

· 1

CAPTIN

2

3

4

5

cooling rate (°C/s) 0

40

80

120

160

200

240

280

Figure 9.Impingement factor m as a function of distancefromthe chill derived byfittingthe model predictions with experimental measurement.

pore radius (/ira)

(b) Figure 8.Comparison between measured (gray columns) and predicted (transparent columns) pore size distribution at different distances from the chill in samples resulted from (a) normal pouring (UBC-NP) and (b) high-surface-area pouring (UBCNSAP) conditions.

Figure 11 shows the model-predicted nucleation site distributions that yielded the best fit to the measured porosity data for the different pouring conditions. Note the same nucleation distribution was used for both the CAPTIN-NP and CAPTIN NPL samples (consistent with the fact that the same pouring procedure was used for both). The results for the different pouring scenarios

Mathematical Model Analysis To help better understand the role of oxidation tendency on nucleation kinetics, the castings have been analyzed using a comprehensive pore nucleation and growth model. The only 'fitting parameters' used in the model are those related to the description of the pore nucleation kinetics and pore/solid impingement; namely ss0, ó and m. A trial and error approach was applied for each casting to determine thefittingparameters that gave the bestfitfor the entire casting. Comparisons between the model predictions and XMT measurements can be found in Figures 3 through 8. In the line graphs, Figures 3, 5 and 7, the model predictions appear as the lines. In the distribution plots (bar charts) the model predictions appear as the un-shaded bars. As can be seen, the model is able to

Figure lO.Optical micrograph of as-cast microstructures in samples resulted from different metal sources at (a) CAPTIN and (b) UBC.

787

Another observation that can be made from Figure 11 is that there is a clear difference between the nucleation population present in the samples cast at UBC and those cast at CAPTIN. The exact reason for this difference is unclear. Based on the PoDFA analysis the UBC metal is the cleaner of the two, which is consistent with the fact that it was produced from virgin ingot (the CAPTIN material contained some revert). A closer look at the difference in metal composition (see Table II) reveals a significant difference in the Sr levels - 0.0092 for CAPTIN material and 0.0030 for the UBC material. Work is currently underway, including a review of the literature, to better understand what potential role Sr, a strong surface tension modifier, will have on nucleation of hydrogen.

indicate both a difference in the potency of the nucleation population, as indicated by a difference in ss0, and a difference in the number of available nucleation sites, as indicated by differences in A dependent on the casting conditions - refer to Equation (1). The effect of cooling rate on pour nucleation can be understood by considering the variation in supersaturation with cooling rate. Analysis with the model has revealed that the highest supersaturations are reached near the chill, hence the largest number of nuclei become activated at the chill. As previously described, this behavior was used to reduce the number of fitting parameters in the model - i.e. the value of A was determined from the pore number density measured at the chill (10mm location). This approach appears to have achieved good results overall, but it should be pointed out that it will work only if the initial hydrogen content is sufficiently high. For example, in the case of the CAPTIN-NPL sample, the analysis with the model revealed that at the lower initial hydrogen content of 0.17ppm, not all of the pores were activated at the chill, hence it was possible to achieve a good fit consistent with the measurements with a single nucleation population.

Summary and Conclusions The effect of varying the tendency to form and entrain oxide films on microporosity nucleation kinetics has been investigated in this work by altering the method in which small directional solidified castings were poured. Changes to the pouring method entailed Ar shielding and increasing the pour stream surface area relative to a base case. The effect of these changes was quantified by measuring the porosity volume fraction, pore number density and pore size distribution as a function of distance from the chill via 3D XMT analysis. By fitting the predictions of a comprehensive pore numerical model to the XMT measurements, differences in the nucleation kinetics have been estimated for the different pouring conditions. The methodology involved adjusting the parameters used to define the nucleation behavior, by trial and error, until good agreement was obtained with the results of the experiments. The results tend to support the hypotheses that oxide films provide good nucleation sites for pore formation in-so-far as increasing and decreasing the tendency to form and entrain oxides in association with the pouring process resulted in a consistent increase and decrease, respectively, in the porosity volume fraction. It was also observed that changes in pour stream oxidation tendency result in both changes in the number of pore nucleation sites and potency (supersaturation) of the sites. The results do not support or refute the bi-film mechanism proposed by Campbell [9-11]. They simply confirm that altering the tendency to form and entrain oxide films alters the tendency to form pores in a manner consistent with oxide films playing a role in the nucleation of hydrogen based pores.

The effect of Ar shielding on the nucleation site distribution can be seen by comparing the results for samples CAPTIN-NP and CAPTIN-ArP in Figure 11. The nucleation potency is reduced when Ar shielding is applied, as indicated by an increase of ss0 from 2.5 mol/m3 to 8.5 mol/m3. The total number of nucleation sites is also reduced for the Ar shielded pouring condition, indicated by a decrease of A from 3.5 per mm3 to 2.5 per mm3. The change is consistent with Ar shielding reducing the tendency to form and entrain oxide films. The effect of pour surface area is shown by comparing the nucleation curves for samples UBC-NP and UBC-HSAP in Figure 11. The nucleation potency is increased when the melt is poured with an increased surface area as indicated by a decrease of ss0 from 0.40 mol/m3 to 0.25 mol/m3. In addition there is also an increase in the total number of nucleation sites from 1.1 per mm3 to 1.5 per mm3. This result is obtained despite a decrease in the AI4C3, MgAl204, (TiV)B2 and TiB2/TiC content as shown in Table III. This is consistent with literature were it has been argued previously that carbides are not effective as nucleation sites for hydrogen porosity [9].

Acknowledgment The authors acknowledge Canadian Auto Parts Toyota Inc. (CAPTIN) and NSERC for the financial support, Rio Tinto Alcan for the in-kind support and the Swiss Light Source at Paul Scherrer Institute for the use of the TOMCAT beamline facilities.

:-0.25) 2 i dAC

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V2^-0.05

References

(AC-0.4)2 20.13 2 CAPTIN-NP/NPL dNmc _ 3.5 { (AC-2.5)2) dAC V27-0.8eXP[ 2-0.82 J

CAPTIN-ArP dNmc_ 2.5 dAC ^.0.7

1. D. Emadi, J. E. Gruzleski and M. Pekguleryuz, "Melt Oxidation Behavior and Inclusion Content in Unmodified and Sr-Modified A356 Alloy-Their Role in Pore Nucleation", AFS Transactions, 104 (1996), 763-768. 2. G. Laslaz and P. Laty, "Gas Porosity and Metal Cleanliness in Aluminum Casting Alloys", AFS Transactions, 99(1991), 83-90. 3. X. G. Chen and S. Engler, "Formation of Gas Porosity in Aluminum Alloys", AFS Transactions, 102(1994), 673-682. 4. K. Tynelius, J. F. Major and D. Apelian, "A Parametric Study of Microporosity in the A356 Casting Alloy System", AFS Transactions, 101(1993), 401-413.

2

\

( (AC-8.5) 2-0.72

0

0.3 0.6 3 6 9 12 Hydrogen supersaturation(mol/m3) Figure 11.Pore nucleation site distribution under various casting conditions derived byfittingthe model predictions with experimental measurements.

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5. D. Emadi and J. E. Gruzleski, "Effect of Castings and Melt Variables on Porosity in Directionally-Solidified Al-Si Alloys", AFS Transactions, 102(1994), 307-312. 6. L. Liu,A. M. Samuel and F. H. Samuel, "Influence of Oxides on Porosity Formation in Sr-treated Al-Si casting alloys", Journal ofMaterials Science, 38(2003), 1255-1267. 7. M. Massoud, Engineering Thermofluids: Thermodynamics, Fluid Mechanics, and Heat Transfer, (Springer Berlin Heidelberg, 2005),637-641. 8. S. F. Johes, G. M. Evans and K. P. Galvin, "Bubble Nucleation from Gas Cavities: a Review", Advances in Colloid and Interface Science, 80(1999), 27-50. 9. J. Campbell, Castings. (Butterworth Heinemann, 2 ed., 2003), 179-181. 10. D. Dispinar and J. Campbell, "Critical Assessment of Reduced Pressure Test. Part 1: Porosity Phenomena", International Journal of Cast Metals Research, 17(5) (2004), 280-286. 11. J. Campbell, "Entrainment Defects", Materials Science and Technology, 22(2) (2006), 127-145. 12. N. W. Lai, W. D. Griffiths and J. Campbell, "Modelling of the Potential for Oxide Film Entrainment in Light Metal Alloy Castings", Modeling of Casting, Welding and Advanced Solidification Process, (2003), 415-422. 13. O. Lashkari et al., "X-Ray Microtomographic Characterization of Porosity in Aluminum Alloy A356", Metallurgical and Materials Transaction, 40A (2009), 991-999. 14. PoDFA-f user's Manual, Revision 1-0, March 1999. ABB Bomen Inc. 15. L. Yao et al., "Modeling of Porosity Size Distribution in A356 Tapered Cylinder Castings", Modeling of Casting, Welding and Advanced Solidification Process, (2009), 385-392.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S CAST SHOP for ALUMINUM PRODUCTION

Grain Refinement, Alloying, Solidification and Casting SESSION CHAIRS

Peter Schumacher University of Leoben Leoben, Austria Arild Hβkonsen Hycast AS Sunndals0ra, Norway

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Hycast 1M Gas Cushion (GC) Billet Casting System 1

Idar Kjetil Steen \ Arild Hβkonsen1 Hycast AS, Industriveien 49, N-6600 Sunndalsora Norway Keywords: Casthouse, Billet Casting, Equipment, HES delivered. Total tonnage of metal casted with the system is more than 2 million tons per year. Figure 1 show a casting machine equipped with the Hycast Gas Cushion (GC) Billet Casting System.

Abstract The Hycast™ Gas Cushion Billet Casting System (Hycast ™ GC) for casting of extrusion billets has been used in Hydro Aluminium casthouses (both primary and remelt) for more than 20 years. The technology is based on a patented mould construction utilizing dual graphite rings for optimal distribution of oil and gas into the mould and increasing graphite ring lifetime. Annual production capacity in existing casthouses equipped with Hycast™ GC is more than 2 million tons per year. The diameter ranges from 152mm up to 405mm (6"-16"). The Hycast™ GC produces excellent and consistent metal quality in combination with high productivity, excellent metal recovery and low maintenance cost. The technology is now available for the aluminium industry also outside Hydro Aluminium. This paper describes the main technical achievements of the technology with focus on operational issues such as Health, Safety and Environment (HSE), metal recovery, metal quality and productivity. The focus on HES in Hydro Aluminium casthouses has been a driving force for developing a reliable casting technology with a high automation level. The casting control system has been designed to minimize manual operation during start and stop of cast, maintaining exact control of all important casting parameters including a proven and documented emergency stop philosophy.

Figure 1. Casting machine equipped with Hycast™ Gas Cushion Billet Casting System in a Hydro casthouse. Equipment design The Hycast™ GC technology is based on a mould construction with patented dual graphite rings (2) for supply of gas and oil to the mould wall. The upper ring distributes the casting oil into the mould while the lower graphite ring distributes the casting gas. Special seals and surface treatment are used between the two graphite rings to eliminate oil penetrating from the upper oil-ring into the lower gas-ring. By the use of this patented dual ring design, we avoid that casting oil disturbs the supply of casting gas to the mould. As a consequence the mould will produce a constant billet surface and sub-surface quality and the need for gas adjustment during or between casts is more or less eliminated. An assembled Hycast™ GC mould can be observed infigure2.

Introduction Hycast AS was established in 1990 as a spin off company from the Hydro Aluminium RTD Centre (Research and Technology Development) in Sunndalsora, Norway. During the years Hycast has developed and commercialized a range of production equipment for aluminium casthouses. Main focus area have been equipment for melt treatment and casting, but the product range today also covers launders for molten metal transfer, rod feeders for addition of grain refining, casting machines, automation systems and special maintenance equipment. Development of a new casting technology for extrusion ingots started in 1985 when ÂSV (now a part of Hydro) bought a technology license from Showa Denko. The patent from Showa Denko (1) covered a technology for hot-top casting of extrusion billets where gas (air) were introduced into the mould. The casting system produced billets with a reduced segregation zone and an improved surface quality compared to the traditional open spout and float system based on open moulds or the hot-top moulds used in the casthouses. The Hycast™ Gas Cushion Billet Casting System was introduced to Hydro Aluminium (and partners) casthouses in the beginning of 1990, and was included in Hycast product portfolio from 1991. Today the casting systems is used in both primary and remelt casthouses, and a total of more than 100 casting tables have been

Figure 2. Hycast™ GC Mould.

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When the mould is completely assembled with all its components, a pretest for oil and gas leakages is performed before the mould is mounted in the module based casting section. Seefigure3. Each casting section is equipped with a pneumatic dam that is automatically (or manually) opened during the start up sequence to ensure simultaneously filling of the moulds. Each casting section has a cabinet where the flow of gas and oil can be individually adjusted on each of the moulds.

Figure 5. ALSIM simulation of temperature fields in distribution launder during casting (5). Additionally the self centering starter blocks are assembled on a module based framework system. The starter block has a special design with a conical center part that effectively eliminates center cracks in the start up phase. (Figure 6)

Figure 3. Module based Casting Section with moulds. To build up the complete casting table, several of these casting sections are assembled on a common water-frame. Figure 4 shows the casting sections mounted on the water frame with metal distribution launder.

Figure 6: Hycast™ GC starter block Due to the high automation level, the Hycast™ GC casting system is easy to use and very operator friendly. In combination with Hycast casting control system, both starting and stopping the casting sequence can be fully automated. Manual work during casting is reduced to inspections. After casting, manual work on the equipment normally includes cleaning of moulds / refractory and inspection of mould condition.

Figure 4. Casting table with casting section mounted on waterframe including metal distribution launder. The distribution launder is used to control the flow of metal into each of the casting sections. This is particularly important on larger tables, where the design of the launder controls the metal flow into each of the casting sections, and thereby minimizes the metal temperature gradient between the moulds. The mathematical model Alsim (3) has been used for optimizing casting parameters and mould design. Alsim has proven to be a powerful tool to simulate critical parameters effect on centre and surface cracks during casting, (4) and also to optimize launder design to minimize temperature gradients during casting startup and steady state. Figure 5 shows an ALSIM simulation of the temperature field in the distribution launder during casting.

The open structure of starter block frame and base plate is designed to minimize metal buildup on equipment if a bleed out should occur. To minimize downtime between casts the moulds can be equipped with an optional internal cooling circuit (6). This allows equipment cooling times down to 5 minutes before the table can be raised and the casting pit can be stripped for logs. Cooling will continue with the equipment in upright position to prevent damages on seals and o-rings. Changing of casting table (from one diameter to another) may be a bottleneck in some casthouses. The Hycast™ GC is equipped with a quick exchange system allowing equipment to be changed within less than 20 minutes. The system is based on a combined lifting of casting sections and starter blockframesby the overhead

794

crane, and a "quick lock" system for fastening of both frames to the casting machine.

Operational experience Since the first implementation of the Hycast ™ GC Billet Casting System at Hydro Sunndal in 1989, the system is now implemented in 17 casthouses (both remelt and primary). The casting system and casting parameters have during the years been continuously optimized with regards to HSE, product quality, component lifetime, metal recovery etc., based on feedback from casthouses and input from Hydro RTD.

The diameter range for the equipment ranges from 152mm (6") and up to 405mm (16"), with standard diameter tolerances as defined in EN 486. Other tolerances can be delivered according to customer specifications. The very dense packing of the mould allows for optimal utilization of the furnace capacity for all billet diameters. The normal setup is to empty the casting furnace in one drop. In some casthouses large capacity increases has been obtained when the number of drops pr furnace charge has been reduced from 2 to 1 by changing the casting equipment to the very dense packed Hycast GC casting system.

Pit recovery is a very important performance indicator for all casthouses. Minimizing casting scrap (surface scrap and internal cracks) has been a focus area for all casthouses in the Hydro group during the recent years. Today the average pit recovery when using Hycast™ GC casting system is above 99,5%, and plugging of moulds due to bleed outs is more or less non existent. Most casthouses has in fact completely stopped the practice of plugging of moulds during casting due to safety concerns.

The highest number of moulds in one table delivered so far is a 152mm (6") casting table with 160 moulds.

The lifetime of refractory and graphite components in a casting system will have a large effect on the casthouse operational costs. It has therefore been an important issue to use component consumption data from the casthouses to optimize the design of all components included in the casting system.

HSE When designing equipment for aluminium casthouses it is important to include "state of the art "knowledge regarding HSE. The recommendations given in "AA Guidelines For Handling Molten Aluminium" (7) have been extensively used to optimize both equipment design and procedures to minimize the potential for incidents during casting.

Based on data from casthouses the average operational cost including all wear components and consumption items (casting oil, casting gas, refractory coating) is today less than 1 USD/ ton. Typical lifetime of hot-top rings is between 400-500 casts, while the graphite rings last for more than 1500 casts.

A main philosophy behind the equipment design has been to minimize operator's exposure to molten metal, and to minimize the risk for "bleed outs" during startup and stationary casting. Traditionally "casting start up" is the phase where most "incidents" take place. To minimize possible problems, it is important to minimize temperature gradients across the casting table, ensure even filling of all moulds and minimize the number of people close to the casting machine.

Operational experience from all users proves that maintenance routines are extremely important for maximizing pit recovery, minimizing operational cost and also for maintaining a consistent product quality. A well functioning mould maintenance workshop with a trained maintenance crew and equipped with the necessary tools is therefore one of the main keys to obtain consistant product quality and high casting line recovery.

Mathematical simulations of the metal flow in the distribution launder and in the basin elements have been the basis for the metal distribution system used today. Validation in casthouses has proven that the short filling channels and the use of automated dams in the metal inlet are very effective to secure minimal temperature gradients during start up and casting. Typical filling times for the casting sections are < 5 seconds and the maximum temperature gradients in the stationary casting phase are reduced to<10°C.

Special training programs have therefore been developed to cover equipment operational practice, maintenance practice and operational trouble shooting. Based on feedback from casthouses these training programs are regularly updated and revised by Hycast. Product Quality

Hycast has developed a standard safety philosophy for extrusion ingot casting which is included in the automation package. Main principles are: • Hard wiring of emergency functions • Block and bleed of the liquid metal • Fail safe function on all valves and critical components. • Maintain cylinder downward movement in case of power outage. • Emergency water from gravity tank. • Use of explosion protective coating on all steelframes below the casting moulds. • Standardized precast checklist a prerequisite before startup.

It is well known that "State of the art" hot-top casting systems utilizing gas has several product quality advantages such as: • Minimized surface segregation zone. • Shallow shell zone. • Even grain structure • Smooth surface appearance • Minimized oxide formation during casting • Minimized/ no pre-solidification. • Low surface porosity. In the early days of hot-top casting, poor casting parameter control could lead to casting defects due to solidification towards the hot-top (Bergmann zones and cold shuts). The casting speed for hot top casting is higher than open mould systems, and therefore the tendency of center cracking could also

795

be higher. With the automated systems for parameter control used at modern vertical DC casting machines in combination with high quality grain refiners and a maximum specification limit for Na contend in the liquid metal, these problems are now more or less none existing. Typical average inverse segregation zone thickness for billet diameter 203mm (8") is -lOOmicrons for a 6xxx-alloy. Due to the effect of the casting speed the inverse segregation zone will increase approximately proportionally with billet diameter. Conclusions The Hycast™ Gas Cushion Casting System has been in use in Hydro and Hydro partner casthouses for more than 20 years. The casting system has been continuously modified and improved based on input from the casthouses. The Hycast™ Casting System produces high quality extrusion billets with an excellent surface quality and with excellent extrudability. Annual production capacity on existing equipment is more than 2 million tons per year. The casting system is characterized by: • Patented dual graphite ring system for supply of oil and gas to the moulds. • Superior safety track records. • Very high pit utilization (Dense packing of moulds) • Excellent pit recovery • Excellent and consistent surface quality. • Fully automated casting sequence is available in combination with Hycast Casting control system. • The Hycast™ Gas Cushion Casting System is now available for casthouses outside the Hydro group. References 1. Patent US 4157728, "Process for direct chill casting of metals" 2. Patent US 5678623, "Casting Equipment" 3. D.Mortensen, " A mathematical Model of the heat and fluid flows in direct Chill Casting of aluminium sheet ingots and billets", Metallurgical and Materials Transactions B, Volume 3OB, (1999), pp. 119-133 4. S. Benum, D. Mortensen, H. Fjaer, H. G. 0verlie and O. Reiso, "On the mechanism of surface cracking in DC cast 7xxx and 6XXX extrusion ingot alloy". In: W.Schneider, editor. Light Metals 2002: The Minerals, Metals and Materials Society, Warrendale, PA, pp967-973 5. E. Sorheim, Institute for Energy Technology, Private communications. 6. Patent. US 6513574, "Water cooling system for continuous casting equipment" 7. The Aluminium Association, "Guidelines for Handling Molten Aluminum", Third Edition, July 2002.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

STUDIES OF FLUID FLOW AND MENISCUS BEHAVIOR DURING HORIZONTAL SINGLE BELT CASTING (HSBC) OF THIN METALLIC STRIPS Donghui Li, Jaspreet Gill, Mihaiela Isac, Roderick Guthrie McGill Metals Processing Centre, 3610 University Street, Montreal, Quebec, H3A 2B2, Canada Keywords: horizontal single belt casting (HSBC), meniscus behavior, strip casting, mathematical modeling Abstract

In order to understand the meniscus behavior and fluid flow mechanisms in HSBC processing, mathematical modeling using COMSOL software was developed, so as to predict the movement of the free air/melt interface and the flow patterns in the near meniscus region. The mathematical modeling of meniscus behavior and fluid flow was later verified by a physical water model of HSBC casting, and was subsequently also validated through HSBC simulator tests using aluminum alloys.

Horizontal Single Belt Casting of strips is a green strip casting technology, potentially capable of replacing current DC and slab caster operations. As-cast strip bottom surface quality is a key factor for the near-net shape casting operations. The meniscus behavior at the triple point of gas, substrate, and liquid metal, where the melt first touches the moving belt is important to surface quality, as is the way in which the melt behaves while in subsequent contact with the chill substrate. In this paper, meniscus behavior and fluid flow mechanisms were analyzed and predicted through mathematical modeling, using COMSOL software. It was found that the backwall gap, combined with melt vertical inlet velocity through the tundish nozzle slot, and the belt speed, dominated meniscus behavior. The backwall gap must be pre-set to be less than the critical gap size in order to prevent melt "back flows" and leaking accidents. The mathematical modeling of meniscus behavior and fluid flow was further supported by physical water modeling, and was validated through HSBC simulator tests using aluminum alloys. Introduction Horizontal Single Belt Casting (HSBC) of strips is a green, nearnet shape, strip casting technology, potentially capable of replacing current D.C. aluminum caster, as well as steel slab caster. This is because of its promising productivity, low energy consumption, and low capital and operating costs [1-3]. The ascast strip bottom surface quality is a key factor before the in-line rolling process. The as-cast strip products must have a satisfactory quality of strip surface, because the high surface area to volume ratio makes it uneconomical for "scalping" the product surface in order to remove surface defects, as carried out in conventional DC casting and rolling operations. Figure 1 provides a schematic view of HSBC process. A small gap must be kept between the backwall of the tundish refractory lining and the moving belt during the strip casting process for free running of the belt, and to avoid possible contamination, or scratching of the belt's upper surface by the tundish refractory. Therefore, a stable melt/air free surface, or meniscus, will be formed during casting. The melt flows down through a simple nozzle slot and changes its flow direction to the horizontal, as it flows and freezes on the water cooled belt. The stable air/melt interface is called the forward meniscus. Near, or at the meniscus, the surface layer of liquid metal meets the air and then the moving chill substrate. It then begins to form the initial solidified shell on the melt's bottom in "contact" with the cooling belt. This initial solidification is important in determining the final strip's bottom surface quality. The oscillation of the meniscus, entrapment of air bubbles at the triple point, or oxidation of the melt at meniscus, would generate strip surface defects, such as air pockets, micro-crack, and a non-uniform microstructure of the strip [4, 5].

Figure 1. Schematic view of the meniscus in the HSBC process. Mathematical model The melt flow and air/melt meniscus behavior in HSBC processing were studied by solving the continuity equation and the Navier-Stokes equations, coupled with the "phase field" method. Since the metal was still in its liquid state near the meniscus region because of its casting superheat, the heat transfer and solidification of the melt were ignored for this simplified model. Also, aluminum alloys cast in ambient air were considered in this paper. Continuity and incompressible Navier-Stokes equations V-u = 0

(1)

P^7 + P(u · V)u = V[-pI + μ(νιι + (Vu)T)] + pg + θνφ (2) The last term on the right hand side of equation (2) is a body force due to melt surface tension, G is the chemical potential (J/m2), and φ is the dimensionless phase field variable, g is the gravity vector. Phase Field Method In order to model the flow of two immiscible fluids and to trace the air/aluminum melt interface, where the exact position of the interface is of interest, the phase field method was adopted. This entails solving the equation of the phase field variable φ [6]. The

797

equations governing the interface dynamics of a two-phase flow can be described by the Cahn-Hilliard equations, shown below, as equations (3) and (4). ^ + ê.νφ = ν . £ ν ø

(3)

Ø = -V · ε 2 νφ + (φ 2 - 1)φ

(4)

substrate laterally under the stationary tundish. Once the substrate starts to move, the tundish nozzle slot opens automatically and melt drops down through the nozzle slot and deposits on the substrate. A high speed video camera was placed to the side, behind the tundish, as shown in Figure 2. The camera lens was adjusted to be parallel to the copper substrate surface, so that we could observe the movements of the melt meniscus through the backwall gap. The camera could record a movie at 300framesper second. This was fast enough to capture the instantaneous air/melt interface oscillation during the strip casting process.

Here φ is the dimensionless phase field variable, varying from 1 to -1, while the variable Ø is referred as the phase field help variable. The Cahn-Hilliard equations force φ to take a value of 1 or -1, except in a very thin region at thefluid-fluidinterface. As such, the position of the air/melt interface would be described by the iso-value curve of φ =0. The volume fraction of the two immiscible fluids can be described by (1 + φ )/2 and (1 - φ )/2. The term ë in equation (3) represents the mixing energy density, while ε is a capillary width that scales with the thickness of the interface. The variable γ is the mobility, which determines the time scale of the Cahn-Hilliard diffusion, and must be large enough to retain a constant interfacial thickness, but small enough so that the convective terms are not overly damped. In COMSOL Multiphysics, the mobility was determined by a mobility tuning parameter χ, which was a function of the interface thickness, shown in equation (5). In the present paper, the mobility tuning parameter χ was set as 1. γ = ÷å2

(5)

The two parameters of ë and ε are related to the surface tension coefficient via equation (6).

Figure 2. HSBC simulator with high speed video camera system.

An aluminum melt was considered in the present work, with a surface tension of 0.914N/m, and a melt density of 2,380 Kg/m3

m.

Numerical method The multiphase flow problem, involving a moving air/moving melt interface in HSBC processing, was simplified to a 2dimensional fluid flow problem, and was solved by COMSOL Multiphysics software. The transient phase field variable was initialized first for setting the phase field variable φ which varied smoothly across the initial interface, and was 1 or -1 anywhere else. Then the momentum equations and continuity equation were solved, coupled with the phase field variable equations. The value of the phase field variable was used to compute the air/melt interface curvature and the surface tension force for the source term in the Navier-Stokes equation. Note that triangular elements were used in meshing the computational domain. The maximum mesh size was less than 0.3 mm in the entire region involving the interface between the air and the melt.

Figure 3. Water modeling system for the HSBC process. A full scale water model of the simulator system was also built in order to model the flow patterns in the HSBC metal delivery system, as shown in Figure 3. An endless rubberized belt, maintained under tension, circulated around a series of rolls continuously. Plexiglas was used to fabricate a tundish with a rectangular slot nozzle (82mm wide x 2.5mm thick) set in the bottom of the tundish. Gravity dominates the fluid flow through the HSBC tundish nozzle system, while inertial forces control the fluid flow pattern developing along the moving belt. As such, the same Froude number and same Reynolds number must be satisfied in order to use the water model to simulate the flow of

Experiments In order to verify the results of the mathematical modeling, experiments were carried out on an HSBC simulator and also on a water model of the HSBC metal delivery system. The HSBC simulator comprised a moving chill substrate, a stationary refractory-lined tundish with a rectangular slot nozzle for melt delivery, and a compression spring system to propel the

798

aluminum in the HSBC simulator. Since the geometry of the water modeling system was 1:1 to the HSBC simulator, the Froude criterion was automatically met. The dynamic viscosity of aluminum is 0.001338 Pas, and its density is 2,380Kg/m3. Water with a density of l,000Kg/m3, and a dynamic viscosity at 48°C of 5.6x1 O^Pa-s, provides a perfect match of Reynolds Numbers between the two systems. As such, the flow patterns generated within the molten aluminum in the HSBC simulator could be studied by the water model system. Discussion Predicted flow patterns in the melt and meniscus behavior Figure 4 displays predicted air/melt interface movements when the aluminum melt is allowed to flow through the nozzle slot and deposit onto the moving belt. The gap size between the backwall and the moving belt was set at 0.8 mm, the belt speed was 0.4m/s, and the average melt inlet velocity was specified as 0.8m/s. At 0.02s, the melt had dropped down sufficiently to touch the belt; the hydraulic impact spread the melt outwards, with most of the melt flowing with the belt downstream. However, a small part of the melt flowed backwards, entering the narrow gap between the backwall of the melt delivery system and the moving belt; at time=0.03s, the melt penetrated about -3-4 mm into the gap, owing to the impact static pressure. A little later, the viscous drag forces generated in the fluid by the moving belt, dragged the penetrated melt back into the main downstream flow. The meniscus profile then oscillated for a few milleseconds, and then became stable. Figure 5 (a) shows the flow pattern near the meniscus region after the melt flow had stabilized. The gap size for the backwall of the delivery system was 0.9 mm and the melt inlet velocity was 0.9m/s. The black curves show the meniscus profile and the top air/melt free surface profile. It was found that the melt flow near the meniscus region was slow, and that most of the incoming melt flowed in the downstream direction close to the upper air/melt free surface; the moving belt only dragged melt in its immediate vicinity downstream by viscous drag forces. A small counterclockwise re-circulatory flow, together with another small upper circulatory pattern, flowing clockwise, were formed between the upper main flow and the belt, The Figure 5(b) displays the static pressure field, related to the flow pattern in Figure 5(a). The pressure near the backwall meniscus region was the highest, owing to the hydraulic impact of the upper incoming flow. The high pressure pushed the melt towards the backwall gap, but the melt surface tension force was able to balance the over-pressure and form a stable meniscus profile.

Figure 4. Predicted images of air/melt interface profile with velocity field at different moments under tundish nozzle, melt inlet velocity=0.8m/s, backwall gap 0.8mm, belt moving speed 0.4m/s.

Backwall gap vs. meniscus behavior For a practical HSBC strip casting process, the vertical distance, or gap size, between the base of the backwall and the belt, may be somewhat variable during belt rotation, owing to slight variations in belt properties, or slight imperfections in the two (or three) rolls of the HSBC caster itself. We therefore modeled the effects of different gap sizes on meniscus behavior, using mathematical modeling. The backwall gaps were chosen to be between 0.8mm ~ 1.3mm. It was found that increasing the size of the backwall gap would lead to deeper penetration length of the melt entering the backwall gap.

For example, when the gap size was increased to 1.1mm under melt inlet velocity of 0.8m/s, and a belt casting speed of 0.4m/s, the melt would backflow, penetrating into the gap for about 10 mm, before the melt retreated back to the main flow, compared to a maximum penetration length of 3 ~ 4mm in Figure 4. When the gap size became more than 1.1mm, the melt would flow upstream, out of the backwall gap, forming a melt puddle in the front of the refractory backwall, as shown in Figure 6. The melt puddle would keep growing, resulting in a leaking accident in strip casting production. The minimum backwall gap size for preventing melt

799

leaking from the bottom of the backwall can be defined as the Critical Gap Size(CSG). Before the strip casting starts on HSBC caster, the backwall gap needs to be adjusted below the CGS corresponding to the casting parameters.

the pressure near the meniscus region increased greatly, and the critical gap size was decreased to 0.8mm. When the inlet velocity was 1.2m/s, the critical gap would reduce to 0.6mm. In order to reduce the melt inlet velocity flowing through the nozzle slot, the melt level should be kept as low as possible to reduce the melt's potential head. 1.2

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Figure 7. Melt inlet velocity vs. Critical Gap Size (CGS), belt moving speed = 0.4m/s. Belt moving speed vs. meniscus behavior The moving belt will drag the adjacent melt to flow in the belt's moving direction by viscous drag forces. This is of benefit in reducing melt penetration under the backwall. The effect of belt speed on meniscus behavior was also studied in the present mathematical modeling work. When the gap size was 1.3mm and melt inlet velocity was 0.8m/s, backflow will happen if the belt speed is 0.4m/s, as shown in Figure 6. However, when the belt moving speed increases beyond 0.6m/s, the penetrated melt under the backwall gap will be dragged back into the main flow after about 0.1~0.2s of the start of a cast. Figure 8 displays the predicted meniscus profile with fluid velocity field, after the melt flow has stabilized, in comparison with the predicted results in Figure 6.

Figure 5. Predicted meniscus and free surface profile, backwall gap is 0.9mm, inlet velocity is 0.9m/s and belt moving speed 0.4m/s. (a)flowpatter; (b) pressure, Pa.

Figure 6. Predicted melt flow pattern when the backwall gap size was 1.3mm, and belt moving speed was 0.4m/s at 0.06s after casting started. Melt inlet velocity vs. meniscus behavior The variation of melt level in the HSBC melt delivery system will change the hydraulic head over the nozzle slot. This, in turn, affects the melt inlet velocity through the slot nozzle and also changes the strength of the hydraulic impact and pressure when the melt touches down onto the moving belt. Different inlet velocities of aluminum melt were tested by mathematical modeling. It was found that when the melt inlet velocity increased, the critical size of backwall gap would decrease accordingly, as shown in Figure 7. For example, when the melt inlet velocity was 0.8 m/s, the critical gap size of the backwall was 1.1mm; when the melt inlet velocity was increased to 1.0 m/s,

Figure 8. Predicted stable meniscus profile and fluid velocity field when the backwall gap size is 1.3mm, and belt moving speed is 0.6m/s, for comparison with the results predicted in Figure 6. Experimental results In the water modeling experiments, the dimension of the nozzle slot was 82mmx2.5mm, the same size as in the casting experiments on the HSBC simulator, and the pilot scale caster at McGill University. The average inlet velocity through the slot

800

the gap, under the backwall, forming a big melt puddle. The melt puddle kept growing and growing until the moving substrate reached the end of its stroke. The experimental results agreed well with the results predicted by mathematical modeling.

nozzle under different water levels in the tundish was measured, and is shown in Figure 9. Since the water modeling had identical Reynolds number and Froude number as the HSBC simulator, the average melt velocity flowing through the nozzle slot could be directly deduced from water modeling experimental results. These flow rates depend on the melt level in the HSBC melt delivery system.

Summary The phase field method, coupled with the continuity and NavierStokes equations were successful in predicting the air/melt interface movement and meniscus behavior for the HSBC process. The backwall gap, combined with the melt inlet velocity through the tundish slot nozzle and the belt's speed, all were important in governing the meniscus behavior. Reducing the size of backwall gap, and/or reducing the melt level, and/or increasing the belt speed, were all beneficial in preventing melt backflow through the gap between the backwall and the belt. The formation of a stable meniscus, and meniscus line, is a necessary condition to the HSBC process. The predicted results are in good agreement with the experimental observations.

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4.

Aluminum melts were next cast on the HSBC simulator, and the high speed video camera, running at 300 fps, was used to record the behavior of the melt meniscus. Superheat of the melt in the "tundish" was kept at 30°C. The melt potential head in the tundish was 0.05m. The casting speed was 0.4m/s. The thickness of the refractory backwall was 10 mm, the same size as that used for the mathematical modeling. Figure 10 gives a photograph of the stable meniscus, as taken through the backwall gap during a casting experiment. The backwall gap was pre-set at about 0.7-0.8mm. The melt did not penetrate through the backwall gap, and form a stable meniscus profile underneath the tundish nozzle. Figure 11 presents photos of how the melt leaking from the bottom of the backwall after the substrate started to move. The photos were taken at 0.06s, 0.10s, 0.25s respectively. The backwall gap was set to be 1.1-1.2mm. It was clearly observed that when the melt dropped onto the substrate surface, the hydraulic impact pushed the melt upstream rapidly out through

5.

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Reference J.Herbertson and R.I.L.Guthrie: U.S. Patent 4,928,748, (1990) J.Herbertson and R.I.L.Guthrie: Canadian Patent 536533, (1992) K-H Spitzer and J.Kroos, "Process Technological Fundamentals for The Production of New Steel Grades by Direct Strip Casting" (1st Chinese-German Seminar on Fundamentals of Iron and Steelmaking, Beijing, 2004), 287. D. Li, S. G. Shabestari, M. Isac and R.I.L.Guthrie, "Studies in the Casting AA6111 Strip on a Horizontal, Single Belt, Strip Casting Simulator" (TMS 2006, 135th Annual Meeting and Exhibition, San Antonio, TX, 2006), 851-856. D. Li, Luis E. Calzado, M. Isac and R.I.L. Guthrie, "Improving the surface of AA6111 sheet material, cast at high speeds, through the use of macroscopically textured substrates" (TMS 2009, 138th Annual Meeting and Exhibition, San Francisco, CA, February 2009), 889-894. P. Yue, C. Zhou, JJ. Feng, C.F. Ollivier-Gooch, and H.H. Hu, "Phase-field simulations of interfacial dynamics in viscoelastic fluids using finite elements with adaptive meshing," J. Comp. Phys., 219(2006), 47-67. T.Iida and R.I.L.Guthrie, The Physical Properties of Liquid Metals (Toronto, ON: Oxford University, 1988).

Figure 10. Photograph of a stable meniscus when the backwall gap was 0.7-0.8mm, and the substrate belt speed was 0.4m/s.

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802

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Development of Alba High Speed Alloy Abdulla Habib Ahmed Ali, Jalal Mohd Ghuloom Hassan, Garry Martin and Kishore Ghosh Aluminium Bahrain B.S.C, (e), ALBA, Kingdom of Bahrain High speed, die pick-up, beta to alpha transformation, work environment, tearing, intermetallic, homogenization, anodizing

Abstract Aluminium Bahrain has developed in-house an alloy called "ALBA HIGH SPEED 6063.10", which is a variant of alloy AA 6063. The driving force behind this has been the ever standing requirement of extruders to attain progressively higher extrusion speeds, thereby boosting productivity. Small and controlled amounts of Mn and Mg additions to alloy AA 6063 were found to help in improving extrusion speeds and reduce die pick-up thereby improving the surface appearance. The reason was attributed to improved beta to alpha phase transformation rates at the microstructure level. This was achieved by having tight control of Mn addition to reduce the band from 0 % - 0.10 % in normal AA6063 alloy to 0.03% - 0.06 % in 6063.10. In addition controlling the Mg addition to 0.45 % - 0.51 % in the new alloy, assisted in better distribution of Mg2Si precipitates in the matrix during the cooling process, post soaking at 580 °C.

Serial No. 1 2 3 4 5 6

Diameter (mm)

Diameter (Inch)

152.4 177.8 203 216 228.6 254

6 7 8 8.5 9 10

No. Of logs/ Cast 128 108 84 68 68 54

Table 1 : Alba billets sizes and configuration

Introduction

Background

Aluminium Bahrain (Alba) is a primary aluminium smelter located in the Kingdom of Bahrain. It has a capacity of about 860,000 mt per annum. The Casthouse-3 of Aluminium Bahrain (Alba) produces billets (about 340,000 mt per annum) in the following alloy groups:

In 2008 some of our customers in the GCC were exploring possibility of extruding our products at higher speeds. In this process, they had encountered some tearing problems that had hindered their efforts. This was investigated by a team of specialists who attributed this effect to the quantum of transformation of beta AlFeSi (plate like monoclinic) phase to alpha AlFeSi (rounded cubic) intermetallic phase, during homogenization [1].

•

6005 (Al-Si-Mg) Typical use: Extruded shapes and tubing for commercial application requiring strength greater than AA6063.

•

6060 (Al-Si-Mg) application.

•

6063 (Al-Si-Mg) Typical use: Architectural extrusion, door & windows.

•

6061 (Al-Si-Mg-Cu-Cr) Typical use: Trucks, pipelines, and other structure applications where strength, weldability, and corrosion resistance are needed.

•

6082 (Al-Si-Mg) Typical use: Extruded shapes and tubing for commercial applications requiring strength greater than AA6063.

Typical

use:

An extensive literature study was undertaken which revealed that small amounts of Mn addition (0.03% - 0.06% by wt) significantly increased the transformation rate from beta to alpha phase and thereby improved extrudability. [1,2] The exact addition rate of Mn was arrived at, through trial and error. Subsequently Alba High Speed AHS 6063.10 was developed after extensive trials at some of our customers' end. Soon after development of this alloy, Alba had conducted technical workshops across the GCC and Middle East to launch the product to a wide customer base. The response was overwhelming and since then, several customers have readily accepted this alloy in lieu of the conventional 6063, which improved their extrusion speeds and productivities by an average of 30 %.

Architectural

Metal preparation and condition

Alba uses direct chill (DC) MaxiCast™ Airslip billet casting technology for manufacture of billets. This method uses enhanced cooling technology feature. Alba produces billets in the following sizes and in the following configuration in terms of log densities as shown in Table 1.

Alba's liquid aluminium metal is treated to minimize hydrogen gas and oxide levels prior to casting into billets by air pressurized (DC) MaxiCast™ Airslip casting. This process is designed to produce high quality billet with consistent extrusion properties.

803

After casting and inspection, ALBA billets are homogenized to 580 °C for two hours to achieve a uniform metallurgical structure with more than 95% ί to a transformation. Homogenizing ensures that billets conform to the extrudability requirements with the end product developing maximum mechanical properties consistent with the composition of the alloy and the specified temper. Features of AHS 6063.10 AHS 6063.10 alloy is a medium strength alloy with excellent extrudability and anodizing characteristics. The alloy can achieve minimum - T5 and - T6 properties specified for 6063 alloy.

Alloy 6063.10 can be used in application requiring medium strength levels. Typical uses include architectural, building, kitchen suite sections and general purpose extrusions. Physical properties of AHS 6063.10 Physical properties are given in Table 2 below. Typical values

Density

2700 kg/m3

Modulus of thermal expansion

69.5 GPa

Coefficient of thermal expansion

23.4 x 10-6 / °C

Thermal Conductivity

209W/(m.K)

Electrical Conductivity

32 MS/m

Approximate Melting Range

615°Cto650°C

i il

T5(MΩ») To (Mût) s Tensile Strength ( M P t ) «Yield StrengA (MFa) Pttip>«»*y

T5ί/im)

Typical ®Elongatiott<«V0> Typical

Figure 1: Mechanical properties of AHS alloy

Application of AHS 6063.10

Property

Mechanicalpropertibs AHS - 6063.10

Field Trials Several Trials were conducted successfully and jointly with customers. The trial conducted at Al Taiseer KSA with 203 mm billets in alloy AHS (6063.10) is discussed in this paper and the result are shown in Table 4: Methodology of trials

1

Table 2: Typical physical properties (T5 temper)

Mechanical properties of AHS 6063.10 With appropriate extrusion and heat treatment practices, the following minimum and typical properties can be achieved as shown in Table in 2. The minimum properties are as specified in the Aluminium Association, USA for 6063 alloy.

The methodology followed at the extruders for the development of optimum extrudability using the AHS 6063.10 alloy billet involved two key steps. Those are outlined below and it was decided that representatives from ALBA's technical team be present to work with the extruder to facilitate the development of optimum extrusion performance using the 6063.10 alloy billet. Stepl. The AHS (6063.10) billets were extruded under the existing process conditions particularly in connection with temperatures (die, billet and exit) as well as extrusion speeds that were being used with the existing 6063 type conventional alloy. All extrusion process conditions as well as the resultant product quality particularly surface condition, recovery and aged (-T5) mechanical properties were recorded for analysis .If anodizing was done by the extruder, then samples from the anodized material were collected for review by the team involved in the optimization work. Step2. Subsequent to a review of the existing process conditions and the resultant section product quality, dies were selected for the optimization work .This included two to three dies used for extrusion of solid profiles as well as two to three dies used for extrusion of hollow profiles. The dies selected were significant usage dies by the extruder and had a good die history recorded. The selected dies were ensured to be well nitrided and in a good condition of maintenance. The selected dies were then heated to the specified temperature for extrusion prior to being loaded to the press. The plan was to extrude the ALBA AHS 6063.10 alloy billet, around 10 billets on each die at progressively increasing extrusion speeds under closely monitored conditions of extrusion temperatures, particularly exit temperature, and speed measurements.

The target was to achieve up to a 50% speed increase on a given die without any deterioration in the as- extruded surface finish (pick- up or die lines) or the onset of tearing or speed cracking. The temperature control in the log or billet heater being used by the extruder was a key requirement in the achievement of the optimum extrusion performance using the AHS6063.10 alloy billet and this was a focus area for monitoring throughout the work. Subsequent to the optimization of maximum acceptable extrusion speeds on the selected dies, the as-extruded material was aged under standard conditions of temperature & time and the mechanical properties. Hardness readings were taken to confirm that the material has been aged to the desired level. If anodizing is carried out then the material was anodized to clear and bronze colors, to again confirm that the required anodized quality level has been achieved.

Die Preheating For extrusion of solid and hollow sections the dies were heated to around. 450 °C .Dies should not be overheated or held for long times in die oven. Results of Trials Samples were sent to external Lab. (SECAT) for evaluation of microstructure before doing field trials at customer end. The mount consisted of three homogenized billet [3]. The summary of SECAT's report is shown in Table 3: 1.

a.

Typical processing conditions

b.

Preheating Billets were preheated to a temperature in the range of 440 °C480 °C to ensure good extrudability and mechanical properties. Difficult sections such as thin walled extrusions or hollows may requite higher billet preheat temperatures (~ 490 °C), while simple sections may be extruded with lower preheat temperatures (~ 430 °C). Press Quenching The cooling rate of the extrudate from exit temperature down to 250 °C should be in excess of 60 °C/min to achieve minimum mechanical properties. For thin sections (less than 3mm), a still air or fan air cooling should be adequate to achieve this cooling rate. For thicker sections (3 to 6 mm), vigorous fan cooling is usually required to achieve the minimum cooling rate. Water mist cooling may be required on thicker sections. Straightening

Homogenization was good with only traces of Mg2Si seen indicating that Mg and Si have been retained in solid solution, refer to Figure 3.

c.

2.

The liquation depth was 90um; it was low and acceptable, refer to Figure 2. The Beta to Alpha conversion was good and greater than 98%; Intermetallic particles were fragmented. The intermetallic particles were relatively small and consistent in size with thickness <1 μπι and length 0.5-10 urn.

In general, billet microstructure was good with no problems expected during the extrusion process. Extrudability was expected to be good. This was based on practical guidelinesfromthe following: a.

The very low levels of Mg2Si

b.

The intermetallic particles were in the form of rounded or small rod like particles with samples showing excellent particle break up and Beta to Alpha conversion.

Stretching to an elongation of about 0.5% is recommended for straightening while stretching beyond 1% elongation may result in orange peel finish on the extrudate. Ageing Treatment Various ageing times and temperatures can be used to obtain good mechanical properties. However a typical aging condition for 6063.10 alloy in production operation is 4 hours at a soak temperature of 185 °C+/5°C . Extrusion Conditions The temperature of the container should be set at least 30 °C below the billet temperature to ensure the billet skin is retained in the butt. The billet preheat temperature and extrusion speed should be controlled so that the exit temperature is maintained in the range of 500 °C to 550 °C .Typical extrusion speeds for this alloy are in the range of 20 to 50 m/min for solid profiles and 15 to 40 m/min for hollow profiles.

Figure 2: Optical microphotograph of alloy AHS (6063.10) sample showing liquation depth

Hardness (Webster) AHS Vs AA 6063 15.5 15

Î

|l4.5

"

£ 14 " §13>5 | 13 χΐ2>5 12 -

Figure 3: SEM images of particle structure alloy AHS 6063.10

1

11

!

ti1

11 S 1

Die# 22676-77

Dier22676-77

1 It

0ie:22676-83

Die:13606-Φl

M ANS 6063.10 IAA 6063.33

Figure 4: Extrusion Speed AHS Vs AA6063.33

Table of Summary of evaluation Alba Billet alloy AHS (6063.10) analysis results [3] C entre >98

ί to a (%) Middle >98

Edge >98

Mg2Si<^m C entre | trace

Middle trace

Edge trace

Compound Size (μιη) Length Thickness 0.5-10 <1 Liquation Over Depth Heated (um) 90

none

1

Table 3: Summary of field trial of alloy AHS (6063.10) analysis results at Al Taiseer extrusion company KSA.

|

Extrusion Speed AHS Vs AA 6063 35

— 1 30

f 25

jJL 2 0 1 15

i™ I

5

D«e# 22676-77

Die: 22676-77

I AHS 6063.10

Customer Feedback The objectives listed earlier were achieved from the trials conducted at customer's end. The development and implementation of the new ALBA HIGH SPEED (AHS) 6063.10 alloy billet was progressed as a result of the visit made. Approximately 10.5 % relative speed increase was achieved on the selected dies (Hollow & Solid) sections being extruded at the time of the trial using the AHS6063.10 alloy billet. [4)

Die:22676-83

Die:là606-01

«AA 6063.33

Figure 5: Hardness (Webster) AHS 6063.10 Vs AA6063.33

1. 2. 3.

4.

The break through pressure increase in ALBA AHS (6063.10) as compared to AA6063 shows evidence of a "stronger alloy". Even the running pressure which is measured after 100 mm of Extrusion shows / support comment # 1. The surface finish appearance during Extrusion / weld joint inspection shows less metal pick up, meaning better flow of metal. This result shows the evidence that ALBA AHS (6063.10) was stronger by approx 10.5 % when compared to AA6063

Conclusion As is evident, the development of alloy AHS 6063.10 has helped the extruders to enhance their extrusion productivity along with improvement in strength. The technical team at Alba has also shown that by working closely with the extruder and monitoring the key extrusion parameters of billet and exit temperature on the press the process can be adjusted to obtain extrusion speed increases of up to approx 30 % over the previous AA6063 type alloy when run on the same die

Die#: 22676-77 Parameters Press # Billet temp °C Wall thickness of profile mm Extrusion speed mts/min Surface finish Ageing Cycle °C -hrs Webster reading( Hardness ) UTS Kg/mm2 Parameters 1 Press # Billet temp °C 1 Wall thickness of profile mm 1 Extrusion speed mts/min 1 Surface finish 1 Ageing Cycle °C -hrs 1 Webster reading( Hardness )

1 UTS Kg/mm2

AHS 6063.10

6063.33

4 4 470 470 0.5 0.5 27.2 18.5 V. Good Good 200 - 4.50 200-4.50 14 15 24.25 23.8 Die#: 24855-17 AHS 6063.10

6063.33

4 470 0.75 21.8 V. Good 200-4.50 15 N/P

4 470 0.75 19.8 Good 200 - 4.50 13 N/P

Die#: 22676-77 Parameters Press # Billet temp °C Wall thickness of profile mm Extrusion speed mts/min Surface finish Ageing Cycle °C -hrs Webster reading( Hardness ) UTS Kg/mm2 Parameters Press # Billet temp °C Wall thickness of profile mm Extrusion speed mts/min Surface finish Ageing Cycle °C -hrs Webster reading( Hardness ) UTS Kg/mm2

AHS 6063.10

This improvement in the extrusion speed, hence productivity, at the extruders is obtained while maintaining a good surface finish on the extruded product for subsequent surface finishing by either anodizing or powder coating. Mechanical properties after standard ageing cycle are also well above the minimum values specified for AA6063 alloy in the -T5 temper and with appropriate cooling of extruded section on the press above the mechanical properties specified for -T6 temper. This development at Alba has therefore contributed significantly to overall optimization of performance for the extrusion industry and into the billets markets that Alba supplies. References 1.

S. Zajac et al. "Microstructure control and extrudability of aluminium-Mg-Si alloys micro alloyed with manganese", JOURNAL DE PHYSIQUE IV Colloque C7, supplement au Journal de Physique 111, Volume 3, Novembre 1993, 251254.

2.

Joseph R. Davis, "Aluminum and aluminum alloys," J. R. Davis & Associates, and ASM International. Handbook Committee, 711

3.

"Evaluation of AHS 6063,10 microstructure" Report-98-10003, SECAT Ine, USA, 2010.

4.

"Report on processing of AHS 6063.10 alloy" (Al Taiseer KSA)6&7July2010.

6063.33

4 4 470 470 0.5 0.5 27.2 18.5 V. Good Good 200 - 4.50 200 - 4.50 14 15 24.25 23.8 Die#: 24855-17 AHS 6063.10

6063.33

4 470 0.75 21.8 V. Good 200 - 4.50 15 N/P

4 470 0.75 19.8 Good 200 - 4.50 13 N/P

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DISSOLUTION STUDIES OF Si METAL IN LIQUID Al UNDER DIFFERENT FORCED CONVECTION CONDITIONS Mehran Seyed Ahmadi1, Stavros A. Argyropoulos1, Markus Bussmann2, Don Doutre3 department of Materials Science and Engineering, University of Toronto department of Mechanical and Industrial Engineering, University of Toronto 3

Novelis Global Technology Center, Kingston, Ontario, Canada Keywords: Si-Al alloys, Dissolution, Forced convection

Abstract

Atomic Patent Silicoü y .......y... ff . y

The dissolution of solid cylinders of Si in a revolving tank of molten AI was studied, under both natural and forced convection conditions. The experiments were carried out at bath superheats ranging from 21°C to 78°C. The dissolution of Si increased with increasing bath superheat, and was accelerated by the forced convection. The increase of the mass transfer coefficient with increasing tank rotation rate clearly shows that the rate controlling factor in the dissolution of Si into Al is mass transfer across the concentration boundary layer, for the range of rotation rates tested.

y y

Introduction The assimilation rates of additions affect the production cost of alloys. The assimilation of a solid metal into a liquid metal can be distinguished as melting or dissolution: the former takes place through heat transfer, while the latter occurs when the solid comes in contact with liquid at a temperature below the melting point of the solid. This study focused specifically on the dissolution of solid Si into liquid Al. Ai-Si alloys have attractive properties for the aerospace and automobile industries. However, the production process is very slow, because the time required for Si to dissolve into liquid Al is long. Industrially at Novelis, chunks of Si are added to molten Al, which is then stirred for up to 30 minutes. During that time, dross forms on the top of the melt that is skimmed off; the longer the alloy making process, the more dross forms. Therefore, we are seeking to accelerate the dissolution process to reduce material loss and energy consumption, by reducing the tap to tap time. Significant savings would be achieved by only a 10% decrease in the dissolution time [1].

Weighs PeccsBt Silke©

Figure 1. Al-Si phase diagram showing the liquidus curve [3]. 2. Transport of dissolved species from the interface to the bulk liquid metal, CSat—>Cb The dissolved species travels from the solid-liquid interface to the edge of the concentration boundary layer, and then from the edge of the boundary layer to the bulk liquid Al. The solute concentration profile next to a dissolving addition is shown in Figure 2. Liquid

When an addition (initially at room temperature) is introduced to a molten bath, the addition and the liquid quickly reach a thermal equilibrium, and then the solid begins to dissolve and its size gradually decreases. Dissolution of a solid addition in a metal bath takes place as two steps [2]: 1. Interface reaction at the solid-liquid interface, CSj—ΚûäÅί

cb

2

Atoms migrate from the solid phase into the melt. The concentration of the dissolving species in the liquid at the interface is given by the liquidus curve on the phase diagram, at the operating temperature, see Figure 1.

3 O

Mass Transfer Boundary Layer

Distance normal to the addition Figure 2. Solute concentration profile in the solid addition, at the solid/liquid interface, and in the liquid metal.

809

When the rate of dissolution is determined by the rate of mass transfer across the concentration boundary layer, the flow in the vicinity of the addition will have a significant influence through the mass transfer coefficient associated with that flow. Examples of dissolution limited by the second step include the natural convection work of Gairrola et al. [4], who studied the dissolution of a vertical Ni cylinder in liquid Al, Shoji et al. [5], who studied the effect of natural convection on the dissolution of solid Cu in molten Sn-Pb alloys, and Niinomi et al. [6], who studied the rate of dissolution of ferrous alloys into molten Al assuming natural convection mass transfer. Their results showed that Fe-Cr, Fe-Cu and Fe-Ni dissolution is controlled by mass transfer within the boundary layer. On the other hand, the same authors showed that the first step, the interface reaction, to be rate limiting for the Fe-C alloy and that both steps (the so-called "mixed effect") affects the dissolution rate for Fe-Si, commercially pure iron, and Fe-Mn. Turning to forced convection, Shoji et al. [7] studied the dissolution of a Cu cylinder in molten tin-lead alloys. The velocity ranged from 0 to 0.754 m/s; the temperature was fixed at 673K. It was shown that the dissolution rate increases under forced convection. To be best of the authors' knowledge, the effect of flow conditions on the dissolution rate of Si into Al has not been studied. The only study related to the Al alloying process using Si is that of Balusis et al. [8], who showed that the use of Si granules over Si chunks decreases the dissolution time of Si. It was found that when the diameter ratio of lumpy to granular is about 5, the required time for dissolution increases by a factor of between 2 and 3. The current study seeks to accelerate the dissolution of Si into liquid Al, thus decreasing the cost of producing Al alloys with high Si levels, by quantifying the impact of agitation of the liquid Al using mechanical stirring (i.e. forced convection). To address this objective experiments have been conducted for different velocities of the liquid Al. This work provides a quantitative comparison of the dissolution rate under natural and forced convection conditions. Experimental Setup Electrical Furnace and Revolving Liquid Metal Tank: An electric resistance furnace was used to melt the Al. The furnace was constructed of refractory brick within a steel enclosure; the central heating area for the crucible totaled 0.018 m2. Two K-type thermocouples located in the furnace hot zone were used to control the temperature. The cylindrical stainless steel tank that rotates within the furnace is shown in Figure 3. The Revolving Liquid Metal Tank (RLMT) is connected to a DC motor that controls the speed to within 1 RPM. The interior diameter of the RLMT is 43.2 cm, the height is 19.1 cm, and it has a capacity of approximately 50 kg (20 1) of Al. A heavily insulated lid is used to avoid heat loss through the top of the furnace. The lid has two holes, as shown in Figure 3; one (diameter of 7.9 cm) is used to insert a thermocouple into the center of the Al bath; another (diameter of 6.3 cm) is used to immerse Si samples. The center of this second hole is located 16.7 cm from the center of the RLMT.

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A Hole to insert a thermocouple 'B Hole to insert a Si sample

Figure 3. Schematic of the electrical resistance furnace including the revolving liquid metal tank showing the holes with corresponding dimensions Si Cylinders The Si additions were made of metallurgical grade Si; the nominal composition is presented in Table 1. Table 1. Nominal composition of the metallurgical grade Si Si(%) Fe(%) Al(%) Ca(%) Ti (%) >98.0


<0.25

~l

~l

The Si we use arrives as large pieces, from which cylinders of 18.75 mm diameter were cut. The length of the cylinders is approximately 10 cm, of which 8 cm is immersed into the liquid metal while the rest is used to hold the cylinder. Both ends of each cylinder are trimmed on a diamond wheel saw. Before each experiment, the cylinders are cleaned with acetone to remove any grease or dirt. Al Bath The Al is of commercial purity (99.87% Al, 0.04% Si and 0.09% Fe) and is received as ingots. The temperature of the Al bath is measured with a K-type thermocouple with tip stationed at the center of the bath about 5 cm below the free surface. Experimental Procedure To initiate an experiment, the electrical resistance furnace is heated from room temperature; the RLMT contains approximately 46 kg of Al. It takes about 12 hours to melt the Al. When the melt temperature is stable, the target rotating speed is set. Then, a Si cylinder of known weight is attached to the holder and immersed into the liquid. The immersion speed is approximately 5 cm/s. After a specified time, the cylinder is withdrawn and detached from the holder. Any attached Al is removed using 38 vol pet HC1 at room temperature. This solution dissolves the Al while the Si remains intact [9]. Then, the cylinder is weighed again. Figure 4 (a) and (b) show the Si samples after immersion for different times, with and without the attached Al.

Table 2. Variation in the dissolution of identical samples under the same experimental conditions (T = 738°C and rotational speed = 10 rpm) 0.39 0.43 0.37 0.45 0.39 0.47 0.29 0.43 0.31

a)

The uncertainty associated with the ratio of the dissolved mass to initial mass can be calculated as

fuM m.

5 min

7 min

where n^ (kg) is the initial mass of the addition and η^ is the dissolved mass. The value of Umd/mi is 0.074 based on these experimental results.

8 min

b)

5 min 7 min

The source of scatter in the data can be largely attributed to the surface of the samples. The Si samples have different macroscopic structure depending on the chunk of Si from which they were cored. Some of the samples have very small pores, some larger and some have directional scratches, related to the method of casting of the Si. For example, the pores are generated due to trapped gas bubbles during the solidification of Si. Nevertheless, these variations reflect the reality of the dissolution process. Put another way, industry uses Si in the same form and so will experience similar variability in the dissolution of Si. It should be noted that the same uncertainty is assumed for all of the data presented here, because it is not possible to repeat a test for each data point several times given the difficulty of the experiments and the required resources. Therefore, the results presented from now on all correspond to a single measurement.

8 min

Figure 4. Si samples after immersion for different times, with and without attached Al, under natural convection condition.

Effect of Bath Superheat on the Dissolution Rate The first objective of this work is to study the effect of bath superheat on the dissolution rate of a Si addition in Al.

Experimental Results and Discussions Using the experimental procedure described above, the dissolution of a Si addition into liquid AI was studied. The bath temperature was set at 681, 703, 715 and 738 °C for different experimental runs. These temperatures correspond to superheats of 21, 43, 55 and 78°C, respectively.

As shown in Figure 5, the superheat of the bath strongly affects the rate at which the sample dissolves. The higher superheat increases the saturation concentration of Si in liquid Al (according to the liquidus curve on the Al-Si phase diagram), which in turn provides a greater driving force for mass transfer.

The RLMT was rotated at ω = 0, 10 and 20 rpm for each melt temperature. The tangential velocity of liquid metal at the center of the Si cylinder was calculated using the following equation, 2π u fl =—xr rot xco È

6 0

rot

(2)

g 0.8

(1)

where rrot = 16.7 cm is the distance from the center of the tank to the point of immersion,. These rotational speeds correspond to tangential velocities of 0, 0.18 and 0.36 m/s, respectively.

Eo,

Reproducibilitv of the Experimental Results We begin by assessing the variability of the results. An experiment was run to measure the dissolved fraction (the ratio of dissolved mass to the initial immersed mass of the cylinder) of cylinders at the same superheat (78 °C), rotational speed (10 rpm) and immersion time (2 min.). The test was performed for 9 cylinders and the results are presented in Table 2. The average dissolved fraction and the standard deviation are 0.39 and 0.023, respectively.

Time (min.)

Figure 5. The effect of superheat on the dissolution rate of a cylindrical Si addition in Al, under natural convection conditions. In this study a range of bath superheats (from 21 to 78°C) was investigated. At the same immersion period (t = 5min.), Figure 5

811

shows a 4.5 times increase in the dissolution rate when the degree of superheat increases from 21 to 78°C.

a)

Extrapolating the experimental data (linear regression of the experimental data points) also reveals that the higher superheat causes the dissolution to begin sooner. As shown in Figure 5, when the superheat is just 21°C, it takes approximately 3 minutes for the sample to start to dissolve in the bath, while this time decreases to about 1 minute at 78°C superheat. The reason could be the higher heating rate of the sample (from room temperature upon immersion to the bath temperature at higher superheat), to reach equilibrium with the bath, and retaining the saturation concentration at that temperature at the solid/liquid interface.

g 0.8

8 0.5

m

E 0.4 o j>0.3|-

^ A

The superheat of the bath has the same effect on the dissolution rate when the RLMT rotates (i.e. forced convection). As shown in Figure 6, for both rotation rates, it takes less time for an addition to dissolve at higher superheats. An increase of superheat from 43 to 55°C yields a 4.5 times faster dissolution of the addition after 3 minutes of immersion at 10 rpm. The same increase in the superheat increases the dissolved mass by a factor of 2.3 at 20 rpm after immersion for 3 minutes.

Natural convection Rotational speed = 10 rpmj Rotational speed = 20 rpml

b)

go.sU

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21 C 43 C 55C 78 C

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1

5

6

Time (min.) I

I

:

4

A

4 I 1

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T

A 1

7

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SPH = 55 C I

at j_

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t

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Time (min.) b)

2

c)

co 0 · 9 (Λ 2 0.8

I

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Time (min.)

Rotational speed = 10 rpm

a)

Natural convection Rotational speed = 1 0 rpm! Rotational speed = 20 rpm1

1 1

i 1

Natural convection Rotational speed = 10 rpm Rotational speed = 20 rpml 1 1 J 1 1

Time (min.)

ff

d) _* SPH SPH SPH SPH

l

I

.

I

.

Natural convection Rotational speed = 10 η Rotational speed = 20 η

= 21 C = 43C = 55 C = 78C

I

Time (min.)

Figure 6. The effect of superheat on the dissolution rate of a cylindrical Si addition into Al under forced convection conditions, when rotating at a) 10 and b) 20 rpm.

E 0.4 σ

Π

§0.3

Effect of Rotating Bath The rotation of the RLMT introduces forced convection near the immersed sample. It is known that forced convection increases the mass transfer rate at the interface by decreasing the mass transfer boundary layer thickness. As a result, higher dissolution rates are expected when the bath rotates (at a constant liquid Al temperature).

_î_ Time (min.)

Figure 7. The effect of forced convection on the dissolution rate of a cylindrical Si addition into Al at a) 21, b) 43, c) 55 and d) 78 °C superheat.

812

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Figure 7 depicts the effect of forced convection on the dissolution rate of a cylindrical addition at different superheat values. As shown, when the bath rotates the addition dissolves more quickly than when the bath is still. This clearly indicates that the mass transfer across the boundary layer is the rate controlling factor in the dissolution of Si into Al, for these rotational speeds.

IE

S

IO" 3

E ~io- 4 e

I I

where Csat (kg/m ) is the saturation concentration of the addition in the liquid at temperature T (°C), and C b (kg/m3) is the bulk concentration.

&8tn1 0

The average mass flux m"ave (kg/s/m2) from the solid addition to the liquid from t = 0 (the moment the cylinder is fully immersed into the bath) to time t (when the cylinder is removed from the bath) can be calculated as

c

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Time (min.)

81oL

ioV

In the above equation m^ (kg) is the dissolved mass of the Si sample and Aave (m2) is the average solid/liquid interface area, expressed as

fι

e

io 7 ,

(3)

A;+Af

I

r

An average mass transfer coefficient km (m/s) can be calculated by the following equation

A„„„ —

A

Ö

8io;

The dissolution of Si into Al is a mass transfer process, that is usually quantified by a mass transfer coefficient that is a function of operating conditions. Invoking some simplifying assumptions the mass transfer coefficient for the dissolution of Si in Al is calculated using the experimental data presented here.

tA„

SPH = 21

i10

Experimental Mass Transfer Coefficient

CL.-C

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,

I 1 1

2

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Natural convection Rotational speed = 10 rpm I Rotational speed = 20 rpm | I I I I I I I 1 I 1 I I 4 5 6

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Time (min.)

c)

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where Aj (m ) is the initial surface area of the immersed part, and Af (m2) is the final surface area. Af can be calculated assuming the following: • •

C

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The density of the Si cylinder is constant (i.e. the variability in the density due to different pore distributions can be neglected). The bottom surface area is small relative to the side area. Based on the cylinder dimensions in our current experiments, the initial ratio is 1.86% which justifies this assumption.

io-7, IO" 3

i

m

2

3

4

I ! I i I i I i I

5

Time (min.)

Λ d)

SPH = 78

*
μ s

Therefore, d

Natural convection Rotational speed = 10 rpm Rotational speed = 20 rpm

L

t * * 11

?

Π

310 =

f

(6) Af =A;X—-x— L df m; where df is the average value calculated from 5 measurements of the diameter of the cylinder. As shown in Figure 4, the dissolution of Si is not uniform; that is, the diameter of the cylinder varies along its length. Therefore, an average diameter was used to estimate the surface area of a cylinder after removing it from the bath.

« 10-

10',

I

I

1 3

* ■ ·■ · ♦ 1 4

Natural convection Rotational speed = 1 0 ipm Rotational speed = 20 ipm 1 1 1 1 1 I 5 6

Time (min.)

Figure 8. Estimated mass transfer coefficients for different flow conditions, using experimental data at a) 21, b) 43, c) 55 and d) 78°C superheat.

813

Figure 8 shows the calculated average mass transfer coefficients at different bath superheats. The mass transfer coefficient ranges 2.9 x 10"6 to 4.8 x 10~5 m/s for natural convection, 1.3 x 10~6 to 9.4 xlO" 5 m/s at 10 rpm, and 1.78 xlO" 6 to 1.4 xlO" 4 m/s at 20 rpm, depending on the bath superheat. It should be noted that the error bars shown in the graphs are associated with the calculated mass transfer coefficient using the following equation,

1 tA AC

U md

Λ2

+

(

-m,

-u

V

t 2 A ave AC t

-m,

(7)

rUA1 tA AC 2

tALAC

For the data points without error bars the calculated uncertainty was higher than the value itself, because the amount of dissolved mass was less than the uncertainty associated with the immersed mass of the cylinder. As can be seen, the mass transfer coefficient increases as the rotational speed increases. This clearly shows that the forced convection causes a faster dissolution of the specimen within the bath. Finally, the mass transfer coefficient can be defined as km=-

l-'Qi/Al

(8)

where D Si/A1 is the binary diffusion coefficient of Si into liquid Al and τ m (m) is the mass boundary layer thickness. This equation in conjunction with the obtained experimental results indicates that the boundary layer thickness decreases when the fluid is agitated in the vicinity of the sample. A decrease in the mass transfer boundary layer thickness implies a lower resistance to transport of species from the solid/liquid interface to the bulk of the liquid, and therefore the dissolution rate increases. Conclusions The dissolution of solid Si in molten AI was studied under natural and forced convection conditions at different superheat values. •

An increase in bath superheat causes faster initiation of the dissolution process, as well as faster dissolution of the Si.

•

Forced convection causes faster dissolution of Si within the Al, which indicates that mass transfer across the boundary layer controls the dissolution process, for tangential velocities from 0 (stationary tank) to 0.36 m/s (the highest tangential velocity tested). The maximum mass transfer coefficient associated with 78C superheat increased from 4.8 x 10"5 to 1.4 x 10"4 m/s as the tangential velocity increased from 0 to 0.36 m/s. This clearly indicates that in order to dissolve silicon faster some type of agitation must be induced within the bath.

Acknowledgments The authors would like to acknowledge the financial contribution of the Natural Sciences and Engineering Research Council of Canada for its financial support through a Strategic grant. References 1. Dutre, D., private communication, Novelis, Kingston, 26 August 2010. 2. Mazumdar D., Evans J.W. Modeling of steelmaking processes (Boca Raton London New york: C R C Press, Taylor & Francis Group; 2009) 3. Massalski T.B., Okamoto H., Subramanian P.R., Kacprzak L., eds. Binary alloy phase diagrams, 2nd ed (Materials Park, Ohio: ASM International, 1990). 4. Gairola P.K., Tiwari R.K., Ghosh A., "Rates of dissolution of a vertical nickel cylinder in liquid aluminum under free convection," Metallurgical Transactions B, 2(8) (1971), 2123-2126. 5. Shoji Y, Uchida S, Ariga T. "Dissolution of solid copper into molten tin under static conditions," Transactions of the Japan Institute of Metals, 21(6) (1980), 383-389. 6. Niinomi M, Ueda Y, Sano M., "Dissolution of ferrous alloys into molten aluminum," Transactions of the Japan Institute of Metals, 23(12) (1982), 780-787. 7. Shoji Y., "Dissolution of copper cylinder in molten tin under dynamic conditions," Welding Journal, 60(1) (1981), 19-24. 8. Baluais G., Brown M., Strydom, J., "Silicon granules for aluminum alloys" (Paper presented at TMS 2, Seattle, TMS, February 2002). 9. Zulehner, W., Neuer B., and Rau G., Silicon, in Ullmann's Encyclopedia of Industrial Chemistry, (Wiley 2005).

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Modification and grain refinement of eutectics t o improve performance of Al-Si castings M. Felberbaum, A.K. Dahle CAST CRC, Materials Engineering, The University of Queensland, Brisbane Qld 4072, Australia Keywords: Al-Si alloys, Eutectic modification, Nucleation tectic grain refinement, which is thought to be essential to improve interdendritic feeding, reduce porosThe formulation of alloy compositions for aluminium ity, and enhance mechanical properties in Al-Si alcastings has not changed significantly for decades. loys. Strontium (Sr) modification is commonly used to obtain a refined fibrous morphology of eutectic in Al-Si Background foundry alloys, and grain refinement can be used to refine the primary Al dendrites. Compared to the During the solidification of the Al+Si eutectic, Si is coarse plate-like unmodified silicon morphology, the believed to be the leading phase and needs to be nurefined fibrous eutectic structure can substantively cleated before eutectic growth can occur [7]. Thereimprove the mechanical properties particularly duc- fore Si nucleation is a key factor for eutectic solidifitility and fatigue. However, a problem often asso- cation. The nucleation frequency and spatial districiated with Sr modification is a change in porosity bution of eutectic cells depend on the available nu- distribution and amount. A new alloy technology clants present in the remaining liquid when eutechas been developed to produce castings combining a tic temperature is reached. Depending on the alloy well-modified eutectic with reduced porosity. This composition and solidification conditions, three main can be achieved by providing effective substrates for eutectic nucleation modes have been distinguished in the nucleation and growth of eutectic cells. This pa- hypoeutectic Al-Si foundry alloys [5] (see Fig. 1). per presents the current status in the development Figure l a illustrates the nucleation of the Al+Si euand commercialisation of the new eutectic grain refiner and modifier. Abstract

Introduction Strontium is commonly used to refine the eutectic structure of Al-Si alloys. Prom a coarse plate-like structure in unmodified alloys, a fine fibrous morphology is obtained by only trace additions of Sr. Such a fine structure can potentially increase the ductility and fatigue life of cast specimens [1]. However modified alloys are observed to contain more porosity than unmodified ones, reason why many industrial companies choose not to use modifiers. Several reasons have been raised to explain this increased porosity due to Sr addition, among others: decreased solid-liquid interfacial energy and increased volumetric shrinkage [2], increased susceptibility of the melt to absorb hydrogen [3], influence on oxide bifilms formation [4], and inhibition of interdendritic liquid due to the equiaxed growth of large eutectic grains [5, 6]. This paper reports a new technique to produce castings having (i) a well-modified eutectic and (ii) a reduced porosity. It is based on the key concept of eu-

Figure 1: Possible eutectic nucleation modes in the Al-Si system [5]: (a) nucleation on primary aluminium dendrites, (b) front growth opposite to the thermal gradient, and (c) heterogeneous nucleation in interdendritic liquid. The coupled Al+Si eutectic growing side-by-side is illustrated in black, while the dendrites are shown in white. tectic (black on the figure) on primary aluminium dendrites (shown in white here). Figure l b illustrates the case where nucleation takes place at or adjacent to the wall and the growth is opposite to the thermal gradient (typical of Na modified Al-Si alloys [7]). Figure lc schematizes heterogeneous nucleation of eutectic on nucleant particles present in

815

the interdendritic liquid (typical of Sr modified Al-Si alloys [7]). It is generally believed that Sr additions deactivate AIP particles that can act as nucleants for eutectic Si [8]. Accordingly, Sr containing Al-Si alloys exhibit a larger eutectic undercooling compared to unmodified alloys, resulting in the nucleation and growth of a few but large eutectic grains [5]. Experimental method Based on our knowledge of known nucleants and lattice parameters, a number of potential nucleating particles for eutectic Si have been suggested [9]. In this study, CrB particles have been chosen and a CrB-bearing master alloy has been manufactured by LSM. The particles contained in this master alloy are thought to promote eutectic Si nucleation. To verify this assumption, thermal analysis combined with quenched experiments, as well as porosity and tensile test experiments have been performed on three different alloys: (i) a base alloy consisting of commercial Al-10wt%Si, (ii) the base alloy modified with 300 ppm Sr (called Sr modified) and (iii) the base alloy modified with 300 ppm Sr and CrB particles (called Sr + CrB modified alloy). The base alloy was melted in a resistance furnace held at 640 °C. After skimming, 300 ppm of Sr using Al-10wt%Sr master alloy was added to the melt together with the CrB-bearing master alloy in the case of the Sr + CrB modified alloy. After 10 min, the melt was stirred manually and after further 10 min waiting, the following experiments were performed. It must be noted that degassing of the melt was never performed in this study. For thermal analysis and quenched experiments, the identical experimental set-up as developed by [6] was used in this study (see Fig. 2a). To sum up, a 640 °C preheated stainless steel cup coated with BN was plunged into the melt of desired composition and then put between two insulating boards. A type N thermocouple connected to an acquisition software recorded the cooling curve of the solidifying sample. A typical cooling rate of 1.5 K/s was achieved just prior to nucleation of the first solid (see Fig. 2b) and the sample was quenched in water at typically halfway through the eutectic arrest (note that the base alloy was not quenched in order to know the extent of the eutectic arrest). The quenched samples were sectioned vertically along the thermocouple line, and after conventional grinding and polishing up to 0.4 micron, they were observed by optical and electron microscopy.

To test the ability of the master alloy to ease interdendritic feeding during eutectic solidification and thus reduce porosity, the previously described alloys were cast in a steel permanent mold that had originally been developed to reproduce the thermal conditions at a spoke-wheel junction of a commercial automotive wheel [10]. The die consisted of three parts, which generated a cast shape with an undifferentiated down sprue and runner feeding perpendicularly into the face of a vertical plate. The junction between the runner and the plate was designed to produce a thermal hot spot as shown in Fig. 3. Large amount of porosity is expected in the plate if interdendritic feeding through the junction is poor and vice versa. The mold was coated with BN and preheated at 400 °C and after casting, each sample was photographed to document its external appearance, with particular focus on the surface of the plate. Finally, in order to investigate the ability of the master alloy to enhance the mechanical properties of cast specimens, ductility in particular, the three considered alloys were cast in a 300 °C preheated stainless steel mold covered with BN. The resulting casting rods (20 mm diameter, 200 mm long) where machined to produce ASTM-E8M standard tensile test specimens. The tensile tests were performed using an Instron 4505 tensile testing machine at a strain rate of 1 0 - 3 . No thermal treatment was performed and the samples were tested as cast.

*Tsski 3 a. m

Tf

«Oî

M

m

Hot spot

V

HE

Ww \/'

Figure 3: Key dimensions of the assembled "elbow shape" mold in mm, and dimensions of the experimental castings [10].

Results and discussion It has often been reported that Sr modification is accompanied by a reduced eutectic growth temperature [7]. In this study, the eutectic growth temperature for the base alloy, Sr modified, and Sr + CrB modi-

816

610

Insulating board (40mm x 40 mm)

TypeN . thermocouple f (tip 15 mm from base)

—C^Al-10wt%Si —»—+300ppmSr A + 300ppm Sr + CrB-bearing master alloy

-U

600Ü

diam

590 1H

^ ^ v

580

diam

570

560

H

100

L . 200

300

Time [s]

(b)

(a)

Figure 2: (a) A schematic (cut-away and exploded for clarity) of the experimental set-up for thermal analysis and (b) cooling curves of the three considered alloys. fied alloys was observed to be at 576.6 °C, 575.7 °C, and 576.3 °C, respectively. These small temperature differences (about 1 °C) are within the typical measurement errors and a eutectic growth temperature of 576 ± 1 °C can be considered in each case. However, whereas the base alloy was obviously unmodified, the two other ones were fully modified (see Figs. 4c and d). This shows that a fully modified eutectic structure is not obligatory accompanied with a reduced eutectic growth temperature. Figures 4a and b show the macrographs of the Sr modified and Sr + CrB modified alloys. Whereas the quench was performed approximately at the same time (see Fig. 2b), the alloy containing the CrB-bearing master alloy contains more and smaller eutectic grains, which shows the ability of the CrB-bearing master alloy to promote eutectic Si nucleation. It has also been often reported that Sr modified alloys exhibit a larger recalescence prior to eutectic growth compared to unmodified alloys. Poisoning of AIP particles by Sr is one of the suggested reasons to explained this increased recalescence. Since the AIP particles are deactivated, a larger undercooling is required so that other particles/nucleants become active. In this study, both Sr and Sr + CrB modified alloys exhibit a much larger recalescence prior to eutectic growth (approximately 4°C and 5°C, respectively) than the unmodified one (about 1°C). This shows that the particles contained in the CrB-bearing master alloys are not as effective for eutectic Si nucleation in Sr modified alloys as AIP particles are in un-

modified alloys, even if the number of eutectic grains is larger in the Sr 4- CrB modified alloy compared to the Sr modified one. However, the increased number of eutectic grains in the Sr + CrB modified alloy compared to the Sr modified alloy seems to have a significant influence on interdendritic feeding. Figure 5 shows the surface of the castings performed using the "elbow shape" mold described in Fig. 3. The amount of surface porosity of the Sr modified alloy (Fig. 5a) is clearly much larger than that of the Sr + CrB modified alloy (Fig. 5b). Such a large porosity difference is very unlikely due to (i) a difference in solid-liquid interfacial energy or (ii) an increased susceptibility of the melt to absorb hydrogen due to Sr modification, since both alloys have the same Sr content. Also, since both samples were cast following the same procedure (and quite roughly moreover), it is believed that the amount of entrapped oxide bifilms is more or less similar in both samples. However, it must be noted that the CrB-bearing master alloy acts as inoculant for primary Al as well (T1B2 particles are likely to be formed, since both Ti and B are present in the master alloy). This can also be seen on the cooling curves (see Fig. 2b), where nucleation of primary Al happens at a lower undercooling for the Sr + CrB modified alloy compared to the other alloys. However even if it is well known that primary Al grain refinement is standard practice to get (among others) uniformly distributed microporosity in castings [11], it cannot explain such a large difference in the amount

817

/

cutecì 1C ecu

(c)

Figure 4: Optical macrograph of a quenched (a) Sr modified and (b) Sr + CrB modified Al-10wt%Si alloy. The quenched liquid is in light grey, whereas the dark grey circles represent the solidified eutectic cells. The bottom pictures show a magnified view of (c) the Sr modified and (d) the Sr + CrB modified alloys. Adding the CrB-bearing master alloy clearly increases the number of eutectic cells and reduces their size. of porosity, since the eutectic fraction in these alloys can reach 55% (using a Scheil approximation). Hence, since only the amount of eutectic grains is significantly changed between both experiments, these results show that the influence of a eutectic grain refiner is critical for interdendritic feeding, and hence for porosity as well.

liminary results that the Sr modified alloy is more ductile than the base alloy, whereas the Sr + CrB modified alloy is even more ductile with a maximum strain of about 3.5% (which is fairly high for an ascast Al alloy). Each of the stress strain-curve exhibits a round shape, which is characteristic of a composite. This is no surprise since the eutectic fraction in these alloys approaches 55%. It was checked that the A reduced porosity should lead to an increased secondary dendrite arm spacing, ΐ2, was similar beductility in a cast specimen. Figure 6 shows the tween all three alloys (X2 ~ 22 ± 2μιη). However, true stress-strain curves of the samples cast using the characteristic length of the coarse plate-like euthe base alloy (open squares), the Sr modified alloy tectic microstructure of the unmodified alloy was at (crossed-open circles), and the Sr -f CrB modified least 3 times larger than the refined fibrous eutectic alloy (black triangles). It is clear from these pre-

818

ana» "

I1

,

·

.

·

.

.

-


S ·..·

y '

:,'

•

; C " V . ',.'· '' ' - · *

• ' ·. *.

;

l •V w v „ .Λ *v' *. « ·« ·

. ' ·' .

^

w

%^'-.'::--

'-_-·.

:

. - -:-^- v::::

'"

:i

i \

1 (b)

(a)

Figure 5: Photographs of the external planar surface of the castings produced with the "elbow shape" mold (see Fig. 3) using (a) the Sr modified alloy and (b) the Sr 4- CrB modified alloy. A better surface finish and a reduced porosity is clearly obtained using the latter. Al grain refinement since (i) the eutectic fraction is about 55% and (ii) the load at high strain is supported almost only by the strong percolated Al+Si eutectic. r-, 120We have then shown that ductility of Al-Si cast eu alloys can be improved with small additions of Sr and 80 the help of a eutectic grain refiner. The latter allows a finer and more distributed porosity in Sr modified -D-Al-10wt%Si —8— +300ppmSr alloys, because the interdendritic liquid can flow more H 40A + 300ppm Sr + CrB-bearing master alloy easily to compensate for solidification shrinkage if the eutectic cells are smaller. 0This brings a new aspect for the simulation of _" 1 ' 1 1 porosity formation in aluminium alloys. This is of 0.06 0.00 0.02 0.04 course beyond the scope of this paper but in addition True Strain [-] to the conventional phenomena that need to be modeled (see [12] for the details), our results show that Figure 6: Influence of Sr and CrB-bearing master althe fluid flow calculations should be coupled with miloy additions on mechanical properties of Al-10wt%Si crostructure simulations, since the interdendritic liqalloys. uid flow depends on the density/type of eutectic cells. 160-

Ý ==

ιο-3 [-]

1 -

T

1

structure of the modified alloys (which had both a eutectic spacing of about 1 μιη). Accordingly, the increase in ductility between the base alloy and the Sr modified alloy is most probably due to the refinement of the eutectic structure, despite of the huge porosity increase in the Sr modified alloy (see Figs. 7a and b). On the other hand, the increase in ductility between the Sr modified and Sr + CrB modified alloy is most likely due to the difference in porosity between both alloys (more distributed and finer pores are found in Fig. 7c than in Fig. 7b), and not due to primary

Conclusion Modification of Al-Si alloys with Sr causes a dramatic reduction of the nucleation frequency compared to unmodified alloys. This, in turn, increases drastically the amount of porosity because of the inhibition of interdendritic liquid due to the equiaxed growth of large eutectic grains, i.e. both feeding and permeability are reduced. In this study, we have shown that by increasing the number of eutectic cells using a eutectic grain refiner, it is possible to reduce poros-

819

(a)

(b)

(c)

Figure 7: SEM images showing the fracture surface of (a) the base alloy, (b) the Sr modified alloy, and (c) the Sr + CrB modified alloy. Development in Al-Si Alloys," Maternais Science and Technology, 26 (2010), 262-268.

ity and improve significantly the ductility of Sr modified commercial Al-10wt%Si alloy. Although more experiments and characterisation are needed to verify these promising preliminary results, the key concept of eutectic grain refinement is thought to be a major advance in the development of Al-Si alloys.

[5] A.K. Dahle et al., "Eutectic Nucleation and Growth in Hypoeutectic Al-Si Alloys at Different Strontium Levels," Metallurgical Materials Transactions A, 32 (2001), 949-960. [6] S.D. McDonald, K. Nogita, and A.K. Dahle, "Eutectic Nucleation in Al-Si Alloys," Acta Materialia, 52 (2004), 4273-4280.

Acknowledgements CAST CRC, established under and funded in part by the Australian Federal Governments' Cooperative Research Centres scheme, is acknowledged for its financial support as well as London & Scandinavian Metallurgical Co Limited (LSM) for providing the CrB-bearing master alloy.

[7] A.K. Dahle et al., "Eutectic Modification and Microstructure Development in Al-Si Alloys," Materials Science and Engineering A, 413-414 (2005), 243-248. [8] Y.H. Cho et al., "Effect of Sr and P on Eutectic Al-Si Nucleation and Formation of /3-A15FeSi in Hypoeutectic Al-Si Foundry Alloys," Metallurgical Materials Transactions A, 39A (2008), 2435-2448.

References [1] Q.G. Wang, "Microstructural Effects on the Tensile and Fracture Behavior of Aluminum Casting Alloys A356/357," Metallurgical Maternais Transactions A, 34 (2003), 2887-2899. [2] D. Emadi, J.E. Gruzleski, and J.M. Toguri, "The Effect of Na and Sr Modification on Surface Tension and Volumetric Shrinkage of A356 Alloy and their Influence on Porosity Formation," Metallurgical Materials Transactions B, 24 (1993), 1055-1063. [3] John Campbell, Castings, (Oxford, United Kingdom: Elsevier Butterworth-Heinemann, 2003), 17-31. [4] J. Campbell and M. Tiryakioglu, "Review of Effect of P and Sr on Modification and Porosity

[9] A.K. Dahle et al., US patent 2009/0297394 Al, 2009. [10] M. Easton, "Grain Refinement Mechanisms in Aluminium and its Alloys and the Effect of Grain Refinement on Cast ability," (Ph.D. thesis, The University of Queensland, 1999), 174. [11] L. Lu and A.K. Dahle, "Effects of Combined Additions of Sr and AlTiB grain refiners in Hypoeutectic Al-Si Foundry Alloys," Materials Science and Engineering A, 435 (2006), 288-296. [12] J.A. Dantzig and M. Rappaz, Solidification (Lausanne, Switzerland: EPFL Press, 2009), 479-514.

820

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Production of Al-Ti-C Grain Refiners with the Addition of Elemental Carbon and K2TiF6 Fatih Toptan, Isil Kerti, Sibel Daglilar, Ahmet S agin, Omer F. Karadeniz, Ay sin Ambarkutuk Yildiz Technical University, Department of Metallurgical and Materials Engineering, Faculty of Chemistry & Metallurgy, Davutpasa Campus, 34210, Esenler, Istanbul, Turkey Keywords: Aluminum; Al-Ti-C Grain Refiners; Elemental Carbon; K2TiF6 Abstract

Experimental Procedure

Al-Ti-B grain refiners are widely used as aluminum grain refiners. Despite the problems in application, Al-Ti-C refiners have an increasing demand in recent years. In the present work, Al-Ti-C grain refiners with different Ti:C ratios were produced by in-situ method with the addition of elemental carbon into the AlTi master alloy (master alloy method) and the addition of K2TiF6 and elemental carbon powder mixture into the commercially pure aluminum (flux method). The effects of production method and Ti:C ratio on the grain refinement process were investigated using Alcoa Cold finger Test and so optimum conditions were determined. The best performance was accomplished with the refiner produced at 1300°C with the master alloy method.

Commercial pure aluminum (AA 1080), Al-10%Ti master alloy, K2TiF6, Elftex 125 Cabot Carbon with average particle size of 1020μπι were used to produce Al-Ti-C grain refiners. Chemical compositions of AA1080 and Al-10%Ti alloy are given in Table I. Experiment sets were prepared by using different carbon ratios ranges changing from 0% to 0.90% as given in Table II and Table III. For each sample Ti ratio was fixed to 3%. Table I. Chemical Compositions of AA 1080 and Al-10%Ti Master Alloy Al

Fι

Si

Cu

Ti

Zn

Bai

(%)

(%)

(%)

(%)

(%)

(%)

(%)

99.80

0.15

0.10

0.03

0.02

0.03

0.10

Alloy

Introduction

AA1080

Grain refinement has been an important technique in aluminum industry for many years to improve the soundness of aluminum products [1], It is well known that a fine equiaxial grain structure can improve the mechanical properties, it reduces the size of defects such as micro porosity and second-phase particles [1,2] The most widely used grain refiners are based on Al-Ti-B which has a-aluminum matrix, Al3Ti and TiB2 particles[3]. On the other hand, the TiB2 particles in Al-Ti-B master alloys lead to a number of problems such as streaking and porosity in foils, scratch-like linear surface, internal cracking in extrusion billets[4], suffering from poisoning in the presence of Zr, V and Cr[3,5]. In order to prevent such as harmful effects of Al-Ti-B, alternative grain refiners have been investigated [6]. The effectiveness of Al-Ti-C master alloys in the grain refinement of aluminum and its alloys have been widely investigated for many years [7]. In recent years, they have been of increasing importance in aluminum casting because, unlike conventional Al-Ti-B master alloys [8,9], they are believed to introduce a smaller volume fraction of insoluble particles into the melt. Al-Ti-C refiner are composed of an aaluminum matrix, Al3Ti and TiC particles. The existence of excess titanium in Al-Ti-C master alloys plays an important role on the refinement efficiency of TiC particles [10]. Al-Ti-C master alloys have been prepared by reaction of carbon with Al-Ti master alloy [11], melting the elemental blends, self propagating high temperature synthesis and reaction of molten Al with K2TiF6 and graphite powder [11,12]

Al10%Ti

Ti

Fι

Si

Al

(%)

(%)

(%)

(%)

9.947

0.179

0.126

Rest

Table II. Grain Refiner Compositions and Experimental Parameters for the Flux Method Carbon and K2TiF6 addition temperature (°C) Exp. Exp. 1 2 900 800

Exp. 1 900

Exp. 2 1300

Exp. 1 20

Exp. 2 20

Al-3%Ti-0.15%C

900

800

900

1300

20

20

Al-3%Ti-0.30%C

900

800

900

1300

20

20

Al-3%Ti-0.45%C

900

800

900

1300

20

20

Al-3%Ti-0.60%C

900

800

900

1300

20

20

Al-3%Ti-0.75%C

900

800

900

1300

20

20

Al-3%Ti-0.90%C

900

800

900

1300

20

20

Grain refiner composition Al-3%Ti

In the present work, Al-Ti-C grain refiners were produced by master alloy method and flux method to compare the performance of refiners. Production parameters such as holding time, different Ti/C ratio and casting temperature were determined for most effective grain refiner.

821

Holding and casting temperature CO

Holding time (min)

Table III. Grain Refiner Compositions and Experimental Parameters for the Master Alloy Method Grain refiner composition

Addition and casting temperature (°C)

Holding time (min)

Al - %3Ti Al-%3Ti-%0.15C Al - %3Ti - %0.30C Al - %3Ti - %0.45C Al - %3Ti - %0.60C Al - %3Ti - %0.75C Al - %3Ti - %0.90C

1300 1300 1300 1300 1300 1300 1300

20 20 20 20 20 20 20

performed by mechanical polishing. Finally, these samples were etched by HF 0.5%. The grain size of each sample was studied to determine the most effective grain refiner. Results and Discussion The best performance was accomplished with the Al-3%Ti alloys prepared by the flux method. The grain refiners produced by the flux method did not give stable results. On the other hand, the results which were obtained by the master alloy method were quite stable. Figure 1 shows all the results for comparison.

Master Al cy Me:hcd

For both methods, graphite crucible and stainless steel rod was coated with boron nitride to ensure that molten metal is purified from impurities. Graphite crucible, mold and foil wrapped carbon powder were preheated at 200°C for 1 hour.

Flux Method

In the master alloy method commercial pure aluminum and Al%10Ti were melted together up to 1300°C. The melt was held at constant temperature for 15 minutes and after this period, carbon was added immediately by light stirring. The stirring was maintained for 5 minutes to obtain homogenous solution (stirring process was skipped for the refiner without elemental carbon (Al3%Ti)). Subsequently, the melt was held at the same temperature (1300°C) for 20 minutes and then poured into the mold which had been preheated to 200°C. The aim of this holding process was to form TiC particles.

0.15

0.30 0.45 0.60 Carbon content(%wt.)

0.75

0.90

Figure 1. Grain size variation versus carbon content for different production method.

Flux method (addition of K2TiF6 and carbon) was investigated for two different techniques which were mentioned as Experimental 1 and Experimental 2 in Table II. For Experimental 1, carbon and K2TiF6 were mixed for 2 hours then powder mixture was preheated at 200°C for 1 hour for humidity elimination in powder mixture. In this method commercial pure aluminum was melted at 900°C, and powder mixture of carbon and K2TiF6 were added to the melt by light stirring. The temperature was held constant at 900°C during adding process. When the stirring process was completed, melt was held for 20 minutes at 900°C and casting was made into the preheated mold. For Experimental 2, the melting and addition temperature of powder mixture in melt was chosen as 800°C as a difference from Experimental 1. After addition of powder mixture, the melt temperature was increased to 1300°C and held stable at this temperature for 20 minutes before casting.

The micrographs shown in Figures 2-8, allow us to compare the effect of the grain refiners used in this study. Structural variables belonging to samples that were produced by both of the elemental carbon and flux methods can also be seen in these figures.

1(b)

f iomm

10 mm

1

In grain refiner performance tests for both methods, commercial pure aluminum was melted in a boron nitride coated graphite crucible by induction furnace. Melt temperature was adjusted to 750°C throughout the experiment. Grain refiner additions were made with 0.2% weight of commercial pure aluminum and inoculated into the melt when desired temperature was reached and then was stirred for 2 minutes with boron nitride coated graphite rod. After stirring, melt was held for 3 minutes then casting was made into a plaster mold which was preheated to 750±5°C as indicated Alcoa Cold Finger (ACF) test procedure[13].

10. mm i" '

The samples which were prepared longitudinally by cutting into half were surface-machined. This surface was macro etched using Poulton's reagent. To measure the grain size, a sample was cut from 25.4 mm distance below the copper chill [14]. Prior to each experiment, the surface pretreatment of specimens were

(II

n

It) Hlpl

Λ

niiinrnmnimi

Figure 2. (a) Non refined AA1080 alloy, AA1080 including Al3%Ti refiner produced at (b) 900°C, (c) 1300°C with the flux method, (d) 1300°C with the master alloy method

822

Figure 3. (a) Non refined AA1080 alloy, AA1080 including Al-3%Ti-0.15%C refiner produced at (b) 900°C, (c) 1300°C with the flux method, (d) 1300°C with the master alloy method

Figure 5. (a) Non refined AA1080 alloy, AA1080 including Al3%Ti-0.45%C refiner produced at (b) 900°C, (c) 1300°C with the flux method, (d) 1300°C with the master alloy method

Figure 4. (a) Non refined AA1080 alloy, AA1080 including Al3%Ti-0.30%C refiner produced at (b) 900°C, (c) 1300°C with the flux method, (d) 1300°C with the master alloy method

Figure 6. (a) Non refined AA1080 alloy, AA1080 including Al3%Ti-0.60%C refiner produced at (b) 900°C, (c) 1300°C with the flux method, (d) 1300°C with the master alloy method

823

et al.[6] have implied in their study that in the grain refiner samples produced at 900°C, had blocky type Al3Ti morphology which is the result of low production and casting temperatures. This morphology changes to needle-like at high production temperatures as stated in [6]. The change of the morphology is related with dissolving behavior of the Al3Ti particles with increasing temperatures. In Figures 2-8 performance of all refiners are compared. Because of the lower amounts of Al3Ti phase and by the absence of TiC particles, the refiners produced at 900°C had a poor refining performance. In the master alloy method, reduction of grain size is detected when the Ti:C ratio of the grain refiner is increased above the TiC stoichiometry. When the refiners produced in 1300°C is examined, the grain refiner becomes more efficient when the carbon amount increases. Increasing TiC amount can be the main reason for this situation. However, at grain refiner production small inefficiencies occurred during carbon addition. Conclusions In this study, Al-Ti-C grain refiners were produced by the master alloy method and the flux method. The grain refiner performances were compared. By these methods, effective grain refinement production is possible. Conclusions are as below: Figure 7. (a) Non refined AA1080 alloy, AA1080 including Al3%Ti-0.75%C refiner produced at (b) 900°C, (c) 1300°C with flux method, (d) 1300°C with the master alloy method

• • • •

• •

It is possible to produce Al-Ti-C grain refiners with the elemental carbon and flux methods, Grain refiners produced by the master alloy method had good performance when compared to flux method, AI3T1 and TiC were detected as effective phases in grain refinement process, Refiners produced with flux method at 1300°C had relatively good performance when compared to 900°C productions by the presence of high amounts of Al3Ti and TiC phases, In Al-Ti-C refiners produced with master alloy method at 1300°C, grain size decreased as the carbon content increased, Al-3%Ti-0.60%C, Al-3%Ti-0.75%C and Al-3%Ti0.90%C compositions produced with master alloy method at 1300°C were determined as the most effective grain refiners.

References 1.

2. Figure 8. (a) Non refined AA1080 alloy, AA1080 including Al3%Ti-0.90%C refiner produced at (b) 900°C, (c) 1300°C with flux method, (d) 1300°C with the master alloy method

3.

The grain refiners produced by flux method at 900°C were not effective in reducing grain size. On the other hand, the grain refiners produced at 1300°C with the same method were found effective, but the grain reduction intensity was not dependant on the carbon content. Figure 1 implies this result. The production temperatures directly affect the grain refiner performance. Gezer

4. 5.

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Z. Zhang, X. Bian, Z. Wang, X. Liu, and Y. Wang, "Microstructures and Grain Refinement Performance of Rapidly Solidified Al-Ti-C Master Alloys," J. Alloys Compd., 339 (2002), 180-188. K. Jiang, X. Liu, "The Effect of Melting Temperature and Time on the TiC Particles," J. Alloys Compd., 484 (2009), 95-101. I. Naglic, A. Smolej, M. Dobersek, and P. Mrvar, "The Influence of TiB2 Particles on the Effectiveness of A1-3TÌ0.15C Grain Refiner," Mater. Charact., 59 (2008), 14581465. X. Ma, X. Liu, and H. Ding, "A United Refinement Technology for Commercial Pure Al by Al-lOTi and Al-TiC Master Alloys," J. Alloys Compd., 471 (2009), 56-59. Y. Birol, "Grain Refining Efficiency of Al-Ti-C Alloys," J. Alloys Compd., 422 (2006), 128-131.

6. 7. 8. 9.

10. 11.

12.

13.

14.

B.T. Gezer, F. Toptan, S. Daglilar, and I. Kerti, "Production of Al-Ti-C Grain Refiners with the Addition of Elemental Carbon,"Mater. Design, 31 (2009), 30-35. A.Cibula, J. Inst. Met., 76 (1951-1952) 321-60 W. Schneider, "Grain Refinement of Al Wrought Alloys with Newly Developed AlTiC Master Alloys," Z. Metallkd./Mater. Res. Adv. Technol., 91 (2000), 800-06. AJ. Whitehead , P.S. Cooper, and R.W. McCarthy, "An Evaluation of Metal Cleanliness and Grain Refinement of 5182 Aluminum Alloy DC Cast Ingot Using Al-3%Ti0.15%C and Al-3%Ti-l%B grain refiners," Light Metals: Proc. Sessions. TMS Annual Meeting, TMS, Warrendate, PA, (1999), 763-72. Z.Q. Wang, X.F. Liu, and X.F. Bian, "Microstructure and Its Influence on Refining Performance of AlTiC Master Alloys," Mater. Sci. Technol., 19 (2003), 1709-1714. B.Q. Zhang, H.S. Fang, L. Lu, M.O. Lai, H.T. Ma, and J.G. Li, "Synthesis Mechanism of an Al-Ti-C Grain Refiner Master Alloy Prepared by a New Method," Metall. Mater. Trans., 34A (2003), 1727-1733. G.S.V. Kumar, B.S. Murty, and M. Chakraborty," Development of Al-Ti-C Grain Refiners and Study of Their Grain Refining Efficiency on Al and A1-7SÌ Alloy," J. Alloys Compd., 396 (2005), 143-150. B. S. Murty, S. A. Kori, and M. Chakraborty, "Grain refinement of aluminium and its alloys by heterogeneous nucleation and alloying, International Materials Reviews 2002 Vol. 47 No. 1 P. Moldovan, G. Popescu," The Grain Refinement of 6063 Aluminum Using A1-5TÌ-1B and A1-3TÌ-0.15C Grain Refiners,"Abi/Inform Tarade Industry Research Summary, 56 (2004), 59.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECT OF MECHANICAL VIBRATIONS ON MICROSTRUCTURE REFINEMENT OF Al-7 MASS% Si ALLOYS Takuya Tamura1, Toshiro Matsuki2 and Kenji Miwa1 Materials Research Institute for Sustainable Development, National Institute of Advanced Industrial Science and Technology (AIST), Nagoya 463-8560, Japan. 2 Yamagata Research Institute of Technology, Yamagata 990-2473, Japan Keywords: Mechanical vibration, Microstructure, Refinement, Solidification solidification in order to refine the microstructure. The authors reported that the velocity of mechanical vibrations was found to be important factor for primary crystals refinement [10]. It is known that vibration energy of the simple harmonic oscillation (sine wave vibration) can be described as follows.

Abstract Mechanical vibration treatment is known to induce microstructure refinement. However, it is not completely understood which factor of vibrations is important for microstructure refinement. Factors of vibrations include frequency, acceleration, velocity and amplitude. In our study, it was found that velocity of mechanical vibrations, namely, the energy of vibrations is important factor for primary crystals refinement. This study aims to investigate effects of casting conditions with the mechanical vibrations on microstructure refinement of Al-7mass%Si alloys. Free space area for vibrating the melt affects the microstructure refinement. Moreover, primary crystal particles become rosette-like and fine when the mechanical vibrations are applied to the melt from about 923 K to 888 K (liquidus - 2 K). Thus, it was found that the mechanical vibrations promote heterogeneous nucleation just under the liquidus temperature.

-mV2 2

(3)

where E (J) is vibration energy, m (kg) is mass of vibration object and V (m/s(0-p)) is velocity of the vibration wave. Thus, the square of the velocity corresponds to the vibration energy. Thus, it was found that the energy of mechanical vibrations promotes microstructure refinement during solidification. In this study, the effects of casting conditions with the mechanical vibrations on microstructure refinement of Al-7 mass% Si alloys were studied. Experimental Procedures

Introduction

Hypo-eutectic Al-7 mass% Si alloys were prepared by using pure aluminum (99.9 %) and Al-24.3 mass% Si master alloys. From cooling curves, the eutectic and liquidus temperatures were found to be 850 K and 890 K, respectively. Iron was found to be contained about 0.5 mass% by ICP analysis. Specimens with weight of about 26-29 g were obtained from the prepared Al-7 mass% Si alloys. The specimen was inserted to an alumina crucible (30 mm in outer diameter, 24 mm in inner diameter and 80 mm in height), and the alumina crucible was sealed with a ceramic cap, as shown in Fig. 1.

Al-Si alloy system is the most common system among the casting aluminum alloys. Controlling the structure of the alloys in the casting is an important factor about the foundry industries. Especially microstructure refinement is one of the most important factors. Thus, various methods have been developed for microstructure refinement. Among these methods, there is application of vibrations such as electromagnetic vibrations [1, 2], ultrasonic waves [3] and mechanical vibrations [4-7]. Especially, application of the mechanical vibrations attracts attention because of a simple system [8, 9]. Most researches about the effect of the vibrations mentioned that the size of primary dendrite particles was decreased by the vibrations imposed during solidification and finally globular particles were formed with increase in vibration intensity. However, it is not completely understood which factor of vibrations is important for microstructure refinement. Factors of vibrations are frequency, acceleration, velocity and amplitude. In the sine wave, two equations hold. V = 2nfD

(1)

A = 2n/V

(2)

Thermocouple

K

Alumina crucible

where D (m(0-p)) is amplitude (0-to-peak value), V (m/s(0-p)) is velocity (0-to-peak value), A (m/s2(0-p)) is acceleration (0-topeak value) and f (Hz) is frequency of the vibration wave. From these two equations, it is found that when two factors are decided, other two factors are decided automatically. The frequency has the different character from other factors, because other factors have waveform. Thus, the frequency and one of other three factors (acceleration, velocity and amplitude) have to be considered as the vibration parameters when the vibrations are applied during

Vibration Figure 1. Schematic illustration of the alumina crucible where the molten specimen was put in. Temperature of the specimen was measured by a K-type thermocouple that was inserted 65 mm from the top of the crucible to the side of the specimen. The crucible which was set as

827

shown in Fig. 1 was inserted into an electric furnace. The specimen was heated up to 923 K and held for 5 min. Then, the crucible where the molten specimen was put in was taken out from the furnace, and clamped by an air cylinder to a vibration testing machine. This vibration testing machine can impose sine wave vibration with given frequency and given acceleration. The vibration direction is shown in Fig. 1. The mechanical vibration was applied to the melt in the crucible from about 923K to 863K, and the crucible was quenched in water at 863K. Values of the acceleration and the velocity were described by 0-to-peak value. However, values of the amplitude were described by general peak-to-peak value. Microstructure of the solidified specimen was observed by optical microscopy after etching.

Schematic illustration of the alumina crucible shown in Fig. 1 has the free space area for the melt with the mechanical vibrations because the alumina crucible was sealed with the ceramic cap. However for schematic illustration of the alumina crucible shown in Fig. 3, there is no free space area for the melt with the mechanical vibrations because the melt was locked by a carbon block which wasfixedto the alumina crucible. Thermocouple

Alumina^ crucible

Results and discussion Effect of free space area for vibrations on microstructure Figure 2 shows microstructures in the specimens solidified during imposition of the mechanical vibrations with samefrequency.The frequency is 50 Hz. The specimen for 87.8 mm/s shown in (a) consists of large dendritic primary particles. Size of the primary particles was decreased with increase in the velocity. And the specimen for 702 mm/s shown in (d) consists of small rosette-like primary particles. Moreover, it was found that the microstructures didn't change even iffrequencyvaried 8 times [10]. However, the microstructure refinement by the vibrations with high amplitude or high acceleration was affected by the frequency [10]. Thus, it was found that the velocity of mechanical vibrations, namely, the energy of vibrations is important factor for primary crystals refinement.

Melt

s^rr^L

Vibration Figure 3. Schematic illustration of the alumina crucible. The melt is locked by a carbon block which wasfixedto the crucible. Figure 4 shows microstructures in the specimens solidified during imposition of the mechanical vibrations. The frequency and the velocity are 50 Hz and 702 mm/s, respectively. The specimen shown in Fig. 4(a) was solidified in the alumina crucible shown in Fig. 1. And the specimen shown in Fig. 4(b) was solidified in the alumina crucible shown in Fig. 3. The specimen solidified during imposition of the mechanical vibrations with free space area shown in Fig. 4(a) consists of small rosette-like primary particles. However, the specimen solidified during imposition of the mechanical vibrations without free space area shown in Fig. 4(b) consists of dendritic primary particles that are similar to those shown in Fig. 2 (b). The specimen shown in Fig. 2 (b) was solidified during imposition of the mechanical vibrations with weak velocity in the alumina crucible shown in Fig. 1. Thus, it was found that the effect of the mechanical vibrations on the microstructure weakens when the melt cannot vibrate freely. From these results, it can be presumed that the mechanical vibrations with higher energy are needed in order to refine the microstructures when a permanent mold with risers is used.

SOOjim Figure 2. Microstructures in the specimens solidified during imposition of the vibrations with samefrequency50 Hz. Figure 4. Microstructures in the specimens solidified during imposition of the mechanical vibrations (50 Hz, 702 mm/s); (a): in the alumina crucible shown in Fig. 1; (b): in the alumina crucible shown in Fig. 3.

However, casting conditions which affect the microstructure refinement by the mechanical vibrations exist with the exception of these 4 vibration factors. For example, free space area for the melt with the mechanical vibrations is. When a permanent mold with risers is used, it is energy-saving that important part of the mold vibrates by a mechanical vibrator because permanent molds are heavy [8, 9]. In this situation, the melt movement is limited. Thus, effects of the free space area for the melt with the mechanical vibrations on the microstructure were studied.

Effect of imposition temperature region on microstructure It is important casting condition that the primary particles were refined in a certain temperature range. Thus, effects of the end

828

temperature of the mechanical vibrations on the microstructure were studied. However, a cooling rate in the liquid state for the specimen solidified during imposition of the mechanical vibrations (50 Hz, 702 mm/s) shown in Fig. 2 (d) is about -3 K/s. This cooling rate was too fast to control the end temperature of the mechanical vibrations accurately. Thus, the alumina crucible shown in Fig. 1 was covered with heat-insulating ceramic boards in order to decrease the cooling rate. The cooling rate in the liquid state for the specimen solidified during imposition of the mechanical vibrations (40 Hz, 702 mm/s) became about -0.2 K/s by heat-insulating ceramic boards. Figure 5 shows cooling curves for the specimens solidified during imposition of the mechanical vibrations with the frequency of 40 Hz and the velocity of 702 mm/s. The end temperature of the mechanical vibrations was varied. The specimen without the mechanical vibrations and the specimen on which the mechanical vibrations were imposed up to 892 K (liquidus + 2 K) showed the supercooling. However, the specimen on which the mechanical vibrations were imposed up to 863 K and the specimen on which the mechanical vibrations were imposed up to 888 K (liquidus - 2 K) didn't show the supercooling.

\ 895

F\ hSuper r cooling

Q.

E

Ë-.*

,

2

imposed in the semisolid region. Thus, it is considered that the rosette-like primary particles grew larger before water quenching at 863 K for the specimen on which the mechanical vibrations were imposed up to 892 K (liquidus - 2 K) shown in (d). As a result, important temperature region was found to be just under the liquidus temperature for primary crystals refinement by the mechanical vibrations. When heterogeneous nucleation just under the liquidus temperature is promoted, the supercooling is considered to disappear. Thus, it was found that the mechanical vibrations promote heterogeneous nucleation just under the liquidus temperature.

I.

Freq.: 40Hz, Ace: é /orn/s , Vel.:702mm/s, Amp.: 5.59mm(p-p)|| No VibrationII To 863K To 892K To 888K ^ — r- Liquidu 3 \?Sk lSuperco«sling

Figure6. Microstructures which correspond to the cooling curves shown in Fig. 5. The end temperatures of the vibrations were; (a): no vibration, (b): 863 K, (c): 892 K, (d): 888 K.

Γ

885

Conclusions 880

i

50

i

i

i

1

.

The effects of casting conditions with the mechanical vibrations on microstructure refinement of Al-7 mass% Si alloys have been investigated, and the following conclusions have been derived. It was found that free space area for vibrating the melt affects the microstructure refinement. The effect of the mechanical vibrations on the microstructure weakens when the melt cannot vibrate freely. The primary crystal particles become rosette-like and fine when the mechanical vibrations are applied to the melt from about 923 K to 888K (liquidus - 2K). Thus, it was found that the mechanical vibrations promote heterogeneous nucleation just under the liquidus temperature.

%

*

100

150

200

Time, f/s

Figure 5. Cooling curves for the specimens solidified during imposition of the mechanical vibrations (40 Hz, 702 mm/s). The end temperature of the mechanical vibrations was varied. Figure 6 shows microstructures in the specimens solidified during imposition of the mechanical vibrations with the frequency of 40 Hz and the velocity of 702 mm/s. The end temperature of the mechanical vibrations was varied and these microstructures correspond to the cooling curves shown in Fig. 5. The specimen shown in (a) was solidified without the mechanical vibrations. The specimen consists of large dendritic primary particles. Moreover, the specimen on which the mechanical vibrations were imposed up to 892 K (liquidus + 2 K) shown in (c) consists of large dendritic primary particles that are similar to those without the vibrations. These two specimens showed the supercooling. On the other hand, the specimen on which the mechanical vibrations were imposed up to 863 K shown in (b) consists of small rosettelike primary particles. Moreover, the specimen on which the mechanical vibrations were imposed up to 892 K (liquidus - 2 K) shown in (d) consists of small rosette-like primary particles. However, size of the primary particles was larger than that shown in (b). The cooling rate for the specimen on which the mechanical vibrations were imposed up to 863 K was faster than those for other specimens on which the mechanical vibrations were not

Acknowledgments The authors thank Mrs. T. Yamaguchi for technical assistance. References 1. A. Radjai and K. Miwa, Metall. Mater. Trans. A31 (2000), pp. 755-762. 2. Y. Mizutani, Y. Ohura, K. Miwa, K. Yasue, T. Tamura and Y. Sakaguchi, Mater. Trans. 45 (2004), pp. 1944-1948. 3. W. Khalifa, Y. Tsunekawa and M. Okumiya, Int. J. Cast. Met. Res. 21 (2008), pp. 129-134. 4. S. Wu, L. Xie, J. Zhao and H. Nakae, Scripta Materialia 58 (2008), pp. 556-559.

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5. F. Taghavi, H. Saghafian and Y.H.K. Kharrazi, Materials and Design 30 (2009), pp. 115-121. 6. F. Taghavi, H. Saghafian and Y.H.K. Kharrazi, Materials and Design 30 (2009), pp. 1604-1611. 7. C. Limmaneevichitr, S. Pongananpanya and J. Kajomchaiyakul, Materials and Design 30 (2009), pp. 3925-3930. 8. N. Omura, Y. Murakami, M. Li, T. Tamura, K. Miwa, H. Furukawa, M. Harada and M. Yokoi, Mater. Trans. 50 (2009), pp. 2578-2583. 9. N. Omura, Y. Murakami, M. Li, T. Tamura, K. Miwa, H. Furukawa, M. Harada, Mater. Trans. 50 (2009), pp. 2604-2608. 10. T. Tamura, T Matsuki, K Miwa, Proceedings of the 12th International Conference on Aluminium Alloys (2010), pp. 658662.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION

ORGANIZERS

Alan Tomsett Rio Tinto Alcan Brisbane, Australia Barry Sadler Net Carbon Consulting Pty. Ltd. Victoria, Australia

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION Anode Baking SESSION CHAIR

Said Al Maawali Sohar Aluminium Oman

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Predicting the Response of Aluminum Casting Alloys to Heat Treatment Chang-Kai Wu, Makhlouf M. Makhlouf Advanced Casting Research Center Worcester Polytechnic Institute (WPI); 100 Institute Rd., Worcester, MA 01609, USA Keywords: Heat treatment, Quenching Factor Analysis, Modeling yield strength and corrosion resistance of wrought aluminum alloys. Since then, the QFA method has proved to be a very useful tool in predicting properties of many cast and wrought aluminum alloys [2]. The QFA method is based on using isothermal precipitation kinetics to predict the results of non-isothermal conditions during continuous cooling. In doing so, it considers the cooling curve to be made up of a series of isothermal transformation steps and adds up the amount of second phase transformed during each of these isothermal steps in order to simulate the overall degree of super saturation of the alloy.

Abstract In this publication we report on the development of a mathematical model that enables predicting the changes in hardness of cast aluminum alloy components in response to heat treatment. This model is part of a more inclusive model that is currently under development and that when completed, will enable predicting the changes in room temperature tensile properties as a function of heat treatment. The model uses the commercially available finite element analysis software (ABAQUS) and an extensive database that was developed specifically for the aluminum alloy under consideration (namely, A356.2). The database includes mechanical, physical and thermal properties of the alloy all as functions of temperature. In addition, boundary conditions - in the form of the heat transfer coefficient associated with each one of the heat treatment steps are obtained from measurements performed with specially designed quenching devices. The database and boundary conditions are used in a thermal analysis module and a userdeveloped module. The user-developed module uses Quench Factor Analysis to predict the maximum attainable hardness that develops in a commercial cast component that is subjected to a standard commercial heat treating cycle. A heat-treated part was used to validate the model prediction.

initial Conditions

Physical and Thermai Properties

Boundary Conditions

•1

ABAQUS Thermal Modute Temperature Dependant Local Heat Transfer Coefficients

\ User-Deveioped Property Module" I Quench Factor ] Heat Treated Properties (T6orT4)

*Quenct» Factor Analysis

Figure 1. Description of the Model.

Assuming that the precipitation transformation follows the Johnson-Mehl-Avarmi-Kolmogorov (JMAK) equation, for continuous transformations, the term t in the JMAK equation may be replaced by the Quench Factor (g) [3]. Rometsch [4] suggested that the development of strength in a precipitation hardened metallic component is proportional to the square root of the volume fraction of precipitate, so that maximum value of the achievable strength for an alloy can be described by Eq. (1). In Eq. (1), ó is the predicted peak strength, 0^ and σ ^ are the minimum and maximum values of the strength achievable for the alloy, and K\ is a constant.

Introduction and Background The mechanical properties of aluminum alloy castings can be greatly improved by a precipitation hardening heat treatment. A typical precipitation hardening heat treatment consists of three steps: (1) solutionizing, (2) quenching, and (3) aging; and is performed by first heating the casting to and maintaining it at a temperature that is a few degrees lower than the solidus temperature of the alloy in order to form a single-phase solid solution. Then rapidly quenching the casting in a cold (or warm) fluid in order to form a supersaturated non-equilibrium solid solution; and finally, reheating the casting to the aging temperature where nucleation and growth of the strengthening precipitate(s) can occur.

-[expi-^g)]^

(1)

In order to obtain the cumulative Quench Factor (Q), incremental quench factors (qfi are calculated for each increment on the cooling curve as the ratio of the time that the material spends at the specific temperature (Δί,) divided by the critical time that is required for a certain amount of transformation to occur at that temperature(Ct.). The incremental quench factor values are then summed up over the entire transformation temperature range in order to produce the cumulative Quench Factor (Q), as shown in Eq. (2).

The objective of this work is to develop a model and the necessary material database that allow predicting these physical and material property changes. The structure of the model is described in Figure 1. A thermal module calculates the thermal history of the part during quenching. The time-temperature output from the thermal module becomes input to a user-developed property module. This is a module and database for predicting the maximum resultant room temperature hardness attainable after aging at each node within the model of the cast component. This is done by a Quench Factor Analysis (QFA) [1].

Q=L«f=Î^

Quench Factor Analysis was first developed by Evancho and Staley [1] in 1971 to predict the effect of continuous quenching on

835

(2)

In order to use Eq. (2), the cooling path taken by the material during quenching must be known. One way of representing the cooling path is via a time-temperature-property (TTP) curve. This curve is often referred to as the ' C curve of the material. The TTP curve is a graphical representation of the transformation kinetics that influences the material's mechanical properties and defines the time that is required to precipitate sufficient solute to alter the strength of the material by a specified amount. The C curve may be defined mathematically by the critical time function (C,), which is given by Eq. (3). In Eq. (3), Ct is the critical time required to form a specific quantity of a new phase. Kj to K5 are constants that depend on the material [5]. Kj is equal to the natural logarithm of the fraction of material which is untransformed during quenching, K2 is related to the reciprocal of the number of nucleation sites, K3 is related to the energy required to form a nucleus (J/mol K), K4 is related to the solvus temperature (K), and K5 is related to the activation energy for diffusion (J/mol), R is the universal gas constant (J/mol K), and T is absolute temperature (K). The main idea of QFA is to transform the TTP curve into a mathematical equation that can be used for calculating the volume fraction of precipitate that form during quenching in terms of loss of strength. Ct =-A'1xΔ'2xexp

KMKJ RT{K,-Tf

X

PiRT]

referred to as a lumped parameter analysis) performed on the system (i.e., the probe + the quenching medium) results in Eq. (4), which yields HTC [8]. In Eq. (4), Λ(Τ) is the quenching heat transfer coefficient of the probe, p, V, Cp, and As are the density, volume, specific heat, and surface area of the probe, respectively. Ts is the temperature of the probe and 7} is the bulk temperature of the quenching medium. The derivative of temperature with respect to time (i.e.,— J is calculated from the measured temperature vs. time data. h(T) = --

pVC s V

s

P

dT

(4)

TJ)dt

In order to determine the HTC with this method, a cylindrical quenching probe, 0.375 inch (9.53 mm) in diameter and 1.5 inch (38.1 mm) in length, and a quenching disk, 1.1 inch (27.94 mm) in diameter and 0.3 inch (7.62 mm) in thickness, were cast and machined from standard A356.2 alloy. Thermocouples were placed in the molds at the geometric center of each casting. During measurements, both the cylinder and disk were quenched from 538°C (1000°F) into three different quenching media: (i) hot water that is maintained at 80°C (176°F), (ii) static room temperature air, and (iii) forced-air obtained by an industrial fan. Figure 2 shows the measured heat transfer coefficients by quenching probe and disk into hot water as a function of temperature. HTC for quenching in static room temperature air ranged between 14-41 W/m2, and that for quenching in forced-air ranged between 168-181 W/m2.

(3)

Materials and Procedures Aluminum casting alloy A356.2 is used to develop and demonstrate the procedure for obtaining the necessary database and modeling the response of aluminum alloy cast components to T6 heat treatment. The data includes mechanical properties and heat transfer coefficients for various process steps as functions of temperature. Other required thermal and physical properties, such as density, specific heat, etc., are obtained from JMatPro Software1. The methodology developed in modeling A356.2 alloy castings can be easily extrapolated to other Al-Si alloys.

~«~~>Querschsng Disk •"—»Quenching Probe

Thermal Conductivity Needless to say, thermal conductivity is an important parameter required for heat transfer analysis. The thermal conductivity of A356.2 alloy in the as-cast condition was measured at several temperatures according to ASTM standard E1225-04. The measured thermal conductivity was compared to values published in reference [6] and was found to be in good agreement in the temperature range between 100°C (212°F) and 400°C (753°F).

Temperature (C)

Figure 2. Quenching heat transfer coefficient measured for 80°C (176°F) hot water. Temperature-Dependant Local Heat Transfer Coefficients For decades, many researchers have tried to determine heat transfer coefficients for various processes analytically [9] and many have tried to determine it experimentally [10]. However, heat transfer coefficients are very much dependant on part geometry, quenching medium and quenching process and this makes their determination difficult and the values obtained are approximate at best. More recently, a computer program has been developed for determining heat transfer coefficients in casting and quenching processes [11]. However, quenching a hot object into a fluid involves complex thermodynamic, fluid dynamic and phase transformation interactions that occur simultaneously and make the necessary simulations require a long time even with the fastest computer processor. For these reasons, an efficient method for obtaining quenching heat transfer coefficients for simulation purposes is needed.

Heat Transfer Coefficients The quenching heat transfer coefficient (HTC) is used by the thermal module in ABAQUS to compute the heat that is transferred out of the casting during quenching. Measurement of the quenching HTC involves quenching hot cylindrical probes that are machined from cast A356.2 alloy and equipped with a k-type thermocouple that is connected to a data acquisition system, into the quenching medium and acquiring the temperature-time curve [7]. Prior to quenching, the probes are heated to the solutionizing temperature for 12 hours. A heat balance analysis (usually 1

Developed and marketed by Sente Software Ltd., Surrey Technology Centre, 40 Occam Road, GU2 7 YG, United Kingdom.

836

These shapes were cast and instrumented with thermocouples. Some castings were machined so as to create the necessary features. The castings were solutionized at 538°C (1000°F) and then quenched into water that is maintained at 80°C (176°F). A thermal module was made for each of the geometries and the exact quenching conditions were used to simulate each part. In calculating the temperature changes during quenching, each surface was assigned one of the measured quenching heat transfer coefficients that were measured either by quenching probe or disk in contact with water or air, depending on the local quenching conditions. For example: in case (2), it is assumed that the cavity volume is completely filled with air from the start of the quenching event until its end, so the cavity surfaces were assigned the measured heat transfer coefficient for static air. The remaining surfaces were assigned the measured hot water heat transfer coefficient obtained by the quenching probe. In case (4), the flat bottom surface was assigned the measured hot water heat transfer coefficient from the quenching disk and the remaining surfaces were assigned the measured hot water heat transfer coefficient from the quenching probe.

Before discussing our effort towards this end, it is necessary to briefly review what happens during quenching. There are three distinct stages during quenching. These are: (1) formation of a vapor blanket around the solid part, (2) nucleate boiling of the quenching medium, and (3) convective cooling of the solid part. Each of these three stages is associated with a distinct cooling regime and heat transfer from the solid surface is very much dependant on small variations in the conditions of the quenching bath and the state of the metal surface. Particularly, the formation of the vapor blanket around the solid surface creates a problem in modeling the quenching process: Due to the large difference between the thermal conductivity of air and that of water, contact of the solid surface with air bubbles decreases the cooling rate while its contact with cold water increases it. In a typical casting, some features may trap the vapor phase and other features may restrict the movement of the quenching fluid causing the fluid in contact with these areas to heat up locally. These effects can reduce the local cooling profile. In this work, an assumption has been made that the air entrapments or restricted air flow can be represented by assigning different heat transfer coefficients in local regions. In order to verify the validity of this assumption, five simple shapes were considered. These shapes allow representation of almost all of the features that may be present in a typical casting. Once the heat transfer coefficient for each one of these shapes is determined, it may be applied locally to the corresponding feature on a complex casting that is to be simulated. There are two important advantages in locally applying the quenching heat transfer coefficients in computer simulations. These are: (1) the boundary conditions for the model become more representative of the physical situation, and (2) the computational time is significantly reduced. The five selected geometric shapes are shown schematically in Figure 3 and they are as follows:

Each of the five castings was quenched and the cooling rate vs. temperature was recorded and compared to the computercalculated cooling rate vs. temperature. In all cases, there is excellent agreement between the measured and computercalculated curves indicating that the developed temperaturedependent local heat transfer coefficients can be used for thermal simulations on different shape castings. The procedure described above provides a useful means of resolving the issues caused by the vapor blanket and air pockets that form unevenly in and around typical castings and cause uneven cooling that results in different cooling profiles from location to location on the same casting. However, details in commercial castings are usually very complicated and it may not be easy to manually assign a heat transfer coefficient to each and every surface on such complex castings. However, a computer module may be easily developed to accomplish this task.

1. A free surface over which bubbles and/or vapor that is produced during quenching can escape freely into the surrounding fluid.

Kinetics Parameters for Quench Factors Analysis

2. A cavity where bubbles and/or vapor that is produced during quenching are trapped and form an air pocket.

The aging curve for A356.2 alloy was needed in order to perform the Quench Factor Analysis. The aging curve was obtained by measuring the Rockwell hardness B scale (HRB) of the alloy. The HRB measurements were performed with a steel ball indenter that is 1/16 inch (1.59 mm) in diameter and a minor and major load that are 98N and 883N, respectively. The result is shown in Fig. 4. In order to obtain this data, small identical samples of A356.2 alloy were solutionized at 538°C (1000°F) and then quenched into ice water. These samples represent the maximum possible quenching rate. Subsequently, the samples were aged at 155°C (311°F) for different periods of time and their hardness was measured. The values were averaged from 20 to 40 measurements and the maximum value was found to be 65 HRB. It was achieved after aging for 19 hours. This number represents the maximum hardness value in the Eq. (1); i.e., ó^^. The value for the minimum hardness in Eq. (1); i.e., σ^η, was obtained by furnace cooling the samples after solutionizing. The cooling rate in the furnace was found to be less than 0.2°C/s, and a^ was found to be 8 HRB.

3. A channel where water flow is restricted and so the water temperature is locally higher than the average bath temperature. 4. A horizontal surface against which the vapor phase that is produced during quenching is trapped giving rise to a smaller heat transfer coefficient. 5. An angled surface that affects the rise of the vapor phase that is produced during quenching through the quenching fluid thus decreasing the magnitude of the heat transfer coefficient (the magnitude of the heat transfer coefficient in this case may be different from that in case 4).

(1)

m

(3)

W

{5}

Figure 3. Five shapes used for determining the quenching heat transfer coefficients. The red lines represent vapor and/or gas bubbles.

Next, the Jominy End Quench test described in ASTM-A255 was used to determine the kinetics parameters. A356.2 alloy Jominy End Quench bars that are 1 inch (25.4 mm) in diameter and 4

837

inches (101.6 mm) long were cast in a permanent mold. The bars were then instrumented with thermocouples at seven different locations along their length in order to record the local cooling data during the end quenching process. The thermocouples were evenly distributed at 0.5 inch (12.7 mm) increments along the length of the bar. The bar was solutionized for 12 hours at 538°C (1000°F) and then quenched from one end by cold tap water while the time-temperature data was being recorded. The unidirectional heat transfer thus created results in a progressively decreasing cooling rate along the length of the bar. The recorded cooling curves and cooling rates vs. temperature are presented in Figure 5Figure 6, respectively. Small flat surfaces were then made along the length of each quenched bar by rubbing the surface with fine sand paper in order to allow for accurate hardness measurement on the flat surface of the bar. HRB measurements were then performed around the perimeter at the thermocouple locations. Because the measured HRB is an arbitrary number with no physical meaning, the HRB values were converted into Meyer hardness [12] for the purpose of calculation and then they were converted back to HRB for presentation.

300 Temperature (C)

Figure 6. Recorded cooling rates for different locations along the length of the A356.2 aluminum alloy Jominy End Quench bar. In order to determine the kinetics parameters Kh K2, K3y KAi and #5 that appear in the Ct function described by Eq. (3), previous researchers [1, 2, 3, 4, 5, and 13] used mathematical equations or other analytical methods which vary the five unknown constants simultaneously in Eq. (3) in order to fit the experimental data. This is a difficult task since different combinations of the five constants could yield equally good fits of the experimental data [14]. In this work, a new approach was adopted wherein three out of the five unknown kinetics parameters; namely, Kh K4, and K5, were fixed, and only K2 and K3 were made to vary. Kx is easily found since it is the natural log of the fraction of material that is untransformed during quenching. K4 and K5 are the solvus temperature of A356.2 and the activation energy for aging the precipitates [17], respectively. Eq. (1) may be re-written as follows,

SOÔÛ

SO0OO

SÔÔ09O

ί = 21n

Aging time (s)

Figure 4. Measured HRB vs. aging time for A356.2 aluminum alloy.

o-om

(5)

Kt

According to Eq. (5), the Quench Factor (Q) may be determined from the measured hardness values (ó, amax and a^n) and Kj. Also, the Quench Factor (Q) may be determined from the local cooling data and the Ct function. The Ct function is given by Eq. (3), which can be re-written as follows,

o-t

Äί,

-J^xA^xexp RT(K4-TJ V

xexp

Mi

(6)

The Quench Factor (Q) calculated from Eq. (5) was plotted against the Quench Factor calculated from Eq. (6). Then the remaining unknown kinetics parameters; i.e., K2 and K3, in Eq. (6) were continuously adjusted until the scatter best fitted a line that passes through the origin and makes a slope that equals to 1. This procedure allowed obtaining all kinetics parameters (K\ through K5) and the results are shown in Table I. With this procedure, the C curve for A356.2 alloy was generated as shown in Figure 7.

Figure 5. Recorded cooling curves at different locations along the length of the A356.2 aluminum alloy Jominy End Quench bar.

838

The computer-calculated cooling curves and cooling rates at the two nodes indicated by (1) and (2) in Figure 9 are reported. For the part quenched in hot water, the recorded and computercalculated cooling data from the two locations are shown in Figure lOFigure 11. For the air quenched part, the results are shown in Figure 12. In all cases the results show excellent agreement between the measured and the computer-calculated cooling data indicating that the developed database of heat transfer coefficients and the method of locally assigning them to regions on a typical quenched part are accurate. Subsequently, two quenched parts were aged at 155°C (311°F) for 19 hours right after quenching, and HRB was measured on four different surfaces, as shown in Figure 13. The measured and computerpredicted hardness results are shown in Figure 14. In all locations, there is excellent agreement between the measured and the computer-predicted HRB values.

Table I. The kinetics parameters for A356.2 alloy. Kj

-0.00501

6.70xl0"17

K3 (J/mol) 8,887

K4 (K) 890.31

K5 (J/mol) 130,000[13]

Figure 7. Generated C curve for A356.2 alloy. Simulation results and comparison to measurements

Yeilow

The part shown in Figure 8 was designed, cast, machined, and equipped with two thermocouples as shown. Two different quenching processes were used to validate the model with this part. For both quenching processes, two identical parts were solutionized at 538°C (1000°F) for 12 hours and then one part was quenched into water that is maintained at 80°C (176°F) and the other part was air quenched by a stream of forced-air. For the water quench, the part was quenched by immersing the front face down into the water so that the blind cavity shown in Figure 8 (a) was filled with air as the part was quenched down to room temperature. For the forced air quench, the air was directly blown onto the front face of the part shown in Figure 8 (a). The timetemperature data was recorded from the two thermocouples. The part was then modeled as shown in Figure 9 and computer simulations were conducted with heat transfer coefficients assigned as follows:

Nodes where compytercalcyiat6
Figure 9. Modeling the part shown in Fig. 8.

~*»~ Measured nt iocstion (1) ....... Predicted at location {D

For water quench: (1) the surface indicated by pink color in Figure 9 together with the face parallel to it were assigned the heat transfer coefficient that was measured by the quenching disk for hot water quenching, (2) the surfaces indicated by yellow color in Figure 9 were assigned the heat transfer coefficient measured for static air, and (3) all the other surfaces were assigned the heat transfer coefficient measured by the quenching probe for hot water quenching. For air quench: (1) the face indicated by pink color in Figure 9 was assigned the heat transfer coefficient measured for air quenching, and (2) the other faces were assigned the heat transfer coefficient measured for static air.

Figure 10. Measured and computer-calculated cooling curves at thermocouple location (1) for quenching in hot water. ~—Measured »t iocatkm (2) - - Predicted at îocatioo {2)

Blind cavity that fills with air when the part is quenched with this face facing down onto the surface of the water

Figure 8. The part used for verification (a)frontview, and (b) back view.

Figure 11. Measured and computer-calculated cooling curves at thermocouple location (2) for quenching in hot water.

839

References \ -^Measured {ôCâtton <1) 2 \ ~*-Measured locatif (2) ð \ »- Predicted location {!} » i s ! -»-Predicted location {2)

, **&$^ j » ^ ^ ^ ^******^ ^ s ^ s ^ C ^

f ι| Q

l

1. J.W. Evancho and J.T. Staley,"Kinetics of precipitation in aluminum alloys during continuous cooling, " Metall. Trans., 5 (1971), 43-47. 2. J.T. Staley, R.D. Doherty and A.P. Jaworski, "Improved model to predict properties of aluminum alloy products after continuous cooling," Metall. Trans. A, 24 (1993), 2417-2427.

^*^^*^^^

3. J.T. Staley, "Quench Factor Analysis of Aluminum Alloys. Material Science and Technology," Material Science and Technology, 3 (1987), 923-935.

*0*r^--

0

100

203

380

400

500

Temperature (Ç)

Figure 12. Measured and computer-calculated cooling rates at thermocouple locations (1) and (2) for quenching in air.

4. P.A Rometch, M.J. Starink and P.J. Gergson, "Improvements in quench factor modelling," Mater. Sei. Eng, A339 (2003), 255264. 5. J.T. Staley and M. Tiryakioglu, "Use of TTP Curves and Quench Factor Analysis for Property Prediction in Aluminum Alloys, in Materials Solutions Conference" (Paper presented at Materials Solutions Conference, ASM International. Indianapolis, IN, 2001) 6. S.I. Bakhtiyarov, R.A. Overfelt, S.G. Teodorescu, "Electrical and thermal conductivity of A319 and A356 aluminum alloys," Journal of Materials Science, 36 (2001), 4643-4648.

Figure 13. Indication of surface sections for hardness measurements. Model Prediction

n 51.9

I Front A

I Measured HM

46,73

5J,t M

Front B

«Jl

48.6

46.72

«.2

S9Ü

64.41

S9.8

$4 3

».«

«4,46

7. M. Maniruzzaman, J.C. Chaves, C. McGee, S. Ma and R.D. Sisson, Jr., "CHTE Quench Probe System - a new quenchant characterization system, in the 5 th International Conference on Frontiers of Design and Manufacturing" (Paper presented at the 5th International Conference on Frontiers of Design and Manufacturing, Dalian, China, 2002)

5$.9

j

1111111 BackC

Air quenched (HRB)

BackD

|; Front A

Fronts

BackC

8. L.S. Tong and Y.S. Tang, Boiling Heat Transfer and TwoPhase Flow (CRC press 1997), 1-5. 9. R.G. Hills and E.C. Hensel Jr., "One-dimensional nonlinear inverse heat conduction technique," Numerical Heat Transfer, 10 (1956), 369-393.

BackD ;

80C Hot water quenched {HRB)

10. K. Ho and R.D. Pehlke, "Metal-Mold interfacial heat transfer," Metall. Trans. B, 16 (1985), 585-594.

Figure 14. Measured and computer-predicted hardness for hot water quenched and air quenched parts.

11. M. Li and J.E. Allison, "Determination of thermal boundary conditions for the casting and quenching process with the optimization tool OptCast," Metall. Trans. B, 38B (2007, 567-574.

Summary and Conclusions A model has been developed using the ABAQUS finite element analysis software to predict the response of cast aluminum alloy components to solutionizing, quenching and aging processes. The necessary database for A356.2 alloy is generated. Both hot water and air quenching were selected for model validation. The thermal module within the model calculates the temperature profile for the quenching process using a database of temperature-dependant heat transfer coefficients and the new method of locally assigning them to regions on the quenched part. The user-developed property module which is based on a Quench Factor Analysis, predicts the local hardness values on the casting. The model predictions were verified by measurements made on heat treated parts, and the model-predicted cooling curves and hardness values were found to be in very good agreement with measured values.

12. M. Tiryakioglu and J. Campbell, "On macrohardness testing of Al-7wt.% Si-Mg alloys I. Geometrical and mechanical aspects," Materials Scence and Engineering, A361 (2003), 232-239. 13. R.J. Flynn and J.S. Robinson, "The application of advances in quench factor analysis property prediction to the heat treatment of 7010 aluminum alloy," J. of Mater. Proc. Tech, (2004), 674-680. 14. P.A. Rometch and G.B. Schaffer, "An age hardening model for Al-7Si-Mg casting alloys," Materials Scence and Engineering, A325 (2002), 424-434.

840

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DETERMINATION OF COKE CALCINATION LEVEL AND ANODE BAKING LEVEL APPLICATION AND REPRODUCIBILITY OF L-SUB-C BASED METHODS Stein Rorvik1, Lorentz Petter Lossius2 and Arne Petter Ratvik3 ^INTEF Materials and Chemistry, Trondheim, Norway 2 Hydro PMT, Primary Metal Technology, Ârdal, Norway 3 Norwegian University of Science and Technology (NTNU), Trondheim, Norway Keywords: Petroleum coke, carbon anodes, L-sub-c / L c crystallite height, crystallinity of carbon Abstract The average crystallite size (L-sub-c or Lc) is an important property of carbon materials for aluminium electrolysis; L c is used for characterizing the petroleum coke calcination level and sometimes also to estimate the baking level of anodes. This paper discusses problems when comparing L c results from different laboratories using precision statements from ASTM and ISO standards. The main cause is peak broadening errors introduced by the XRD instrument and sample preparation. The L c standards ASTM D5187 and ISO 20203 neglect these errors. Two ways are demonstrated to minimize the peak broadening effect to improve the standards, 1) by using thin sample thickness and 2) by embedding the coke in a high absorptive medium. Using L c to determine the anode baking level is discussed and three practices are discussed; measurement on the anode directly or two methods for using a reference coke that is baked with the anode. It is shown that precision is better for the latter methods. Especially for underbaked anodes a baking level estimated from measurement of the anode L c can be misleading.

Figure 1: Illustration of Franklin's model of a graphitizable (but non-graphitic) carbon, from [2]

Introduction

Application of L c Measurements

One of the earliest models of the structure of disordered carbons, based on X-ray diffraction, was published by R.E. Franklin in 1950 [1] and 1951 [2]. She proposed that graphitizable carbons were built up of hexagonal sheets of carbon rings, which formed small individual stacks, connected using cross-links. The stacks in the graphitizable carbons tend to pack with a preferential orientation, enabling them to merge into bigger stacks upon further heat-treatment. Petroleum coke, which is used for production of carbon anodes in aluminum production, belongs to this category. There are several other models that also have been proposed for the description of the structure of disordered carbon. But for the structure of petroleum coke, Franklin's model seems adequate and is currently the most commonly used.

The most common expression for crystallinity in calcined coke used in anodes for aluminium metal electrolysis today is the L c value. The development of L c with calcination temperature is illustrated in Figure 2 for five cokes with different sulfur levels. The L c growth is nearly linear in the temperature range of interest for green coke calcination. Figure 2 also illustrate that L c growth differs between cokes, with cokes D and E reaching an L c of 30 ΐ at 1200 to 1210°C calcination temperature and the lower sulfur cokes A, B and C reaching 30 Â at 1250 to 1260°C.

WtÈËÊÈËÈÊÊÊ

The parameter used to describe the quality of the coke structure from this model is the crystallite size, described by the distances L c and La. These are estimations of the size of each individual crystallite (= "graphite stack") contributing to the diffraction of the X-rays. L c is the distance along the c-axis, perpendicular to the graphitic planes (= "crystallite height"); while L a is the distance parallel to the planes (= "crystallite diameter"). These distances are drawn upon Figure 1. Of these two parameters, L c is the one that is most relevant to carbon properties, and is also the easiest one to measure. It can be calculated by measuring the broadness of the main graphite peak in the XRD pattern. The L c distance is usually expressed in angstrom (Â) and increases upon heat-treatment.

35.0

?o

- -Coke A, 1.3% S ^ C o k e B , 1.5% S Coke C, 2.0 % S Coke D, 3.9 % S -•-Coke E, 4 . 6 % S

σ> 30.0

Calcination Temperature (°C)

Figure 2: L c development with calcination temperature for five cokes with different sulfur levels [3].

841

L c is also useful for checking uniformity of heat treatment of a coke. Due to the nature of the process, rotary kilns will cause different L c development for fine, medium and coarse grains as the tumbling action sends coarse grains closer to the fire [4,5]. The coarse grains then undergo higher L c growth e.g. reaching an L c of 34 ΐ while medium grains are 32 Â and fine 30 Â [4].

Precision Discussion As a control on calcination, the L c within-laboratory repeatability of 0.5 Â is sufficient. For the same coke it corresponds to control of the calcination level within 8°C, estimated from Figure 2. However, the between-laboratory reproducibility of 1.8 ΐ or 1.9 Â is not satisfactory as a precision limit for comparisons and a control of a specification. With this level of precision, the coke buyer's laboratory will have problems running a reliable monitoring program on coke shipments. The uncertain precision for the between-laboratory reproducibility is problematic.

These aspects of coke quality illustrate the importance of having accurate and precise L c measurements. This paper shows that there are several sources of error in the ASTM and ISO standards for measuring L c in coke.

A Critical Review of the Current Standards

Precision of Current Standards

The current paper will in the following section argue that there are many problems with the current ASTM and ISO standards for measuring L c in coke. Since the relevant sections of these standards are identical, a reference to "the standard" in this paper will apply to both ASTM D5187-91 and ISO 20203. Quotations are emphasized in italics with section number from the ISO 20203 standard.

ASTM and ISO method precision is expressed by the withinlaboratory repeatability, r, and the between-laboratory reproducibility, R, usually at 95 % confidence level. This is the r&R statement. It is obtained through an interlaboratory study (ILS), also termed a round robin (RR). It should be noted that the precision values obtained tends to be the best case as the voluntary participation attracts a good quality class of laboratories.

9.1.2 Determine the average low and high backgrounds (Points A and B, respectively) on the diffraction scan and connect them with a straight line.

ASTM D5187 The method ASTM D5187-91 (2010) - Standard Test Method for Determination of Crystallite Size (Lc) of Calcined Petroleum Coke by X-ray Diffraction was recently revised with a new precision statement/1} The revised precision limits are

This is not as easy as it sounds, and the illustration to how to do this (Figure 3) is not accurate. The starting point A should have been set at a lower angle to include the entire peak, perhaps at 15° in this case. The broader the peak is (lower Lc), the harder it is to decide where the peak actually starts. A better way to do this would be to use a mathematical algorithm to curve-fit the background and use either a derivative of the backgroundsubtracted curve to determine where the peak starts and stops, or use a fixed angle window (which probably is a more reproducible approach).

Repeatability, r = 0.5 ΐ Reproducibility, R = 1.9 Â The earlier r&R(2) precision used in ASTM D5187-91 (2007) and previous revisions was L c dependent, with r equal to 0.021 *LC and R equal to 0.11*LC. Especially between-laboratory comparison precision was improved with the 2010 revision. For example, for an L c of 30.0 Â, the old R value of 0.11 *LC was 3.3 Â, significantly higher than the new R value of 1.9 Â. In a comparison between two laboratories, the R value states that determination on two samples of the same material the difference in L c should be within 1.9 A for 95 out of 100 such comparisons

9.1.3 Construct line CD parallel to line AB, and going through the apex of the peak at point G f(hkl (002) at 0,335 run]. Draw the line such that, if the peak is irregular, it will pass through the average of the irregularities.

ISO 20203-2005 The method ISO 20203-2005: Carbonaceous materials used in the production of aluminium — Calcined coke — Determination of crystallite size of calcined petroleum coke by X-ray diffraction was based on ASTM D5187-91 (2002) and had the same r&R precisions statement with r of 0.021*LC and R of 0.11*LC described in the preceding section. A 2010 revision with new r&R precision limits is being voted on.(3) The new r&R limits are Repeatability, r = 0.5 Â Reproducibility, R = 1.8 ΐ The ISO r&R is comparable to the ASTM r&R. Five laboratories (Hydro & Slovalco) participated in both the ISO and ASTM round robins. ()

ASTM 2010 r&R: The research report is available from ASTM as RR: D02-1690 (D5187). Eleven laboratories analyzed two duplicates of each of seven samples spanning 21 to 34 Â (angstrom). ASTM 1991 r&R: Ten laboratories analyzed two duplicates of each of six materials. 3 ISO 2010 r&R being voted on: Eleven laboratories analyzed three duplicates of each often materials.

842

It is not mentioned in the standard what an "irregular" peak is, which leaves it to the analyst to subjectively determine good data from bad, and allows a possible arbitrary correction in the analysis. For FWHM determination, the three lines AB, CD and EF should be parallel, in the illustration they are not. Also, the theoretical do02 peak of graphite at 0.335 nm corresponds to an angle 2Θ at 27.38° (for the typical Cu Κα radiation at ë = 1.54056Â commonly used in XRD equipment). This angle is illustrated by the authors of the current paper as a dotted line in Figure 3. Petroleum cokes usually have the main peak shifted a couple of degrees lower, because Franklin's model is not exact for cokes. An alternative model [6] is where the structure of disordered carbons is viewed as graphite sheets with a various amount of interstitial carbon atoms. From this model, a soft carbon where all possible interstitial sites are present has an interlay er spacing of 0.344 nm. Petroleum cokes have a do02 spacing fairly close to this. The illustration in the standard show the usual peak which is shifted, but the standard do not explain why the peak is shifted nor explain how to deal with the shifted peak.

Lc=-

0.89A 2(sin θ2 - sin θγ )

This equation is a derivative of the Scherrer equation. It is easy to use, but not accurate. Frank R. Feret has described [7] the modifications Alcan has applied to the ASTM method, and explains why. Feret's comment on the ASTM method's use of the derivative equation is as follows: "This approximation is valid only when È = (È] + 0J/2 and â are both small. Most of the calcined coke peaks are asymmetric and some are very asymmetric. Therefore, È Ö (Oj + È2 )/2 and 20j - 2È2 is not small. Moreover, it seems that because the original Scherrer equation itself is simple, there is no needfor the approximation. " The original Scherrer equation (which Feret prefers) is given as - _ 0.89/t Lc — âïïæθ -

14

18

22

26 i

30

34

38

where β is the peak integral breadth (IB) or full width at half maximum (FWHM) and Θ is the angle at the peak position. The authors of the present paper assumes that the reason for using the derived equation in the standard is that it is slightly easier to use, since the input is Θι and Θ2, which is measured manually on the XRD scan. However, modern XRD software using a built-in function for the FWHM measurements would instead report β and Θ directly, which can then be input to the more correct Scherrer equation. Hence, there is no reason to use the derivative equation instead of the original equation.

X

Figure 3: Calculation of the graphite peak broadness by the ASTM D5187-91 / ISO 20203 method. Y axis is intensity and X axis is 2Θ angle. 9.1.4 Determine the full-width half maximum (FWHM) of line AB. Construct line EF such that it intersects the peak at half of its maximum value. The points at which EF intersects the peak are 20j and 2È2, respectively. This instruction is good for its simplicity, but modern computerized X-ray software will usually already have a built-in function for determining the half-width of a peak automatically, which will be easier to use and much more accurate. A potential problem with this approach is that the actual curve-fitting procedure used inside the software may be unknown and may vary between software packages. The standard should ideally specify how to do this mathematically or algorithmically. 9.1.5 For computer simulation based on the intensities recorded at 0.2° intervals, produce a mathematical representation of the diffraction curve. Determine the baseline, peak, peak height, and half-peak height to produce the half-peak height angles, 2È] and 2È2, as above. The 0.2° step size is a hangover from older generation diffraction equipment, where low count rates required that compromises in resolution were made to allow data collection at a reasonable speed. Modern PSD detector systems offer in the order of 100200x the counting rate of old point detectors, making such compromises unnecessary. The choice of such a large step size is also very problematic when collecting data on highly carbonized materials, which may exhibit a peak width as narrow as 0.05°. 9.2

It is mentioned in the standard that the use of the constant 0.89 is "arbitrary". Feret also comments on this: "The Scherrer constant depends largely upon the crystallite shape, the (hkl) indices and the definitions taken for â and Lc. Various investigators have assumed values from 0.70 to 1.70 for this constant. For cokes it is set equal to 0.89 for the sake of uniformity in published results. " The value of 0.89 is not truly arbitrary, as the standard suggests it actually derives from the assumption that the sample is comprised of spherical particles of cubic symmetry. Further, it is related to the method used to measure the peak width. The value of 0.89 is actually derived for the case where the integral breadth of the peak is used, whereas the standard stipulates that the FWHM is used. In this case a value of 0.94 should be used. While this makes little difference to the repeatability and reproducibility of results, it does lead to a systematic error in the absolute values measured and their comparability with values obtained by other techniques. The ISO standard has this closing remark in section 9.2: The above equations make the assumption that the true line width is equal to the measured width, the contribution of instrumental line broadening is negligible. This assumption is manifestly wrong. Every diffractometer has an inherent linewidth which is a function of many variables, including the measuring radius, detector system, the radiation spectrum used and the choice of beam collimating slits and optics. The differences between instruments are not insignificant and ignoring inherent resolution characteristics results in a number of errors to the reproducibility of L c values. The most significant contributors are the radiation spectrum and the instrument optics.

Determine the mean crystallite height Lc (derived from Scherrer equation)

843

However, one could argue that the petroleum cokes usually have a FWHM much larger than the instrument contribution, so that the latter can be neglected. A potential problem with the "instrumental line broadening" is that XRD software usually has a function for removing the secondary (Ka 2 ) X-ray peak, which will be present unless the instrument is equipped with a monochromator. This function may be undocumented, and its use can have unpredictable results on the very broad coke peaks. The instrumental broadening is also influenced considerably by the setup of divergence slits in use, so two different instruments can have a considerably different resolution.

silicon single crystal, which is oriented in such a way as to give no X-ray reflections in the scanned angle range. Sample holders made for this use are commercially available. Upon evaporation of the solvent, a very thin layer of sample powder will be distributed across a surface which has no X-ray reflections. The lack of reflections from the silicon crystal will ensure a good signal-to-noise ratio, which may be necessary because of the lower X-ray intensity from a sample prepared this way. For 2), when mixing the carbon with something of higher X-ray absorption, these requirements must be met:

Earlier Proposals to Improve Methods for Measuring L c

a)

The paper by F. R. Feret [7] provides important understanding of the relation of the do02 peak shape to the degree of graphitization of carbon, and suggest to use computerized profile fitting methods to calculate the L c values. Iwashita et.al. [8] published in 2004 a proposal for a new standard procedure of X-ray diffraction measurements on carbon materials. Currently this is by many scientists regarded as "state of the art" for L c measurements. The main differences to the ASTM and ISO methods are 1) Silicon is added as an internal reference to the peak position and 2) The Xray pattern is corrected for angle dependent factors. These include the Lorentz factor (L), Polarization factor (P), Absorption factor (A) and Atomic Scattering factor (Fc).

b) c) d)

the mixing compound must have a much higher X-ray absorption than carbon the mixing compound must not have reflections near the doo2 peak of carbon the mixing compound must be inexpensive and not be hazardous to the environment (since it would be used as a disposable) the compound must have a similar grain size/shape so that it mixes well with the carbon powder

Preliminary investigations has shown that while it is relatively easy to find compounds that satisfies requirements a), b) and c), this is not the case for requirement d). Carbon powder consists of flaky particles that form a low density powder. The best candidates satisfying both a), b) and c) are pure metals and their borides and carbides. These compounds have high densities, and tend to form rounded grains when milled. This makes it difficult to mix such powders homogeneously with carbon, lowering the reproducibility of the procedure.

The effect of using an internal reference is largest for high crystalline carbons. The effect of the angle dependent corrections is largest for low crystalline carbons, and is therefore relevant for cokes. The application of the angle dependant corrections to the X-ray pattern will increase the measured L c value. Whether or not the application of the angle dependant corrections will improve the between-laboratory precision has not yet been investigated. It is relatively straightforward to perform such a study since it only requires a purely mathematical treatment of the collected X-ray patterns and no instrumental changes.

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The Effect of Sample Preparation Carbon has a very low X-ray absorption coefficient(4). A consequence of this is that the X-rays will penetrate relatively deeply into the sample used for X-ray diffraction, giving diffracted reflections also from a volume below the sample surface. This results in a broadening of the diffraction lines, so that the FWHM calculates to a lower L c value than the actual value. The amount of broadening (and consequent lowering of the measured L c value) will depend on the sample preparation method and the sample holder in use.

Angle (2-Theta) Figure 4: The effect of X-ray absorption The effect of the X-ray absorption and the improvements proposed in this paper are illustrated in Figure 4. The patterns shown are of a graphite material, where the relative contribution from the absorption broadening is large. This example represents therefore a worst case situation. The red line shows the pattern using a standard sample holder, where the depth of the sample is about 1 mm. The blue line shows the pattern using a thin sample on a single crystal sample holder. The green line shows a pattern collected from an absorptive mix of 50% carbon and 50% copper powder in a standard sample holder. The absorption of the copper(5) reduces the X-ray penetration depth into the sample, giving a narrower peak. But the peak is still not as narrow as with

There are two ways to reduce the absorption effect: 1) 2)

Prepare a very thin sample where the penetration depth can be neglected Mix the carbon with something that absorbs X-rays better than carbon, without introducing extra peaks in the graphite peak interval of carbon

For 1), a commonly used method to prepare a very thin sample is to suspend the powder in a volatile solvent (e.g. isopropyl alcohol) and place some droplets of this mix onto a polished surface of a 4.576 cm2/g at 8 keV, data from http://physics.nist.gov/

5

844

52.55 cm2/g at 8 keV (11 times higher than carbon)

Table 1: Comparing the between-laboratory precision for the three methods for estimating anode baking level. Method Range of Between-lab Rel% precision values 1.8 to 1.9 ΐ 1. L c anode 28 to 34 A 31 1.8 to 1.9 ΐ 17 to 37 Δ 2. L c reference 9 3. Equivalent 1000tol400°E 14°E 3.5 temperature

the silicon sample holder. The FWHM of the pattern collected on the silicon sample holder is about 50% of the normal holder, resulting in a much higher L c value. Using a silicon sample holder seems to be the best method, but may be less reproducible, as the distribution of the powder of the carbon-solvent mix may be too unpredictable to be used as a standard procedure for quantitative measurements of L c in carbons. At the time of writing, no comparative studies have been performed on this issue.

The Equivalent Baking Temperature Scale The scale is based on a calibration set of eleven calcined reference coke samples heat treated from low (underbaked) level to high (overbaked) level. This ensures a range for all baked anodes. On the scale, normal baking level is around 1230°E, underbaked anodes are below 1150° and overbaked above 1330¸. The unit ¸ instead of °C is meant to emphasize that the measured temperature is an equivalent heat treatment of the reference coke.

The Anode Baking Level - Methods Using L c Three methods will be discussed: 1. 2. 3.

The direct measurement of the anode L c An indirect method using a reference coke that is calcined with the anode during baking The indirect method using the reference coke, with the addition of a calibration linking the reference coke to a temperature scale

Measuring Lr on Anodes Directly A reasonable criticism of the equivalent method is that establishing and using the method is more complex and costly than method 1 (the anode Lc). Analyzing anode L c on the anode cores is simpler and less expensive than other methods and would be a method of choice. Hydro has investigated this possibility, but extensive testing showed the relationship given in Figure 5.

It has been shown that between-laboratory precision for L c is 1.8 to 1.9 ΐ in the ISO and ASTM methods. This means that laboratories can report Lc-values that are systematically 1.5 Â off each other even when using exactly the same raw materials. For this reason, between-laboratory comparisons based on methods 1 and 2 can be troublesome to interpret. However, successful between-laboratory comparison is possible, and can be achieved if the laboratories use a common reference, a calibration. This is the basic assumption of the equivalent baking level method. By using a common temperature scale, the equivalent scale, the L c analysis for anode baking level can be harmonized at any laboratory. The key is the use of the same reference coke and its calibration.(6)

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The method is standardized as ISO 17499 (2006) — Carbonaceous materials used in the production of aluminium — Determination of baking level expressed by equivalent temperature. Central to the method is establishing the analytic relationship between the individual laboratory L c measurements and the temperature scale for the reference coke heat treatments.

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Between-laboratory Precision The major advantage of this method compared to using an uncalibrated reference coke is the improvement in comparisons between laboratories. Note that, as the equivalent scale value is a calculation from the L c of the reference coke, the repeatability, or within-lab precision, will not be better than for the L c measurement itself. The use of the calibration flushes out most of the interlaboratory differences in the L c part of the analysis giving a between-lab precicion closer to the within-lab precision. The precision statement of ISO 17499 is based on a 2003-2004 interlaboratory study.(7) The r&R limits are

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Figure 5: Anode L c development when plotted versus the equivalent baking temperature method. [9] The most obvious weakness is the horizontal part of the curve, which will make the detection of underbaked anodes difficult, if not impossible. The normal baking level is 1230°E(8), and the chart indicates poor detection from that baking level and downwards. That is problematic, and adding to that uncertainty is the spread in values for one L c measurement, given the withinlaboratory repeatability of 0.5 Â. Taking a measurement of 33.0 ΐ as an example, the baking level can be any value from 1050 to 1300°E. And then comes additional noise such as incidents of

Repeatability, r = 9°E Reproducibility, R = 14°E The gain in precision is considerable, see Table 1, which presents the relative precision of the three methods as the ratio of R versus the expected range of measured values. A good reference coke is a low sulfur single source sponge petroleum coke such as cokes A and B in Figure 2. Inter laboratory study with ten laboratories; three duplicates of each of ten materials.

Based on acceptable specific electrical resistance, strength and carboxy dusting levels.

845

highly under- or overbaked butts, the calcination level of the cokes, the coke blends, recipes and type of pitch.

For the direct anode L c method it is shown that precision is comparatively poor, and especially for underbaked anodes there is an inherent risk in the method of reporting with low accuracy.

To illustrate this issue in a more controlled study, a series of designed pilot anodes were made with three types of coke. Each batch was baked to three baking levels, slightly underbaked (1150°E), normal coke calcination level and slightly overbaked (1330¸). The L c values are plotted in Figure 6. The development of the anode L c is difficult to quantify below baking level 1230°E - the change in L c is small compared to the repeatability of the L c analysis method.

For Hydro, the gain in precision and the necessity of comparing anodes from different baking furnaces is sufficient reason for using the somewhat more complex equivalent method. And it has been shown that the equivalent temperature yield baking level information that is significant for anode quality in electrolysis, specifically carboxy dusting and current efficiency [10]. Acknowledgements The authors wish to thank Rain CII and Hydro for releasing data that made this paper possible, and Julian Tolchard at NTNU for providing comments to the X-ray diffraction discussion. References 1.

Rosalind E. Franklin: The interpretation of diffuse X-ray diagrams of carbon, Acta Crystallographica 1950, 3, 107121.

2.

Rosalind E. Franklin: Crystallite growth in graphitizing and non-graphitizing carbons, Proceedings of the Royal Society of London, 1951, A209, 196-218.

3. Figure 6: Anode L c plotted versus the equivalent baking level, for pilot anodes made using different types of coke.

L. Edwards, Coke Property-Temperature Data, Rain CII Carbon LLC, Internal Report, 2010.

4.

Les Charles Edwards, Keith J Neyrey, and Lorentz Petter Lossius, A Review of Coke and Anode Desulfurization, Light Metals 2007.

Conclusions

5.

Lorentz Petter Lossius, Keith J Neyrey, and Les Charles Edwards, Coke and Anode Desulfurization Studies, Light Metals 2007.

6.

Jacques Maire, Jacques Mering, Graphitization of soft carbons, in Chemistry and Physics of Carbon, PL Walker, Ed., Marcel Dekker, New York, 1971, Vol. 6, p. 125

7.

Frank R. Feret: Determination of the crystallinity of calcined and graphitic cokes by X-ray diffraction, Analyst, 1998, Vol. 123, 595-600

8.

N. Iwashita, C.R. Park, H. Fujimoto, M. Shiraishi, M. Inagaki: Specification for a standard procedure ofX-ray diffraction measurements on carbon materials, Carbon 42 (2004) 701-714

9.

Inge Holden, Korrelasjon Lc-Tekv Ârdalsanoder 1994-2000 Hydro, Ârdal Carbon Internal Memo.

The L c value development during calcination of coke is suitable for quantifying the heat treatment of the coke. However, L c analysis as described in ASTM D5187 and ISO 20203 has some errors and weaknesses leading to poor between-laboratory reproducibility. The paper discusses the improvements and also potential drawbacks: •

Correct the errors in ASTM / ISO standards - a drawback is that new values cannot be compared directly to old Improve sample preparation using a Si sample holder or an absorptive mix - // will be a challenge to gain acceptance for this across labs

•

Harmonize the computerized curve-fitting and calculation of L c - // will be a challenge to have software vendors standardize these methods

It is suggested to ASTM Committee D02.05 and ISO Technical Committee 226 that they look into these proposals for improving the L c analysis, thereby improving its commercial relevance. The paper also discusses use of L c analysis to determine the anode baking level. Three practices are presented; measurement of L c on the anode directly and two methods for using L c of a reference coke that is baked with the anode. It is shown that between-laboratory comparison is best for the equivalent baking level method, ISO 17499, as it avoids some of the current weaknesses in D5187 and ISO 20203.

846

10. T.E. Jentoftsen, H. Linga, I. Holden, B.E. Aga, V.G. Christensen and F. Hoff, Correlation Between Anode Properties and Cell Performance, Light Metals 2009.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

OPERATION OF AN OPEN-TYPE ANODE BAKING FURNACE WITH A TEMPORARY CROSSOVER Esteban Cobo1, Jorge Rey Boero^Luis Beltramino1, Juan Pablo Artola^Jean Bigot2,Pierre-Jean Roy2 Carbon Dept. - Research and Development Area - Aluar SAIC, 2 Rio Tinto Al can, Aluval, Centr'Alp ,BP7 ,38341 Voreppe, France Keywords: Anode baking furnace, refractory construction, fire operation Abstract

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As part of the Puerto Madryn smelter expansion project, Aluar Aluminio Argentino has built and now operates an open type furnace designed with Rio Tinto Alcan AP Technology. The furnace was built and put into service in two stages, each of which consisted of 34 sections and 2 fires. In order to allow the erection of the second stage and the connection between both halves of the furnace, a method that had proved to be successful on other projects, was applied. A temporary crossover was used to connect two sections at the end of the first stage. This paper describes the experience gained during the operation under these conditions and the procedure and process control modifications that have been necessary to maintain the anode baking quality.

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E 3 Dead sections [251 Temporary crossover Figure 1. Schematic view of the furnace during temporary connection The installation of a temporary crossover was one of the main steps for connecting the two halves of the furnace during construction. Twice before Aluar, this type of equipment and methodology was used successfully for anode baking furnace expansions: in 1979 at Sabart, and in 1986 at St Jean de Maurienne, both in France.

Introduction Aluar Aluminio Argentino operates the only aluminum smelter in Argentina. It commenced production in 1974 with a nominal capacity of 140000 t Al/year. After the retrofitting of the original Montecatini cells, using an in-house developed technology, and two expansion projects, the plant has reached its present capacity of 430000 t Al/year (2010). During the last expansion project, a new open type baking furnace was built in order to fulfill the anode requirements from the pot rooms.

More recently this technique was used in 2008 at Tomago Aluminium Company in Australia to rebuild the oldest anode baking furnace, which was started up in 1983. The concrete tub was retained with some strengthening. The heat insulating and refractory were changed. The 74-section furnace was completely rebuilt in a record four months, from the beginning of demolition to laying the last top block. The project was carried out in two steps with one half of the furnace being rebuilt while the other half remained in operation with the use of a temporary crossover. This reduced the need for inventory and external purchases of baked anodes.

The furnace was built in two stages. The first 34 sections, 8 pits, 192 anodes per section stage, was dried out during September 2007. The anode handling equipment and the firing control system, for this two fires furnace, were commissioned during the drying out process. In the second stage, the furnace was extended by a further 34 sections. The drying out process of these sections started in May 2009. The result is a furnace with 68 sections and 4 fires.

Although this type of equipment and methodology were used in these projects, some adjustments and improvements from the originally considered procedures were developed to reduce the impact in the furnace operation. The purpose of this paper is to describe the experience of operating an open type furnace with a temporary crossover during 6 months. As some impacts in the baking process behavior have been observed under these operating conditions, the actions which should have been taken to minimize their effects are described.

The end sections of the first stage of the furnace have been modified taking into account the work required for the erection and connection of the second stage. Therefore, the end headwalls, heat insulated sidewall in sections 52 and 17 (see furnace scheme in Figure 1) have been designed to reduce and facilitate the modifications required for the furnace extension. The construction of stage 2 began at the "new" end of the complete baking furnace, namely sections 34 and 35 (Figure 1), and progressed towards the center of the furnace, in order to minimize furnace operation disruption. The stage 2 works, while the stage 1 furnace was in operation, created many operating constraint which will be described in the following paragraphs.

847

Before installing the temporary crossover, the stage 1 furnace needed to be prepared and modified. Sections 16, 17, 52 and 53 were fully loaded with baked anodes and packed with coke. Therefore, four sections of the first stage furnace were not used in the anode baking process during the operation with the temporary crossover. The second top blocks of section 16 and the third top blocks of section 53, as well as some bricks were removed, in order to allow the vertical ducts of the temporary crossover to be installed.

Temporary crossover In an open type furnace, the gases flow along each flue wall line, from the blowing ramp and even the cooling ramp, up to the exhaust ramp. The regulation of the negative pressure downstream of the heating zone and the high pressure in the blowing zone allow the control of the flow inside the flue walls and the fire progression speed. The typical design of the open type furnace has two metallic ducts, with the inner part lined with insulating materials, which are located at each end of the furnace. These devices, called crossovers, are used to collect the gases from each line of flue walls and to conduct themfromone furnace half to the other.

Once the brickwork modifications were finished, the temporary crossover was installed. The metallic connections with the flue walls were not in direct contact with the bricks, allowing free expansion and movement of the different materials. Pressed refractory fiber was used to ensure the tightness between the vertical metallic connections and the flue walls. Figure 3 shows the erection of the temporary crossover after finishing the brickwork modifications.

Before the erection of sections 18 and 51, i.e. the "new" sections adjacent to the "old" ones, it was necessary to remove the crossover between sections 17 and 52, and to demolish the end insulation. Modifications of the headwalls between the "old" and "new" sections were also required. In order to allow the normal operation of stage 1 furnace, a temporary crossover was installed between section 16, second flue wall peephole and section 53, third flue wall peephole. Figure 2 shows a scheme of this configuration Fire direction

The gas passage at the central baffle was closed with dense bricks, separating in this way the end sections from the rest of the furnace. In order to avoid any flow of gasses into the furnace, expandable baffles were located at the headwall windows, close to the fourth line of peepholes of section 16 and the first line of section 53.

15

Figure 3: Erection jobs of the temporary crossover. Figure 4 shows a sketch of a flue wall configuration in section 53. The gas flow is indicated and the brickwork and expandable baffles "blocks" are shown.

Fire direction »- 54 Figure 2: Schematic view of the location of the temporary crossover on the dead sections.

It is important to point out that many actions were taken in relation to the interlocks of the movements of the central conveyors and with the furnace tending assembly, in order to ensure a safe operation while the temporary crossover was in operation.

The temporary crossover was made of four metallic separated square section segments. These parts were joined with expansion seams and internally insulated, in order to protect the metal sheet from the high temperatures and reduce the heat loss. The segments were laid with metal legs on the corbel casing and the flue walls. They were connected to the flue walls by rectangular ducts.

848

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Decreasing of soaking temperaturefrom1180°C to 1160°C. As the total soaking time was increased in 12 hours, in order to compensate this "extra" energy input the soaking temperature was reduced in 20°C. The parameters of the original and modified cycles are shown in Table I. No impacts on the baking level or other properties of the baked anodes were observed.

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Figure 4: schematic view of the modifications at a flue wall (downstream side) Steps before installation Before installing the temporary crossover, several modifications to the operation and process control routines were made. While the anode baking furnace worked with 34 sections, the nominal fire permutation cycle was 24 hours, with 2 fires of 17 sections each. In order to remove the four end sections from the process, it was necessary to create afirewith 15 sections and another with 19 sections. This meant it was necessary to "separate" the end of one firefromthe first section of the other. To do this, the cycle time of the two fires was changed to 26 and 28 hours respectively, during approximately 45 days. Once the "long" fire acquired its configuration and reached the end corner of the furnace, the temporary crossover was installed. The cycle time of the two 15 sections fires was then set at 28 hours.

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Updating of the Fire Control System program. The SCADA system which managed the firing control system was modified in order to operate with a 30 section baking furnace with two "short" fires. In normal operation, this system, which was supplied by Innovatherm, uses special control algorithms to adapt to process requirements during "corner and pre-corner" situations. The purpose is to have an even temperature distribution along the flues just after the crossover channel. During the operation with the temporary crossover these modules were disabled, until knowledge of the impact of these devices on the process behavior was acquired. Operation with the temporary crossover

In normal operation, when the exhaust ramp passes from one side of the furnace to the other, draught set point values higher than normal are required, in order to compensate for the energy used to heat the crossover insulating lining as well as the additional pressure drop. In some furnaces the exhaust ramp is moved directly to one section after the crossover, where it remains during two cycles, increasing the heating time of the section after the crossover. This practice is not normally used at Aluar.

As a consequence of the operation with such short fires, the available area for the load/unload and routine maintenance was significantly reduced. Some measures were therefore taken in order to ensure effective cooling and improve the working conditions in those sections where refractory maintenance personnel had to work: Removal of one "blowing zone section", obtaining an extra empty section. The two fires configuration was kept with three sections for natural preheating and three sections for forced heating, but the blowing zone or the first cooling area was reduced from four to three sections. As a result, the total cooling time (blowing + forced cooling) was reduced by four hours. Zero point ramp control adjustment. The shortening of the controlled cooling area required the adjustment of the control parameters of the zero point ramp, decreasing the over pressure set point in the section upstream of the last heating ramp. In addition, the lids of two peepholes lines next to the blowing ramp had to be removed. After this change the maximum temperature measured on the surface of baked anodes during the unloading process did not exceed 350380°C, as in normal operation.

849

In the natural preheating sections, the flue thermocouples used by the firing control system are placed on the fourth peephole line of the first section, and connected to the exhaust ramp local panel. As was described above, an automatic temperature control algorithm follows the crossover situation adjusting the draft setpoint values with the object of minimizing the temperature deviation between the flue walls. As a result of the action of this module, at the starting of the baking process in the first section downstream the crossover, usually the temperature is around 280°C, increasing up to 800°C by the end of the cycle. In concordance with this temperature evolution, the draught set points in the exhaust control loop value varies from -140 Pa to 170 Pa.

First movement through the temporary crossover.

Visual inspections into the flue walls showed an "abnormal" behavior of the degassing front. Normally the progression of the fire front along the flue wall, following the internal "channel" defined by tie bricks and baffles, is observed. In this case, a "quasi instantaneous" ignition of the pitch vapors along the wall occurred. As a result, only the upper part of the flue walls showed smallflameswhile the bottom part remained "dark".

As soon as the operation with the temporary crossover was started, some important differences with the normal crossover situation were observed: Very low heat-up rate in the section after the temporary crossover. The mean flue temperature measured upstream and downstream the temporary crossover showed a difference between 450 and 500°C The pressure drop measured with manual devices between both locations was 90-120 Pa. The position of all the dampers in the exhaust ramp remained almost completely open, resulting in a poor draught control. High temperature deviation between flue walls was observed Maintaining the negative pressure and flow from the Fumes Treatment Center (FTC) was difficult..

Figure 5 shows the fume temperature evolution (average of the nine flue wall temperatures). The steps in red indicate the number of the section where the exhaust ramp was placed (on the right axis). As a reference, the data from the cycle before the temporary crossover are included. It can be observed that just after the finish of the fourth consecutive fire advance, the fumes temperature reached standard values. At the end of the baking, the maximum fume temperature in the head walls on the section downstream the temporary crossover was in the order of 950°C. This indicated a very poor baking quality of the anodes baked in that section.

The initial attempts to improve the process conditions were directed towards the decrease of the "fresh" ambient air infiltrations from the "dead sections". The upper part of the packing material was removed from all the pits, and a layer made of pressed ceramic fiber was installed and covered with packing coke up to the original level. In addition, the ring main fume collector draught set point was increased at the FTC control loop. Visual inspections of the flue walls in the section where the crossover was located, showed that a very large part of the heat coming from the sections under fire, was being used to heat the "dead sections", before reaching the operating sections. Excessive heat loss through the crossover was discarded as a cause due to the low temperature measured on the outer metallic walls. At the time for the next fire advance, the temperature measured in the first section was around 350°C. This value was approximately 450°C below the target. In spite of this abnormal situation, the fire advance was performed as usual.

Figure 5: Exhaust temperature evolution during the first advance through the temporary crossover Second and subsequent advances through the temporary crossover.

It was evident that the heating rate downstream of the temporary crossover was extremely low. To increase the heat up gradient in the preheating section, a ramp using the start-up burner kit was installed in the fourth line of peepholes of section 54, downstream the temporary crossover on the second section under preheating. For safety reasons, those burners had a UV detection device to monitor the flame. They were manually controlled.

To mitigate the difficulties in following the heating curve in the natural preheating sections, several actions were adopted for subsequent baking cycles: The Section downstream the temporary crossover was not unloaded. It was evident that the anodes in this section were baked at very low temperature. In addition, keeping the already heated anodes and packing material inside the sections, could help to reduce the heat requirement for the next cycle.

It was decided to delay the next fire advance until the anode degassing process started on section 54. This situation was reached 55 hours after the previous movement. At that time, even when the temperature measured in the flue walls was lower than expected, the fire advance was performed (see Figure 5).

Reinforce the sealing of the "dead sections". The upper part of the packing coke and the ceramic wool placed on top of sections 16 and 53 were removed. The packing material was leveled with the top blocks and a metallic cover was placed on the top of each pit. The seams between these covers and the top blocks were sealed withfiberwool.

The "extra" gas injectors helped to heat the flues in the preheating sections, but increasing the draught in order to keep a minimum negative pressure at the heating ramps was necessary. It was evident that these actions should be carefully balanced. In fact, the preheating sections should be heated at a reasonable velocity but not so fast as to cause excessive temperature increase at the exhaust ramp. It was found that once the degassing started, it was very difficult to control the fume temperature increase. The only available action in case of excessive temperature at the exhaust ramp was to move it to the next section earlier than scheduled.

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Increase the flow of gas injection. A manifold with the nine start-up burners was assembled on the temporary crossover. The injectors were located on the first line of peepholes of

partially compensated by the additional combustion at the supplementary burners and the heat remaining in the load.

section 53, which was out of the baking process. The picture in Figure 6 was taken during the assembly of these burners. Another set of nine start-up burners was located in the fourth line peepholes of section 54, as before. Both lines of injectors were then on sections already loaded with baked anodes. They were controlled manually. These start-up burners were part of the equipment available for the drying out of the furnace.

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Double move of exhaust ramp. The fire advance procedure on the temporary crossover was modified. In the new procedure, the exhaust ramp was moved directly from section 15, before the temporary crossover, to section 55. Then, the next exhaust ramp advance would be delayed by two cycles. The other firing control elements were moved as usual. The effects of this procedure were described earlier.

A recommended procedure for anode baking level control is to measure the anode temperature during the gas heating stage. Figure 8 shows the evolution of gas and anode temperature in section 55, this is to say the one in which the anode baking quality could be more severely affected by the use of the temporary crossover. For comparison, the values for section 2, which was in an equivalent position to the normal crossover, are shown. As can be seen, there are no significant differences between both sections. In this figure, the "gas temperature" is the average of the nine fume temperatures measured in each flue wall, while the "anode temperature" is the average of the three anodes located on the top layer of three pits.

Control routine monitoring: Many unusual control routines to be performed by the furnace operators were programmed, including follow-up of start/stop events of the start-up burners kit, manual adjustment of the draught set point in the exhaust ramps, temperature measurements between crossover ends. This set of actions gave the expected results and the baking process quality in the preheating sections showed an important improvement. Figure 7 shows the average fiimes temperature measured at the exhaust ramp control during four fire advances through the temporary crossover. The values correspond to the second and fourth passage through the temporary crossover. The only section with a final temperature significantly lower than in normal baking is the one downstream the temporary crossover, which was loaded with baked anodes. The baking curves of the other sections normalized progressively and for the fourth advance through the temporary crossover they were completely normal, reaching 800 °C as expected. These values, as well as the shape of the curves, confirm that the preheating of these sections was normal.

-Sec. 55-Gastemp - Sec. 55 - Anode temp

- Sec. 2 - Gas temp ~ Sec. 2 - Anode Temp

Figure 8: Evolution of gas and anode temperature in sections 55 and 2.

As was mentioned before, section 53 was not unloaded and remained with baked anodes. No degassing occurred in this section, the lack of energy from the burning of pitch vapors was

851

Baking temperatures

Conclusions 1

Aluar measures the baking temperature using the Lc method [l] . Each section is loaded with two samples of green coke stored in a graphite crucible, one on the anode bottom layer and the other on the anode top layer, both on one of the outer pits. After the commissioning of the firing control system, the higher temperatures were obtained on the top layer.

The experience of an anode baking furnace operation with a temporary crossover was described. It could be concluded that after some adjustments in process control parameters as well as in the operative routines and procedures, there is no significant impact on the quality of the anodes baked in the sections downstream of such device.

In order to compare the impact of the crossover operation during the complete baking process, the baking temperature of the anodes baked in the sections downstream of the normal and temporary crossovers are shown in Figures 9 and 10 (average ± standard deviation) . In the top layer, the difference between both groups of sections was not significant.

In the case described in this paper the procedure allowed the construction and start-up of a four fires baking furnace in two separate stages. The second half of the furnace was erected and connected to the original one with a relatively small loss in production. The good results that have been obtained allow consideration of the use of similar equipment for other applications, for examples, large brickwork repair jobs on part of a furnace or complete furnace rebuilding in stages as was carried out for the oldest baking furnace at Tomago. The main problems that have been experienced were due to the increased flow resistance at the corner and to the energy required for heating the sections on which the temporary crossover was installed. Although it has not been experienced, the connection of the temporary crossover to the peephole lines closer to the active headwalls, may help to reduce these problems.

Figure 9: Baking temperature at the top layer

References

On the contrary, in the worst location (the so called cold position of the section) an important difference can be observed for the second section, where the impact of the temporary crossover is clearly visible. However, the values for the other sections are quite similar. Then with the exception of the cold part of the second section, all the baking temperatures fulfilled the smelter quality requirements.

1. J.F.Rey Boero, H.Alcantara, "Mediciσn de Lc de coque y temperatura de cocciσn por difracciσn de rayos X " Aluar internai reports ID/CL 450 and 451,2003 2. L.P. Lossius, I. Holden, and H. Linga, "The Equivalent Temperature Method for Measuring the Baking Level of Anodes", Light Metals 2006, 609-613.

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The relationship used by Aluar to convert Lc in Tb (baking temperature) was determined in such a way that the resultant value is close to the real maximum temperature reached by the anode [1]. Even when the procedure is similar, Aluar values are different to the "equivalent temperature" as described in the literature [2], which is based on a calibration with shorter soaking times. The numerical values of equivalent temperature are approximately 100-120° higher than the corresponding Tb values.

852

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

RECENT DEVELOPMENTS IN ANODE BAKING FURNACE DESIGN Dagoberto S. Severo1, Vanderlei Gusberti1, Peter O. Sulger2, Felix Keller2, Dr. Markus W. Meier2 *CAETE Engenharia. Rua Caetι 162, Porto Alegre RS, CEP 91900-180, Brazil 2 R&D Carbon Ltd. P.O. Box 362, 3960 Sierre, Switzerland Keywords: Anode Baking, Bake Furnace Design, Furnace Optimization furnace in such a way that all boundary conditions for optimum furnace operation as listed above are fulfilled. A similar approach allows identification of possible improvements for existing furnaces not performing as expected. Note that the specific energy consumption is not a characteristic furnace parameter that can be chosen freely but a result of the furnace design and operating parameters [2 and 3].

Abstract Today, furnace design still proceeds mainly by extrapolation from existing furnaces. Investigating existing furnaces shows the dangers of underestimating the impact of apparently small modifications: e.g., larger pits to accommodate higher anodes can result in furnaces with substandard performance. Side effects such as soot creation have then to be accepted. This paper presents an approach to bake furnace design which completely eliminates extrapolation from existing furnaces. The same approach can be used to estimate the optimization potential for existing furnaces.

Boundary Conditions to be Observed For typical anodes produced from calcined petroleum coke (CPC) and coal tar pitch, experience shows that a furnace should be able to fulfill the following boundary conditions in order to be able to properly bake all brands of raw material that may be expected: o Maximum anode heat-up rate of approximately 15 °C/hour, in order to avoid anode cracking.

Introduction This paper discusses the design of open top ring type furnaces (Figure 1) with typically six to seven sections in heat-up and 16 to 18 sections per fire. A modern and well-designed anode bake furnace should fulfill simultaneously the following goals [1]: o Produce anodes with high and uniform quality, fulfilling the requirements of the potroom o Production at design capacity even at the highest baking temperature. o Soot free combustion. o Lowest possible NOx generation. o Low maintenance cost, high flue wall life time.

o Maximum final anode baking temperature of 1150 °C, measured by the Xylene density method according to ISO9088. Note however that 1150 °C is an extreme baking temperature, possibly resulting in "over baking" with poor air reactivity properties as a consequence. Depending on the coke properties, a lower baking temperature in the range of 1050 °C to 1100 °C is most often sufficient without any significant quality loss but with the advantage of a lower energy consumption. However, the furnace has to be able to reach the maximum temperature level mentioned as certain CPC's may call for a high baking temperature level in order to reach the required quality level. Applying the method ISO-9088 is recommended, as experience shows that this method gives the most accurate results. The method is based on the fact that the Xylene density (or true density) correlates strongly with the baking temperature. In the practical application the method has to be calibrated by baking bodies of the same green formulation as the anodes in pilot plant furnaces to different baking levels. o Experience shows that for soot free combustion a minimum oxygen level of 8 % has to be maintained in all flues at all times and at all places. o Experience shows that for acceptable temperature homogeneity over the length and depth of a pit a "soaking time" (i.e. the time where the combustion gas temperature is held constant at the maximum level) should be not less than 1.5 times or preferably two times the nominal cycle time duration.

Figure 1 : Open top ring type furnace under construction

o For refractory construction stability reasons and for anode deformation reasons, pit length and depth is limited to approximately 6 meters each. o For the sake of a high refractory service life, most refractory material suppliers ask for a maximum refractory surface temperature in the burner area in the range of approximately 1250 °C to 1320 °C. In a standard arrangement with a

In this paper, two different situations will be discussed: o Design of new furnaces. o Optimization of existing furnaces. At TMS 2010 a paper has been presented discussing the factors influencing the specific energy consumption of anode bake furnaces [2]. This document now presents a method to design a

853

thermocouple one baffle in front of a burner, a maximum refractory surface temperature of 1300 °C will typically be reached with a maximum combustion gas temperature somewhere in the range of 1150 °C to 1250 °C. Note that the acceptable refractory surface temperature is significantly lower than the maximum service temperature of the bricks applied as given in the data sheets. The maximum service temperature will typically be in the range of 1400 °C to 1500 °C.

calculation results of fluid flow patterns, anode temperature distribution and gas temperature in the fire section. The 2D global model was developed by CAETE to study the entire baking process simultaneously. Balance equations for heat transfer, species concentrations and pressure distribution are solved by a dedicated software, built up in a Fortran platform. The anode baking furnace can be understood as a counter-flow heat and mass exchanger. The gases flow from the cooling sections to the preheating sections at a certain velocity determined by local pressure, temperature and infiltration conditions.

o To keep the required ring main under pressure within reasonable limits, the under pressure in the flues in the first peephole just upstream of the exhaust manifold is typically limited to approximately 400 Pa. From the customer, the following information has to be submitted: o Maximum and minimum output in tons per year, defining the shortest and longest cycle time to be considered. For a given furnace and anode load the output is a function of the cycle time only, with the highest output determined by the shortest cycle time.

Gas flow

o Anode size, nominal and maximum. Experience shows that more often than not, the reduction plant will ask for longer and/or higher anodes. If this has not been taken into account when designing the furnace, experience shows that increasing the anode size will have a significant negative impact on pitch volatile matter combustion and on refractory maintenance cost. Proposed Furnace Design Approach Today, furnace design is all too often still done by extrapolating from existing furnaces, complemented by Computational Fluid Dynamics (CFD) calculations for the flue cavity design [4] among other aspects [5] and [6]. With a trend to higher anodes and thus wider pits simply adapting existing designs may result in unsatisfactory furnace behavior. This document describes a new approach for the determination of all key dimensions required to design a new furnace. It is assumed that the anode size, the required output and the desired maximum anode baking temperature is all that is given as a "starting point" for the calculations aimed to design the furnace. Unfortunately, the number of variables is too high to calculate all possible combinations of key dimensions. Extensive test calculations on existing furnaces allowed to prove that the job could be massively simplified by: o Treating several variables as constants. o Determination of furnace parameters by iteration. After years of development and validating the calculations on a significant number of anode bake furnaces with different designs, the approach can now be launched for application in the industry. Furthermore it could be proven that a simplified approach also allows an estimate of the optimization potential of existing furnaces.

Figure 2: Gas flow streamlines and anode temperature distribution calculated by the 3D model

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Numerical Models Used The numerical models used in this paper are a combination of a 3D detailed model (1 section) with a global (all sections) 2D mass and energy balance model. The 3D model was presented in [4] using a computational commercial code. It is useful to study gas flow patterns, pressure drop and thermal insulation options, considering detailed geometry. Figure 2 and Figure 3 present examples of 3D

The solids "flow" in a discrete step equal to 1 section per cycle time in the opposite direction of the gas flow. This model is inspired in the models presented R.T. Bui et al. [7] and R. Ouellet et. al [8], incorporating some improvements. The real transient effect of the fire and manifold movement is taken into account as the model is a true transient model inside each fire cycle and the fire movement occurs in a discrete way. The finite volume method is used for the evaluation of the balance equations.

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The global model uses a 2-dimensional mesh representing a horizontal slice of the sections directly involved in the process (preheating, firing, cooling). Heat conduction inside the solids is calculated in 2D, but heat balance, oxygen concentration (and also CO2, H20), inside the flue is evaluated in one-dimensional form, as it is considered that variations are small in the flue width direction. Heat losses to environment occur at furnace top, as well part of the heat is lost to the ground. These losses are taken into account in the modeling by imposing appropriate heat loss coefficients.

necessary cycle time and soaking time for a given baking temperature. In a second step, the availability of the draught required to produce soot free combustion under all circumstances is checked. The fuel consumption for this situation is also calculated. Soot free combustion is generally achieved if the oxygen concentration is 8 % or more in all heating sections at all times. A second calculation loop may be required, including the first step, if the oxygen check shows that the required minimum concentration cannot be achieved with an acceptable level of draught or if the fuel consumption is too high. Graphically, the calculation sequence is shown in Figure 5.

Treating Variables as Constants Preliminary calculations have shown that a significant number of variables can be treated as constants, greatly reducing the number of alternatives to be considered. As an example, it can be shown that specific energy consumption decreases when increasing anode pack length (Figure 4) and height. Therefore, these variables are limited by the maximum allowable value compatible with refractory stability. The variable will thus be set to the limit given by the refractory restrictions.

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Similar considerations are valid ibr parameters such as pit height, top packing material layer thickness, headwall thickness, maximum combustion gas temperature and maximum refractory surface temperature. Furnace Dimensioning by Iteration Within certain limits, the number of sections in heat-up and the number of pits per section can be varied. Typically, six to eight sections in heat-up with three to four sections equipped with burner bridges will be found. The number of sections in heat-up will have an impact on properties as, e.g. the specific energy consumption, the optimum flue cavity width and the required under pressure. Separate calculation runs will be required for different flue designs and flue cavity widths. As the amount of computer time per run is within reasonable limits, this approach is considered adequate. In a first step it will be possible to identify the optimum configuration regarding the number of sections in heat-up, the required flue cavity width for a given flue design and the

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Figure 5: Basic calculation flow sheet As an example, the significant impact of the baffle design and of flue cavity width on the minimum O2 concentration in the flues for different under pressure levels has been calculated. Regarding baffle design, typically one of the two arrangements shown in

855

Figure 6 will be found. In the design on the left side, at the first and third baffle an opening with the height of one brick (approximately 0.09 m) is arranged. In the design on the right side no such gap is provided. Figure 7 shows the impact of the baffle arrangements depicted in figure 6 on the gas flow pattern.

Depending on the furnace construction the O2 concentration may vary heavily over the duration of the fire cycle. Figure 9 shows an example of the 0 2 distribution over the total heat-up area and over the full cycle time in intervals of 6 hours, for a configuration with 6 sections in heat-up. In the example shown in figure 9 the target flue gas temperature is reached at the end of the cycle time and in the section equipped with the back burner bridge only (i.e. between the points 5 and 6 on the x-axis). The extra energy required in this section is the reason for the steeply decreasing oxygen concentration. The fast rate of oxygen level decrease observed between the points 1 and 2 on the x-axis is a result of the combustion of pitch volatile matter released from the anodes. In the first section upstream of the exhaust manifold (i.e. between the points 0 and 1 on the x-axis) most often an increase of the oxygen concentration is observed, due to fresh air infiltration. If pitch volatile matter combustion is already observed in the section mentioned (typically in the last few hours before the fire change) a decreasing oxygen content may result, as depicted in the 24 hours line shown infigure9. Similar calculation runs will be made for different combinations of parameters until a satisfying combination has been found.

Figure 6: Typical baffle arrangements; left with gap on top of first and third baffle, right with closed baffles

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Figure 9: Oxygen concentration over heat-up area and one fire cycle, for a given baffle construction and draught level. Sections counted from the exhaust manifold Impact of Higher Anodes on the Furnace Design In the past, reasonable results regarding furnace behavior have been achieved by extrapolating the dimensions from existing furnaces. Such an approach is of course fast and cheap. There is however a significant risk of a pitfall. For example, in the last few years a trend to higher and higher anodes can be observed. In most furnaces the anodes are set in the pits with anode top and bottom facing the flue walls. In order to accommodate such anodes, the furnace design results in larger pits. Analysis of existing furnaces has shown that the impact of larger pits on the cycle time is often underestimated. Figure 10 shows the massive impact of the pit width on the shortest possible cycle time required to reach a given anode temperature. In wider pits more time is needed by the heat wave

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856

to penetrate into the center plane of the pit. Increasing the cycle time is then the only option to reach the required baking level. Such an action has however as drawback a lower production as the furnace output is inversely proportional to the cycle time. In order to reach simultaneously a sufficient time for the heat wave to penetrate the pit and to maintain an acceptable output (i.e. to operate the furnace with a "short" cycle time) increasing the number of sections in heat up from six to seven or even eight sections instead of increasing the number of fires may be a solution. The different configurations will also differ in the achievable specific energy consumption. It is a wise and cost effective approach of the furnace buyer to request from the supplier of the furnace and/or of the firing system during the bid phase a numerical furnace simulation to prove that the furnace is functional under the current and future conditions.

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Anode temperatures inside the furnace were measured during the process by thermocouples installed inside the packing coke between fluewall and anodes. The Figure 12 shows the comparison between the averaged measured and model temperatures. Note the excellent agreement between the curves including the volatiles peak (after ~60h).

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Figure 10: Impact of pit width on the required cycle time; six sections in heat-up Discussion of Predicted and Achieved Results The furnace operation data and measurements from several plants cooperating with the authors have been compared with the behavior predicted by numerical models in different aspects. These data were used for the fine tuning of the models. The Figure 11 shows measured underpressure inside the flue cavities versus calculation results. The calculated underpressure profile was obtained by the 3D model using appropriate resistance parameters between flue chamber and atmosphere.

0 24 48 72 96 120 144 168 192 216 240 264 288 Time [h] Figure 12: Comparison of measured and calculated temperatures inside the packing coke during the baking process Designing New Furnaces Using the calculation tools available today, it is possible to design and to predict the behavior of a furnace in such a way that the expected results can be achieved as well as regarding output, quality, energy consumption and combustion quality. The supplier of the furnace and/or the firing system should be requested in the bid phase to prove with a numerical furnace simulation the correct functionality under the current and future conditions.

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8. R. Ouellet, Q. Jiao, E. Chin, C. Celik, D. Lancaster and D. Wilburn, Anode baking furnace modelling for process optimization, Light Metals (1995), 653-662.

Optimizing Existing Furnaces If a furnace does not perform as expected, the tools developed allow pinpointing of the critical items and identifying possible actions to improve the situation. As the "starting point" is an existing furnace, the number of variables is extremely small. Measurements on the furnace can be used as input values, reducing heavily the amount of calculation work to be done. As an example, in designing a new furnace, the under pressure profile from the exhaust manifold to the zero pressure point upstream of the back burner bridge has to be calculated. In an existing furnace, this profile can easily be measured and applied as input for the calculations. Knowing the possible improvements makes it much easier to justify the required optimization tests [9].

9. Vinicius Piffer et al., Process Optimization in Bake Furnace Light Metals (2007), 959-964

Conclusions A new approach regarding the design of new anode bake furnaces has been developed. Although the baking process is quite complex, it has proven possible to simplify the calculations by treating some variables as constants. Key dimensions can then be optimized within a reasonable number of iterations. In doing so, a furnace can be designed, to fulfill all important goals, i.e. output, quality, energy consumption, soot free combustion and operating cost at the same time. The supplier of the furnace and/or the firing system should be requested in the bid phase to prove with a numerical furnace simulation the correct functionality under the current and future conditions. Finally, this same approach can be used for the estimation of the optimization potential of existing furnaces. Acknowledgements The authors thank all the plant managements allowing to perform measurements on their furnaces in order to validate the furnace calculation program described. Taking into account the secrecy agreements to be observed, naming the plants is not possible. References 1. Felix Keller and Peter O. Sulger, Anode Baking (Sierre, Switzerland, R&D Carbon Ltd., 2008) 2. Felix Keller, Peter Sulger, Markus Meier, Dagoberte S. Severo, Vanderlei Gusberti, Specific Energy Consumption in Anode Bake Furnaces, Light Metals (2010), pg. 1005-1010. 3. Markus Meier, Influence of Anode Baking Process on Smelter Performance (Aluminium 1-2/2010) 4. D.S. Severo, V.Gusberti, E.C.V. Pinto, Advanced 3D Modeling for Anode Baking Furnaces, Light Metals (2005), pg. 697-702. 5. Felix Keller, Ulrich Mannweiler and Dagoberto S. Severo, Computational Modeling in Anode Baking, 2 nd International Carbon Conference, China (2006). 6. Frank Goede, Refurbishment and Modernization of Existing Anode Bake Furnaces, Light Metals (2007), 973 - 976 7. R. T. Bui, E. Dernedde, A. Charette, T. Bourgeois, Mathematical simulation of horizontal flue ring furnace, Light Metals (1984), pg. 1033-1040.

858

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SOHAR ALUMINIUM'S ANODE BAKING FURNACE OPERATION Said Al Hosni, Jim Chandler, Olivier Forato 1 Franηois Morales2, Jean Bigot, Christian Jon ville3, 1 Sohar Aluminium, PO Box 80, Postal Code 327, Sohar Industrial Estate, Sohar, Sultanate of Oman 2 Rio Tinto Alcan, LRF, BP 114, 73303 Saint Jean de Maurienne Cedex, France 3 Rio Tinto Alcan, Aluval, Centr'Alp, BP7, 38341 Voreppe, France Keywords: Anode baking furnace, energy consumption, refractory performance Abstract The Sohar Aluminium anode baking furnace was successfully commissioned in 2008, and furnace performance since has remained at excellent levels in terms of gas consumption, baking level, fire productivity, tar emissions and firing cycle range. Ten fluewalls per section result in furnace productivity levels of 70 kt per fire group. Gas consumption levels under 1.9 GJ/t have been maintained for a baking level (Lc) of more than 34 angstrom. Operation has been demonstrated across a fire cycle range from 24 to 36 hours. Refractory condition is excellent, and first generation refractory life is expected to achieve >170 fire cycles due to a thorough maintenance program and a very low anode sodium concentration of less than 200 ppm. Some of the challenges in achieving these results are discussed. These results stand for a combination of design, process control and operation which place it amongst the benchmark furnaces of its type today.

These design characteristics in turn deliver a productivity level of 210 kt baked anode with three fires. They reduce the ground surface area to approximately 150 m by 35 m and minimize the capital expenditure. The 9 pit configuration and the optimized pit sizes lead to a high anode/refractory ratio. This ratio is a measure of the refractory impact on energy consumption [1]. Figure 1 presents the evolution of this ratio from 1980 until today for open type AP baking furnaces. Sohar has the highest ratio which contributes to low energy consumption in the furnace.

Introduction Sohar Aluminium, a joint venture between Oman Oil (40%), Abu Dhabi Water and Electricity Authority (ADWEA) (40%) and Rio Tinto Alcan (20%) operates a greenfield aluminium smelter in the Sultanate of Oman. This smelter, started in 2008, operates 360 reduction cells using AP 35 technology and produces an Aluminium tonnage of 374 kt per annum at current operating amperage of 370 kA.

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Over 200 kt per year of carbon is needed to supply the 360 reduction cells. The carbon plant employs the latest AP technology and includes a 52 section gas fired horizontal baking furnace comprising 3 fires equipped with an Innovatherm firing system and operating at a 24 hour fire cycle at full capacity.

Figure 1 Ratio anode weight/refractory and insulation weight The 9 pit configuration also results in a long headwall which increases the risk of distortion in the event of uncontrolled expansion. An innovative headwall expansion joint design [2] as shown in Figure 2, limits the elongation of the headwalls, while at the same time allowing for brick expansion. Moreover, these joints are designed to protect against the infiltration of packing coke.

Furnace design characteristics The Sohar furnace stands out from other furnaces through a number of innovative features that have led to excellent results. The Sohar furnace is characterized by the following specific points, namely: • • • • • •

Sohar

A section design with 9 pits and 10 flue walls per section.. A very efficient pit packing geometry to maximize the amount of carbon in the pit. Innovative headwall expansion joints, Afluewall designed using proven modeling techniques, The ability to achieve required baking levels by the use of a 4th burner ramp on shorter cycles when necessary. The ability to operate at a fire cycle range of 24 to 36 hours by the application of a new process control methodology that favors optimum combustion

Figure 2 Headwall expansion joint design concept

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The fluewall design was developed using a computational fluid dynamics model [3] to ensure baking level homogeneity. Figure 3 shows the modeled flow patterns.

Figure 5 View of the Sohar baking furnace Operational Overview The Sohar Anode Baking Furnace is one of the most efficient and environmentally friendly furnaces in operation. Application of a new fire process control methodology ensures quicker and more long-lasting achievement of performance levels, superior to that of other furnaces with the same characteristics.

Figure 3 Sohar Modeled Flue Design Furnace Start-up Furnace drying commenced in the April 2008 and was immediately followed by the start-up of the 1st production fire. The drying fire was then converted into the 2nd production fire. The final fire was started in November 2008 in accordance with line needs and at the end of 2008 all three fires were in operation. The program is illustrated in Figure 4.

The new methodology developed at Sohar ensures a complete and effective combustion of the volatile matter and injected fuel. The entire energy potential generated by volatile matter and injected fuel is recovered regardless of the cycle (24 to 36 h).

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This methodology demonstrates significant advantages when compared to furnaces applying a conventional process. These advantages are detailed as follows.

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In a conventional horizontal baking furnace process, irrespective of the technology used, part of the volatile matter from the anodes and, in some cases, injected fuel partially escapes with the fumes. Residue of the unburnt matter may then be deposited on the FTC collector walls, and there is a loss of potential heating value requiring compensation by the amount of injected fuel necessary to reach final anode temperatures and soaking times.

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Figure 4 Planning of the furnace start-up An overall view of the furnace is shown in Figure 5.

Optimum operation of the conventional process depends on a sensitive balance of the combustion of the volatile matter in the preheat zone and the injected fuel in the forced heating zone of the furnace The methodology applied to the Sohar furnace allows the control of both combustion zones separately. This separation makes it possible to: 1. Ensure sufficient oxygen level for both combustion zones at all times regardless of the duration of the firing cycle 2. Ensure complete combustion of the volatile matter released by the anodes from a temperature of > 200°C by having the circulating gas in the flue walls at temperatures > 700°C in the preheat zones.

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A comparison on the conventional heating curves to the new methodology is shown in Figure 6.

leading to the fume treatment centre after two years of operation with no requirement to perform regular cleaning.

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Figure 6 Comparison of heating curves The sustainable complete combustion of the volatile matter achieved by the new methodology (shown in Figure 7), as opposed to a conventional methodology, results in the following benefits: Improved furnace energy efficiency, leading to a reduction in the fuel quantity needed to achieve the same anode baking level. Improved environmental performance by ensuring that volatile matter is completely burnt inside the flue walls, thus reducing emissions limiting carry over and condensation of volatile matter in the collector duct work and fume treatment centre. Figure 9 Internal condition of collector duct Firing Cycle Flexibility The baking cycles have been adapted to satisfy the reduction line start up needs from the 1st to the 360th cell. Figure 10 shows the evolution of the firing cycle times since start up.

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A complete inspection of the fume treatment centre completed every 6 months confirms this satisfactory control of the process. No trace of deposits, combustion residues or soot has been detected since the start up of the furnace. Figure 8 shows the good opacity control measured at the exhaust ramp on one of Sohar fires, Figure 9 shows the condition of the collector duct

Sohar operation is not limited by firing cycle duration with the flexibility to operate at either long or short firing cycles. The degassing front position was perfectly controlled while guaranteeing the combustion of the volatile matter and of the fuel injected by the ramps.

861

This is due to be able to combine the new process control methodology and to make targeted use of a 4th burner ramp for a few hours at the end of the cycle for the shorter cycles to ensure that final anode baking temperature and corresponding baking levels are achieved.

theoretical gas consumption calculated with a thermal balance model [l1. The theoretical gas consumption considers that all volatile matter has burnt and has recovered this energy to bake the anodes.

Refractory The current flue wall age currently is in the vicinity of 40 fire cycles (August 2010). The condition is shown in Figure 11. At present there are no signs indicating any early deterioration of the furnace, leading to a predicted life expectancy beyond 170 fire cycles.

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A number of factors contribute to the increasing flue wall life span. The first factors are furnace design and the quality of the bricks used. This is followed by the low sodium level in the anodes, which has been consistently maintained below 200 ppm (Figure 12), homogenous combustion in the flue walls eliminating localized over heating, and sealing carried out after each fire move.

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Figure 13 Furnace energy consumption Technical Performance Overview The Sohar furnace has achieved world class performance on a number of technical indicators detailed as follows: Baking level: Baking levels (shown in Figure 14) have been maintained at a level comfortably above the minimum required to ensure a high anode quality resulting in good performance in the cell.

Figure 11 Refractory condition after 40 fire cycles Figure 14 Baking level results Anode reactivity: Anodes exhibit excellent reactivity. Carbon dioxide (Figure 15) and air reactivity residues are above 92% and 80% respectively.

Figure 12 Sodium content in baked anodes Energy consumption All these elements have resulted in the Sohar furnace achieving an average energy consumption level of 1.9 GJ/tonne baked anode despite a high baking level (Lc > 34 A). Further improvement in 2010 has reduced the consumption to 1.8 GJ/tonne baked anode as shown in Figure 13. The current gas consumption is close to the

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The electrical resistivity is lower than 54.5 uD.m and the air permeability (shown in Figure 16) is very good and had an average of < 1 nanoperm.

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Figure 18 Anode ahead of schedule results

^v# ^ν ^ >>w o°w° ^v# ^ν ^ ^ Figure 16 Air permeability

Conclusions Baking homogeneity is excellent. The modeled flow patterns for the flue design shown in Figure 3 results in a homogeneous pit profile expressed in real density from the operating furnace (see Figure 17).

A combination of design, operational and process performance has demonstrated the capability of the Sohar Aluminium anode baking furnace to:

DRfLc 20BD 2 Œ Q - 2Œ5 11 2035- 2090 2 0 9 0 - 2095 2096- 2100 2 1 0 0 - 2105 2105- 2110 1 1 2 1 3 0 " 21B ■ 2115- 2123 B 212D- 2125 2125

4 ROW

Deliver a furnace capable of productivity level of 70 kt baked anode per fire group Control both combustion zones to ensure complete combustion of all volatile matter coming from the anodes and the injected natural gas at long and short fire cycles. Deliver a high standard of baked anode quality with benchmark gas consumption. The quality of the baked anode supplied to the reduction line has ensured that no dust generation has occurred so far in the cells. This excellent performance and production capacity opens a new way forward to meet the needs for large and efficient furnaces necessary for high amperage reduction cell anodes.

References

Figure 17 Baking level pit profile

1: J.Bigot, M.Gendre, JC. Rotger, "Fuel Consumption: a key parameter in anode baking furnace" Light Metals 2007, p 965 2: International Patent Application WO 2007/006962, "Chamber setting with improved expansion joints and bricks for making same". 3: J.C Thomas, P. Breme, J.C. Rotger, F. Charmier, "Conversion of a closed furnace to the open type technology at Aluminium Bahrain", Light Metals 1999, p 567

Reduction line performance The anodes have performed very well in the reduction lines with an absence of dusting since the beginning at high amperage and very low ahead of schedule rates < 1% (Figure 18). Net carbon consumption is consistently < 415kg/t.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

MEETING THE CHALLENGE OF INCREASING ANODE BAKING FURNACE PRODUCTIVITY Franηois Ordronneau1, Magali Gendre1, Luc Pomerleau2, Nigel Backhouse2, Adam Berkovich3, Xin Huang3 1 Rio Tinto Alcan Laboratoire de Recherches des Fabrications, St Jean de Maurienne, France 2 Rio Tinto Alcan Arvida Research and Development Centre, Jonquiθre, Quebec 3 Rio Tinto Alcan Pacific Technology Centre, Brisbane, Australia Keywords: Carbon, Anode Baking Furnace, Optimization, Modeling, Productivity. ramp be necessary? Will it always be possible to bake and cool different anode formats sufficiently? Would it be possible to reduce the concrete casing insulation thickness? • As part of slowing down furnace baking cycle time what would be the adequate baking curve to maintain an anode baking level equivalent to the standard target one? • As part of design of a new furnace, which furnace dimensions should be chosen to satisfy the customer's production requirements at minimum cost and guarantee a reliable process, bearing in mind that pit width is fixed by the size of the largest anodes and that this furnace must be able to bake different anode formats?

Abstract The need to support amperage creep in smelters requires an increase in anode baking furnace productivity. The furnace operation must be adapted to provide more anodes while maintaining adequate baking performance, minimizing energy consumption and assuring anode quality. To do so, the essential factors to optimize include the firing strategy to ensure the adequate baking level for a given fire cycle, fuel combustion efficiency and adequate cooling capability even at accelerated fire cycles and high ambient temperatures. Through the operation of 28 baking furnaces of differing technologies, tools have been developed to support this process. Simulations of firing and baking as a function of the fire cycle and key furnace design parameters such as anode size, pit width and flue design; injector combustion efficiency and cooling capability are now routinely used. Industrial examples are shown from a number of sites with varying baking technology that demonstrate the gains achieved.

Thermal models have been developed to help solve these problems. These models guide customers in their choices and propose solutions to satisfy their requirements. They are applicable for all open furnace and heating equipment technologies. Model theoretical equations

Introduction Rio Tinto Alcan and its joint venture partners operate 28 anode baking furnaces of varying technologies and firing systems throughout the world producing more than 15 different anode formats. The continual increase of reduction line amperage requires increased anode mass and/or increased anode production to support the corresponding anode change cycles. This in turn requires increases in baked anode productivity (unit mass and production). Therefore anode baking furnace design and process criteria need to be optimized to ensure correct anode baking levels, baking homogeneity, optimum combustion and effective cooling. A variety of simulation tools have been developed by Rio Tinto Alcan to ensure that optimal solutions are selected to deliver the highest value in terms of investment and operating cost. These tools have been progressively enhanced during the course of their application at different operating sites with effective results. This paper describes these tools and gives examples of their utilization.

The model theoretical equations are general thermal equations: • The equation governing thermal conduction in materials (Laplace equation):

p(nCp(T^

= Ä.^T

+

^ί)(VTj (l)

Where: o p(T) : material density (kg/m3) o Cp(T) : material heat capacity at constant pressure (J/kg.K) o T : temperature (K) o t : time (s) o ë : material thermal conductivity (W/m.K) • The equation governing heat exchanges by convection (Newton's law) the origin of which is motion of a fluid in contact with a solid wall:

Heat transfer study

φ = h.{Tf - Ts)

A number of different situations can lead to the need to perform a heat transfer study. Examples of challenges that can arise are listed as follows:

(2)

Where: o φ : heat flow (W/m2) o h : exchange coefficient by natural and/or forced convection (W/m2.K) o Tf : fluid temperature (K) o Ts : solid temperature (K)

• What is the fastest baking cycle at which the furnace can operate? Can the production be increased further and in what conditions? • Anode height increases are required to support reduction line amperage increases and it is necessary to enlarge the furnace pits: Can the modified furnace continue to operate with the same baking cycle as at present? Will one additional heating

• The equations governing heat exchange by radiation based on the black body radiation equation (Stefan Boltzmann law):

865

(3)

Where: o o o

Increase of production needs through firing cycle optimization

φ : heat flow (W/m2) σ : Stefan Boltzmann's constant (=5.67.10φ W/m2.K4) T : body temperature

The following study was completed for Sohar Aluminium when the baking cycle was accelerated from 26 hours to 24 hours. The 24 hour cycle operation was studied by modelling to confirm and optimize the process and operating parameters. As the Sohar furnace has a relatively high cooling capacity, the only point to be checked was whether the furnace can effectively maintain an adequate anode baking level with at a 24h cycle.

Figure 1 shows an example of the heat transfer from fumes flowing through the flue walls to the anodes. Material layers

Knowledge of process characteristics at a 26 hour fire cycle allows the simulation of the 24 hour case. The capacity to burn pitch volatile matter was also considered in the study. Model inputs included flue wall, pit and anode dimensions and a gas temperature curve based on the process target temperature curves. The model then calculated the resulting anode temperature. Figure 3 shows the baking simulation results at 26 and 24 hour cycles for the same furnace. Final anode temperature corresponds to the temperatures measured on the site.

Fumes

Refractory brick

Packing material

Anode

Figure 1 : Schematic of heat transfer from fume through to anode during heating

/\ * /f^ y* *

1000

The following is an example relating to increasing anode size at Aluminium Dunkerque. The feasibility of increasing pit width by reducing the bake furnace concrete casing insulation thickness was studied. The main issue was whether or not furnace insulation would need to be replaced. Insulation samples were taken and sent for analysis to measure the thermal characteristics of the worn material. The resulting values along with the material thickness and a defined gas temperature curve were then used in a thermal model to ensure that relatively low temperatures were maintained at the interface between the insulation and civil works to avoid concrete deterioration.

2 800

//

3

I 600

A

y / A - - Fume temperatures - 26h Cycle L

400

■ - Anode temperature - 26h Cycle Fume temperature - 24h Cycle

200

r

Anode temperature - 24h Cycle 80

100

120

140

160

18

Time - Hours

Figure 3: Simulation of anode baking in the Sohar furnace for 24h and 26h cycles

Figure 2 shows the good agreement between the measurements taken at the insulation/concrete interface and the results of the model. It was then possible to reduce insulation thickness in the model to assess the potential savings in terms of pit width, and thus prove the feasibility of increasing furnace capacity without reconstructing the concrete casing or fully restoring the insulation. The design change has been incorporated into the upcoming major furnace rebuild. This has provided the potential for Aluminium Dunkerque to increase the baked anode production by 4.2%.

For a peak firing temperature of 1200°C, final anode temperature dropped by 20°C when the cycle was accelerated. To increase final anode temperature by 20°C with at a 24h cycle, either the peak firing temperature can be increased, or an additional heating ramp can be added to prolong peak firing, provided that exhaust capacity is sufficient. Table 1 shows the results of different simulations. In order to maintain an adequate baking level, the peak firing temperature needed to be increased from 1200°C to 1225°C, or the peak firing time needed to be extended by 8 hours at 1200°C with an additional heating ramp. The extended fire solution was used as it exerts less stress on the refractory materials and ensures greater flexibility for the process.

Measurements - Fire 1 Measurements - Fire 2 Model results

Peak fire Final anode temperature temperature 26h 52h 1200°C 1115°C 24h 48h 1200°C 1096°C 24h 48h 1225°C 1115°C 24h 56h 1115°C 1200°C Table I: Final anode temperature simulation results

Cycle

200

L '« ^^\^itjf^ ^^ * ^^*

jsf

Furnace design: pit enlargement case study

300

Time - Hours

Figure 2: Insulation - civil works interface temperature simulation

866

Peak fire time

The simulations determined the parameters required to ensure optimized operation at a 24 hour cycle and the solution has since been implemented [1].

1100 1000 900

Impact of reducingfluewallbrick thickness In another case study the impact of fluewall brick thickness on anode temperature was investigated for Tomago Aluminium Company. Reduction of the fluewall brick thickness by 10 mm was considered to allow final baking temperature to be achieved at faster fire cycles. The furnace was limited in achieving the peak fire temperature at the required soaking time. The site did not have the space to add a 4th burner ramp. The resulting increased pit width also allowed for potential future increases in anode size within the existing furnace footprint. There was also an added benefit of reducing energy consumption. Modeling was used to calculate an optimal heating curve to enable an equivalent final anode temperature to be reached during the baking cycle,

500 » Measurements 400 300

80

100

120

140

" Model after adjustment

Blowing 2

Blowing 3

Cooling 1

Figure 5: Cooling model anode temperature validation The simulation and field measurements show good agreement. The model was then used to investigate improved cooling configurations to deliver the best outcome for the client. Parameters such as ramp position, peephole opening/closing configuration, peephole dimension, ramp blowing power and flue wall brick thickness were studied. This in turn allowed a preselection of the most promising cases, thus considerably limiting the tests to be conducted on the site. An example of the Alma furnace cooling simulation is shown in Figure 6. It shows the calculated anode temperature profiles through the different zones for the initial case of a 27 hour fire cycle.

- Fume temperature - 24h Cycle - Brick =110 mm - Anode temperature - 24h Cycle - Brick = 110 mrr Fume temperature - 23h Cycle - Brick = 100 mm —Anode temperature - 23h Cycle - Brick = 100 60

* Model before adjustment

160

Time - Hours

Figure 4: Calculated temperatures for reduced fluewall brick thickness. Such a solution would allow for a production increase of around 4%. However these gains would need to be weighed up against the potential risk of reduced refractory life due to the narrower bricks. Cooling Optimization Increased furnace production is possible either by accelerating the baking cycle and/or increasing anode mass per pit for an equivalent baking cycle. The capacity of the furnace to cool the anodes adequately before unloading must also be determined. To assist in determining the cooling behavior for a given furnace an anode cooling model was developed.

Figure 6: Simulated anode temperature for the initial case Table II shows the results of different simulations at other fire cycles and cooling configurations on anode unloading temperature. The simulation suggested that a 30% increase of the blower flow would reduce anode unloading temperature by 30°C without impacting downstream process control. This solution was tested on site and validated by site measurements confirming an anode unloading temperature reduction of 25°C. As a result Alma was able to increase production of its furnace by 4%.

Alma works requested a study on the anode cooling phase to enable the acceleration of the baking cycle from 27h to 26h, which had led to an increase in anode unloading temperature of between 20°C to 40°C. Initially, the cooling model was validated by comparing the simulation and site measurements results in the initial cooling configuration. Figure 5 shows the correlation between the model results and the measurements. The anode cooling zone covers the upstream furnace sections from the heating ramps to the anode unloading zone.

Cycle 27h 26h

Cooling configuration

Calculated anode I unloading temperature 390°C 420°C

Initial Initial 30% increase of 26h 390°C blowing flow Table Π: Cooling simulation results summary

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In another case study for Tomago Aluminium Company, the cooling efficiency of one of the baking furnaces was studied to determine if sustained operation at very short fire cycles could be achieved. Of particular interest was the impact of different ambient air temperatures (summer versus winter) on cooling efficiency. Modelling was used to develop optimum cooling configurations including ramp and peephole positions. A number of the modelled cooling configurations were trialled while operating on a 21 hour fire cycle. Configurations included 1 or 2 cooling ramps and different numbers of cooling and unloading sections. Measurements conducted during the trials were used to further develop and validate the cooling model.

Figure 8: Velocity field flow simulation After optimization, velocities were homogenized in the flue wall by the addition of tie-bricks which also enhanced mechanical strength. Flow is globally better distributed in the fluewall zones and fewer dead spots remain. Normally addition of the extra tiebricks would have led to a 30% increase in pressure drop but optimization of the position of these bricks made it possible to maintain an equivalent pressure drop to that of the original design. Enhanced flow in this flue wall has improved heating homogeneity and combustion, thus increasing process flexibility.

5.3S*«Î2 5.65e-62 5.35e»82 « l e . 82

l§f

.«^

WÊÊÊÊÊÊÊÊÊÊÊÊÊÊÊÈËÈËÈm

Combustion: optimization of injector design

4.<ί»^2 Ole-82 Y 4.15.-K μ *

ut··»

An example from a study conducted for the Aluchemie site is shown in Figure 9. Simulations of the initial and optimized cases of continuous natural gas injection show the volumetric concentration of methane in air, which is the image of the mixture between methane and air.

Figure 7: Anode temperature u\ varying ambient air temperatures. The modelling indicated that increase in anode temperature would have a corresponding increase in packing coke temperature. The maximum manageable packing coke temperature for the bake furnace crane would be exceeded during summer conditions. Therefore, with the current cooling ramp capability sustained operation of the furnace at 21 hour fire cycle is not feasible during summer conditions, Figure 7.

Fire direction

Additional modeling was performed to investigate options for improving cooling efficiency, these included additional cooling ramps and increasing cooling flow. The results suggested that sustained operation at short fire cycles would be possible if the blowing and cooling systems were upgraded. Fluid flow and combustion studies

Figure 9: Gas injection simulations

Internal flow characteristics of the furnace flue walls are important parameters to take into account as they have a significant impact on baking homogeneity and combustion quality. Pressure drop is also critical as there needs to be sufficient suction in the flue wall lines to ensure proper transfer of volatile matter. However, it must not be too high to cause unwanted air infiltration into the furnace.

The initial case shows a poor gas / air mixture resulting in large areas of poor combustion at the top of the flue wall. The gas is not burnt in optimum manner and homogeneous heating of the flue wall is not achieved. In reality this results in the formation of incomplete combustion products and excessive energy consumption. The optimized injector design markedly improved the air / gas mixture; the methane is better dispersed in the injection zone and no poor combustion zones are present.

Models have been developed to ensure proper understanding of the related phenomena and resulting furnace optimization. These models allow definition of flue wall designs with optimized flows and gas injector performance. They are applicable for all open furnace and heating equipment technologies.

These simulations were followed by tests of the injector on site. Figure 10 shows a pit profile of anode crystalline length (Lc) in nm of the original and optimized injectors.

Flue wall design: Optimization of homogeneity and pressure drop The model that used was described in a previous study [2]. Figure 8 shows a comparison of initial and optimized flow characteristics for the Saint-Jean-de-Maurienne furnace fluewall.

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paper outlines a number of process tuning initiatives that have contributed to this outcome.

Le (nm)

m «= 2,350 mm<= 2,500

Burner injection cycle optimisation

Ί Η <= 2,650 <= 2.800

Using a flue gas analysis system the combustion quality within the fluewall was measured downstream from the first burner ramp for different gas impulsion cycles. Figure 12 shows the comparison a pulse cycle of 30 seconds compared to that of a pulse cycle of 15 seconds for an equivalent burner power. The measurements showed that the long injections resulted in higher amplitudes of oxygen concentration with a higher risk of insufficient oxygen being available for gas and volatile matter combustion while the shorter pulse reduces the amplitude and ensures that there is always sufficient oxygen available for combustion of the natural gas and the downstream volatile matter. The shorter pulse has since been implemented.

«= 2.950

mm«=3,100 «= 3,250 mm

mm<= 3,400

mm> 3,400 Existing

Figure 10: Results of optimized burners at Aluchemie For equivalent natural gas consumption levels the mean baking level increased from 2.54 nm to 3.27 nm. The site has implemented the optimized injectors on all furnaces confirmed the results which in turn have allowed acceleration of furnace baking cycles.

12

~ - Burner Injection Cycle = 30 s — Burner Injection Cycle = 15 s

A similar study was conducted for the Alma site following the identification of similar problems. The model was used to recommend changes to the burner dimensions and operating pressures. Trial burners were tested and the results are demonstrated in Figure 11 which shows a pit profile of anode crystalline length (Lc) in nm of the original and optimized injectors. As seen in the figure there is a significant improvement in the baking level for an equivalent quantity of gas injected due to improved combustion and resulting heat transfer. Time

Lc (nm) 1 «= 2.350

Figure 12: Pulse length impact on oxygen concentration

M i <= 2,500 WM <= 2.650

Burner ramp temperature target optimization

mm «= 2,800 mm <= 2,950

Further combustion measurements at different times during the fire cycle showed that at the start of the cycle following fire change the oxygen concentration remained low (0-2%) when compared to the middle or the end of the cycle (5-10%). In addition to this the first burner ramp struggled to follow the temperature target. Finally the corresponding gas injection rates were very high on burner ramps 2 and 3. All of these items resulted in excessive injection of natural gas, incomplete combustion, excessive energy consumption, process instability and increased fire and explosion risks.

« I «=3,100 »

<= 3,250

mm <= 3.400 ■ i > 3,400

Existing MâWIWMΚli Optimized Figure 11 : Results of optimized burners at Alma Firing system tuning- Alma works example Since the start up of the Alma plant in 2000 the reduction line amperage has continued to increase requiring larger and denser anodes in larger quantities. However in the baking furnace area the process control system and anode baking performance was not adapted in parallel to ensure stable and optimum operation. Over time these factors plus some others led to a deterioration of furnace condition and performance, a number of duct fires, loss of production and limits to furnace production output which led to a need to externally purchase anodes.

To address the problems the firing curves were tuned to reduce gas injection of burner ramps 2 and 3 following fire change so that burner ramp 1 were supplied with more oxygen. This resulted in a much better combustion performance and better conformance to the temperature target as shown in Figure 13 while still achieving target final anode temperatures. Preheating temperature target changes

Since 2008 a number of actions, including those already outlined earlier in the paper, have been implemented to turn this situation around with the result of increased furnace productivity, improved baking quality and reduced fire risk. This turnaround has eliminated the need for anode purchase and has set Alma up with sufficient baked anode production to sustain upcoming amperage increases with minimal investment. The following section of the

Historically fluewall temperatures achieved in the natural preheating zones remained too cold for the full length of the cycles. This in turn resulted in a late pitch burn with a significant amount of partially combusted material and high levels of opacity.

869

consumption levels as shown in Figure 15 and duct fires have been eliminated.

Temperature Target of burner ramp 1 — Actual temperature of burner ramp 1 before tuning — Actual temperature of burner ramp 1 after tuning

970 Increase of measured temperature under burner ramp 1 for an equivalent target

2 .

at)

= 52 .2 §E 30 çOc 25

S o 20

950 930

IE

Reduction of injected gas quantity by burner ramps 2 and 3 during first baking hours -Injection rate of burner ramps 2 & 3 before tuning -Injection rate of burner ramps 2 & 3 after tuning 4 5 6 7 8 9 10 Time during baking cycle - Hours

11

12

13

jö

Figure 13: Natural gas distribution and temperature behavior Again this was a factor increasing the risk of fires as well as higher energy consumption and a limiting factor for the maximum furnace production level by limiting the firing cycle. The preheating targets have been tuned to rectify this situation and ensure an early, hotter pitch burn. The target curves are shown in Figure 14.

co

Figure 15: Energy consumption and baking level results Conclusions To support the ongoing amperage increase initiatives at a number of sites, a series of models have been developed and validated to allow anode baking furnace productivity to improve. These models have been validated and used to predict the process and design impact of test cases for heat transfer, fluewall flow characteristics, gas injection combustion and anode cooling. In addition to the models, process tuning exercises have also been successful in unlocking extra capability. Implementation of initiatives coming from different studies have resulted in concrete productivity improvements at a number of Rio Tinto Alcan sites and in many cases have had the added benefits of reducing fire and explosion risks, improving energy consumption and anode baking levels. References :

10 15 20 Time in baking cycle (hrs)

[1] S.A1 Hosni, J.Chandler, O.Forato, F.Morales, J.Bigot, C. Jon ville. Sonar Aluminium's Anode Baking Furnace Operation. Light Metals 2011,2011.

Figure 14: Natural preheating changes to improve pitch burn The change to this curve in combination with the other initiatives mentioned earlier in the paper now ensure more complete combustion in the natural preheating section. This is essential due to the fact that along with sufficient oxygen availability, the higher preheat temperatures ensure that the ignition temperatures within the fluewall are always above the lower limit of around 600°C require to ensure the volatile matter will completely burn.

[2] J.C Thomas, P. Breme, J.C. Rotger, F. Charmier. Conversion of a closed furnace to the open type technology at Aluminium Bahrain. Light Metals 1999, pages 567-572, 1999. [3] APR patent (will be published for the 11th of November)

Benefits have included no partially combusted material carry over and reduced fire risk. Energy consumption has improved due to the significantly higher anode temperature prior to the forced heating section of the furnace. The concept has been pushed even further by the implementation of an innovative process control methodology developed in collaboration with Sohar Aluminium [3] that will further allow improvement in productivity. The methodology has been successfully tested at a 25 hour fire cycle. Combined with the other initiatives detailed in this paper this will safely allow the overall furnace productivity to improve by 7%. In addition to the productivity increases, all these improvements have resulted in improved baking level and reduced energy

870

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

WIRELESS COMMUNICATION FOR SECURED FIRING AND CONTROL SYSTEMS IN ANODE BAKING FURNACES Nicolas Fiot1, Christian Coulaud1 folios Carbone, 32 rue Fleury Neuvesel, BP24, 69702 Givors Cedex, France Keywords: Carbon, Anode, Baking, Furnace, Wifi, Wireless, Communication, Safety, Solios. Abstract Anode baking requires firing and control systems which move every day, as the fire progresses around the furnace. Wired connections between moving equipment and the central control unit have always been an operation and maintenance concern. Wireless networks became the logical modern solution with standard PLCs. Modern furnace control requires new safety loops between firing equipment. However, whenever safe communication between safety PLCs are used for secured control of the baking furnace, wireless communication has encountered numerous drawbacks due to the nature of safety communication protocol and the interference with other WiFi systems in the baking furnace area.

The Firing and Control System is composed of several mobile ramps that are grouped by Fire (1 Exhaust Ramp (ER) + 1 Temperature & Pressure Ramp TPR) + 3 Heating Ramps (HR) + 1 Zero Point Ramp (ZPR) + 1 Blowing Ramp (BR)) - The Fires are located on top of the Furnace and are distributed over the various firing sections. They are controlled and monitored by two hot redundant computers (Central Control System) located inside the ABF control room. The master computer makes calculations based on data that it collects from each ramp through the communication network and sends commands to the ramps. All commands are sent to each ramp using the same Communication Network. These data are also displayed on the supervisory computer screens (Real Time Supervisory) for operator follow-up and stored in the Data Management computer (Data Management System).

Extensive development work was completed with major PLC suppliers to find the right combination of modems and antenna and to fine tune the PLCs and WiFi systems so that operation performance and safety requirements are fully met. WiFi is now available for the secured baking of anodes.

As part of the normal operation, each Fire moves one section forward every day. For a 4-Fire Furnace with a 24 hour baking cycle time, 20 ramps (4 ER + 4 TPR + 4 HR + 4 ZPR + 4BR) are relocated inside the building every day. Consequently, the Programmable Logic Controller (PLC) controlling the ramps to be moved is stopped before the move and restarted once the ramp is set in its new location. At each Fire moving, as part of normal operation, the ER is always replaced by a new ER and sometimes a ramp can be changed by a spare one, for maintenance purpose.

Introduction Aluminium is produced through Alumina electrolysis by means of carbon anodes. Prior to use in the pot lines, the green anodes produced from petroleum coke and coal tar pitch need to be baked in an Anode Baking Furnace (ABF) fitted with a Firing and Control System (FCS).

In addition to the ramps, one PLC named Auxiliary Equipment (AE PLC) ensures the interface between the ramps and the Furnace Fume Treatment Plant (FTP), the Furnace fuel supply loop and some other furnace utilities (for example, emergency stop and explosion vents). For user-friendly management of spare ramps, Solios Carbone is also using this PLC to manage communication between ramps. Under normal operation ramps exchange process and safety data with the AE PLC and between each other through the AE PLC. All these exchanges take place on the Communication Network. The Communication Network must have the same performance on all the Furnace sections, because at some point, each section will host a ramp.

The ABF is made of refractory bricks walls built in a concrete casing located inside a metallic building. (See Figure 1) For 66 sections: the building is around 260 m long x 40 m wide x 25 m high, one or two Furnace Tending Assembly (FTA) cranes are moving above the FCS pieces of equipment located on top of the furnace, to load and unload the anodes in/out of the furnace pits.

Communication Network Few industrial networks with a good bandwidth allow hot connecting and disconnecting of a User without trouble. One of the best available nowadays is Ethernet. The Wired Communication Network for a Firing and Control System needs each section of the Furnace to be equipped with a communication plug to connect the ramp. A Wired Ethernet Network is a star network topology requiring a heavy infrastructure (Enough Ethernet switches dispatched inside the Furnace building to have one port for each section and wiring up to each section). Moreover, this Network will suffer from the same problems as

Figure 1 - Anode Baking Furnace

871

other Wired Networks, plugs, sockets and cables are cumbersome and cannot endure cyclic change-over. Consequently, wireless connections appear to be the logical modern solution. (See Figure 2)

[Main ABF Control Ro

{Control Room

RTS-A

RTS

3

*

fSfL

WiFi is a radio network using frequency bands, named channels, located around 2.4 GHz (up to three distinct 20 MHz wide channels) and 5 GHz (up to twenty distinct 22 MHz wide channels). These frequencies have been opened to free usage in most countries in the world. Although, sometimes an authorisation may be required from local regulation authorities regarding frequency usage and/or for radio modems themselves. Moreover for 5 GHz channels, it is necessary to check which are non DFS (not subject to apply the Dynamic Frequency Selection procedure as described in relevant standards, i.e. in Europe EN 301 893) because the channels that must apply the DFS procedure might have to switch off from time to time in case of radar detection.

RTS-B

iir

?

L "|

LEVEL 2 - Ethernet Network Bub-Station

(Compnter Room DMS *f

t^MPn

a -£*·

CCS-A CCS-B

*

LEVEL 1 -.WiFi Network L\BF

A WiFi Network is controlled by the Access Points. They are radio modems, that follow the WiFi standard (IEEE 802.11 a/b/g/n) and they are connected to the field Ethernet Wired Network linking together the Central Control Computers and Auxiliary Equipment. According to the purpose of the WiFi Network (such as public WiFi Free Zone, domestic network), the settings will follow different practices for Service Set Identifier (SSID - commonly the network name), security management, channel allocation and roaming parameters. This publication describes one kind of WiFi Network suitable for the Solios Carbone Firing and Control System for an Anode Baking Furnace.

Ö ft

Figure 2 - ABF Control System Architecture Control Systems All the ramps are locally controlled by a PLC. Depending on the smelter projects and customer specifications these PLCs may either be standard or safety.

The Access Point sets the essential parameters of the WiFi Network: Standard 802.11a (5 GHz), 802.1 lb/g (2.4 GHz) or 802.11η (2.4 or 5 GHz), Channel, SSID and encrypting key. Because of the size of the building several Access Points are located inside the Furnace. Adjacent Access Points must use non overlapping channels so there is no interference with each other. However they share the same SSID and the same encrypting key.

The safety PLC manages local safety loops (Input and Output on the same PLC) but also manages safety loops across the Network (Input on one PLC and output on another PLC). Both kinds of PLCs use the same Wireless and Wired Ethernet Network but the communication mechanisms are different when safety PLCs exchange safety tags through the Network.

Ramps are connected to a radio modem named "Client". Client and Access Point can have the same hardware but they could also be built around different ones. Each Client must be parameterized with the network Access Point channels, SSID and encrypting key so that it can connect itself to any Access Point within the WiFi Network deployed inside the Furnace.

Over the past years Solios Carbone has accumulated successful experiences regarding installation of Baking Furnaces using Wireless Network either with standard (3 projects) or safety (3 projects) PLC.

When a Client modem is started, it connects itself only to one Access Point and they remain connected. However, the Client can choose to disconnect, to connect to another Access Point with a better signal level (Roaming phenomena) offering a natural redundancy and enhancing process continuity. Indeed, the Anode Baking Furnace radio coverage offers all the time, at least two Access Points, close enough for the Client to connect to. Radio exchanges occur only between the Access Point and the Client. Client to Client communications are routed through the Access Points.

Feedback from Qatalum During the recent start-up of Qatalum Smelter, new communication problems occurred that requested extensive development work in collaboration with the PLC supplier to eliminate the communication disruptions. The aim was to find the right combination between WiFi hardware installation and fine tuning of PLCs and WiFi systems parameters, so that operation performance and safety requirements were fully met. Qatalum is an Aluminium Smelter in Qatar. It is equipped with two Anode Baking Furnaces based on an Aluminium Pechiney technology: One of 66 sections with 4 Fires and one of 50 sections with 3 Fires. Both furnaces are end to end in the same building with a common Communication Network and the Firing Ramps can be used on both furnaces.

872

Siemens would use "unicast" messages for their safety exchanges whereas Rockwell Automation would use "multicast" messages. The method using "multicast" messages is a good solution for Wired Network but not easily handled by Wireless Network. In the coming months, Rockwell Automation is to release a new firmware to allow the use of "unicast" messages for their safety communication mechanisms.

At the request of the customer, the ramp PLCs are « Guardlogix » safety PLCs from Rockwell Automation. During the project engineering phase, the PLC manufacturer was not proposing any proprietary solution for the Wireless Network. Consequently, Solios Carbone decided to use the WiFi Network equipment from another supplier already proven in service in past projects. Even though, the platforms tests done before dispatch were successful, quickly during commissioning phase untimely ramp stoppage occurred due to problems linked with safety data exchanges between PLCs.

The possibility to extend the timeout duration for the safety communication is different and is limited by PLC firmware, which can be subjected to supplier modification. Indeed, thefirmwarerelease specified for Qatalum project has a reduced timeout adjustment range.

On the one hand, all indicators (such as coverage, bit rate, retry rate, ratio signal/noise) normally used to characterise a radio network performances did not show any problem. On the other hand, when the Wireless Network was substituted by a Wired Network, there were no stoppages. Consequently, it was concluded that the WiFi Network and the PLC were working properly individually and that the problem was linked to the PLC safety communication mechanisms and the way it was handled by the WiFi Network.

For the Qatalum project, it was not possible to return to the previous firmware release to adjust the safety communication timeout parameter as high as it was successfully set on in the past on other Furnaces. Also, it was not possible to use the other communication mode of the standard PLC using "unicast" packets because the project had specified that safety data exchanges had to be done using the safety communication mechanisms of the PLC.

In theory a properly installed WiFi Network should not be different from a wired Ethernet Network. There are, however, some differences, such as:

Modification of the WiFi system

The WiFi Network introduces latency time as an Access Point communicates to only one Client at a time.

Rockwell Automation recently signed a partnership with CISCO and said that they will warranty their PLC architecture only with CISCO Wireless Network. In order to solve quickly the issue this project was facing, it was decided to follow the PLC manufacturer recommendations and to deploy a CISCO Network: all the ramp Client modems and furnace Access Points were changed.

The WiFi Network doesn't handle "multicast" and "broadcast" traffic as well as a Wired Network. Effectively, most of the time, the time slot reserved to transmit these packets is reduced and they have to be forwarded systematically to all associated Clients. Moreover, for this kind of traffic, there is no reception acknowledgment between the Access Point and the Client, as for "unicast" traffic. (A "broadcast" message is sent to all network devices without distinction. Multicast messages are basically broadcast messages that can be routed, using specific functions of the manageable Ethernet network switches, only to the devices that have requested them. Whereas a "unicast" message is sent only to one designated network device.)

A new "Site Survey" and "Spectral Analysis" were conducted to confirm the number of Access Points, their location inside the furnace, type of antenna to be used and modems settings (for example channels, emission power). The "Site Survey" and "Spectral Analysis" are two main tools enabling a check of the conditions required for a successful Wireless Network installation: Availability of channels in particular non DFS.

The lapse of time before a Client chooses to disconnect from an Access Point when they can not communicate anymore with each other must be set carefully, because it could introduce additional unexpected latency in the wireless communication. From our experience, a WiFi Network has more retry or even lost packets than a Wired Network. Because of these drawbacks, it was empirically found that communication timeout (maximum time expected between two valid packets) on the WiFi Network should be significantly increased compared to the Wired Network. The PLC software must be optimized to take this into account in order to avoid any impact on the process.

External disturbances (Such as other Networks, meteorological and army radars).

Wireless

Number of Access Points and type of antenna to have a good coverage with the requested Ethernet bandwidth using lower emission power in order to avoid wave reflection inside the metallic building. Clients emission power and type of antenna. The preliminary and post-installation analyses are even more important for the Firing and Control System as the Furnace building is a real challenge to the operation of a high performance Wireless Network: A metallic building (walls, roof and even most of the time floor) favours wave reflection,

Safety communication mechanisms are different from one manufacturer to the other: The duration and the way the PLC handles the reconnection after a timeout is different.

873

Only the longer lateral walls are accessible to fit the Access Points, (Below or above the FTA rails) The anodes waiting on the mobile trolley in the middle of the Furnace make an obstacle to wave transmission,

Carbon dust settling on antennas,

equipment is compatible and validated with the specific communication mechanisms used by the chosen PLCs.

One or two big metallic FTA cranes that are moving on top of the Firing Control Equipment,

A single common Wireless Network would also call for a strict management of the Ethernet bandwidth because the Firing and Control System, with its process and safety exchanges, requires a high availability of the Network due to the nature of its communication.

Interference with other WiFi systems For Qatalum, during commissioning of the CISCO solution as it was defined during the preliminary study, a new problem arose. From the four required non DFS channels available in Qatar and that were reserved for the Firing and Control System for coverage of both Baking Furnaces only two were still available because the FTA initially using channels in the 2.4GHz band was moved to the 5GHz frequency band to solve communication problems. As more than two Access Points were required for both baking furnace coverage, in this particular case, extensive on site engineering managed to limit channel overlapping by adjusting the number, location and emission power of Access Points.

Conclusions With the experience of the Qatalum start-up, Solios Carbone has been able to understand better the complexity of combining safety PLC controls with wireless communication for firing systems on ABF. Most of the issues have been solved. A similar solution will be implemented for the Hindalco projects of Mahan and Aditya in India. Further improvements will come from the management of a common wireless communication for the whole ABF. However, as of today Solios Carbone can provide Firing and Control Systems using complete Wifi solutions with Siemens and Rockwell Automation safety or standard PLCs. WiFi is now available for the secured baking of anodes.

Moreover, even though the CISCO Network was commissioned with the parameters defined during engineering phase, some communication problems still persisted. It was necessary to fine tune all the parameters again to achieve operation performance and safety requirements: Modification of PLC software to optimise exchange between PLCs. Optimisation of PLC recovery duration in case of timeout on a safety tag. Optimisation of Access Point coverage using the two channels available, only. Optimisation of multicast traffic, (Ethernet manageable switches, Access Points and Client modems parameters). The Qatalum project shows that several radio networks can work at the same time inside the same building. However, it also highlights that for future project, it would be better for all the suppliers to share a common Wireless Network deployed inside the Anode Baking Furnace building. This common Network could be designed and specified by the EPCM. Firstly, it would allow a check that conditions are met for a Wireless Network (such as enough channels available, no interference) before the Smelter is erected. Secondly, it would help in reducing conflicts between the various suppliers for sharing the channels available taking into account that 2.4 GHz band only has 3 usable channels that are most of the time polluted by an increasing number of wireless devices (for example, cell phone Bluetooth headset, WiFi smartphone, printer, laptop, microwave oven) and sometimes, only a small number of channels are available within the 5 GHz band (depending on the local radio regulation authority). The WiFi equipment must be selected carefully during engineering phase for an industrial Wireless Network such as the one used for the Anode Baking Furnace. Because, even if all of the WiFi equipment follow the same standard and suit most day to day general public usage (such as access to internet in public space), all equipment do not have exactly the same ways to handle the communication. The project must check that the chosen WiFi

874

Acknowledgments The authors wish to express their grateful thanks to Qatalum and Rockwell Automation who authorized the publication of this paper.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

FULL CONTROL OF PITCH BURN DURING BAKING: IT'S IMPACT ON ANODE QUALITY, OPERATIONAL SAFETY, MAINTENANCE AND OPERATIONAL COSTS Detlef Maiwald, Domenico Di Lisa, Hans Peter Mnikoleiski, innovatherm GmbH + Co KG, Am Hetgesborn 20, 35510 Butzbach Germany

pack has to be over heated. Consequently energy will be lost at a cost, associated with the green house gas problem which is an important factor these days.

Abstract The quality of a baked anode is, amongst other criteria, defined by the heat treatment. Each anode in a pit has to reach a specific temperature overall in a certain time. The heat transfer is determined by the temperature versus time curve of the surrounding flues and the furnace geometry.

This paper will summarize already known practices and will describe new advanced control philosophies to further optimize the anode quality to a premium baking level by following the physical process limits and providing a homogenized product. This task is not just providing positive aspects within the baking process but influences the potline production as well.

Due to the physical design of the open top ring furnace, less energy is introduced to the outer flues in the preheat area. Further, the pitch burn starts and ends at a later point than at the inner flues. As a primary result, the homogeneity of the anode quality will be affected. Secondarily, other disadvantages will occur as well as which are:

This paper assumes that the general furnace design is known and will therefore focus on the baking process and its specific details. The anode baking process is running basically as a two convection heat exchanger with a firing zone in between, thus forming three areas. These areas are the preheat, firing and cooling area. Therefore, the basic requirement of the Firing Control System is to control each flue wall as an individual control loop in these three areas. This paper will basically concentrate on the preheat and the firing area, normally called a "6-section-fire" as shown in Figure 1.

unburned volatile fractions which condense downstream in the furnace, exhaust pipe, ducts and FTC facilities higher emissions, especially CO, C0 2 and NOx parts

hydrocarbons (PAH16),

high operational costs for extensive maintenance to clean equipment, ducts and FTC facilities impact on operational safety due to accumulation of ignitable fractions downstream to the FTC higher running and energy costs This paper will describe the different steps and strategies to optimize the baking process, achieving full control of the pitch burn and will show the results as remarkable improvements on anode quality, emissions, fuel consumption and running costs. Introduction

Figure 1 - Typical 6 section fire

The physics of pitch burn is quite complex. Once the volatile degasing in the anode has started it behaves exothermically as a gas generator and reacts very sensitively on any variation of the controls. The question is how to establish complete pitch combustion throughout the operation cycle in compliance with the final baking temperature requirement in the upstream burner area?

In most cases a "6-section" fire arrangement (3 preheat and 3 firing sections) is being used. Due to capacity reasons some application are enhanced by an additional burner ramp forming a seven section fire where the following theories can also be adopted.

The final quality of a baked anode is defined by several parameters in the manufacturing process. Within the heat treatment, a specific gradient limits the procedure as well as the maximum refractory temperature. The major task is to energize the cold spot of the anode pack within a section above a certain baking level. To realize this task, many times, part of the anode

Below is a cross section (Figure 2) of a standard triple baffle flue wall with 4 peepholes (1-4) based on a given firing direction for common understanding. This arrangement will be used as a base for all provided sketches and diagrams.

875

Specific individual flue designs will have influence on the following procedures and should be taken into consideration separately. Physics of the pitch occurrence in the carbon area

Typical section arrangement 1 2 3 4

_

The anode is formed out of a paste consisting of pet coke fractions and liquid pitch as a binder. Since the raw material prices are increasing dramatically, trials on new recipes have proven that pitch contents may be dropped down to a range of 13.5 - 14.0 %, while still achieving comparable anode properties.

u

ir1...T1.

H N

H

During the heat-up to final anode temperature of 1100°C, the following Figure 4 illustrates the thermal conditions in the preheat area where the treatment of the pitch burn takes place. ¸ÅÂÎ

: Ï E Œ •=m H

iFiring directioi

Figure 2 Section arrangement

The equipment placed onto the furnace surface occupies the shown peepholes (PH) as follows: Exhaust Ramp: Extracting out of PH 1, temperature (TC) measurement in PH 4. Former furnace designs utilizing head wall openings for the fume extraction are becoming uncommon and may vary from the following described advantages due to the very high dynamic flow resistances.

0

20

40h Time

Figure 4 Volatile Kinetics in the preheat area

Measurement Ramp (MR): In normal production (Fig. 1), TC and draft measurement in PH 3.

Modern anode baking processes provide ignition temperature and the required oxygen content at the right moment necessary to establish complete volatile combustion within the baking furnace.

Burner Ramps 1-3: Fuel injection into PH 2 and PH 4; TC measurement in PH 1.

The major task is to burn pitch volatiles within the baking process to almost 100% to reduce the consumption of high cost fuel down to a minimum. It is therefore also essential to focus on the so called "early volatiles" of the PAH 16 group, and find a strategy to burn them as well. Former control strategies and practices (double jump practices, opening peepholes) have had limited success in improving pitch burn. If gradients are too low in the preheat area, conditions are not conducive to proper combustion.

Figure 3 Location of sensors and actuators

Quality of pitch burn Incomplete pitch burn generates numerous problems in the process and the downstream facilities. Figure 5 indicates inefficient pitch combustion at a level of 97% or less.

876

Figure 7: Bottom view into a flexible Exhaust boot

Figure 5: Bottom view of a flexible exhaust boot (97% pitch burn)

Control of the Preheat area for full pitch burn

There are 'experts' around the world calling these negative results the 'price to be paid' achieving high quality anode properties. Over the last 10 years a continuous development to a state of the art control strategy replaced these rumors.

The baking process is a temperature oriented process. Former production crews were manually focusing on equalized draft values within the preheat area assuming homogenized heat development. Due to different flow resistances within the individual flue walls variations in gradients could be observed.

The following Figures 6 and 7 show typical inside views of an exhaust ramp with two years of service at full pitch burn and excellent anode properties as shown later in this paper. Soot and tar residues are non-existent.

Since the relevant process value is the temperature gradient, a temperature-draft cascade control loop for the preheat area is implemented into the firing control system as shown in Figure 8.

Figure 6: Inside of an Exhaust Ramp after 2 years operation

Figure 8: Preheat control T-P cascaded loop For the production of a consistent anode quality, the firing area is not the main influence on the final temperature homogeneity. The final temperatures of the anodes are also highly influenced by the treatment in the preheat area. In order to achieve the best baking consistency, the flue gas volume for defined temperature gradients in the preheat area needs to be controlled. Figure 9 shows the temperature development in the preheat area (Section 2) after implementation of this preheat control strategy.

877

Exhaust Profile Design

X

Figure 9: Temperature development in the preheat area

Higher Oxygen Requirement '

Cyc,e l26h]

Figure 11: Volatile volume over the cycle time

The baking process in the preheat area especially on the top anode layers is negatively influenced by ingress of false air through the packing coke material. To improve this bad situation, PVC covers as shown in Figure 10 have been introduced on most furnaces for the first 2 sections including the headwalls coverage. As a result the ingress of false air is minimized and the preheat temperature is more homogenized from top to bottom and higher overall in addition.

All process target curves should feature this basic knowledge. Overall, a process quality measurement e.g. opacity meter (optical or electrostatic principle) as shown in Figure 12 or a CO analyzer has to be installed to indicate the quality of the pitch combustion. Only this measurement ensures a transparency of the pitch combustion behaviour, and the development of the pitch front can be optimized and synchronized in relation to the fire cycle time. The corresponding opacity curve is also shown in Figure 9.

Figure 10: PVC covered Preheat area Intensity of pitch burn through the fire cycle Figure 12: Highly reliable Opacity sensor (electrostatic principle)

To provide ignition temperature and the required oxygen content at the right time it is essential to know when the volatile volume is reaching its peak. The following diagram (Figure 11) indicates the standard volume appearance over the cycle time comprising an optimized preheat development:

The degree of pitch combustion is directly linked to the quality of the preheat control, the oxygen residue out of the burner area and the zero point control upstream. Based on the optimized preheat control it is now necessary to introduce the required fuel (oil or gas) with a high speed, high pressure burner equipment in order to achieve an optimized fuel combustion. It is absolutely essential to maximize the oxygen level in the flue gas out of the burner area since this feeds the pitch burn area.

878

The data in Table II illustrates the same baking parameters at full control of pitch burn. Results The control strategies have been implemented in carbon plants at various smelters in the world for full control of pitch burn. The differences in performance compared to an incomplete pitch burn are measurable. Please find the results and comparison on different data as follows; Impact on anode quality Table 1 shows the characteristic values before implementation of a full pitch burn control: Table 1 : Data at incomplete control of pitch burn Parameter

Unit

Average

Table II: Data at full control of pitch burn Parameter

Average

Unit

1

Firing Cycle time

h

26

2

Peak T Exhaust Flue wall.

°C

370

3

Peak T Burner Rue wall

°C

1150

4

Soaking time

h

39

5

Anode Pitch content

%

13,6

6

Energy Consumption

7

Sulphur in green Anode

%

2.42

GJ/t

baked

1,89

1

Firing Cycle time

h

26

2

Peak T Exhaust Rue wall.

°C

300

8

Sulphur in baked Anode

%

2.135

3

Peak T Burner Flue wall

°C

1180

9

Desulphurization rate

%

0.285

4

Soaking time

h

35

10

Baked App. Density

g/cm3

1,582

5

Anode Pitch content

%

13,6

11

CO2 Reactivity Residue

%

89,5

6

Energy Consumption

GJ/t baked

2,3

12

Net Carbon consumption

kg/t

402

7

Sulphur in green Anode

%

2.42

8

Sulphur in baked Anode

%

1.98

9

Desulphurization rate

%

0.44

10

Baked App. Density

g/cm3

1,573

11

C0 2 Reactivity Residue

%

87

12

Net Carbon consumption

kg/t

418

Tremendous improvements can be shown not only in the direct baking parameters like peak flue temperatures, specific energy consumption, soaking time etc. but also in the parameters influencing the potline performance like C02 reactivity and net carbon consumption. Impact on operational safety Operational safety has become a top priority to all smelter plants in the world. Specially developed components for the anode baking process as well as improved process philosophies result in optimized fuel and pitch combustion which improve the process emissions drastically. Full control of pitch burn avoids accumulation of unburned soot and tar in the exhaust ramps, side main ducts and FTC. Consequently it eliminates the occurrence of wild fires, which is a critical hazard in the baking furnace area. Impact on emissions Full control of pitch burn has a direct impact on emissions because it solves the problem at the source. Emissions that are not produced will not be emitted to the environment and will not harm the downstream process or equipment (e.g. the FTC). As a result, the reliability and availability of the FTC will increase

tremendously, costs for regular cleaning (up to 2 times per 12 months) and extensive repair of dampers, actuators and sensors will be eliminated.

Conclusions This paper discussed the impact of a full pitch burn control on several technological parts of the anode baking process. The results demonstrate very impressively the improvements that have been achieved.

Impact on maintenance cost The same advantages mentioned above do have a direct impact on related maintenance costs as well. The frequent cleaning of exhaust ramps, side main ducts and FTC ducts and structures can be eliminated. Flexible exhaust legs will last much longer and therefore minimize maintenance attention. Reduced peak temperatures result in less maintenance for burner equipment, thermocouples, cables, peephole covers and reduced refractory costs. The overall cleaner anodes and their improved behavior in the potline process will reduce anode cleaning and handling. . Every prevented wild fire saves additional repairs, manpower and unknown insurance increases. Cost savings observed at various smelters are in the range of 200.000 US$/annually.

The anode baking process as a link within the carbon production does play its roll in the anode quality and is normally the highest cost portion of anode production. Optimizing strategies on the process to achieve at least the same if not improved anode quality on a lower cost level are always worth looking into it.

References [1] W. Leisenberg, Flue Gas Management, Light Metals (1999)

Impact on operational cost

[2] J. Ameeri; K. M. Khaji; W. K. Leisenberg, The Impact of the Firing and Control System on the Efficiency of the Baking Process, Light Metals (2003) 589-594.

Full pitch burn control also has a direct impact on operational costs. The pitch energy introduced by the green anode will be released, ignited, totally burned and its energy results in being able to reduce the peak temperature later in the process as well as shortening the overall soaking times. These advantages are leading to an energy saving of quite a high extent. A smooth baking process is not just providing less heat stress to all related equipments it also reduces the packing material consumption due to the mentioned lower peak temperature and will also play its part in reducing the running costs. Finally all reduced maintenance and running activities will not require the same amount of manpower anymore.

[3] W. Leisenberg, Firing and Control Technology for Complete Pitch Burn and its Consequences for Anode Quality, Energy Efficiency and Fume Treatment Plant, 9th international Conf. on non ferrous metals, July 8-9, 2005, Pune. [4] W. Leisenberg: "Modellierung thermischer Prozesse" Ergebnisbericht Forschungssemester 1999, University of Applied Sciences Giessen-Friedberg, FB ET II, 2000, [5] D. Maiwald, D. Di Lisa "Conversion of the firing and control system and the impact on the efficiency of the baking Process " Aluminium World 2008 [6] D. Maiwald, D. Di Lisa "Innovatherm Advanced Control Modules and Control Philosophy - The Success for High Quality Anode Production in Open Top Ring Furnace" Aluminium World 2008

The energy consumption as shown in Figure 13 prove the significant impact of a full pitch burn after implementation in February 2010.

[7] D. Di Lisa, H.P. Mnikoleiski, D. Maiwald "Homogenized Anode Baking Quality" Light Metals 2009 [8] Demedde et.al.:"Kinetic Phenomena of the Volatiles in Ring Furnaces", Light Metals 1986 [9] Hameed Abbas, Khali Khaji, Danial Sulaiman "Desulphurization control during baking", Light Metals 2010

Figure 13: Energy consumption

880

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

HIGH PERFORMANCE SEALING FOR ANODE BAKING FURNACES Pierre Mahieu1, Sιbastien Neple1, Nicolas Fiot1, Ismael Ofico2, Manuel Eufrasio2 ^OLIOS Carbone, 32 rue Fleury Neuvesel, BP24, 69702 Givors Cedex, France 2 MOZAL Aluminium Smelter, Parque Industrial de Beluluane, Caixa Postal 1235, Maputo, Mozambique Keywords: Carbon, Anode, Baking, Furnace, Flue Wall, Sealing, Solios,

Abstract Anode baking requires important volumes of air and fumes for combustion and thermal exchange. Efficient refractory port sealing is an important condition required to reach rated flows inside flue walls. Conventional sealing techniques have many problems which can lead to process inefficiency and degraded operating conditions. A new technology has been developed and then tested at the Mozal smelter to enhance this sealing function: it is based on the use of advanced materials that combine flexibility, resistance to the extreme conditions in the furnace and offers a new potential of process improvement in anode baking furnaces.

The problem with this technology is that the sealing contact of the cloth with the refractory is not perfect and sealing capacity may vary with the gate position, flue wall deformation or gate cloth wear. These defects may cause incomplete combustion due to a lack of draught inside the flue walls and can generate condensation of hydrocarbon volatiles in the ring main due to cooling effect of added ambient air at the exhaust manifold. Another main issue of cold air leaking at the exhaust is the formation of corrosive agents, mainly sulphurous and hydrofluoric acids, when moist air meets hot fumes, leading to rapid deterioration of exhaust pipes and ducts.

A flexible membrane is held between two rigid plates and encloses the inner refractory port when inflated. As a consequence, cold air incursion into the fume exhaust ramp is reduced. The results obtained, including a limited condensation of exhausted fumes and an increased furnace thermal efficiency, are presented in this paper.

Draught from FTC fans

ttttt Exhaust Ramp main duct

KN

Introduction In order to direct only the exhaust gases into the external Fume Treatment Center and prevent cold air from the upstream end of the adjacent fire group from flowing backwards into the exhaust, a seal is placed inside the flue wall against the fume path opening (Figure 1). The design commonly used in open type furnaces, known as « folding shut-off gate », is aflat,flexible,high temperature cloth that is placed across the internal rectangular opening, also called « Port », in the headwall (Figure 2).

Figure 2: Air leakage at port shut off gate A sealing technology based on an inflatable membrane The new solution proposed by Fives Solios is an alternative to folding shut-off gates dedicated to improve sealing of flue walls. It is registered as a patented invention. The Port Sealing Ramp (PSR) is equipped with an air blown inflatable sealing membranes. The inflatable membrane is surrounding a metallic structure that can be easily inserted and removed from the peephole (Figure 3). The membrane is inflated and deflated by mean of an air fan. Switching from inflation to deflation modes is simply done by operating a hand valve. The advantage of an inflatable membrane is that the port is closely sealed even if it is deformed or it contains deposits.

Figure 1: Standard "folding shut-off gate'

881

»·"v '.

V

Figure 3: Inflatable gate in position in the port

Figure 5: PSR in position for FTA loading

The inflatable membrane material is designed to be flexible enough to fit the port inner surface as well as to sustain temperature, abrasive and corrosive conditions. The elasticity of the material selected allows a total expansion of the membrane at low pressure, generating a negligible stress on the refractory port. The membrane is protected from abrasion and impacts by a cover that can be easily replaced.

Technical textiles to sustain furnace environment The membrane material must be able to face simultaneously: thermal constraints due to the temperature level in the port, - corrosive and aggressive ambient conditions due to some fume components, severe and various mechanical constraints due to inflation / deflation, abrasion along bricks of the peephole when set in / out, or shocks when handled and rested on the PSR

Integrated solution designed considering ergonomics and safety principles

But it also requires to remain: flexible enough to retract properly between plates, to allow an easy insertion through the peephole, and to fit properly against the refractory walls. air proof in order to stay properly inflated during thefirecycle without significant air consumption, - light for ergonomie reasons, - easy to manufacture (by gluing or thermoforming rather than expensive sewing), a cost effective material.

The PSR is an independent structure that holds the inflatable shutoff gates and the air pressure distribution circuit (Figure 4). The Ramp is equipped with walkway platforms allowing the operator to install the shut-off gates safely even when pits of the preheating zone of an adjacent section are not loaded with anodes.

The technical specification was very challenging and in order to select appropriate materials to sustain furnace conditions, several life duration tests were conducted in real operation at MOZAL smelter ABF n°l from August 2009 to June 2010.

Figure 4: PSR with inflated membrane in position in the port The inflatable shut-off gate, designed with consideration of ergonomie principles, has a total weight lower than 10kg. The membrane is retracted inside 2 plates when deflated, allowing the inflatable shut-off gate to be installed and removed easily through the peepholes. The PSR is designed to be positioned along the exhaust ramp during the FTA loading operation in the sealed section (Figure 5). Figure 6: Temperature measure on the inflatable gate cover

882

The first consideration was to identify the maximum temperature reached inside the port at the end of the cycle. Temperature probes positioned on the membrane cover recorded a maximum of 152°C for an exhausted fume temperature of 309°C at the end of the preheating process (Figures 6 and 7).

10

selection of the membrane textile. The membrane material shall mainly remains flexible, heat-proof and airtight. Following temperature measurements and chemical analysis, several combinations of materials for membrane and cover have been selected and tested at the MOZAL smelter (Figure 9).

15

Time in cycle (h)

Figure 7: Cover surface temperature recorded during a baking cycle

Figure 9: Tested covers

Another issue is related to the exhausted fumes composition and particularly the presence of gaseous fluorides, sodium and sulfur oxides that are very corrosive agents. Gaseous fluorides and sodium are mainly coming from the anode butts that are part of the recipe of green anodes, Sulfur oxides are mainly originating from coke and fuel combustion.

In total four materials for cover and three materials for membrane were selected among thermoplastic polymers. The best material "M3" for the membrane is still operating normally after 140 days. The longest cover lifetime duration recorded with material "C4" is 115 days. It is important to note that lifetime duration of both the membrane and the cover are dependent on many factors like flue wall conditions, handling operation, baking parameters. The lifetime duration of the membrane and cover materials recorded in real operation are summarized in Table I and Table II.

Analyses have been performed to determine the main factors of corrosion of the gate materials after using it several weeks in the refractory ports. Scanning Electron Microscope (SEM) observation and Spectrometry were used. The analyses revealed that a chemical degradation mainly from sulphurous and fluoride compounds occurred on the cover, causing modifications of the properties of the material (Figure 8). Such chemical degradation is catalysed under the thermal effect.

With the combination of material C4 and M3, the lifetime duration of the inflatable shut-off gate is 3 months on average. Table I: Results of cover life duration tests C4

Material ref.

Cl

C2

C3

DAYS (min/max)

39

30/60

30/80

75/115 1

V V V

V

V

V

Origin of damage chemical attack Mechanical wear temperature

V

Table II: Results of membrane life duration tests Material ref.

Ml

M2

M3

Figure 8: SEM picture of degraded material matrix

DAYS (min/max)

2/37

11/40

79/>140

No satisfactory potential material was found to meet every requirement and it was finally decided to split the list of requirements by using a cover around the inflatable membrane. The role of this cover is to protect the membrane from abrasion and chemical corrosion, thus relaxing some constraints in the

Origin of damage

V

V

mechanical wear temperature

883

V

Draught capacity rμse due to membrane sealing

Improvement of combustion quality in ageing furnaces

The trials performed in operation at MOZAL aluminium smelter showed an increase of draught inside the flue wall up to 30% compared with a folding shut-off gate in the same conditions, proving the excellence of an inflated membrane at eliminating the ingress of cold air into the fumes exhaust line (Table III).

It is assumed that combustion occurs as a balanced stoichiometric reaction. For example, the complete combustion of 1 Nm3 of Natural Gas requires 2 Nm of oxygen that represents approximately 10 Nm3 of air. If the volume of available air drops below the volume required for stoichiometry, incomplete combustion starts leading to unburnt hydrocarbons and typically Carbon Monoxyde (CO) formation (Figure 11).

Table III: Records of draught increase with inflatable gates Date

Time Section

fw

PI (Pa) FG

IG

A(Pa

Δ(%)

1

•f 16

| 10/08/2009 15:35

43

6

-78

-91

| 10/08/2009 15:37

43

8

-73

-93

-20

+ 27.4% |

1 10/08/2009 15:38

43

9

-85

-95

-10

+ 11.8% 1

| 11/08/2009 10:15

44

8

-117

-133

-16

| 11/08/2009 14:10

44

2

-100

-111

-11

•f 13.7% 1 + 11.0% 1

| 11/08/2009 13:57

44

9

-60

-74

-14

+ 25

| 12/08/2009 13:45

45

9

-75

-90

-15

+ 20

| 12/08/2009 13:45

45

7

-145

-160

-15

PI FG IG

Natural gas 1 m3

Combustive air

Compiete Incomplete combustion combustion

2m302

hLO

H,0

C0212%

CO210% CO

h^ftμ%* 1 Additional air from improved sealing

Static negative pressure recorded in the first preheating section Folding Gate Inflatable Gate 1 Lack of air 10%

In most of flue walls tested, whatever original pressure level, draught improvement was recorded showing that most of standard shut off gates are not positioned correctly along the port or are damaged (Figure 10).

+ 10% air

Figure 11: Stoichiometry of natural gas combustion During trials in the ageing furnaces with deformed and even obstructed flue walls at MOZAL, additional air volume from membrane sealing allowed to switch from incomplete to total hydrocarbons combustion, especially in sections located at the end of the ring main where draught capacity was very poor, less than half of the requested level.

Because air leakage at the port is very difficult to identify, sealing problems with standard shut-off gate remain hidden but have a large impact on baking process performance. For example, a 30% reduction in draught due to cold air leakage at the port represents a 15% drop in the volume of fumes and air flowing through the flue walls, impacting thermal efficiency and in some cases combustion quality.

300

^,.

This phenomenon is illustrated by the significant drop of Carbon Monoxide (CO) content in the exhausted fumes when using the PSR instead of standard shut off gates (Figure 12) CO content drop after PSR installation ABF1 01/03/2010 Section 20

Standard shut off gate replaced by inflatable gate ABF1 01/03/2010 Section 20

—CO(ppm) Pipe Temperature (eC)

250

170 160

2000

200

150 140

1500

150

130 120

100

110

50

100

500

90 12:43

13:55

15:07

16:19

17:31

80

18:43

12:43

Figure 10: Draught increase with inflatable gate (fwl to fw9)

13:55

15:07

16:19

17:31

18 43

Figure 12: Combustion improvement with inflatable gate

884

Furthermore, as port sealing improvement tends to raise combustion quality, particularly in ageing flue walls, the following production costs will be cut when using PSR equipment in the same operating conditions: maintenance expenses dedicated to exhaust pipes, ring main and FTC duct cleaning, energy required to bake anodes. Finally, by ensuring stable under-pressure in the flue walls, the PSR solution also promotes baking uniformity that impacts positively on potline operation and consequently metal production costs.

The direct consequence of combustion quality improvement is the reduction of unburnt hydrocarbons inside ducts as well as a thermal efficiency increase. This was confirmed at MOZAL by a 15% increase on average of the preheating gas temperatures recorded at the end the baking cycle, only 2 cycles after PSR was installed in the fire group (Table IV). Table IV - Preheating gas temperatures (°C) recorded at the end of the baking period in the first preheating section peephole 1 and in the Exhaust pipe. ABF1 Fire 1 from 01/03/2010 to 03/03/2010

fwl fw2 JO fw3 fw4 8 fw5 OD .S fw6 te fw7 ou JS fw8 a fw9 Avg (fw) ΛR pipe

e

1-H

without PSR 560 640 725 725 510 625 670 560 280 588 250

With P S R - 2 cycles later 625 740 740 675 710 725 730 740 480 685 300

Evaluation of the benefit is not available yet, but the balance between the new PSR equipment extra costs and baking furnace operational cost savings thanks to port sealing improvement is expected to be largely positive.

Ä(%) +10% +14% +2% -7% +28% +14% +8% +24% +42% +15% +17%

From prototype to first industrial version Increasing performance of sealing in anode baking furnace thanks to PSR has been demonstrated during the trials at Mozal smelter. But using a prototype in real operation also reveals downsides of the equipment. A noticed disadvantage of the current version is that it requires a crane to be lifted during the fire move operation, impacting negatively the FTA availability. By reducing the PSR structure weight, the first industrial version of the ramp will be lighter and movable directly by furnace operators. Another operational issue identified is the difficulty to move the PSR along the exhaust ramp, limiting the FTA access to pits during anode loading operation. The PSR structure of the first industrial version will be lowered and resized to comply with the FTA filling pipe movement. This problem is largely minimised when front section is loaded before the fire move as per normal practise.

Therefore, improvement of flue wall sealing by means of membranes may contribute to breaking the vicious circle "lack of air" -> "hydrocarbons deposit -> "pressure drop in exhaust pipe and ring" -> "air volume reduction" that causes process instability, baking level inconsistency and fire risks particularly in ageing furnaces. Further advantages of the proposed solution In new operating furnaces, the Port Sealing Ramp promotes process stability by avoiding any air ingress from the wall port and thus maintaining permanently draught at its full capacity. As explosion risk is mainly related to loss of draft or very low flue gas rate in anode baking furnaces [1], PSR also contributes to improve safety during baking process.

Conclusion Sealing of fire groups in open type anode baking furnaces has been underestimated, mainly due to technological barrier. With the recent progress in technical textiles, a high performance seal for flue wall port was designed and patented. A test campaign was held at the MOZAL smelter on a prototype Port Sealing Ramp (PSR). The first trials have demonstrated the capability of the new equipment to enhance process efficiency and to promote safety. It also confirmed that flue wall port sealing, all too often neglected, is a major operation in anode baking furnaces.

Contrary to the standard folding shut off gate, when an inflatable gate is damaged, the loss of draught in the flue wall is apparent and so high that membrane replacement is necessary to maintain process operations inside the flue wall. In this way, hidden infiltrations degrading process performances all along baking cycles are avoided. Another advantage resulting from the improvement of flue wall port airtighntess is to avoid corrosive attacks of exhaust ramp and ring main metallic structures that occur when moisture from infiltrated cold air meets hot combustion gases, forming highly corrosive agents.

References [1]

PSR, a long-term cost effective solution PSR solution requires higher maintenance costs than standard shut-off gates, mostly because of membrane design and material. But the gap represents a minor part of the total costs required to bake anodes.

885

Inge Holden, Olav Saeter, Frank Aune, Tormode Naterstad, "Safe Operation of Anode Baking Furnaces", Light Metals 2008, 906, 908

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION

Anode Raw Materials and Green Carbon SESSION CHAIR

Frank Cannova BP Coke Huntington Beach, California, USA

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PROPERTY PROFILE OF LAB-SCALE ANODES PRODUCED WITH 180 °C METTLER COAL TAR PITCH Winfried Boenigk, Claudia Boltersdorf, Falk Lindner, Jens Stiegert RUTGERS Basic Aromatics GmbH; Kekulestraίe 30, 44579 Castrop-Rauxel, Germany Keywords: high-temperature mixing, low toxicity, anode optimization Abstract

100 i 90 \

The PAH-induced toxicity of coal tar pitch decreases with increasing softening temperature. The use of high- melting pitches is however restricted due to operational challenges. One critical limitation is high-temperature mixing. Temperatures up to 300 °C were now realized using a newly developed EIRICH mixer. A 180°C Mettler binder pitch resulted in lab-scale anodes having a higher baked density (+0.04 g/cm3). Electric resistivity and air permeability are reduced whereas compressive strength and Young's modulus are increased. The significantly lower baking loss allows a faster carbonisation process in the low temperature range, potentially increasing the throughput of baking furnaces.

80 j

s ^

70 \

^^

: :

Ψ

%

1

yS

s^r

P" 60 J E* 50 3 40 (0 30 20 10

—; *"

^ ~ -

ίT Ψ Γ/ -ff f f

oJ

,

0

1

, 2

,

,

,

1

3 sieve [mm]

4

5

6

Figure 1 : Coke particle size distribution

Introduction

The characteristics of the trial pitches are shown in Table 1. As expected the high softening point of HM 180M pitch results in a carbon yield of 76% instead of 58% of the standard CTP Bx95. The CANTOX-toxicity of HM180M pitch is reduced by 87% compared to CTP Bx95.

Aluminum is a versatile engineering material and can contribute significantly to reduce energy consumption when used for car production. An unwelcomed by-product of aluminum production is C0 2 released during electrolysis. The carbon footprint of aluminum production needs to be improved e.g. by optimizing carbon anode performance Coal tar pitch (CTP) will remain the predominant anode binder for the foreseeable future despite its content of carcinogenic polycyclic aromatic hydrocarbons (PAHs). The potential for CTP to contribute to an even brighter future of aluminum production is not yet exhausted. Reduction of CTP toxicity and an increase in anode density are main improvement targets in aluminum production. The best established method to measure pitch toxicity has been introduced by Alcan in cooperation with CANTOX [1,2]. One approach to lower pitch toxicity applicable to dry-paste Soederberg plants is blending CTP with petroleum pitch [3]. Obviously the removal of low molecular weight PAHs is more efficient than dilution [4]. One challenge for plant-scale use is the increase of CTP viscosity with decreasing toxicity. For pitch qualification RUTGERS operates an R&D Carbon bench scale plant [5]. It consists of an oil-heated mixer and a floating- mould press limited to 250 °C: sufficient for state of the art pitches but limiting investigations going beyond. In 1992 first attempts were made to produce anodes with high-softening pitch (170 °C Mettler) [6] expecting denser anodes and correspondingly lower specific electrical resistivity. A higher amperage would increase cell efficiency. EIRICH has recently introduced a new mixing system capable to operate up to 300 °C by introducing the inductive heating concept. This new EIRICH mixer allows for evaluation of pitches which soften up to 200 °C. Lab anodes were produced and analysed using a commercial Bx95 CTP and a 180 °C Mettler CTP at comparable viscosities.

Table 1 : Analytical properties of the investigated pitches HM 180M

Bx95 Softening point Mettler

DIN 51920

[■C]

118

183

Coking value (Alcan)

DIN 51905

[%]

57,9

76,4

Toluene Insoluble Quinoline Insoluble

DIN 51906 DIN 51921

[%] [%]

29,9 7,9

43,1 13,3

Distillate 020 - 230°C

GC

0,04

GC

[%] [%]

0,01

Distillate 230-270°C

0,09

0,00

Distillate 270 - 360°C

GC

[%]

1,10

0,01

Distillate 360 - 440°C Distillate 440 - 500°C

GC GC

[%]

4,81 12,53

0,13 1,88

Distillate >500¼

GC

[%]

14,06

13,37

Sum of Distillates

GC

l%]

32,60

15,44

GC-analysis

24

0

0

0

Acenaphthylene

0,000

0

0

0

0

Acenaphthene

0,000

488

0

o

0

Fluorene

0,000

116

0

0

0

Phenanthrene

0,000

3276

0

0

0

Anthracene Fluoranthene

0,000 0,034

658 10018

0 341

0 81

0 3

Pyrene

0,000

8666

0

89

0

Benzo[a]anthracene

0,033

7748

256

235

8

Chrysene

0,260

8253

2146

282

73

Benzo[e]pyrene Benzo[b]fluoranthene

0,050 0,100

8345

417

1111

56

12090

1209

1198

120

Benzo[k]f 1 uora nthene

0,010

6045

60

599

6

Benzo[a]pyrene

1,000

10307

10307

1608

1608

Dtoenz o[a,h]anthrace ne

1,400

2141

2997

721

1009

Benzo[ghi]perylene

1,000 0,100

8097

8097

462

462

7702

770

3490

349

26600

9876

3694

Average toxicity index

889

BaP equiv |

0,000

Total

For the lab scale anode production a mixture of two coke fractions is used (S = 2,26 %, Na = 143 ppm, Ca = 12 ppm). The sieve analysis of the dry recipe is shown in Fig.l.

RPF(Canto PAH[ppm] BaP-equiv PAH[ppm]

Naphthalene

lndeno[1,2,3-ccQpyrene

Raw materials

[%]

93974

1

0,13

EIRICH inductively heated mixer [7] To realize high heat flows into the mixture / material to be heated high temperature gradients between the product and the wall are necessary. In wall heated, batch wise operated heating systems the speed determining factor is usually given by the energy supply through the external heating system, which determines if and by how much the wall temperature drops after the addition of cold material and especially how fast the temperature rises again to the set point. When using a liquid heat transfer medium to heat the outer wall of a mixer the limitation is mainly given by the heat transfer between heating medium and chamber wall. The heat transfer will improve when using turbulent flow but is usually limited by the double-shell geometry and high pressure drops. A particularly efficient method in terms of heat transfer, uniformity of heat distribution on the heating surface and regulation of the heating power provides the newly developed and patent pending induction heating system of the rotating mixing chamber in an Eirich mixer. Due to the high energy output between the inductor and the rotating mixing chamber the user can realize extremely fast heating curves, high wall temperatures, short drying times as well as high product temperatures up to 250-280 °C. The inductively heated mixer consists of a rotating mixing chamber moving the process material into the zone of the eccentric quick-running mixing tool. The mixing chamber as well as the mixing tool are equipped with their own drives enabling the adaptation of the speeds to the individual application and to the process material. The combination of the mixing chamber and the mixing tool rotation as well as the inclined arrangement of the mixing chamber causes the generation of an intensive threedimensional fluid bed-like material flow in the mixer interior in connection with a stationary material deflector at the chamber wall. The constructive design of the mixing tool with knife-type mixing blades, high peripheral speeds as well as the opposed rotating direction to the mixing chamber, enable an intensive circulation and the surface of the process material is "ripped" open. In the granulation phase, the opposing movement results in a mechanically generated fluidized bed which presents the condition for maximum heat and mass transfer. The stationary material deflector also reliably avoids cakings on the wall and the bottom and facilitates the discharge of the material through the discharge flap in the centre of the rotating mixing chamber bottom.

Figure 3: Opened lab scale EIRICH mixer with stirrer removed Production parameters The EIRICH intensive mixer used for the experiments has a volume of 5 L with a diameter of 235 mm. The mixing vessel (42 rpm) and the stirrer (900 rpm) rotate in the same direction. The mixing time is 9 min to reach 280 °C with a pre-heated coke and cold solid pitch for the high melting pitch and 4 min to reach 210 °C for the standard binder pitch. The pitch content in the batches was selected in the range from 14 to 18 w.-%. From each batch 13 anodes with a diameter of 50 mm and a height of 100 mm are prepared. Therefore nearly 350 g of the mixture are necessary. The anodes are pressed with 120 kN corresponding to 600 bar. Results After production of the anodes the following characteristic data were collected: apparent green density, apparent carbonized density, baking loss, coking yield, volume expansion, compressive strength, Young's modulus, air permeability, specific electric resistance, C0 2 reactivity. The anode properties are shown in Tables 2 and 3. Using the high melting pitch an increase of 0.05 g/cm3 of the green density is possible for all examined pitch contents (Figure 4) which translates to a similar difference in density after baking (Figure 5). Using the standard pitch leads to a density of 1.59 g/cm3 whereas the high melting pitch results in a significant increase in density of 1.63 g/cm3. The baking loss of the anodes is less than 4 % for the high melting pitch while it is more than 5% for the standard pitch (Figure 6). In correlation to this the coking yield increases from 67 to more than 80% (Figure 7). This is a direct result of the higher coking value of the high melting pitch, HM 180 M. The significantly lower baking loss of the anodes produced with the high melting pitch and the reduced volatiles of HM 180M (Table 1) very likely allow a faster carbonization program for anode preparation. This was tested subsequently and confirmed on lab scale. The carbonization program sequences used are shown in Tables 4 and 5.

Figure 2: Mixer with inductive heating

890

Apparent density, green

Table 2: Properties of bench scale anodes

Apparent [g/cm3] Apparent carbonized density [g/cm3] Baking loss [%]

Coking yield [%]

Volume expansion [%]

Compressive Strength [MPa]

Pitch Content [%] 14 15 16 17 18 14 15 16 17 18 14 15 16 17 18 14 15 16 17 18 14 15 16 17 18 14 15 16 17 18

Eirich Mixer BX95 Normal Fast Carbonization Carbonization 1.644(0.004) 1.664 (0.003) 1.672 (0.003)

1.667(0.004)

1.583(0.004) 1.591 (0.003) 1.582 (0.004)

1.589 (0.003)

4.9(0.1) 5.3(0.1) 5.9 (0.2)

5.9(0.1)

67.2 (0) 66.7 (0) 65.2 (0)

63.2(0)

-1.2 (0.2) -1 (0.2) -0.6 (0.3)

-1.3(0.2)

39(2) 40(3) 42(2)

35(1)

-Bx95 ~-&~~HM180M normal carbonization —♦—HM180M fast carbonization Eirich HM Normal Carbonization 1.656 (0.014) 1.686 (0.007) 1.713(0.004) 1.726 (0.004) 1.624 1.645 1.637 1.629

(0.013) (0.004) (0.010) (0.009)

Mixer 180

I | Fast I Carbonization 1.653(0.018) 1.679(0.007) 1.688(0.019) 1.721 (0.009) 1.624(0.018) 1.627 (0.003) 1.63(0.007) 1.615(0.012)

2.7(0.1) 2.9(0.1) 3(0.1) 3.1 (0.2)

2.8(0.1) 3.0(0.1) 3.3(0.1) 3.5(0.1)

81(0) 81(0) 81.1 (0) 81.7 (0)

80.1 80.1 79.6 79.3

(0) (0) (0) (0)

-0.6(0.1) -0.5 (0.3) 0.6 (0.9) 1.7(1.3)

-1.0(0.2) 0.2 (0.3) 0.4 (0.4) 2.8 (0.4)

51(3) 54(0) 55(0) 54(0)

44(3) 54(0) 54(0) 53(0)

Figure 4: Apparent green density versus pitch content Apparent density, baked - HM 180M normal carbonization

- HM 180 M fast carbonization

Values in brackets represent standard deviations.

Table 3: Further anode properties Pitch Content

f%l

Young's Modulus [GPa] AirPermeability [nPm] Specific Electric Resistance

pom]

cor

Reactivity residue [%] COr Reactivity dust [%] CO r Reactivity loss [%]

14 15 16 17 18 14 15 16 17 18 14 15 16 17 18 14 15 16 17 18 14 15 16 17 18 14 15 16 17 18

Eirich Mixer BX95 Normal Fast Carbonization Carbonization 2.2(0.1) 2.3 (0.2) 3.4 (0.4)

2.2 (175)

0.4 (0) 0.5(0.1) 0.7(0)

0.4 (0)

74.3 (7.6) 60.9 (3.5) 54.7 (0.5)

61.2(1.4)

84.5(1.1) 85.2 (2) 86.4(1.1)

82 (0.9)

6.5 (0.7) 6.0(1.2) 4.8 (0.5)

8.1 (0.8)

9.0 (0.4) 8.8 (0.8) 8.9 (0.6)

9.9 (0.4)

Eirich Mixer HM 180 | Normal Fast I Carbonization Carbonization 2.3 (0.4) 2.1 (0.4) 2.8 (0.2) 2.9 (0.2) 3.5 (0.2) 3.1 (0.3) 3.4 (0.5) 3.3 (0.5) 0.4 (0.3) 0.1 (0) 0.2 (0) 0.3 (0.1 )

0.3 (0.2) 0.17(0) 0.5 (0.2) 0.4 (0)

62.2 (4.9) 52.4(1.9) 46.9 (0.6) 47.3 (4.3)

63.4(6.1) 52.8 (2.7) 49 (1.3) 45.8 (0.8)

82.0 (2.0) 84.8(1.8) 84.4(3.1) 83.2 (2.7)

84.8 (2.7) 86.5 (2.6) 87.2(1.2) 89.3 (0.8)

7.9(1.3) 6.3(1.0) 6.8 (2.0) 7.1 (1.8)

6.4(1.7) 5.4(1.7) 4.7 (0.6) 3.2 (0.4)

10.1 (0.7) 8.9 (0.8) 8.8(1.2) 9.7(1.0)

8.8(1.1) 8.1(1.0) 8.1 (0.6) 7.5 (0.4)

Figure 5: Apparent carbonized density versus pitch content Baking loss

-

Bx95 ■■·*··· HM 180M normal carbonization - ♦ - H M 180 M fast carbonization

Figure 6: Baking loss versus pitch content Coking yield \-m— Bx95 .-.*■·· HM 180M normal carbonization ~-»~-HM 180M fast carbonization I

Values in brackets represent standard deviations.

15%

16% Pitch content

Figure 7: Coking yield versus pitch content

891

Table 4: Norma carbonization program 1 Temp. Beg Temp. End Gradient °C/h °C °C 20 600 50 600 100 1,100 1,100 Total Carbonization Time

Time h 11.6 5 5 21.6

Table 5: Fast carbonization program Temp. Beg Temp. End Gradient °C/h °C °C 20 250 100 250 600 75 600 1,100 100 1,100 Total Carbonization Time

Time h 2.3 4.7 5 5 17

Specific electrical resistance

1

- B x 95 -~&-~~HM 180M normal carbonization —♦— HM 180M fast carbonization

70,0 „

65,0

a 60,0

14% Pitch content

Figure 10: Specific electrical resistance versus pitch content Air permeability \-m— Bx95 ~«~-HM 180M normal carbonization - ♦ — HM 180M fast carbonization I

With the high melting pitch the carbonization time is reduced about 4.6 h (21%) because the temperature gradient was increased in the lower temperature range. The faster carbonization program has only a small influence on the densities of the anodes. It might be possible to use faster heating programs for anode production. This could be an interesting aspect for anode producers because they can raise the throughput of the baking furnace.

Figure 11 : Air permeability versus pitch content

Compressive strength

Anode producers put much emphasis on thermal shock resistivity of their anodes. This parameter should be high when the anodes have a high compressive strength and a high Young's modulus. Figures 8 and 9 show these properties for the prepared anodes. If one uses the high-melting pitch both parameters could be raised about 35%. A very important quality parameter for anode producers is the specific electrical resistance (Figure 10). This parameter should be as low as possible to ensure low energy consumption during alumina electrolysis. This aspect is particular important for new high-amperage pots. For lab scale anodes the use of high softening pitch results in a decrease of the specific electrical resistance by more than 10 %.

-m— Bx 95 ~T&— HM 180M normal carbonization — · — HM 180M fast carbonization I

55

~~*

^-—-^?^*~

To'

——i

/ /

Q.

4Ü ; r

"

""

15%

Pitch content

16% Pitch content

Figure 8: Compressive strength versus pitch content

The other parameters are commented as follows: •

Young's modulus - B x 95 — Ä — H M 180M normal carbonization —♦—HM 180M fast carbonization |

•

Pitch content

Figure 9: Young's modulus versus pitch content

892

Air permeability: The higher baked density translates into a lower air permeability of the anode (Figure 11). C02-reactivity: This parameter is mostly influenced by the characteristics of the coke used for anode production. Therefore there are no big differences detected by using different pitches with the same coke for anode production.

References

Conclusion An interesting new property profile of laboratory carbon anodes is achieved by replacing a standard binder pitch by a high softening point binder pitch: • • • • •

The baked density of the anode is improved by +0.04 g/cm3 The specific electrical resistance is lowered by 15 % reducing electricity consumption. The pitch content in the anode is reduced from 16 to 15%. The pitch toxicity is reduced by 87 %. The baking process for the anodes can be expedited because of a lower volatile content in the pitch therefore the throughput of the baking furnace can be raised.

Of particular note is that new high-amperage pots will require anodes of much higher quality and the expected quality improvements associated with the use of high softening point pitches will be appreciated.

[1]

A. Mirtchi, L. Noel "Polycyclic aromatic Hydrocarbons (PAHs) in pitches used in the aluminum industry" Proc. Carbon (1994) 794

[2]

R. F. Willes et al. "Application of risk assessment to point sources of polycyclic aromatic hydrocarbons (PAHs)", Proc. 5th conference on toxic substances, Montreal, (1992), 75-100

[3]

W. Boenigk et al. "Production of low PAH pitch for use in Soederberg smelters", Light Metals (2002) 519

[4]

W. Boenigk, J. Stadelhofer "Coal-tar pitches with a reduced low-molecular PAH content", Proc. Carbon (1992)30

[5]

A. Alscher et al., " Evaluation of electrode binder pitches for the production of prebaked anodes using a bench scale process" Light Metals (1987) 483

[6]

W. Boenigk, A. Niehoff, R. Wildfφrster, "A highmelting coal-tar pitch as binder for anode production? A bench scale approach, Light Metals (1992) 581

[7]

S. Gerμ, "Granulation by drying from pasty phase in an inductively heated mixer drier", in Association Franηaise de Sιchage pour l'Industrie et l'Agriculture AFSIA, (2009) 36

Further investigations Next steps are planned to verify our results on full scale anodes. In addition high softening point pitches will be evaluated for other applications: •

For the production of graphite electrodes (in particular electrode nipples), carbon and graphite cathodes and specialty graphite, high softening point binder pitch would most likely improve the properties of these products significantly. The expected higher density and higher strength may make certain impregnation steps redundant.

•

For the production of dolomite and magnesia-carbon refractory bricks the use of high softening point binder pitch would most likely improve the brick properties and may make the tempering process redundant. Acknowledgement

This paper is based on many experiments. We thank Rüdiger Titze, who prepared and analyzed many lab scale anodes. We also express our gratitude to Mario Orozco who is a student of the University of Monterrey, Mexico. He attended an internship in our German plant and worked very hard to achieve the project objectives in a challenging short period. Last, but not least we thank EIRICH for their cooperation to introduce the lab scale mixer.

Light Metals 2011 Edited by: Stephen J. Lindsay IMS (ºÌ MLmml^ Metals & Materiate Society 2011

QUALITY AND PROCESS P E R F O R A M N C E OF ROTARY KILNS A N D SHAFT CALCINERS Les Edwards Rain CII Carbon LLC, 2627 Chestnut Ridge Rd, Kingwood, Texas, 77339, USA Carbon, Petroleum Coke, Calcination, Anode during anode baking and 3) transform the structure into an electrically conductive form of carbon. Green petroleum coke typically contains 9-13% volatile matter (VM). Real density (RD) is the most common measurement for tracking development of coke structure. Most anode producers prefer coke with an RD in the range of 2.04 - 2.08 g/cc. The industry trend is towards lower RD and several papers have been published on this recently [5,6]. Rain ΟΓ s experience with shaft calcined coke is based on coke with a typical RD of 2.03-2.04 g/cc.

Abstract Rotary kilns have been used successfully for many years to produce calcined coke for the aluminium industry and they offer a high level of automation, performance and flexibility. Shaft calciners make a high bulk density, coarse particle size product and several papers have been published recently highlighting these benefits. This paper presents a comparison of the merits of these two different calcining technologies from a product quality and process performance perspective. It addresses several misconceptions about the technologies related to operability, product quality and their ability to handle a wide range of green coke qualities. Both technologies will continue to be used in a complimentary manner in the future.

Rotary kilns are large diameter, sloped, refractory lined steelshelled cylinders which rotate during operation. Green coke is fed continuously in one end and calcined coke is discharged from the other end at 1200-1300°C. The coke bed loading in the kiln is low (7-10% of the cross-sectional area) as depicted in Figure 1. Heat is transferred to the coke bed predominantly by radiative and convective heat transfer from the counter-current gas stream and refractory lining. 40-50% of the VM is combusted inside the kiln and the rest is combusted in a pyrsocrubber upstream of the kiln. The VM combusted in the kiln provides most of the heat for calcination but natural gas, fuel oil and/or pure oxygen can be added to provide additional heat.

Introduction Several papers have been published over the last 5 years comparing rotary kiln calciners to shaft calciners [1,2,3,4]. Shaft calciners are common in China but there are very few operating outside of China. Rotary kilns on the other hand, have been the technology of choice for most of the rest of the world and the technology is generally well known and understood. Most of the papers referenced above do a good job of describing the differences between the two technologies and the generic differences in product quality. What is not covered in much detail, are differences in operation and some more subtle differences in product quality. This has led to a number of misconceptions about the pros and cons of each technology. Rain CII has operated rotary kilns for many years and has been a supplier of calcined coke to the aluminium industry for over 50 years. Rain CII's first experience with shaft calcined coke was in 2001 when it imported small volumes into the US for blending with rotary kiln product. Through a marketing arrangement, Rain CII has also supplied Chinese shaft calcined coke to the smelting industry for ~6 years. One such smelter, the Tomago smelter in Australia, has extensive experience using shaft calcined cokes.

Figure 1. Rotary kiln calciner & coke bed loading in kiln

In August 2009, Rain CII purchased a small 20,000 ton/year shaft calciner in China. The aim was to gain more direct operating experience with the technology. The combination of supplying shaft calcined coke to the aluminium industry and operating a shaft calciner has given the company a unique perspective on the merits of the two technologies.

A shaft calciner has multiple vertical refractory shafts surrounded by flue walls. The green coke is fed into the top and travels down through the shafts and exits through a water cooled jacket at the bottom, Figure 2. The movement of coke is controlled by opening a slide gate or rotary valve at the bottom of each shaft to discharge a small amount of coke. The discharge is intermittent (-every 20 minutes) and green coke is added to the top to maintain the feed.

The purpose of this paper is to present a suppliers perspective on the two technologies. Both have their pros and cons and these will be covered in detail in the paper. Combining the two technologies probably makes the most sense for the industry in the future.

The VM in a shaft furnace travels up through the coke bed and enters flue wall cavities at the top of the furnace. It is mixed with air at this point and then drawn down through a set of horizontally orientated flues. VM is combusted inside the flue walls and heat is conducted to the coke indirectly from the flue walls in an analogous manner to heat transfer in an anode bake furnace.

Brief Review of Rotary Kiln and Shaft Calcining The primary goals of calcining green coke are to: 1) remove volatile matter, 2) densify the structure to avoid shrinkage of coke

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Green Coke

In a shaft calciner, the slow heatup gives more time for thermal cracking and polymerization reactions and can almost be considered an extension of the delayed coking process. The typical mercury apparent density of shaft calcined coke is 1.791.82 g/cc. The same coke calcined in a rotary kiln might yield a mercury density of 1.73-1.77g/cc depending on how it is calcined. The differences in density between shaft calcined coke and rotary kiln coke are relatively small at low VM (9-10%) but increase as the green coke VM increases. Some high VM (>13.5%), fine particle size green cokes calcined in a rotary kiln can give mercury apparent densities as low as 1.67-1.70 g/cc. Such low densities may also be driven by structural differences between cokes (degree of isotropy) as well as VM content. An unfortunate quality trend for green cokes is increasing VM. The trend towards heavier, sour crudes increases vacuum resid. volumes and refineries have to increase the throughput of their cokers. This is typically achieved by lowering cycle times and reducing coking severity. Higher throughputs usually mean lower coker feed temperatures and both of these favor higher VM. Historically, the calcining industry has used cokes with VM in the 9-12% range. This range is growing, and cokes with VM >12% are used more and more frequently in blends.

Figure 2: Shaft furnace cross-section and shaft outlets With this brief overview, the major fundamental differences between the two technologies are as follows: • There is a large volume of counter-current gas flowing inside a rotary kiln over the top of the coke bed. The air or oxygen required for combustion of VM is added through primary, secondary and tertiary air fans. • The large gas volume in the kiln causes entrainment of fine green coke. Typically, -10% of the green coke is entrained in the flue gas stream and combusted in a pyroscrubber along with residual VMfromthe kiln. • There is very little counter-current gas flow inside a shaft calciner. The only flow is VM traveling up through the coke bed before it enters the top flue walls. • There is virtually no carryover of fine green coke into the flue walls in a shaft calciner. • The yield of calcined coke per ton of green coke is significantly higher in a shaft calciner. A typical yield in a rotary kiln (dry basis) is 77-79% which represents the yield after loss of VM and fines. The comparable yield for a shaft calciner is 85-88%. • The coke residence time is much shorter in a rotary kiln than a shaft calciner (-50 minutes vs 28-36 hours). This translates to a significantly higher heat-up rate for green coke in a rotary kiln (~50-100°C/min versus ~l-2°C/min in a shaft calciner).

Shaft calciners are much better at dealing with this problem than rotary kilns but there is a caveat to this which is discussed in more detail later. When high VM cokes are used in a shaft calciner, they must be blended with lower VM green cokes and calcined coke to maintain the average VM in a narrow range, typically 11-12%. Particle Size Another commonly reported benefit of a shaft calciner is coarser particle size. This is generally true with one significant qualifying comment. Shaft calcined coke has high levels of -75μιη fines. This can result in significant dusting problems that have not been mentioned in previous papers and which is not obvious when one first looks at the coke. Shaft calcined coke is very coarse when it comes out of the bottom of the shafts and +4.75mm (4 mesh) levels above 60% are common. The lower heatup rate minimizes explosive shattering of large particles caused by rapid VM release. Mechanical attrition of coke particles is also dramatically lower in a shaft calciner. The coke bed moves very slowly through the shafts. In a rotary kiln, the coke is tumbled to improve heat transfer and consistency of calcination and this reduces the average particle size. Mechanical handling of the coke in screw conveyors, bucket elevators, conveyor belt transitions and silo's also contributes to particle attrition in a rotary kiln calciner as previously reported [9].

Impact of Differences on Calcined Coke Quality

When the product from a shaft calciner coke is inspected more closely, a high level of agglomeration is evident. Coarse and fine particles are stuck together in odd shapes and sizes. This has been reported previously as a benefit rather than a problem [4,10]. The agglomerated coke pieces are usually very friable and easy to break by hand. The more it is handled after this, the more it breaks down and the more dust is generated. The scanning electron microscope images below show this quite clearly, Figure 3. Agglomeration occurs on both a macro and a micro scale.

Bulk Density & Apparent Density The differences outlined above have a significant impact on some calcined coke quality parameters. The most universally reported difference is the higher bulk and apparent density achieved with a shaft calciner. This is due to the slower heat up rate of green coke. The loss of VM creates porosity in coke. Lower VM gives lower porosity (and higher density) so lower VM green coke is always preferred. Porosity is also a function of the heat-up rate of the coke and this is a well known and documented phenomenon [7,8]. Faster heat-up rates create higher porosity and lower bulk density.

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removed by dust collectors and disposed of. It does not remove all the dust however, and additional dust is generated during vessel loading, unloading and use in the carbon plant. Ironically, the manual handling of calcined coke at many shaft calciners exacerbates the dusting problem for end-users. The coke typically receives very "gentle" treatment after calcining relative to what happens at a rotary kiln calciner where mechanical handling of the coke contributes to breakage and attrition. Real Density. Coke Reactivity and Impurity Levels There is no difference in the ability of a rotary kiln or shaft calciner to meet RD targets. This is contrary to what has been reported in some other papers and is based on Rain CIFs experience with multiple rotary kiln and shaft calciners. Both technologies are capable of making coke with RD's ranging from 2.01-2.10 g/cc and it is a matter of changing operating conditions to achieve this. When green coke quality changes significantly, it is easier and quicker to adjust a rotary kiln due to the high level of automation and multiple control variables available.

Figure 3: SEM images of shaft calcined coke agglomeration at low (lOx) and high (300x) magnification. The root cause of this goes back to one of the fundamental differences between the two calcining technologies. In a rotary kiln, most of the fine green coke (<250 μηι) is blown out the back of the kiln with the flue gas. In a shaft calciner, there is no mechanism to remove fine green coke and it stays with the product. The fine coke attached to the surface of coarser particles is easily abraded when the coke is handled and dust is generated.

There are also no fundamental differences in coke C0 2 and air reactivities between the technologies. Coke reactivities are dependent on impurity levels in the coke (S, Ca, Na, V etc) and the level of calcination. Many aluminium companies have moved away from coke C0 2 and air reactivity specifications because they typically show no or little correlation to anode C0 2 and air reactivities [11]. The latter are dependent on a range of factors including coke impurity levels, butts cleaning efficiency (which controls anode Na levels), anode forming and anode baking level and consistency. Calcination technology also has no influence on coke impurity levels. These depend on green coke impurity levels.

This phenomenon was not known by Rain CII when the company began its experience with shaft calcined coke. Similarly, Tomago was unaware of the issue until they increased their use of shaft calcined coke. Dusting problems become more evident during vessel unloading and subsequent handling in the carbon plant. Tomago uses a modern vacuum unloading system and they noticed a significantly higher -75μπι dust content in shaft calcined coke relative to rotary kiln coke. A comparison of the -75μιη dust content of typical shaft calcined coke and rotary kiln calcined coke after vessel unloading is shown below for nine different shipments, Figure 4.

Operation of Rotary Kilns Versus Shaft Calciners Operabilitv with High VM Cokes There is a general belief that shaft calciners handle high VM cokes much better than rotary kilns. This is the caveat mentioned earlier that needs some clarification. It is very important when operating a shaft calciner, to control the average VM of the feed within a narrow range, typically 11-12% (these numbers are not absolute since VM values depend on the method used).

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Figure 4: -75um fines content of vessel unload samples. Samples of shaft calcined coke taken before vessel loading do not typically show a high -75 urn content. Breakage and attrition starts occurring during vessel loading and thereafter. Shaft calcined coke can also "hang-up" on the inside of the cargo holds at very high angles of repose (>60°). This is believed to be due to breakage of agglomerated coke during shipping resulting in higher packing densities. The higher density and more irregular particle shape of shaft calcined coke probably also plays a role. Shaft calciners that export coke typically need to screen and crush oversize lumps (at 30mm or 50mm). The crushed oversize coke is then recombined with undersize coke and de-dust oil is added. This process generates significant amounts of dust which must be

The need for careful VM control of the feed is driven by the processes occurring inside a shaft calciner. High VM coke produces more condensable hydrocarbons when heated. The coke softens and becomes "sticky" before the residual tars thermally crack into lower molecular weight volatile species. The coke bed is relatively densely packed and when the coke becomes "sticky" it acts like a glue that causes particles to stick together. This is what causes the particle agglomeration that is evident both visually and microscopically. If the average VM of the feed is too high, the coke bed will literally fuse together inside the furnace and result in a solid plug

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very labor intensive and a 200,000 MT/yr plant may require 150 employees vs 30 for an equivalent rotary kiln calciner.

that blocks the shafts. If this happens, the coke must be physically removed before operation can continue. This is hazardous and damaging to the refractory shafts and must be avoided.

Automation of the green coke feed and calcined coke discharge systems at shaft calciners is straightforward and becoming more prevalent in China. It does add some additional cost to the construction and requires development of maintenance systems to maintain equipment such as belt conveyors. The high number of individual shafts requires a lot of replication of equipment. Automation can help avoid problems such as those associated with allowing feed hoppers to run low which can cause visible emissions of VM into the work environment.

The best way to illustrate this phenomenon is to show a photograph of two green coke samples after a VM test. In a VM test, green coke is crushed to -0.25 mm, dried and then placed in a crucible which is heated at 950°C for 7 minutes. Most low VM cokes go into the crucible as a powder and come out as a freeflowing powder. High VM cokes (>~13%) come out as solid, fused, disk. Some particles appear to melt and assume the same flat profile as the crucible bottom, Figure 5. The higher the VM, the more deformation that occurs and the harder the pellets are.

One aspect of a shaft calciner operation which is more difficult to automate is the frequent cleaning required inside fluewalls. Tar builds up inside thefluewallsdue to inefficient combustion of VM and the small amount of fines carried into the flues. The buildup must be removed manually by scraping with a rod. Cleaning can be as frequent as every 2 hours and is dictated to some extent by the green coke type, VM content and particle size. To eliminate this, a change in the sophistication of the combustion control and draft systems would be required. Today, combustion air is added via manually operated sliding metal dampers or in some cases, refractory bricks. A main guillotine damper on the flue gas duct entering the stack can also be adjusted to change the overall furnace draft. Buildup of tar inside the fluewalls changes the draft level and must be removed to maintain a stable operation and ensure proper combustion of VM.

Figure 5: Samples after green coke VM test. In addition to VM content, green coke particle sizing also influences coke bed agglomeration. Shaft calciners do not operate well with fine green coke. It reduces the bed permeability for VM release and the high surface area causes more agglomeration which increases the risk of shaft plug-ups. Shaft calciner operators are all very aware of these problems and they are controlled by increasing recycle rates.

Waste Heat Recovery, Safety and Environment All rotary kilns built in the last 20 years (and many built prior to this) have sophisticated waste heat recovery systems. The waste heat is used to fire boilers and steam is either sold to neighboring process plants or used to drive turbines for power generation. The two kilns at Rain CIP s Visakhapatnam plant in India (300,000MT/yr each), generate enough waste heat to drive a 50MW turbine generator. At current energy prices, waste heat is valuable and considered a "green" form of energy since it avoids the need for thermal power generators to burn coal or other fuel.

Rotary kilns are more forgiving in their ability to operate with a wider range of green coke VM. Higher VM cokes will however, more negatively affect the coke bulk density - much more so than in a shaft calciner. This is the major benefit of a shaft calciner relative to a rotary kiln. It can still make a dense product when high VM cokes are used in the blend.

Waste heat recovery systems are also being used more and more with shaft calciners. Smaller plants use waste heat for lower grade heat applications such as preheating pitch and paste if they are attached to an anode plant. The principle of application is identical to a rotary kiln but waste heat production is lower (per ton of coke) due to the lower fines carryover. This represents lost value in terms of waste heat recovery but must be compared to the lower green coke cost per ton of calcined coke produced

If the average VM level of the feed to a rotary kiln gets above -13.5%, it can also be problematic. A layer of coke can build up on the inside of the refractory lining at the feed end of the kiln (known as a coke ring) and this reduces the effective diameter of the kiln and eventually reduces the production rate. An excellent paper was published on this problem over 30 years ago [12]. Automation Differences in the level of automation between rotary kilns and shaft calciners have been well covered in previous papers. Modern rotary kilns are highly automated and can be operated with a small labor force. There is no physical handling or contact with green and calcined coke and the entire process can be operated from a central control room. Issues occur from time-to-time occur that require manual intervention but they are not common.

Waste heat recovery and flue gas desulfurization systems are very advanced and easy to operate with a rotary kiln. Controlling the flue gas flow and draft in one large rotary kiln producing 300,000 MT/yr of coke is easier in principal than controlling drafts in multiple furnaces and fluewalls in a shaft calcining plant of similar capacity. A shaft calciner producing 300,000 MT/yr comprises -350-400 shafts each with its own set of fluewalls and manual damper systems to control airflows. It should be possible to increase the capacity of individual shafts by making them longer but the maximum width will always be limited by heat transfer just like an anode bake furnace. The number of shafts will continue to remain relatively high even in newly built furnaces.

Many shaft calciners in China are operated manually and green coke is lifted to the top of the shafts in small hoppers with electric hoists and then added to the shafts using an overhead crane. Calcined coke is discharged into carts and then hand wheeled and dumped into storage areas. Operating a shaft calciner this way is

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Construction, Startup, Operation and Shutdown A major benefit of a shaft calciner (at least in China) is that it can be built cheaply and quickly at a very small scale. A company may choose to build a 20,000 ton/yr or a 200,000 ton/yr calciner because the technology is readily scaleable. By comparison, modern rotary kilns are large: 250,000-300,000 MT is a standard size today and necessary to achieve economies of scale. The volume and mass of refractory required to build a shaft calciner is significantly higher than a rotary kiln of comparable size (~2.5x the mass). This does not present a problem in China where refractory costs and refractory labor costs are low but it can be problematic elsewhere. In an internal study completed at Rain CII several years ago, the estimated cost of building a shaft calciner in the US was ~1.5x the cost of a rotary kiln on a $/ton of installed capacity basis. This assumed the same level of pollution controls as a rotary kiln. In 2007, Rain CII engaged a Chinese engineering company to estimate the capital cost of building a large shaft calciner and rotary kiln calciner in China with waste heat recover and flue gas desulfurization. The capital costs estimates were very similar.

Figure 7: Expansion tie-rods with adjustable tensioning system. This problem is easy to manage with a stable green coke supply and a low level of automation. It becomes more problematic as automation is added to the calciner. A fully automated shaft calciner with belt conveyors and a waste heat recovery system requiring pumps, fans etc. now becomes more dependent on the reliability and maintenance of ancillary equipment. Contingencies for unplanned events like power failures could get quite complicated when there is no ability to shut a furnace down.

All shaft calciners in China are built with high silica (>95% Si02) refractory bricks for the shaft and flue walls. These bricks are readily available at low cost in China but their use has one significant drawback. Silica brick undergoes a very high thermal expansion compared to the 60-70% alumina bricks used in rotary kilns or the 40-60% alumina bricks used in baking furnaces. The coefficient of thermal expansion (CTE) is shown in Figure 6. As a result, shaft calciners have to be started very, very slowly and startup typically takes ~2 months.

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Rotary kilns on the other hand, are easy to startup and shutdown. It is typically a 3-4 day process to start a cold kiln using a natural gas or oil fired burner. The green coke feed to a rotary kiln can be stopped easily at any time without the risk of refractory damage. This is partly due to the higher grade (and higher cost) bricks used and their lower expansion rates but also the very simple refractory geometry and design used in a rotary kiln. The brick linings in rotary kilns are very stable. Bricks in the lower temperature, feed-end can last >20 years. Bricks in the hotter discharge-end, need to be replaced more frequently (typically every 4-5 years) and the repairs are staggered so small sections are replaced at a time. Most kilns will take a 2-3 week shutdown every 12-18 months for refractory and other plant maintenance work. The highest refractory wear components are tertiary air nozzles (for kilns that have them) which protrude into the kiln. These typically drive the maintenance cycle. Refractory in the cooler is usually also repaired every 12-18 months.

1500

Temperature (C)

During operation, there are no refractory maintenance costs with shaft calciners and this is a very attractive feature. Small shaft calciners with little automation and no waste heat recovery can be built and operated very cheaply. The furnaces are generally run to failure and this can be anywhere from 5-10 years. Replacement after 5 years is costly and a more typical life is 8 years. Replacement of the furnaces typically takes 6 months so there is a lot of incentive to run the furnace as long as possible. The life of a shaft calciner can be increased by running the furnace with a stable green coke feed and good operating practices. Maintaining clean fluewalls, drafts and temperature profiles is important for prolonging the life of the furnace.

Figure 6: CTE of high silica brick vs 60% alumina brick. Expansion tie rods with a spring assembly must be adjusted carefully to avoid cracking and damage during the critical expansion range from 25-700°C, Figure 7. Once a shaft calciner reaches its operating temperature, the high silica bricks are very stable and have a low thermal expansion coefficient above 750°C. High silica brick is an excellent material for any refractory application like this involving continuous high temperatures. The thermal expansion of the bricks is reversible however, and once a shaft calciner is started, it cannot be shut down. This is generally not well known outside of China but all shaft calciner operators in China are very aware of this. It would be very difficult, probably impossible to cool a shaft calciner uniformly enough to avoid cracking the high silica flue-wall bricks. Coke oven batteries used to make metallurgical coke use the same type of refractory brick and suffer the same problem.

Discussion The objective of this paper is to present a balanced view of the merits of rotary kiln and shaft calcining technology. Both technologies are widely used (albeit one predominantly in China) and both will continue to be used in the future.

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technologies have a strong future. The pros and cons of each need to be recognized.

Shaft calciners make a high bulk density product and this is very appealing for a smelter and anode plant. The coke can be very dusty however, and may require modification of dust collection systems in carbon plants if it is used in high volumes. Shaft calciners do a better job of making a high density product with high VM green coke but high VM cokes are typically softer and finer in particle size and this can compound dusting problems. The VM of the feed to the furnace needs to be kept constant and whilst this is not difficult, it ultimately translates to lower production. This is easily solved by building more shafts.

Shaft calciners are easy and fast to build but are more suited to regions with low labor rates and low cost refractories. They make an excellent bulk density product but dusting can a problem. Automation of shaft calciners is increasing but there are some practical limits which will constrain the level automation relative to a rotary kiln. The aluminium industry needs both of these calcining technologies in the future. The products from each can be blended together in a complimentary manner. The addition of 10-20% shaft calcined coke to a rotary kiln product can provide a good density boost if needed and will minimize the impact of dusting problems. This could serve as a good model for the industry to pursue as green coke VM continues to increase.

The trend towards higher VM green cokes makes shaft calcining attractive in terms of density but the technology does not deal as well with the finer particle size that comes with higher VM. Rotary kilns have a built in mechanism to remove fine green coke via entrainment in the flue gas stream. As long as the kiln has waste heat recovery, value will be recovered from these fines. The level of automation is improving with shaft calciners but there are some fundamental problems related to tar buildup that will be difficult to overcome without improvements in combustion control. This ultimately means a higher capital cost if more sophisticated control systems are used. A shaft calciner is unlikely to ever compete with a rotary kiln in terms of manpower requirements, ease of operation and flexibility. It is very easy to startup and shutdown rotary kilns. The more a shaft calciner is automated, the higher the capital and maintenance costs and the more it starts resembling a rotary kiln. The inability to shut a shaft calciner down becomes more of a problem when the level of automation is increased.

References Sun Yi, Xu Haifei, Wang Yubin, Cui Yinhe and Liu Chaodong, "The Comparison Between Vertical Shaft Furnace and Rotary Kiln for Petroleum Coke Calcination," Light Metals, 2010, 917-921. 2. Kenneth Ries, "Enhancing Coke Bulk Density Through the Use of Alternate Calcining Technologies," Light Metals, 2009, 945-949. 3. Guanghui Lang, Rui Liu, Kangxing, "Characteristic and Development of Production Technology of Carbon Anode in China," Light Metals, 2008, 929-934 4. Raymond Perruchoud, Timea Tordai, Ulrich Mannweiler and Liu Fengquin, "Coke Calcination Rotary Kiln vs Shaft Calcining," (Paper presented at 2 nd International Carbon Conference, Kunming, Sept 17-19, 2006). 5. Jeιrιmie Lhuissier, Lailah Bezamanifary, Magali Gendre, Marie-Josιe-Chollier, "Use of Under-Calcined Coke for the Production of Low Reactivity Anodes," 2009, 979-983. 6. Marie-Josιe-Chollier, A. Gagnon, C. Boulanger, D. Lepage, G. Savard, G. Bouchard, C. Lagacι, A. Charette, "Anode Reactivity: Effect of Coke Calcination Level," Light Metals, 2009, 905-908. 7. Paul Rhedey, "Structural Changes in Petroleum Coke During Calcining," Transactions of the Metallurgical Society of AIME, 1967, 1084-1091. 8. D. Kocaefe, A. Charette, L. Castonguay, "Green Coke Pyrolysis: Investigation of Simultaneous Changes in Gas and Solid Phases," Fuel, 74 (6), 1995, 791-799. 9. R.M. Garbarino, R.J. Brown, M.F. Vogt and S.A. Vogt, "Particle Degradation During Coke Handling," Light Metals, 1995, 454-548. 10. Liu Fengqin, "Chinese Raw Materials for Anode Manufacturing," R&D Carbon, Sierre, 1st Edition, 2004 11. B. Desgroseilliers, Lise Lavigne and Andre Proulx, "Alcan Approach for Evaluation and Selection of Coke Sources," Light Metals, 1994, 593-596. 12. D. DuTremblay, P.J. Rhedey and H.Boden, "Agglomeration Tendency of Petroluem Coke," Light Metals, 1979, 607-621 1.

A refinery building a calcining plant as an off-take for green coke would be unlikely to ever build a shaft calciner. A case in point is the rotary kiln calciner commissioned in mid-2010 at the CNOOC refinery in Huizhou, China. It is a modern rotary kiln with waste heat recovery and is much more in line with the refinery's standards for automation, manpower requirements, environmental and safety performance, flexibility and so on. Perhaps the most significant benefit of shaft calcining is the low cost to build and operate small calciners in some parts of the world. Rotary kiln calciners are better suited to large volume applications which make the most of their economy of scale advantages. This makes the combination of a shaft calciner and a rotary kiln potentially very attractive. A relatively small amount of shaft calcined coke (10-20%) can provide a significant density boost to a rotary kiln product without creating too many dusting problems. This can be considered a recommendation to the industry when considering these two calcining technologies. Both have their pros and cons and some of these are regional in terms of capital and operating cost. Rotary kilns are very efficient at producing high volumes of coke and they will continue to be built and used in many parts of the world. Shaft calciners produce a high bulk density product that can be used in a complimentary way with rotary kiln product to increase average bulk density. Blending the two cokes together makes a lot of sense for the industry. Conclusions Based on experience supplying both rotary kiln and shaft calcined coke to the aluminium industry, Rain CII believes that both

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SUBSURFACE CARBON DIOXIDE REACTION IN ANODES Donald P. Ziegler Alcoa Canada Primary Metals, Aluminerie de Deschambault, 1 Boulevard des Sources, Deschambault-Grondines, Quιbec, G0A ISO, Canada Keywords: anode reaction, carbon dioxide, permeability, diffusion, tortuosity, dusting dominant influence on internal mass transport in the sub-surface zone above the working face of the anodes.

Abstract Formation of carbon dust in electrolysis is linked to reaction of the anodes with carbon dioxide. The general understanding is that the reaction takes place inside the anodes, below the bath surface and rather more toward the sides of the anodes than the bottom. Given this, a relevant question is the relative importance of transport through the anode compared to its intrinsic reactivity. To provide quantitative answers, a transport-reaction model has been developed. The key finding is that the extent of reaction is insensitive to the permeability. This is because the reaction produces two moles of CO for every mole of C0 2 consumed, so that there is a net flow away from the reaction locale. Consequently, fresh reactant must be supplied by diffusion rather than convection. Since these two processes are governed by different material properties, this finding opens the possibility of new approaches to optimization of anode structure.

This concept, if valid, may have important implications for the interpretation of anode quality measures. A rather simplistic assessment of the situation might be that given sufficient time for the reaction to take place, the gas exiting the anode would be in equilibrium with the carbon, i.e. the reaction would be under mass transport control. Alternatively, if the reaction has not gone to equilibrium, then a change in the intrinsic reaction kinetics will have an important effect on the overall reaction rate. In laboratory anode quality measurement, these two aspects are measured separately: permeability and reactivity are measured independently of each other. It is not clear at first glance which of these might be more important to the behavior of the anodes in the cells. (The reactivity is measured with a test that exposes the outside of an anode sample to a stream of carbon dioxide.) A mistake in interpretation could have serious consequences: if the reaction in the cells is under mass transport control, then a strong focus on the intrinsic or chemical reactivity might be misplaced.

Introduction Carbon dust in Hall cells is an operational problem in itself, is associated with anode deformations, and is a loss mechanism resulting in increased net carbon consumption. Dust formation is thought to occur due to preferential oxidation of the more reactive pitch coke, allowing for the disintegration of the rest of the anode structure for which the pitch coke forms the glue. The purpose of the present study is a preliminary analysis, using simple models, of several aspects of the reactions leading to dust formation.

In the work cited, Sadler and Algie attempted to remedy this situation by measuring the reaction rate using an apparatus in which carbon dioxide was made to flow through a sample of anode carbon by imposition of a pressure head. As that apparatus is time consuming to operate and so has not been developed into a quality control tool, the desirability remains for some means to link and weight appropriately the contributions of the intrinsic reactivity and anode permeability, particularly under conditions relevant to operating anodes. This attraction was the basis of this study, which started as a model investigation of the reaction situation described in the Sadler quote above.

Based on observations of industrial anodes, the air-burning reactions are not believed to be strongly responsible for formation of carbon dust. Rather, either the Boudouard reaction or an electrochemical reaction has been determined to be culpable. The focus of the present study is the former: C + C02=2CO

Our first models in this study were developed on the scale of the anodes, with sizes of the order of half a meter. As will be described, this model did not properly represent the reported effect of the permeability on the amount of internal reaction. Consequently, another effort was made in the development of a model of the Sadler-Algie experiment, which is a time-dependent version of the first model, and on the smaller geometric scale of a few centimeters.

(1)

Near cell operating temperature, the equilibrium of this reaction is far to the right. That is, if there is some carbon dioxide available, most of it will react with any excess carbon to form carbon monoxide. The initial motivation for the present study was an observation published by Sadler and Algie [1]:

Through the use of these models, an attempt has been made to gain some additional understanding of the corrosion of anodes leading to dust formation. The models are of a preliminary nature, and perhaps will serve to frame and guide additional research in this area.

Mass transport of carbon dioxide through the pore structure of anodes will predominantly be via viscous flow and hence related to anode permeability. The driving force for this flow is a hydrostatic pressure of around 200 mm water gauge generated at the electrolytic face of the anode ... This hydrostatic head is likely to be the

Finally, an alternative hypothesis has been proposed in the literature to explain carbon dust formation: preferential electrochemical attack of the binder, rather than preferential oxidation by carbon dioxide. This hypothesis was advanced by

901

Euel Cutshall [2] and found some supporting evidence from FarrWharton [3]. For a number of reasons beyond the scope of this paper, the present author is inclined to support the hypothesis of chemical reaction with C0 2 rather than the electrochemical, and additional discussion is limited to the chemical mechanism.

burn

burn

The Model Equations: This model is based on flow of carbon dioxide, and any carbon monoxide formed, through the anode under the influence of a pressure gradient, simultaneously with their diffusion under a concentration gradients. For each gas, this is modeled with a convection diffusion equation:

V {ü [CO]) = V (Deff V[CO\)A 2R

(3)

The convective velocity is modeled using Darcy's law:

(5)

The pressure in turn is calculated from the concentrations of the two gases using the ideal gas law:

(6)

The equilibrium constant for equation [1] is about 50 atm. at 950 C. [4]. This means that at 1 atm. total pressure, the reaction will proceed until 98% of the gas phase is CO. As will be seen from the results below, in the reaction zone the CO2 concentration remains close to the concentration at the surface, so that the neglect of the equilibrium is appropriate for the present level of detail. On the other hand, the kinetics of this reaction are known to follow what is known as the Langmuir-Hinshelwood model in which the reaction is inhibited by both carbon monoxide and carbon dioxide (e.g. [5, 6]). In view of the exploratory nature of this study and the difficulty of finding a suitable set of rate constants for anode materials, this complication has been neglected. However, this may introduce some inaccuracy and should be considered as a topic for further study.

eD

(9)

The gas properties: The gas density is from the ideal gas law. The diffusivity is an extrapolation from C0 2 data in CRC Handbook of Chemistry and Physics (85th ed., CRC Press, 2004). (No distinction is made between the diffusivities of CO and C02.) The viscosity is uniform, and that of a 50-50 mixture of CO and CO2. The values of all of the properties are evaluated at 960 C using the formulas given below the list of symbols.

K=

As presently implemented, the effective diffusivity is modeled with no dispersion contribution:

Deff =

MWC 4 . F 1 0 0 0 . / > App

The anode parameters need to include some microstructural information. Generally, the permeability and specific surface area are related, especially for materials with simple structures. The initial intention was to use the Kozeny-Carman relationship, equation (10) between the two to eliminate one parameter from the model. To apply this equation to the anodes, we have considerable permeability data and a few measurements of specific surface area. A typical value of the permeability is 0.5 nPm (nano-perms, equivalent to 10"13 m2). For the specific surface area, Cutshall reported values in the neighborhood of 1.5 x 106 m2/m3, but some other measurements, particularly of the component coke particles can range higher. Using these values for the surface area and the usual voidfraction,it appears that the Kozeny-Carman relationship does not fit well, at least this version of the relationship. This is indicative of a more complex microstructure. Other forms of the equation are available in which the material length scale is provided by a pore or particle size; all of these approaches require a suitable choice of a mean value of the length scale parameter to make the equation work. This might be a fruitful area for further study, but within the present preliminary scope, no additional work was done to relate the permeability to the other properties.

The gas density varies as the gas composition varies throughout the anode. The reaction rate, àί, is modeled as a simple first order reaction:

R=ks\co2]-[coe2q]i

(8)

Parameters, Domain, Boundary Conditions and Solution:

(2)

P = ([CO\- [CO2])RT

R

Equations (2), (3) and (8) comprise the equations for the three field variables in the model with the other equations being the auxiliaries needed for closure. These were solved using the FlexPDE finite element package, generally with no problems.

V {ü [C02]) = V (Deff V[C02])- R

(4)

MWC 1000. p,real

where the downward anode velocity is given by

Steady-State Anode Reaction Model

ά = -—VP ì

d€ dy

196. g 3 S2(l-£)2

(10)

Two parameters are needed in the chemical reaction rate expression, the rate constant and the specific surface area. Rate constants for this reaction have been measured for a variety of types of carbon, however, in view of the scarcity of data related to anode carbons, the use of the simplified first-order kinetics and the present scope, the rate constant has been taken as a variable parameter. Also, without a relationship linking the surface area to the permeability, the surface area only appears in the model when multiplied by the rate constant. Under the circumstances, the product can be taken as a single parameter, which is the

(7)

The final field equation describes the evolution of the porosity or void fraction:

902

approached used for our first results. By fitting the results to some measured reaction rates, we hope to be able to approximate a suitable value for this joint parameter.

General Characteristics: A very important feature which is essential to the interpretation of the results is the net volume change associated with the Boudouard reaction: two volumes of CO are produced for each volume of C0 2 reacted. There will be a net volumetric outflow from regions where the reaction is taking place. For the one dimensional model, when the flow resistance of the anode is high enough (thick anode or low permeability), there will be little flow through the anode. Under these circumstances, the reaction will cause a convective flow out of the reaction zone and the reaction is sustained by fresh C0 2 which must diffuse upwind against the convective flow outward. The reaction causes a high pressure zone inside the anode. For a case near the middle of the range of rate constant-surface area and permeability, the maximum pressure is 103760 Pa, it occurs at 0.17 m above the anode bottom, and the flow is outward, both through the bottom and top of the anode, with the superficial velocity equal to -2.E-5 m/s at the bottom and about 1.5E-5 m/s at the top. The porosity increases from the initial value of 0.238 to 0.251 at the bottom. The carbon consumed by this reaction is 1.7% of the electrolytic consumption, in the range of what has been estimated for this reaction, (e.g. 5 to 6% is reported in [7], but that would include the higher extent of reaction on the sides of the anode.)

The final microstructural parameter is the tortuosity which appears in the relationship for the effective diffusivity. This parameter relates primarily to the shape of the pores and the sinuosity of the paths that they make for diffusion. A value of three was used for the initial calculations, subsequently allowed to vary. The apparent and real densities of the anode are set to 1600 and 2100 kg/m3, respectively, which gives a void fraction at the top of the anode of 0.238. The current density, used for the anode disappearance at the bottom, is set to 9000 A/m2. Together these parameters give a burn rate or descent velocity of the anode of 1.75 xlO"7 m/s. This model has been solved on two geometries. The first corresponds to a one-dimensional column through an anode. The bottom of the anode is immersed 0.12 m into bath of density 2100 kg/m3. Atmospheric pressure is taken as 105 Pa, so that the C0 2 pressure on the bottom of the anode is 102470. Pa. At this surface, the gas is assumed to be 100% C0 2 ; the concentration is calculated from the ideal gas law. At the top of the 0.60 m thick anode the C0 2 concentration is set to zero and the CO concentration corresponds to 1 atmosphere total pressure. These boundary conditions may not correspond exactly to the physical condition. At a boundary where the gas is flowing outward a different condition may be appropriate, perhaps one with no concentration gradient. However, it is not immediately evident how this should be specified, and for the ranges of the parameters that seem appropriate, the results are relatively insensitive to these boundary conditions.

The sensitivity of this picture to the rate constant-surface area product is shown in Figure 1. Here we see that when the reaction rate is very low, the amount of gas produced is not enough to counterbalance the hydrostatic head imposed by the immersion, and all of the gas flows upward and out the top of the anode. As the rate increases, the maximum pressure becomes linear in the rate constant, and gas escapes from the bottom as well as the top at increasing velocity. Also from this sensitivity analysis, we can get a perspective on the maximum void fraction and the net rate of Boudouard parasitism: with the rate constant-surface area product near 50.E-4 s"1, the void fraction on the bottom face is about 30% and the totalfractionreacted is 10-12%. These values indicate the upper limit on the value of the rate constant-surface area parameter that we should consider.

The second geometry corresponds to an extension of the first to include the corner and side of the anode. The width of the domain is 0.3 m, which, with the plane of symmetry in the center would correspond to an anode of 0.6 m width. The immersed corner of the anode is a circular arc with a radius of 0.04 m. The boundary conditions remain essentially the same: hydrostatic pressure and 100% C0 2 for the immersed parts, and 0% C0 2 and atmospheric pressure for the parts above the bath. Preliminary results with this geometry indicated the possibility of higher extent of reaction than with the 1-D model, and so the model was also extended to include the evolution of permeability and specific surface area.

t:

- Pmax

"-«a

:

·^".ν:;;;:;

VbOtt

/

'"?

The specific surface area is assumed to evolve along a parabolic trajectory as a function of the void fraction from its initial specified value through a terminal value of 0.0 when the porosity reaches unity. As the parabola has three coefficients, an additional piece of information is required to complete its specification. For the present, this additional information is the slope of the line fit to the data in Table III of Cutshall and Bullough [8], 1.44E7 m2/m3.

/

/

/...^.

\

/ _ kr x S x IO4 (s 1 ) Figure 1: Effect of the rate constant-surface area product on the maximum pressure and superficial velocity at the bottom of the anode.

The permeability is assumed to evolve according to an equation of the form:

Effect of Permeability - 1-D Model: The effect of the permeability is a main theme of this paper. The effect calculated

Results and Discussion:

903

with this model is rather small, which is at variance with the experimental results of Sadler and Algie [1]. Consequently, the question of why the discrepancy needs to be answered. We are placing considerable reliance on the data of Sadler because his experiments were intended to mimic more closely the conditions experienced by the anodes, with the possibility of forced flow of C0 2 through the anode driven by the higher hydrostatic pressure on the working face. This is in contrast to the usual measurement of CO2 reactivity, wherein a cylindrical sample is exposed on its outside to a stream of the reacting gas. Also, the anodes used by Sadler were made using a simple recipe of multiple size fractions of the same coke and a single pitch, eliminating a range of other influences. Other experimental results, for example Ross, et. al. [9], which show a correlation between a conventionally measured C0 2 reactivity and permeability, indicate a much lower sensitivity of the reaction rate to the permeability. (This is not entirely a clear comparison, however, as Sadler's experimental conditions and the standard reactivity tests are different.)

1.09 1.08 _ Ό *

a Cu

E

On

1.07

Permeability x 1013 (m2): ···<·)

:

1.0

0.5

1.06 1.05 1.04 1.03 1.02

O.l

Figure 2 shows the effect of the rate constant-surface area product on the maximum pressure inside the anode for two levels of permeability. Similar plots showing the effect on the velocity leaving the bottom of the anode, on the maximum void fraction and on the total reaction rate would show almost no effect of the permeability. Thus we see that changing the permeability does very little except to change the internal pressure in the anode.

1

10 4

krxSxl0 (s

_1

100

)

Figure 2: Maximum pressure in the anode vs. rate constantsurface area product for different values of permeability. So, what we see is that of the two mechanisms for mass trasnsfer, diffusion and convection, the reaction rate depends much more strongly on the parameters that control the diffusion. The tortuosity may or may not be correlated with the permeability, depending on the nature of the microstructure. Some versions of the Kozeny-Carman relationship include it, in which the permeability is inversely proportional to the square of the tortuosity [11]. Generally, it is associated with the distance that material needs to diffuse, or the length of the pore per unit length of the matrix in the flow direction, although other effects such as constrictions can play a role. Tortuosity is generally not measured in anodes. Sufficient data on effective diffusivity of hydrogen is given in a pair of papers by Walker et. al. [12, 13] to estimate the tortuosity of some carbon materials, which works out to range from 4.5 to 6.8. These are rather large compared to other materials, which more typically are around two and ranging to 4 [14].

At this stage, it might appear that the present results indicate kinetic control, whereas the experiments indicate mass transfer control. With the present model, it would be difficult to change to a mass transfer controlled regime: either the rate constant would have to be increased considerably, or the permeability decreased. The first of these would result in much higher net reaction rates, representing a large fraction of the total carbon consumption, while the second would require permeability much smaller than the measured values. However, a closer look indicates that the diffusive flux is considerably larger than the convective flux, and that discussion of mass transfer control should include the diffusive effects. As a next logical step we examine the geometric effects, moving away from this 1-D representation of a thick anode. Two-dimensional Results: The evidence linking the formation of carbon dust to the sides of the anodes that are submerged in the bath appears solid. In addition to Cutshall and Farr-Wharton results already mentioned, macrographic examination on anode butts in an earlier work by Sadler and Algie showed deeper extent and more intense attack on the anode sides than on the working face [10]. Figure 3a, 3b and 3c show the pressure, reaction rate and porosity distributions from the model now extended to 2-D. Figures 3b and 3c show that the reaction rate and porosity are higher at the side of the anode than on the working face. This is the expectation, supporting the hypothesis that this mechanism is responsible for the increased formation of carbon dust on the sides of the anode, and perhaps indicating that we are arriving at a quantification of this phenomenon which will allow us to attach a proper weighting to the measured anode properties. However, similarly as the 1-D results, a plot of the integrated reaction rate vs. the reaction rate constant would show overlapping curves for differing values of the permeability. That is, the permeability has very little effect on the reaction rate. On the other hand, Figure 4 shows a marked effect of the tortuosity, here for the maximum void fraction, although the net reaction rate plot is quite similar.

2 4 6 8 10 12 14 Reaction Rate Constant (m s"1) Figure 4: Maximum void fraction vs. reaction rate constant, showing the effect of the tortuosity.

904

Transient Model - Sadler Experiement The above results fit the reports in the literature that carbon dust forms mostly on the sides of the anode below the bath surface. However, Sadler's experimental results show a strong relationship between the anode permeability and the reaction rate, which does not appear in the model results. But since Sadler's experiment was run under different conditions than an operating anode, there is the possibility that some aspect of this difference could be the cause of the differing sensitivity to permeability. Consequently, the model was modified in order to more closely duplicate Sadler's experiment. Transient terms were added to each equation, an additional equation like (2) was added for the argon diluent gas, and the convective term with v t ^ in (8) was set to zero. The domain boundaries and boundary conditions were modified to fit the samples of Sadler's experiment..

1.03E5

. ,

^1 ^

St "

1.00E5 0-t

0.2

03

A first model run was completed corresponding to the initial permeability of 6.7 cD, or 0.66 x 10"13 m2, of Sadler's laboratory anode 2. The initial porosity was calculated from the measured density of 1557. and an assumed real density of 2100. kg/m3. The initial specific surface area was calculated from equation [10] to be 9.66 x 106 m2/m3, based on the specified initial permeability and porosity. The rate constant was adjusted to 33. x 10"10 m/s to give an overall fraction reacted of 0.1051, which compares with the measured value of 0.103. (from Table 3 of (1)).

0.4

Figure 3a: Pressure distribution for reaction rate constant = 1.33 E-10, S0=1.5E6, Ko=0.5E-13 andx=3.

2.4E-3

Using this value of the rate constant, but adjusting the permeability to correspond to the measured value of Laboratory Anode 1, 15.5 cD, or 1.53 x 10"13 m2, results in an overall calculated fraction reacted of 0.1056, practically the same as for Anode 2, although the permeability increased by a factor of 2.3. In comparison, the measured fraction converted increased to 0.239, nearly exactly proportional to the increase in permeability. Thus we see that even with the model of the exact experiment, the measured sensitivity to the permeability does not show up in the model results.

0.4E-3

Scale ~ £ ò

-0.1

0.

0.1

0.2

0.3

Both of these runs used a value of 3.0 for the tortuosity. Decreasing the tortuosity in the second run to 2.0 increased the fraction reacted to 0.112. While this is appreciable, it is not sufficient to account for the difference between the measurements and calculations.

0.4

Figure 3b: Reaction rate distribution g-mole/(m -s); same conditions as Fig 3a.

-0.1

0.

0J

02

0.3

0.260

These calculated results are quite similar to those described for the real anode model described above:. Because of the increase in volume due to the reaction, there is generally a convective flow outward from the reaction zone. In order for fresh C0 2 to enter, it must diffuse against this flow. The convective flow itself is relatively insensitive to the permeability; if the permeability decreases, the pressure gradient can increase to provide about the same flow rate, which is ultimately determined by the diffusion rate. Of course, there are secondary effects, which account of for the calculated minor changes due to the permeability.

0.238

Summary and Conclusions The objective was to build a numerical version of a model for subsurface carboxy reaction in anodes, so that it would be possible to assess the impact of changes in intrinsic reactivity and permeability of the net rate of reaction and the maximum porosity generated. The results are that the model accurately describes the reaction localization on the submerged sides of the anode;

0.4

Figure 3c: Void fraction distribution for conditions as Fig.3a.

905

μοο = ( 49.725+0.47868*T-0.00014113*T2)*l.E-7

however, there is almost no sensitivity to the permeability. This is at variance with published experimental results, especially those by Sadler and Algie. The reason for the insensitivity is that because of the reaction stoichiometry, there will always be a net flow of gas away from the reaction so that new reactant gas must be supplied by diffusion. The net rate of gas outflow is determined by this upwind diffusion, and the role of the permeability is only to determine the pressure gradient necessary to drive the convective flow. The important parameters that control the mass transfer are those that control the diffusion, namely the porosity and tortuosity. To some extent these are correlated with the permeability, but generally, the functional dependence of the permeability on these parameters are not linear, so that focus on the true parameters rather than the surrogate, permeability, is likely to be a better practice for anode property optimization.

μπûχ=0.5*(μ<:ο2+μα))

Dab = 3.6009*1.E-9*T3/2

References 1. B. A. Sadler and S. H. Algie, "Sub-surface carboxy reactivity testing of anode carbon," Light Metals 1992, ed. Euel R. Cutshall, TMS, Warrendale, PA, 1991, p 649. 2. E. R. Cutshall, "Influence of anode baking temperature and current density upon carbon sloughing," Light Metals 1986, ed. R. E. Miller, AIME, Warrendale, PA, p 629. 3. R. Farr-Wharton, B. J. Welch, R.C. Hannah, R. Dorin and H. J. Gardner, "Chemical and electrochemical oxidation of heterogeneous carbon anodes," Electrochim. Acta, 25(1980)217. 4. E.g.: http://gasifiers.bioenergvlists.org/reedboudouard 5. Pao-Chen Wu, "The kinetics of the reaction of carbon with carbon dioxide," D.S. Thesis, MIT, Dec. 20, 1949. 6. Shabi Ulzama, "A theoretical analysis of single coal particle behaviour during spontaneous devolatilization and combustion," Dissertation Dr. Ing., Otto-von-GuerickeUniversitat, Madgeburg, April 2, 2007. 7. Kai Grjotheim and Halvor Kvande (eds.), Introduction to Aluminum Electrolysis, 2nd edition, Aluminium-Verlag, Dusseldorf, 1993, Table 4.10, pl20. 8. Euel R. Cutshall and Vaughn L. Bullough, "Influence of baking temperature and anode effects upon carbon sloughing," Light Metals 1985. ed. H. O. Bohner, TMS, Warrendale, PA, p 1039. 9. Tony Ross, Ken Krupinski and Marilou McClung, "Plant statistical evaluation of the effect of aggregate composition and pitch level on anode distortion and predicted performance," Light Metals 1998, ed. Barry Welch, TMS, Warrendale, PA, 1998, p 637. 10. B.A. Sadler and S.H. Algie, "Macrostructural assessment of sub-surface carboxy attack in anodes, ' ' Light Metals 1989, ed. Paul G. Campbell, TMS, Warrendale, PA, 1988, p 531. 11. Hideki Minagawa, Yasuhide Sakamoto, Takeshi Komai and Hideo Narita, "Relation between pore-size distribution and the permeability of sediment," Proc 19th (2009) Intl. Offshore and Polar Eng Conf., Osaka, June 2009, p 25. 12. P. L. Walker, Jr. and F. Rusinko, Jr., "Gasification of carbon rods with carbon dioxide," J. Phvs. Chem., 59(1955)241. 13. P. L. Walker, Jr., F. Rusinko, Jr., and E. Raats , "Changes of macropore distributions in carbon rods upon gasification with carbon dioxide," J. Phvs. Chem.. 59(1955)245. 14. G. H. Geiger and D. R. Poirier, Transport Phenomena in Metallurgy, Addison-Wesley, Reading, Massachusetts, 1973, Figure 13.2, p 469.

What of the experimental results? We might observe that the samples were fabricated by mixing various size fractions made from the same parent material; this introduces an interesting correlation into the results. The fraction reacted correlates strongly with the permeability, but it correlates just as strongly with the average particle size in the sample. Furthermore, any property that correlates with the particle size would have produced a similar correlation with the extent of reaction. The correlation with the permeability may be happenstance, with the causal variable being something else altogether. The models described here are very simple, based on homogeneous microstructures. Anodes are more complex, and more accurate modeling will probably depend on taking that into account. Models that include the effects of multiple microstructural length scales exist, such as those describing reactions in beds of catalyst particles. Adaptation of such a model to anodes may be a useful next step. List of Symbols D F j K k MWx P

% S u v

burn

μ

Pg [X] X

ε

Preal pApp

τ

m2/s C/eq A/m2 m2 m/s g/mole Pa moles/(s m3)

diffusivity Faraday constant current density permeability reaction rate constant molecular weight of species x pressure chemical reaction rate

m2/m3 m/s m/s kg/(m s) kg/m3 g-moles/m3

surface area concentration gas superficial velocity anode consumption velocity viscosity gas density gas phase concentration of species

kg/m3 kg/m3

-

(T in K)

void fraction real density of the solid apparent density of the solid tortuosity

Viscosity and Diffusivity Relationships μ€Ο2=(-10.382+0.56503*Τ-0.0(Ì6364*Τ2)*1.Å-7

906

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PASTE QUALITY IMPROVEMENTS AT ALCOA POΗOS DE CALDAS PLANT Beatriz Vry, Ciro Kato, Jeronimo Araujo, Fabiano Josι Ribeiro, Andrι Luis Abreu Alcoa; Rodovia Poηos de Caldas Andradas, km 10; Poηos de Caldas, Minas Gιrais, 37719-900, Brazil Keywords: Soderberg, Fines, Mixing. Oversize particles go to the Hammer Mill and return to the screens while coarse and intermediate particles go to the designated storage tanks. The coarse tank overflow goes to the Hammer Mill and consequently to the Vibrating Screens again. The intermediate tank overflow goes into the tank used to feed the Ball Mill (Figure 2).

Abstract Alcoa Poηos de Caldas Soderberg Carbon Plant began its operation in 1965 and after a series of upgrades including a mixer heating system upgrade in 1989, fines control in 2000, mixing temperature optimization in 2004 and installation of coke preheaters in 2009, the plant now has twice the original production capacity with near world-class paste quality. The benchmark paste quality characteristic is 1.55 g/cc of baked apparent density (BAD), while Poηos results are about 1.45 g/cc. The Paste Plant is now challenged to meet Potroom requirements for anode performance (higher loads and dry anode top technology). This paper describes the enablers chosen to improve the baked properties, which includes the optimization of fines production, mixing process, and recipe. The quality management system which includes a carbon laboratory with Soderberg baking furnace, sample preparation, and baked analysis (BAD, Air Permeability, C0 2 Reactivity and Electrical Resistivity) are also described.

The Ball Mill is the equipment responsible for producing the fines fraction, where, according to the feed rate and air classifier adjustments, it is possible to produce a different fineness. In Poηos, the feeding rate is adjusted according to the power necessary to rotate the Ball Mill. As the mill empties, the force required for rotation increases, consequently increasing the feed rate. When the mill is overloaded, the force required drops, and the feed rate either decreases or is manually interrupted when there is evidence clogging. Having consistent Ball Mill operation is fundamental in order to produce a fines fraction that has consistent sizing. Dust collected in the plant is added in the Ball Mill outlet at an established rate according to the dust tank level (measured twice per shift). Fines particles are the most important coke fraction because [2]: i. It must not be in excess which will require additional pitch to coat the increased surface area; ii. It fills the spaces between coarser coke particles to produce a maximum dry bulk density; iii. It fills the open pores, resulting in a denser anode.

Introduction Different from a Prebaked Plant, the Poηos Smelter operates with Soderberg technology which requires additional control from the potroom area to form and bake the anode using the Electrolytic pot. Anode top control is needed with plasticity and leveling as a requirement for stubbing operation. In the Paste plant a control of paste flowability is needed with continuous adjustments on aggregate pitch demand. The paste production process can be divided into three different areas: dry aggregate, mixing, and extrusion processes. Dry Aggregate: Incoming coke consists mostly of coarser particles, with which it is not possible to produce a dense anode due to the high level of inter-particle porosity. Classifying, crushing and milling the incoming coke will result in different particle sizes capable of forming a denser aggregate and consequently a denser anode [1]. This step consists of separating the material into three fractions: coarse, intermediate, and fine in order to produce the densest packing (Figure 1). The number of fractions depends on the plant facilities (number of sieve decks and tanks) and the particle size for each fraction is obtained through a ternary diagram that takes into account, among other things, raw material quality and the particle size.

Figure 1. Dry Aggregate Flow

Upon receiving and characterizing, the coke is unloaded and sent to the Vibrating Screens where it is separated into three products: oversize (over 0.5 mesh), coarse (between 4 and 0.5 mesh) and intermediate (under 4 mesh).

907

INTERMEDIATE

*aί*sίmsί

Figure 3. Mixing and Extrusion Process Flow Extrusion: Figure 2. Ball Mill Flow Mixing Process:

After mixing, the paste is transferred to the extrusion box and extruded. Cooling the briquettes is necessary after extrusion to ensure they do not deform or stick together.

The Mixing process can be divided into three steps: weighing, preheating, and mixing.

While the paste is not completely cooled, binder (pitch plus fines) is still soaking into coke pores.

Each fraction is fed individually to the Preheater where it is weighed (22% coarse, 35% fines and 43% intermediate respectively).

Pasting on Anode Top: The briquettes are loaded every 48 hours in each of the 288 pots using a special vehicle, respecting a time frame between the stubbing and tapping operations.

Since 2009, Poηos has used a preheating facility where, after weighing the batch, the coke is heated through electrical induction.

After a daily top condition inspection, three different briquettes can be added with the aim of keeping anode top plasticity under control: wetter (around + 1.0% pitch), dryer (around -1.0% pitch) or typical.

After the preheating process, the aggregate is transferred to the mixer. Poηos has three Preheaters and consequently three Batch Mixers working in parallel (Figure 3). The mixers are heated with thermal oil (kept at around 200°C), which guarantees that the coke will not lose temperature while the dry aggregate is mixed [3]. After 6 minutes, the total programmed pitch is added. The paste continues to be mixed until the final temperature reaches 188°C and with a minimum of 45 minutes of total time mixing. Each batch is equivalent to 3.6 tons of paste.

Solution Approach The strategy of working with higher pitch content in the paste attends to the anode top requirements and gives good operational conditions for anode formation, but impacts negatively the anode quality. An anode with a higher pitch percentage shows higher porosity, permeability and reactivity, causing higher dust generation and anode consumption in the potrooms. Based on this restriction and consequences to the potroom, training seminars were held involving the entire team in search of practical solutions.

Pitch wets the surfaces of all particles, fills pores in and between the particles, and is added in liquid form (at around 185°C). Optimized pitch content is necessary in order to avoid cracking and poor strength, with a minimum of air permeability.

In order to pursue paste quality improvements, a step by step action plan was developed in Paste Plant in Poηos de Caldas: i. Fines production optimization; ii. Mixing process optimization; iii. Recipe optimization. This approach was established based on the following theory: i. Having fines enough to fill open pores; ii. Use of a higher mixing energy, to force binder (fines plus pitch) into coke pores; iii. Formulation of a recipe containing a minimum quantity of coarser material (more porous material) and still

908

keeping the strength characteristic necessary for the anode operation. Experimental approach

Table I. Test 1 set ups. Pitch Fines Intermediate Content Fraction Fration (%)

Test

(%)

Fines Production Optimization;

c

This strategy refers the best fit fines fraction; in other words, it works with a fines target with the highest Blaine* number sustainable, using the smallest fines fraction generating a denser baked paste [4]. * Blaine equipment is an indirect measure for fineness offines,by measuring dry particle air permeability. It is a technique widely used in the cement industry.

(%)

27.7 27.7 27.7 27.6 27.6

Typical A B D

35 35 32 35 32

Blaine Number 4000 4500 4500 4500 4500

43 43 46 43 46

The baked analysis results from each test are shown in Table II. Table II. Test 1 results. Baked Test

The main objective of this test was find the best configuration for pitch content and fines fraction running with 4500 Blaine, ensuring paste quality improvements. The tests were arranged as follows:

Apparent

Air

Eletrical

Permeability Resistivity

C02

C02

C02

Reactivity -

Reactivity -

Reactivity Loss {%)

Density {g/cc)

(nPm)

(nOm)

Residue (%)

Dust (%)

A

1.476

2.166

69.217

88.527

2.437

9.037

B

1.484

2.063

79.660

73.105

9.005

17.715

C

1.489

2.143

88.560

80.675

6.363

12.955

D

1.486

3.599

70.905

84.405

4.055

11.540

Test 2 (December/2009):

Test 1: Evaluate impact of each change (increasing fineness, decreasing fines fraction, decreasing pitch content). Test 2: Assess Paste Plant productivity capacity and stability; Test 3: Practical application (potroom results).

During 2 days, the Paste Plant was run with the set up developed in Test 1 D. A total of 90 batches were produced with no operational instability.

Test 1 (November/2009): i. After ensuring Blaine machine calibration, four recipes were produced and sampled following the typical Paste Plant set up. Fines Fineness regular target = 4000 Blaine. ii. Keeping the same pitch and coke supplier, 25 recipes were produced while the fines fineness was adjusted through increasing the air classifier speed from 22% to 29%. The Ball Mill product was sampled and analyzed each cycle and followed (4500 ± 200 of Blaine number). While the results on the weigh belt feeder had not reached 4500 (sampled each 2 hours), all the produced batches were considered as transitional material and were not analyzed. iii. Following the configuration below, 25 recipes have been produced in order to evaluate the gains from each change, see the explanation below: • Test A: keeping the old recipe (22% coarse /35% Fines /43% Intermediate), increasing fines fineness (from 4000 to 4500) and keeping the original pitch content (27.7%); • Test B: adjusting the recipe (22% coarse /32% Fines /46% Intermediate), increasingfinesfineness and keeping the pitch content; • Test D: adjusting the recipe and pitch content in order to reach flowability target (131%) and increasingfinesfineness; • Test C: keeping the pitch found on test D and fines fineness, but using the old recipe.

Test 3 (April/2010): After a flowability test to ensure top plasticity and tofindthe pitch target, 24 pots received briquettes produced with 4500 Blaine for 3 months. These briquettes were produced at the beginning of the month during 2 or 3 days, and as a safety factor, 30% of total production scheduled was produced with plus or minus 1% pitch for top correction. Comparison of a typical month's data with test data is presented in Table III. Table III. Test 3 results. Month Typical Paste April-10 Test Blaine 4500 April-10 Typical Paste April-10 Test Blaine 4500 April-10 Typical Paste July-10 Test Blaine 4500 July-10 Average Test Blaine 4500 Average Typical Paste Difference

Baked Apparent Density (g/cc) 1.448 1.453 1.454 1.456 1.460 1.463 1.457 1.454 0.003

Air Permeability (nPm) 1.738 4.645 1.671 5.785 3.795 5.215 2.401 2.814

Electrical Resitivity (HQm) 86.11 77.39 79.49 82.40 61.10 65.83 75.203 75.568 -0.364

C02 Reactivity Residue (%) 72.94 75.28 75.94 77.68 92.85 94.24 82.397 80.577 1.820

C02 Reactivity Dust (%) 16.790 9.980 7.640 9.580 0.878 0.190 6.583 8.436 -1.853

Mixing Process Optimization: Using the same mixer and aiming to increase the mixing energy, the strategy adopted was to simulate what happens during the prebaked paste mixing process. By adding part of the total pitch scheduled, the mixer blades found it more difficult to promote the aggregate mixing, consequently mixing energy was raised (measured by monitoring the mixer amperage) and forcing the fines penetration into the open coke pores.

The test set up are further summarized in Table I.

Pitch addition is made in one step, 6 minutes after the beginning of the mixing process. In the new strategy, this addition was made in two phases, where the first occurs at 6 minutes, with 70% of the total mass, and the second phase after about 30 minutes of mixing, adding the remaining 30%.

909

Therewith, the first phase will work to promote the binder penetration (pitch plus fines) and the second phase will guarantee theflowabilitycharacteristic necessary for the stubbing operation. This new way to manage this process has been named as the New Mixture Strategy (NMS).

Results and Discussion Fines production optimization Table II demonstrates the test data resulted from each change (increasing fines fineness, adjusting recipe and adjusting pitch content). Tests A and B were not applicable due to dry top plasticity. While Test C showed slightly higher baked apparent density, Test D showed better C0 2 Reactivity Residue and Dust, while keeping top plasticity under control. Due to that Test D was taken as the standard for the subsequent tests.

The test was arranged as follows: Test 1: Evaluate the operational strategy; Test 2: Develop the NMS control logic; Test 3: Test the NMS control logic; Test 4: Evaluate paste quality gains with the new logic.

Compared to typical paste results (Table III), there was a potential to increase C0 2 Reactivity Residue by 1.82% and BAD by 0.003 g/cc. During almost all tests, there was no necessity to add briquettes to correct the top plasticity (> 95% under control), showing higher top plasticity stability.

Test 1 (February and March/2010): Based on Paste Plant production flow, the operability of this strategy was tested manually by the operating team. As Poηos has only one pitch scale, there was an observed delay on the total mixing time when at the same time different mixers require pitch addition.

Mixing process optimization: Table IV shows the mixing strategy test results compared to typical paste results. There was a potential to reduce pitch by 0.60% while the baked apparent density increased by 0.008 g/cc.

For process safety, there were some conditional circumstances to transfer the paste for extrusion when the team was developing the new mixture control strategy program: i. Both phases completed; ii. Minimum of 15 minutes mixing in phase 2;

Connecting both strategies During an excursion of low coke quality (from 0.940 g/cc to 0.910 g/cc of Vibrated Bulk Density), the need to increase pitch content in the typical paste arose in order to recover top plasticity (more than 20% of pots with dry top condition). Instead, the Process Team has opted to use the potential savings of the NMS and Fines Fineness to improve top condition.

Tests 2,3 and 4 (from April to August/2010): Four tests have been completed with one, two, and three mixers respectively to evaluate the new program and the operability of this approach.

Introducing the NMS gave an opportunity to reduce pitch by 0.6% and the increasing Fines Fineness provided top plasticity stability, thereby improving paste quality.

After testing the program, there was a flowability test to find the optimum flowability target. Three different briquettes were produced with 130%, 143% and 150% flowability targets respectively. A total of 6 batches were produced for each recipe and added into three different groups of pots (4 pots per group) within a week.

The first step, in September 2010, was to change the mixing strategy, and the second step in October 2010was to increase the Blaine number from 4000 to 4500. Even with these strategies, in order to recover top plasticity, pitch content had to be slightly increased from 29.0% to 29.4%.

During this test, the typical paste had 29.0% pitch and a flowability of 143%, while the briquettes produced with the NMS optimum condition had 28.4% pitch and 143% flowability.

A further increase in mixing energy and maintenance of constant paste temperature was required to best mix the dry aggregate with binder. The preheating time was increased from 17 minutes to 19 minutes and the thermal oil reduced from 295°C to 255°C. The final results are provided in Figures 4, 5, 6 and 7.

The test results are compared with a typical month paste results in Table IV: Table IV. Test results. Month

Baked Apparent Density (g/cc)

Typical Paste March-10 Test NMS March-10 Typical Paste April-10 Test NMS April-10 Typical Paste June-10 Test NMS June-10 Typical Paste August-10 Test NMS August-10 Average Test NMS Average Typical Paste Difference

1.443 1.438 1.454 1.459 1.454 1.484 1.450 1.455 1.459 1.450 0.008

Air Permeability (nPm) 1.314 3.037 1.671 1.718 3.863 4.809 5.286 6.805 4.092 3.033 1.059

Electrical Resitivity (Hum) 84.36 86.83 79.49 85.46 67.76 60.71 66.02 71.40 76.10 74.41 1.693

C02 Reactivity Residue (%) 85.76 89.26 75.94 67.19 91.95 92.61 91.66 91.41 85.12 86.33 -1.211

C02 Reactivity Dust (%) 5.453 3.050 7.640 15.160 1.035 0.660 1.425 1.205 5.019 3.888 1.131

Recipe optimization In both strategies presented, the recipe configuration and adaptations were essential to make the quality improvements possible.

910

Figure 7. Mixers Amperage (A) after changes. November 2010. (red line: mixer 5, yellow line: mixer 6 and blue line: mixer 7). By increasing coke temperature, reducing thermal oil temperature and expanding the NMS, the mixer amperage increased approximately 3.67 amperes, with no change on paste flowability (Figures 6 and 7).

Figure 4. Temperature (°C) during mixing process before changes. September 2010. (red line: mixer 5, green line: mixer 6 and white line: mixer 7).

According to previous experiences in the plant, reducing 0.040 g/cc on coke vibrated bulk density it was expected to decrease baked apparent density by 0.030 g/cc.

f PASTA - f ß MPSRÂTÜPA MäS

Looking for paste quality evolution on Figure 8, paste quality, as measured by BAD: • Decreased at first by 0.038 g/cc due to the low coke quality (from 1.445 to 1.407 g/cc throughout September); • Increased by 0.015 g/cc (from 1.407 to 1.422 g/cc), after the NMS and fines fineness strategy implementation; • Increased by 0.022 g/cc (from 1.422 to 1.445 g/cc), after reducing the thermal oil temperature and increasing the coke temperature. Baked Apparent Density

Figure 5. Temperature (°C) during mixing process after changes. November 2010. (red line: mixer 5, green line: mixer 6 and white line: mixer 7).

Aparent Density (g/cc)

~*~VBD

1460

i

Z ^ 1,*48

|

1450 ΐ 1

As shown in Figures 4 and 5, comparing temperature standard deviation there was a significant reduction after the change were implemented.

1,440 - j - l

3s |

l,4bï>^ 3 J

M54

11

• 0,950

,

1 * *** 1

|

MSO 0,940 ■ 0,930 "y

! 1

1.42S

1,423

1

1,420 4 - 1

• 0,920

g

- 0,910 ■ 0,900

1,390 4 - 1

S

29,0%

S

3,0% 1

29,0%

8

i

S

§

s

Setembro 2(

1

S

Julho

1

§

29,0%

'S

3

29,0%

Î

Ì

29,0%

s

Maio

«

i

Abril

§

28,2%

Ü

Abril

1380

H è E

1

S

S

S S

s

i

*- ι

S a" s s o

S

_ 1

o

Figure 8. Baked Apparent Density compared with Vibrated Bulk Density (VBD) evolution in 2010. Other baked paste analysis evolution during 2010, can be seen in the following Figures: Figure 6. Mixers Amperage (A) before changes. August 2010. (red line: mixer 5, yellow line: mixer 6 and blue line: mixer 7).

911

Higher mixing energy (mixer amperage) provides a more homogeneous mixture, due to the filling of coke open porosity and better mixing of the dry aggregate. For Batch technology it is even more important to combine both approaches to be able to capture the gains and not to dry the paste.

Air Permeability ^

Air Permeability

~*~VBD

ο

90

"i

0,950

g 0,940

10,0 0,930 _ 1 8,0

1

hrs

|

11 1 m 1111111 ».74 «z m m m M.....M.. Setembro 2010 29,0%1

1

Julho 2010 29,0%

I

Abril 2010 29,0%

1

Maio 2010 29,0%

fi

Agosto 2010 29,0%

Ella

1,03

Abril 2010 28,2%

2,0 0,0

3,79

6.25

9 ' i | I

3

!

1

Conclusion

3 ! • 0,920 |

m Outubro 2010 29,4%

3,86 2,39

Junho 2010 29,0%

4,0 -

\

á9ΐο 0,900 0,890

Outubro 2010 29,4% Thermal Oil

f 6,0 -

Figure 9. Air Permeability compared with Vibrated Bulk Density (VBD) evolution in 2010.

Baked results have demonstrated that combining the New Mixture Strategy with Fines Production Optimization was an excellent strategy to produce higher paste quality (mainly higher BAD and C 0 2 Reactivity Residue) while reducing pitch consumption. These strategies are a good approach to minimize the negative impacts caused by lower coke quality on baked properties and anode condition. Future data will prove whether it was possible to fully compensate for coke quality gaps using preheater and thermal oil changes, or whether some additional adjustments are necessary. Acknowledgments

Electrical Resistivity ■■Electrical Resistivity «&~VBD

90,0 84.36

85,0

P

81

PCΜ9^*X.

1^

rη

1 !

70,0 -

66,02

\i

6

6630

65,85

• 0,940

64,9*

1

Paste Plant Staff at Poηos de Caldas were involved in making this project a success and their commitment was essential to the project completion. Special thanks to Chin Woo and Ciro Kato for contributing with their ideas and knowledge.

0,930 . |

References

• 0,920 *

i

i

!

!

|

[

1

!

Hü

jgM Setembro 2010 29,0%1

5

Agosto 2010 29,0%

40,0

1)

Kirstine L. Hülse, Anode Manufacture (Sierre, CH: R&D Carbon Ltd, 2000), 77 - 156.

2)

Woo, Chin, "Fines Control" (Paper presented at Green Mill Training in Deschambault, Quιbec, 14 September 2009), 26.

3)

Kirstine L. Hülse, Anode Manufacture (Sierre, CH: R&D Carbon Ltd, 2000), 207 - 236.

4)

Woo, Chin, "Fines Control Initiative" (Paper presented at the 1st Paste Plant Workshop, Poηos de Caldas, Minas Gerais, 28-29 January 2010), 16.

• 0,900

|H

45,0

■ á9Àï

jj||

1 !j

1 ì

M

55,0

Julho 2010 29,0%

60,0

50,0

::

0,950

mm 1

I 65,0

6637

Junho 2010 29,0%

75,0 ·

Maio 2010 29,0%

80,0 -

■ 0,960

"

*L

1 1 1

1

S

S 1E

s I !-

• 0,890

2 ™

1

o

1 î1 o

Figure 10. Electrical Resistivity compared with Vibrated Bulk Density (VBD) evolution in 2010.

Figure 11. C 0 2 Reactivity Residue compared with Vibrated Bulk Density (VBD) evolution in 2010. Due to the more porous coke (lower VBD) the most impacted parameter since September was Air Permeability (Figure 9), even so, initial results have already shown that the adopted strategy is a good way to minimize coke impacts.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PREBAKED ANODE FROM COAL EXTRACT (2) - EFFECTS OF THE PROPERTIES OF HYPERCOAL-COKE ON THE PREFORMANCE OF PREBAKED ANODES Maki Hamaguchi, noriyuki Okuyama, takahiro Shishido, ^oji, Sakai, nobuyuki Komatsu, 2Toshinori Inoue, 3Jiro Koide, and 3 Keisuke Kano ^obe Steel, Ltd., Takasago, Hyogo 676-8670, Japan. 2 Kobelco Research Institute, Inc., Nishi-ku, Kobe 651-2271, Japan. 3 Sumitomo Corporation, Harumi, Chuo-ku, Tokyo 104-8610, Japan Keywords: Coal, Solvent Extraction, Coke, Anode. Kobe Steel, Ltd., have been developing a new non-hydrogenative solvent extraction process for coal with the aim of applying the coal-extract as a binder for the metallurgical coke production. The process and the product (the ash-free coal-extract), is Hypercoal (HPC) [1-4]. A flow diagram of the Hypercoal process and a photo of the Hypercoal pilot plant are shown in Figures 1 and 2, respectively. Since the content of ash in HPC is usually extremely low, HPC is considered to be a promising candidate for alternative feedstock of anode cokes. In a previous paper, we reported the first attempt to obtain a prebaked anode test specimen using HPC as a feedstock of coke [5]. It was demonstrated that prebaked anodes prepared from HPC coke have various advantages such as extremely low impurities such as sulfur, vanadium and nickel, high apparent density, and low air- and CCVreactivity compared to those from anode grade calcined petroleum cokes.

Abstract The preparation of prebaked anodes utilizing coal solvent extraction technology is reported. A steaming coal was extracted with methylnaphthalene-based solvent under pressurized nitrogen atmosphere at 653K, and ash-free coal extract (Hypercoal, HPC) was obtained. The HPC was further heat-treated at 633 to 673 K in the presence of the methylnaphthalene solvent to alter the carbonizing properties of HPC. It was found that adjusting the H/C atomic ratio of HPC in the range 0.6-0.65 suppressed excess dilatation ability of the HPC and produced anode-quality coke. The effect of heat treatment temperature of the HPC on the resultant properties of the HPC coke (HPCC) was also investigated. It was confirmed that HPC is a suitable source of anode coke due to the low content of impurities such as sulfur, nickel, vanadium, and good chemical stability.

In this study, we have investigated the post treatment of HPC in detail, since the as-prepared HPC does not necessarily exhibit suitable carbonizability for anode coke. The coke produced from the as-prepared HPC is generally too porous for anode manufacture. Thus, it is required to improve the dilatation characteristics of HPC to utilize it as anode coke feedstock.

Introduction In the prebaked anode industry, it is recognized that the quality of anode coke has continuously declined, namely, higher sulfur and impurities, higher volatile matter content, and lower density, mostly due to the deterioration of crude oil quality. Thus, technologies for alternative sources for anode coke are of great interest.

i

Extractor

·

r V · 1 \ì

g

Spray dryer

<^HPC^)

Spray dryer

Figure 1. Schematic flow diagram of Hypercoal process

Figure 2. A photo of the Hypercoal pilot plant

913

amounts (13-17wt% in the paste), and hot-pressed at a pressure of 400 bar at 393K. The pressed green molds were baked at 1373K for 20 hours. HPCC and prebaked anodes were evaluated in terms of the ISO standard test methods for anode cokes and prebaked anodes. A schematic of the prebaked anode preparation from HPCC is shown in Figure 4.

Experimental Materials Three bituminous coals from Australia were used as the starting material. Proximate and ultimate analyses of the coals and their extract (Hypercoal) are shown in Table I. Solvent extraction of the coal was carried out under the following conditions: solvent, a mixture of 1-methyl and 2-methyl-naphthalenes; solvent/ coal ratio (v/v), 4; temperature, 673K; initial nitrogen pressure, 2MPa; duration, 20min. The coal extract (HPC) was recovered from the extraction slurry by filtering at high temperature and evaporating the solvent from the extract solution. Proximate and ultimate analyses of the extract (Hypercoal) are also shown in Table I. It should be noted that the ash content of the HPC was less than 1 wt%.

Table II. Properties of the binder pitch used Properties Method Value Unit 1 Softening Point MettleijlSO 5940 113 °C Insoluble in Quinoline ISO 6731 8.4 % 27.7 % Insoluble in Toluene ISO 6376 Viscosity 433K ISO 8003 1555 m Pas Density in Water 1.312 kg/dm3 ISO 6999 Ash content ISO 8006 0.25 % ISO 12980 0.47 % S

Na Ca Si Fe Zn IPb

For fabrication of anode specimen, a commercially available coal tar pitch was used as the binder. The properties of the pitch are listed in Table II. Post treatment of Hypercoal

II II II II It

Quartz testing tube v i (16mmID) ^ 1

Since the as-prepared HPC possesses extremely high dilatation during carbonization, the HPC was subjected to post heat treatment. Detailed investigation of the post heat treatment for the HPC from Coal A and B was carried out in this study. A 100mlautoclave was used for the experiment. Twenty grams of HPC and 20 ml of methylnaphthalene solvent were charged into the autoclave, and the content was heat-treated under the prescribed reaction conditions. The heat-treated HPC was recovered by distillating the solvent in vacuum. It was then subjected to elemental analyses and the carbonization test.

202 ppm 59 ppm 106 ppm 277 ppm 311 ppm 232 ppm

II

Compressive strength test (Indirect tensile test) (after carbonization)

Table I. Proximate and ultimate analyses for the coal and HPC Product Coal-A Coal-A HPC Coal-B Coal-B HPC Coal-C Coal-C HPC

Moisture [wt%] 2.9 0.0 2.2 0.3 1.8 0.1

Ash

VM

[wt%db] 12.2 0.1 12.4 0.1 7.24 0.06

41.3 41.5 39.7 41.2 40.9 43.0

H [wt% daf] 82.5 5.5 87.8 5.5 84.1 5.7 5.4 86.6 83.6 5.8 85.7 5.5 C

N

S

Oa)

2.0 2.2 1.9 1.9 1.9 1.9

0.6 0.6 0.6 0.6 0.7 0.7

9.5 3.8 7.7 5.5 8.1 6.2

Figure 3. Schematic of carbonization and compressive strength test for HPC. Solvent extraction

Post treatment

|Rawcoal|E=C>|HPC|CZZ^>

a) By difference

Coking

I

Molding & Baking

>lHPCC| I

> Prebaked anode

fBinder pitoh~|

Coking of HPC and preparation ofprebaked anodes

Figure 4. A schematic of prebaked anode preparation from coal.

A laboratory scale carbonization test of the post treated HPC was carried out according to the following procedures. An induction heating carbonization furnace was used. Four grams of Hypercoal powder (-1mm) was charged in a quartz tube of 16mmID, and was heated to 1273K at a rate of 3 K/min under inert atmosphere with 0.05kgf/cm2 load. The sample was held at 1273K for 30 min. The compressive strength test of the HPC coke tablet was measured to assess the quality of the coke. The carbonization test procedure is shown schematically in Figure 3.

Results and Discussion Effects ofpost heat treatment on the properties of HPC The change of H/C atomic ratio in the HPC produced from CoalB during the heat-treatment in methylnaphthalene solvent is shown in Figure 5. It is clear that the H/C ratio of HPC decreases with increasing reaction time. This indicates that a dehydrogenative polymerization reaction of HPC molecules takes

Prebaked anode specimens of rectangular shape of 27mm x 50mm x 120mm were fabricated according to the conventional methods. The crushed HPCC was mixed with the binder pitch of various

914

smaller than 0.6, the compressive strength of coke again decreased. The solid-state carbonization is the likely reason for poor carbonizability of HPC obtained by excess heat treatment. Figure 8 shows the relationship between the compressive strength and the porosity of HPC coke. The maximum strength was achieved when the porosity of the HPC coke was adjusted to around 40%. No significant difference in the H/C ratio, coke strength, and porosity relationship was observed for the two kind of coals examined. Figure 9 is a polarized light microphotograph of cross section of Coal-A HPC coke. It is a characteristic of HPC coke that a medium to fine mosaic texture is the main component.

place under the reaction conditions. At higher temperatures, the H/C ratio decreased much faster. A similar reaction behavior was observed for the HPC produced from Coal-A. The Arrhenius plot of the H/C ratio decrease is shown in Figure 6. The activation energy, Ea of this reaction was calculated to be 172, and 169 kJ/mol for Coal-B and Coal-A HPC, respectively. These values are in good agreement with those reported for polymerization of coal-tar pitch, 150-190kJ/mol [6], and petroleum pitch 200-220kJ/mol [7]. 0.68

14

0.66

12

0.64

OCoahAHPC • CoahB HPC 1

10

1.62

+

430°C

0.60

Ä

450°C

o 8

~ r~

#

L

440°C

:

6 4

460°C 0.58 20

40 Duration, rrin

60

06» 0.50

80

Figure 5. Changes of H/C atomic ratio of Coal-B HPC during the post heat reatment in methylnaphthalene at various temperature.

0.65 0.70 0.75 0.55 0.60 H/C atomic ratto of heat-treated HPC

0.80

Figure 7. The relationship between the compressive strength of HPC coke and the H/C atomic ratio of the post heat-treated HPC.

Figure 8. The relationship between the compressive strength and the porosity of the HPC coke.

Figure 6. An Arrhenius plot of H/C decreasing rate during the post heat treatment of Coal-B HPC. The effect of the post heat treatment on the carbonization properties of HPC was assessed in terms of the appearance and the mechanical properties of HPC coke. The compressive strength of the HPC coke was represented as a function of H/C atomic ratio of the heat-treated HPC (before carbonization) in Figure 7. It should be noted that the H/C ratio of the starting HPC was the highest and that the ratio decreases with increasing severity of the heat treatment. As mentioned earlier, HPC with the higher H/C ratio, i.e. as-prepared HPC and HPC obtained from very mild post heat-treatment resulted in coke with poor compressive strength. This appears to be due to excess dilatation ability of HPC: The HPC has expanded too much during the carbonization. On the contrary, with decreasing H/C, the compressive strength of coke increased, and reached a maximum, where the H/C ratio of the HPC was in the range of 0.60-0.65. For HPC with the H/C ratio

Figure 9. Polarized light microphotographs of polished cross section of cokefromCoal-A HPC

915

Effects of the carbonization temperature on the properties of HPC coke and theprebaked anode The properties of HPC coke calcined at 1273K (1000°C) and 1573K (1300°C) and a typical petroleum coke (PC) calcined at 1473K are compared in Table III. The content of volatile matter is very similar for HPCC and PC, irrespective of the calcination temperature. The density of the HPCC is substantially lower than PC when the HPC is calcined at 1273K, but are similar at 1573K. The crystallinity of HPC coke may be slightly lower than PC as judged by the X-ray diffraction parameters. The optical texture of HPC and the petroleum coke was quantitatively compared by polarized light microscopic measurement, and is listed in Table IV. The most significant difference is the higher content of fine and/or medium mosaic components in HPC coke, whereas the PC contains coarse mosaic and fibrous texture at a higher amount. This suggests that graphitizability of PC is higher than HPC. Table III. Properties of HPC coke and Petroleum coke WK

w%

Density, g/ml

( W nm

Lc, nm

HTT1000 HTT1300 1HTT1000 HTT1300 HTT1000 HTT1300 HTT1000 HTT1300 Coal-A HPC

0.26

0.18

1.855

2.031

0.3487

0.3481

23

Coal-C HPC

0.31

0.06

1.848

1.997

0.3513

0.3509

15

Ref. PC

2.025

0.39

24 15

25

0.3474

PC was calcined at 1473K

Table IV. Optical Texture of HPC coke and Petroleum Coke Content, % Source

Fjne

Isotropie μ

Coal-A HPC Coal-C HPC Ref. PC

mosaic 41 44 0

0 0 0

c..

Medium mosaic

Coarse mosaic

Fibrous

Leaflet

Inert

55

4

0

47

9

0

15

43

24

0 0 19

0 0 0

The properties of the prebaked anode prepared from HPC coke prepared with different calcination temperatures were investigated. Table IV compares the results with that of anode from a conventional petroleum coke. As a general tendency, HPC coke calcined at 1273K has a slightly poorer performance for feedstock for prebaked anode, as compared to PC, i.e. lower apparent density, higher electrical resistivity, and higher coefficient of thermal expansion (CTE). However, when the HPC coke was calcined at 1573K, the properties of the anodefromHPC coke are almost the same as those from petroleum coke, except the higher CTE. Table V. Properties of anode specimen prepared from HPC coke and Petroleum coke Source

Apparent density, g/ml

CTE. 10"6/°C

Resistivity, μÙιτι

HTTIOOO

HTT1300

HTT1000

HTT1300

HTT1000

HTT1300

Coal-A HPC

1.338

1.380

62.0

53.5

5.08

5.64

Coal-C HPC

1.335

1.306

61.1

71.7

5.25

Ref. PC

1.365

5^4

Conditions for anode specimen preparation Pitch content in the mold: 17.8 wt% Molding: 45MPa, 2min; 27mm x 50mm x 120 mm Calcination: 1000°C

Conclusions Results of this study can be summarized as follows:

5.47 4.56

1. A prebaked anode sample for aluminum smelting was successfully prepared from coal by utilizing Hypercoal coke technology. 2. It was demonstrated that the controlling the H/C atomic ratio of HPC to 0.6 -0.65 by post treatment to produce acceptable anodegrade coke, whereas the as-prepared HPC exhibited excess thermal dilatation. 3. The effect of heat treatment temperature of HPC on the properties of the final HPC coke (HPCC) was investigated. It was confirmed that HPC is a quite suitable source of anode coke in terms of coefficient of thermal expansion, electric resistivity, density, and contains a low content of impurities such as sulfur, nickel, vanadium and has good chemical stability. References [1] Okuyama, N., Komatsu, N., Kaneko, T., Shigehisa, T. and Tsuruya, S. (2004) Fuel Processing Technology, 85 947. [2] Okuyama, N., Furuya, A., Komatsu, N. and Shigehisa, T. (2005) International Conference of Coal Science and Technology Proceedings. [3] Komatsu, N., Okuyama, N., Furuya, A. and Shigehisa, T. (2006) International Pittsburgh Coal Conference. [4] Hamaguchi, M., Komatsu, N., Okuyama, N., Furuya, A. and Shigehisa, T. (2007) International Conference of Coal Scinece and Technology Proceedings. [5] Hamaguchi, M., Okuyama, N., Komatsu, N., Koide, J., and Kano, K (2010) Proceedings ofTMS 2010 Annual Meeting. [6] Honda, H., Kimura, H., Sanada, Y., Sugawara, S., and Furuta, T. (1970) Carbon, 8, 181. [7] Greinke, R.A. (1986) Carbon, 24, 677.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

NEW GENERATION OF VERTICAL SHAFT CALCINER TECHNOLOGY Jingli Zhao, Qingcai Zhao, Qingbo Zhao Jinan Aohai Carbon Corporation Ltd., Jinan, Shandong Province, 250400 China Keywords: green coke, CPC, vertical shaft calciner, heat recuperation, power generation Abstract

Most of CPC is produced by either vertical shaft calciners or rotary kilns in China. The rotary calciners are usually applied in the anode manufacturers belonging to the large scale integrated smelters, while the shaft calciners are commonly used by the specialized and independent anode manufacturers and all the graphitized electrode producers in China.

The vertical shaft calciner has been widely applied in China for CPC calcination. Its application, however, is restricted by its lower capacity, less automation and no waste heat recuperation. A new generation of energy saving shaft calciner technology with higher capacity and power generation system has been developed recently and illustrated in this paper. 75 kt of CPC annual production can be achieved by only one new shaft calciner, the waste heat from which can be recuperated to generate electricity of about 28 million kWh. The major invention is calciner structure optimization by computer simulation for better volatile combustion and calciner heat balance, which brings about better CPC quality, higher capacity and provides more energy for power generation. The new shaft calciner technology is flexible to the size distribution and volatile content in the green coke and applicable for the pulverous coke and the coke with high volatile content.

A characteristic comparison between shaft calciners and rotary calciners will be made and the development of the new generation of shaft calciner technology will be outlined in this paper. In addition, its application and great achievements obtained in Chinese anode industry will be introduced. Comparison between the shaft calciner and rotary calciner technology Γ21 Figure 1 and Figure2 show briefly the technology scheme of rotary calcining and vertical shaft calcining technologies respectively.

Introduction China is the largest production country of calcined petroleum coke (CPC) in the world with about 100 CPC producers. The total CPC capacity reached more than 10 million tons in 2009, of which about 80% was used for anode manufacturing in aluminum reduction industry and about one million tons of CPC was exported abroad. Only a small part of high quality CPC was applied for the production of graphitized electrodes and acieration agent. The green coke produced in crude oil refineries that meets quality requirements for size, impurities, and coke structure is calcined before production of anodes for the aluminum industry in order to remove the volatiles and to form the coke grain structure [1].

Figure 1 Rotary kiln calcining technology scheme

The calcination technology has a significant impact on CPC quality, especially such properties as its bulk density, grain size distribution and hardgrove index. To produce high quality CPC it is essential to keep the optimized control of calcination parameters including calcination temperatures, duration and volatile burning

È with Baffles

Cooling Sleeves

The calciner operation cost for CPC is closely related to coke recovery and energy consumption during the process and to requirements for green coke quality, for which the extra coke burning control and waste heat recuperation play a very important role. Both the calcination process control and waste heat recovery mainly depend on the calciner itself and the affiliated facilities.

Longitudinal Sectional View |

Figure.2 Vertical shaft calcining technology scheme

917

Table 1 The technical and economical comparisons of both calcining technologies Calciner types

Rotary calciner

Old vertical shaft calciner

capacity control operation cycle

larger higher automation level shorter, less than 85% operation

middle or smaller more manual operation much longer

heat loss environment

more heat loss better waste gas treatment

less heat loss waste gas emission issue

1 1

working conditions heat recovery

better waste heat boiler

worse heat recovery for oil heating

1 1

coke recovery ash content in CPC

lower, about 75-78% more ash from calciner lining

higher, more than 82% less ash content

1 1

CPC bulk density CPC reaction resistance

lower lower

higher higher

1

strict for green coke size

wider tolerance for GPC size

1 requirements for GPC*

1

GPC *-- green petroleum coke Rotary kiln calciners are widely used for CPC production in the world, while the vertical shaft calciners are used in most Chinese CPC plants. Vertical calciners are stationary calcining furnaces.

set of the relevant design, construction and operation technologies has been developed and optimized.

The technical and economical comparisons between the rotary calcining and the old vertical shaft calcining technologies are illustrated in the Table 1.

Technology development of the new generation vertical shaft calciner There are two aspects for extending the shaft calciner capacity: increasing the unit number and changing the calciner structure. A new structure design concept has been developed and an optimized calciner layout completed by using computer simulation.

There are many advantages for rotary calciners: larger capacity, higher automation level and better working conditions. But in the major aspects on CPC properties and coke recovery, which related more closely to the CPC application and its production cost, the advantages come up for vertical shaft calcining technology.

For the structure design a better heat transfer, insulation and the flue dimensions should be considered carefully and the refractory expansion during furnace baking and long term operation is of great importance.

The adaptability of the calciner to handle dusty coke becomes more and more important since the green coke grain size distribution has deteriorated. The vertical shaft calciner will be more suitable to the calcining process for this kind of GPC.

The high temperature waste gas can be used for a surplus boiler, which supplies steam for power generation. The pyrology simulation is applied for higher heat transfer efficiency.

Major considerations for developing new shaft calciner technology are that the old shaft calciner technology should be modified and improved in the following aspects:

The different technical solutions have been compared for sulfur removal from the waste gas. A technical solution has been selected for the higher efficiency and the best economical results.

(1) Extend its capacity by new design concept and simulation; (2) Set up a modern inspection and control system; (3) Install sets of operational machines instead of original manual operation;

1.

(4) Install a waste heat boiler and power generation system to recover surplus heat from the waste gas;

The green coke feed stock usually comes from various suppliers with quite different chemical and physical properties in large scale CPC plants. Coke blending should be implemented for CPC quality stability and uniformity.

(5) Remove sulphur from the environment technology.

waste gas by the advanced

Green coke blending system [3]

The coke blending facility is installed in the calcination plant. Cokes are blended based on the property analysis of the different sources.

A new generation of vertical shaft calciner with 14 shafts and 56 units has been successfully developed and applied in Jinan Aohai Carbon Corporation, Ltd., China, to replace the original old shaft calciner. The calcine, which has increased capacity and improved energy efficiency based on the above mentioned considerations. A

The Chinese GPC produced in the refineries in different areas in China has quite different content of the elements, such as S, Na, Ca, V and Ni, greatly impacting on the coke reactivity. Some

918

elements will also catalyze coke oxidation, while some elements like sulphur will inhibit the oxidation. Based on the regulation mentioned above a blending technology is developed for improving the coke reactivity. Some high sulphur content coke is blended with low sulphur content and high calcium and vanadium cokes. The blended coke mixture will have better properties than the original cokes. 2.

Green coke feeding system

There are many more calcination units and feeding entrances for the new generation shaft calciners. An automatic green coke distribution facility has been developed and installed above the calciner for more homogeneous feeding and less manual operation. Every unit is fed with an equal amount of green coke during calcination, which keeps a constant operation for all the calciner units.

Figure 4 Display of the online inspection system 5.

Considering the process parameters of the new shaft calciner for heat recovery the technical and engineering scheme for the power system is determined as follows:

Distribution facilities are specially designed and installed over the calciner units with multiple outlets that can be moved forward and backward. A crane is installed to charge the GPC into the distribution facility. 3.

Power generation by the waste heat recovery [4]

A surplus heat boiler is installed at the outlet of the waste gas from the calciner for generation of the superheat vapor with 3.82Mpa and 450°C. The superheat vapor is used for electricity generation in the power generator.

CPC discharge and dust recovery during discharge

The discharge of CPC for the old shaft calciners was labor intensive and generated high levels of dust.. Based on our research, a new CPC discharge system and a dust collection system have been developed. No manual intervention is required for CPC discharge and the dust emission issue is solved. The coke dust collected can be used for dry aggregates in anode production. Figure 3 shows the CPC discharge system of the new shaft calciner.

The exhaust vapor from the generator is condensed and pumped into a deoxidizing facility for oxygen removal, from which the water is recovered and recycled for the surplus heat boiler. The facility configuration of the surplus heat recovery and electricity generation will provide a reliable and flexible operation platform to combine the shaft calcination system and electricity generation system for maximum surplus heat recovery and utilization. 6.

Sulfur removal from the waste gas

A sulfur removal technology from the new calciner waste gas has been applied in the new shaft calciner plant. The key technology is to absorb sulfur oxides from the waste gas by ammonia to form ammonia sulfates as a fertilizer. With this process, the achievable sulfur recovery can be greater than 90%. The waste gases meet all local sulfur emission limits. Technology characteristics of the new generation shaft

Figure 3 CPC discharge system 4.

calciner

Online inspection and display system

This new generation shaft calciner technology has the following characteristics:

An online inspection and display system has been developed in the new generation shaft calciner. The temperature and pressure at the key points in the calciners are monitored. The system also includes an alarm system which highlights any out of range parameters. Figure 4 shows the display in the online parameter inspection system. The online inspection system shows the temperatures in various calciner units. The color of the digits shows the deviation from the target temperature.

919

•

Greater height and length of the units for a longer calcination stage;

•

More and larger exit shutes for more efficient volatile emission;

•

Optimized flue design for larger flue profile squares;

•

Larger refractory expansion slots to effectively prevent calciner deformation;

•

Improved heat insulation design for the upper and side areas

Owing to the stable feeding and operation in the new calciner CPC quality becomes more stable and more homogeneous.

of the calciner to maintain the flue temperatures; •

Utilization of automatic green coke feeding and calcined coke discharging facilities;

•

Fully airtight and automatic conveying of the materials to andfromthe calciner;

•

Application of the integrated temperature inspection and display by computer.

S ·

Applications and achievements of the new generation shaft

82

>;i2 = 10

-

v*VA Ë í ^

4 1

calciner

3

5

7

9

11

13

15

17

19

21

M>nths

The new generation shaft calciner technology has been applied in Jinan Aohai Carbon Corp. Ltd. An installed capacity of 150 kg/hr per unit and 73.6 kt of annual yield for one calciner with 56 units is achieved.

It is predicted that the calciner life will be much prolonged due to the structure design and stable operation.

The operation has been kept stable and highly efficient. A reduction in manpower is required and the working conditions are much improved. There has also been a reduction in the coking and blockages observed in the calciner flues, further reducing manual interventions..

The new shaft calciner technology can be applied for the new calcination systems or to retrofit engineering projects with lower investment, higher capacity, more energy saving, less emission and higher productivity. This will lead to lower calcined coke production costs.

Most waste gas heat is recovered and the power generation goal is achieved. The environment in the plant is much improved and the sulfur emissions are greatly reduced.

The higher CPC quality from the new shaft calciner will benefit anode manufacturing and anode quality improvement.

Figure 6 C0 2 reactivity of CPCfromthe new shaft calciner

Conclusion

The CPC quality has been improved. Compared with the rotary calciners the vertical shaft calciners are able to produce better quality CPC as follows:

1.

0The CPC real density from shaft calciner can reaches to 2.082.10 g/cm3 and is increased by 0.04 g/cm3 as shown in Figure 5.

A new generation of vertical shaft calciner technology has been developed and applied in China instead of the old shaft calciner.

2.

The key technology for the new generation shaft calciner includes a new shaft calciner structure design, new feeding and discharge systems, heat recovery and power generation technology from waste gas and sulfur removal technology from exhausted gas.

3.

The major technical achievements are: much larger scale; great energy saving; mechanized feeding and discharge systems; better environment and working conditions; power generation from the waste gas; much lower sulfur emission.

4.

The CPC quality is much improved by using the new shaft calciner, especially the CPC bulk density, air and C0 2 reactivities etc., which greatly benefit anode quality improvement.

Η 2 . 12 » 2.1

^-2.08 Î>2 . 06 10

S2· 04

°2. 02 2 1.98 1

3

5

7

9

11 13 Months

15

17

19

21

Figure 5 Real density of CPC from the new shaft calciner

References

©The CPC from shaft calciner has lower powder electrical resistance and is reduced by 30-50μ Ω .m. ©The CPCfromshaft calciner has better C0 2 reactivity of 8-12% compared with rotary calciner of about 15% due to its higher bulk density as shown in Figure 6. © About 3% higher recovery rate of coke can be reached.

[1]

Sun Fang, "Delayed petroleum coke and calcining measures". Light Metals (in Chinese) (10) (2005), 60-62

[2]

Fengqin Liu, Chinese Raw Materials for Anode Manufacturing, (Switzerland: R&D Carbon Ltd., 2004)

[3]

Fengqin Liu, Yexiang Liu, Mannweiler U, Perruchoud R, "Effect of coke properties and its blending recipe on performances of carbon anode for aluminium electrolysis", Science & Technology of Mining and Metallurgy, 2006 Volume 13 No.6 : 647-652

[4]

Li Xiuli, "Feasibility study of using waste heat to generate electric power which comes from the gas of

©There is no need of extra fuels for shaft calciner technology. The C0 2 reactivities of CPC can also be much improved by use of green coke blending technology.

920

pot-type calciner",. Applied Energy Technology (in Chinese). (1) (2010), 36-39

921

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION Petroleum Coke VBD SESSION CHAIR

Angιlique Adams Alcoa USA

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Historical and Future Challenges with the Vibrated Bulk Density Test Methods for Determining Porosity of Calcined Petroleum Coke Jignesh Panchal, Mark Wyborney, Jeffrey Rolle A. J. Edmond Company, 1530 W 16th St. Long Beach, California 90813, USA Keywords: Calcined Petroleum Coke, Vibrated Bulk Density, Porosity density measurements or with mercury porosimetry values. In 1988-89, Belitskus and Danka (5) reported that anode density correlated well with coke vibrated bulk density and coke porosity below 5 μπι in diameter. From the previous studies, it clearly appears that, coke porosity is a key factor in optimizing pitch demand and anode density. The easiest and fastest way of measuring coke porosity is Vibrated Bulk Density. A literature search reveals that 1) the precision statement for ASTM D 4292 and other VBD methods are inadequate, and 2) correlation between macro and micro porosity and VBD has not been published.

Abstract Over several decades, calcined petroleum coke producers and the aluminum industry have been using various techniques such as Vibrated Bulk Density (VBD), Mercury Intrusion Porosimetry (MIP), and Mercury Apparent Density to predict the porosity of calcined petroleum coke. Better knowledge of the coke porosity allows a more accurate estimation of coke quality and pitch demand for fabricating anodes of optimum performance. Industry has had limited success in accurately predicting porosity using the existing VBD methods. Direct measurement of porosity by MIP is an alternative way to measure porosity accurately. Currently there is limited correlation between VBD and MIP. Any improvements to the VBD test method should demonstrate improved correlation to the results obtained from MIP. This paper covers the historical and traditional approach to predict calcined coke porosity, correlation study of VBD test method and porosity by MIP, and input to the development of superior test methods to meet industry requirements.

Review Of Porosity And VBD Measurement Techniques At present, VBD (28x48 Tyler Mesh (TM)) or VBD (30x50 US Mesh), VBD (8x14 TM), VBD (9x16 TM), VBD (28x48 TM) Natural, AD (10x20 TM) or Pechiney AD, Mercury Intrusion Porosity (16x20 TM) or (10x20 TM), test methods are used widely to determine coke quality for the purpose of estimating the pitch demand. Aluminum Industries are using one or more of the above methods, or their own proprietary methods, to estimate pitch demand. Calcined coke producers are also using one or more of the above tests methods as specified by their customers.

Introduction Relationship Between Porositv/VBD And Anode Properties

The actual optimum pitch demand observed by the aluminum industry is not universally well predicted by VBD. However, Mercury Intrusion Porosimetry (MIP), which measures porosity directly, works universally to quantify porosity, and therefore can be used to reliably estimate pitch demand. A. J. Edmond Company has performed multiple design experiments and has established relationships of MIP to optimally pitched anodes. Most of the data is governed by confidentiality agreement, and thus can not be published. For many years, VBD has adequately predicted pitch demand between a single buyer and seller. Each VBD method uses a different particle size, and a different preparation method, but they all predict porosity and packing density.

The methodologies for qualitative analysis of calcined petroleum coke and the effects of coke quality on the properties of anodes are well known and have been well researched over several decades. The porosity of calcined coke is one of the important properties to determine the quality of coke in terms of pitch demand, and optimum pitch quantity is very important for optimizing anode performance. Collier Carbon introduced an indirect measurement of porosity by measuring Bulk Density of coke. In 1961, Alcan adopted the method, and in 1976 modified it, with a slight change in granulometry to the present version. During the years from 1971 to 1978 several efforts were made to improve the method, to gain worldwide acceptance. In the mean time, industries adopted their own granulometry and bulk density methods. In 1978, ASTM was requested to form a study group and ASTM D4292 Standard Test Method for Determination of Vibrated Bulk Density of Calcined Petroleum Coke was introduced in 1983.

Current Challenges With VBD The ASTM D4292 VBD method has been questioned several times in the past twenty years, and it is still in question. The current methods are not very specific and can be interpreted in several ways. A decade ago, there were few producers of calcined coke and the problem with the determination of VBD and correlating VBD with pitch demand was low. In the last 10 years, the number of calcined coke producers has doubled, which has increased the likelihood of misinterpretation of VBD test method(s) and the results. In the last several years the coke and pitch quality has deteriorated, and producers are not able to provide coke of consistent quality. Aluminum industries are blending different quality cokes together in the manufacturing of anodes, and with the drawbacks of the VBD test methods they are not getting enough insight into the porosity of the coke, to

In the past, many attempts have been made to correlate different test methods to reveal the effectiveness of each test and to try and correlate them with anode pitching demand. During 1971-1980, several publications (1-6) indicated satisfactory correlation between important anode properties and VBD tests. In 1977, S.S. Jones (4) indicated that porosity of a certain pore diameter range (1-10 μπι) is particularly significant to carbon anode consumption efficiency. In 1980, Gehlbach (3) reported that an increase in anode baked apparent density correlated well with an increase in the -20+48 mesh coke vibrated bulk density, but not as well with other coke

925

optimize the anode properties. ASTM has been working to improve the precision of the methods, both inter-laboratory (repeatability) and intra-laboratory (reproducibility).

The correlations were not very promising, but it inspired us to initiate a project to determine Mercury Apparent Density and Porosity by Mercury Intrusion Porosimetry of a single size fraction produced by a single method. We chose a sizefractionof 8x14 TM, and 28x48 TM prepared in accordance ASTM D4292 for the study.

As an independent laboratory, A. J. Edmond Company receives coke from all over the world for quantitative and qualitative analysis. Out of all the ASTM and ISO test methods performed on calcined coke, the VBD method reproducibility is among the worst. The strength of the VBD method is that it is a rapid and easy way to estimate the coke porosity. With the current methodology, the coke may meet the specifications set by the marketing and technical groups when it is produced or sold, but it fails in universally predicting pitch demand. Every once in a while, the calcined coke producers and/or aluminum companies, audit the VBD test method D4292 with the hope of resolving this discrepancy. But there has been limited success, due to the poor precision of the test method and the variability of the results obtained. Also, with the increased use of near anode grade calcined coke (shot, under pyrolyzed, higher impurities) in making anodes, it is very important to improve the precision of the test method.

VBD 8x14 vs Porosity 10x20 @ 55000 psi F

O)

t

0.18

VBD 8x14 vs Porosity 10x20 @ 55000 psi

0.17

P 3 file O

>Φ

o 0.1b

0.

— &'0.14 o o

jo 0.13 ®

*

8

This paper examines correlations between MIP and VBD and provides information to the study group to improve existing and future methods.

0.12

o 0.11

Q_

0.10

0.755

Initial Investigation One focus of our study was to find an alternate test method or a back up method to support VBD. From the historical data, we performed a correlation study of all VBD test method(s) with Mercury Apparent Density and Total Porosity by MIP as shown in Figures 1, 2 and 3. The data used in Figures 1, 2, and 3 were collected over 6-7 months by several technicians. Even though the methods have different particle size and sample preparation, the data showed some correlation, although poor, between the methods. We believe the poor correlations are due to the precision (repeatability) of the VBD test method, different particle size, and different preparation methods. This also reflects that analysis of different particle sizes does not always reveal the actual characteristics of the coke.

R2 = 0.5717! 0.775

0.795 0.815 0.835 VBD8x14(grrVml)

0.855

0.875

Figure 2. Correlation of VBD (8x14) and Porosity (10x20) VBD 28x48 vs Porosity 10x20 @ 55000 psi 0.15 V BD 28x48 vs Poros ity 10x 20 @ 55000 ps i

VBD 8x14 vs AD 10x20 VBD 8x14 vs AD 10x20 R2 = 0.692 0.11 0.840

0.855

0.870 0.885 VBD 28x48 (grrVml)

0.900

0.915

Figure 3. Correlation of VBD (28x48) and Porosity (10x20) Experimental Commercial calcined cokes produced by BP and Rain CII, and Laboratory Pilot Plant Rotary Kiln Calcined Coke produced at A. J. Edmond Company were used for the study. The samples for VBD (8x14 TM and 28x48 TM) were prepared in accordance with ASTM D4292-92 (2007) at A. J. Edmond Company, Long Beach. A split of the prepared material was analyzed for VBD as prescribed in the method and another split was analyzed for Porosity using a Micromeritics Autopore IV Mercury Intrusion Porosimeter.

R2 = 0.7359

1.68 4

0.755

0.775

0.795 0.815 0.835 VBD8x14(grrVml)

0.855

0.875

Figure 1. Correlation of VBD (8x14) and AD (10x20)

926

The sample preparation for the MIP analysis was modified for this study, the common practice accepted by industry for porosity involves deoiling of the sample to remove any dedust oil. But here, the goal was to determine the effectiveness of VBD test method for porosity measurement, so the samples were not deoiled and the same particle size was used to determine porosity. All samples were run in duplicate to ensure the quality of analysis.

R2 of VBD and Fbrosity vs Pore Diameter R-essure (psi) 111

2197

6182

14473 23186 32889 52950

1 0.9

10.8

Initially, 12 samples produced by different commercial calciners using Rotary Kiln technology were studied. The data of VBD and Porosity were correlated against each other. The selected samples had VBD (8x14) values ranging from 0.766 to 0.847 g/cc and total porosity ranging from 19% to 25%.

20000 (psi) R2 = 0.8 Average, 92% of Total Porosity

o0.7 ca | 0.6 Φ

"S 0.5 Q o 0.4

The second experiment was carried out on a second set of 12 samples produced by different commercial calciners using Rotary Kiln technology. Some of the coke samples were blended with non anode grade coke. The selected samples had VBD (28x48) values ranging from 0.833 to 0.980 g/cc and total porosity ranging from 22% to 35%.

I 0.2

ü

0.1

|

Average, 30% of Total Porosity

j

at 15 psi

I

,

i

0.0091 (μτη)

133.1062.529 0.113 0.038 0.016 0.010 0.007 0.005 0.003

Results and Discussion

Pore Diameter (μπι)

The results from the VBD (8x14) analysis were correlated against total porosity obtained by MIP. Figure 4 shows that the VBD data correlates well with the total porosity with a coefficient of determination (R2) > 0.89. This shows that the VBD method is effective enough to estimate the total porosity. Figure 5 shows a plot of the coefficient of determination of VBD and Porosity at different pore diameters and pressures.

— — R2 of VBD and Porosity vs Pore Diameter

Figure 5. Coefficient of determination of VBD 8x14 and Total Porosity vs Pore Diameter Figures 6 and 7 show the pore size distribution of samples with different VBD. The samples having high VBD are showing low porosity and vice versa. The samples with VBD 0.820, 0.833 and 0.846 showed a similar pore size distribution pattern with a moderate difference in porosity of macro pores up to 0.5 μιη pore diameter. The porosity pattern for samples with VBD 0.766 and 0.794 was very different. The sample of VBD 0.794 showed high macro (>0.05 μπι pore diameter) porosity compared to the sample of VBD 0.766, but the latter sample showed high Meso and Micro (<0.05 μιη pore diameter) porosity. The total porosity of sample with VBD 0.766 was higher than sample with VBD 0.794.

The data indicates that the correlation is better when porosity obtained at high pressure is included. At 20,000 psi 92% of total porosity is filled by Mercury and the coefficient of determination (R2) is above 0.80. While at atmospheric pressure MIP fills on average 30% of total pores and the coefficient of determination (R ) is below 0.60. MIP at atmospheric pressure produces identical results as Pechiney Apparent Density, but it does not reveal any details useful for optimizing total pitch demand, since it does not include meso-micro porosity. The comparison of MIP and Pechiney Apparent Density is not included in this study.

Pore Size Distribution, 8x14 TM

VBD 8x14 vs Porosity 8x14

28 VBD 8x14 vs Fbrosity 8x14

27 26 25

i24 2 23 22 21 20

19 0.750

133.11 89.864 32.828 5.004 0.765

0.780

0.795 0.810 VBD(grrVml)

0.825

0.840

0.855

0.554

0.045

0.010

0.004

Pore Diameter (μιη)

Figure 6. Pore size Distribution chart for samples of different VBD 8x14

Figure 4. Correlation of VBD 8x14 vs Porosity 8x14

927

Pore Size Distribution, 8x14 TM

Sample 1, VBD 0.820 and Sample 3, VBD 0.766 showed a very similar pore size distribution up to 0.10 μπι with a delta porosity (difference of cumulative porosity above 0.10 μπι between two samples) of +0.5%, but below 0.10 μπι the delta porosity (difference of cumulative porosity between two samples below 0.10 μπι between two samples) is -4%. Due to this difference in porosity below 0.10 μπι, sample 3 showed a lower VBD compared to sample 1. This shows that the micro porosity has a significant influence on VBD. Table 1 shows the delta porosity at different pore diameter between sample 1&2, sample 1&3 and sample 3&4. Delta

Delta

Delta

Pore

Porosity

Porosit

Porosity

Diameter (ìçé)

0.01

0.00

0.00

133.106

0.04

0.00

0.12

111.231

-0.04

-0.02

0.17

89.864

-0.38

-0.01

0.32

-0.50

0.01

-1.45

45.122

-0.95

0.00

-0.59

0.05

-2.10 -1.90

11.319

-2.80

1.051

Sampls 133.1189.864 32.828 5.004 0.554 0.045 0.010 Pore Diameter (ìçé)

0.004

Figure 7. Pore size Distribution chart for samples of different VBD 8x14 Figure 8 shows a comparison of pore size distribution data for samples having 0.820 and 0.766 VBD 8x14. These four samples are from different calcined coke producers. The difference in porosity between both sets of samples is 1.7% and 2.6% respectively. Sample 1 and Sample 2 have identical VBD values of 0.820. The VBD results are same, but the pore size distribution is very different. Both samples showed a delta porosity (difference of cumulative porosity above 0.10 μιη between two samples) of -2.77% up to 0.10 μπι pore diameter, but below 0.10 μιη, delta porosity (difference of cumulative porosity below 0.10 μπι between two samples) is +1.08%. A similar trend was observed for sample 3 and sample 4.

T"Δ

~

Cumu& Delta

Cumulative Delta below 0.1

Fbre Size Distribution, 8x14 TM Sample 1, VBD = 0.820, Porosity = 20.64%

32.828 5.004

-0.41

0.47

0.07

-0.01

-0.29

0.554

0.25 0.10 0.10 0.21 0.17 0.11 0.14 0.01 1.08

-0.12 -0.41 -0.85 -1.40 -0.55 -0.64 -0.08 0.04 -4.02

0.26 0.48 0.92 1.60 0.75 0.90 0.39 0.04 5.33

0.095 0.045 0.023 0.010 0.007 0.004 0.003 0.003

Table 1. Cumulative Delta Porosity (%) for Samples 1, 2, 3 and 4.

Sample 2, VBD = 0.820, Porosity = 22.38%

The above observations showed that the VBD correlates well with overall porosity but it does not reveal the proportion of macro and meso-micro porosity.

Sample 3, VBD = 0.766, Porosity = 24.23% Sample 4, VBD = 0.763, Porosity = 26.81%

A similar set of experiments and comparisons were carried out on calcined coke commercially produced from different sources and calcined in rotary kilns. Some of these coke samples were blended with non anode grade coke. In this set of experiments, VBD and Porosity was analyzed on 28x48 TM size fraction prepared according to ASTM D 4292 (2007). The VBD results were correlated against the total porosity obtained by MIP as shown in Figure 9. The VBD 28x48 data correlated well with the Porosity with coefficient of determination R2 >0.84. Figure 10 shows a plot of coefficient of determination of VBD and Porosity at different pore diameter and pressure. At 800 psi 87% of total porosity is filled by Mercury and the coefficient of determination (R2) is above 0.80. This data indicates that the VBD correlates well with the total porosity.

133.1189.864 32.828 5.004 0.554 0.045 0.010 0.004 Pore Diameter (ìðé)

Figure 8. Pore Size Distribution chart for samples of same VBD 8x14

928

Pore Size Distribution, 28x48 TM

VBD 28x48 vs Porosity 28x48 35.0

VBD 28x48 vs Fbrosity 28x48

•VBD = 0.833

2

R = 0.8372

-

*

VBD = 0.862

·

VBD = 0.893

— - VBD = 0.926 VBD = 0.980

0.825

0.850

0.875

0.900

0.925

0.950

0.975

1.000

134.8 89.75 32.84

VBD(gnVml)

4.9

0.554 0.045 0.01 0.004 0.003

Pore Diameter (um)

Figure 9. Correlation of VBD 28x48 vs Porosity 28x48 TM Figure 11. Pore size Distribution chart for samples of different VBD 28x48

R2 of VBD and Porosity vs Pore Diameter Pressure (psi) 1

112

2-97

6181

14471

23186

32887

52950

Average, 87%of Total Porosity

Pore Size Distribution, 28x48 TM

02280 (urn) 134.793 2.534

0.113

0.038

0.016

0.010

0.007

0.005

0.003

Pore Diameter (urn) — — R2 of VBD and Porosity vs Pore Diameter

Figure 10. Coefficient of determination of VBD 28x48 and Total Porosity vs Pore Diameter Figure 11, shows the pore size distribution pattern for samples with different VBD 28x48. The samples having high VBD are showing low porosity and vice versa. Figure 12 shows the comparison of three VBD 28x48 samples having similar VBD results. These samples have the same VBD results but different total porosity. All three samples; A, D and E showed significant difference in the porosity above 0.5 μπι. and below 0.5 μπι the difference in porosity was minor.

0.0 134.8 89.75 32.84

4.9

0.554 0.045

0.01

0.004 0.003

Pore Diameter (μητι)

Figure 12. Pore size Distribution chart for samples of same VBD 28x48 TM

929

particles generated with greater crushing energy. Gehlbach (8) mentioned in his study that the coke isotropy, coke porosity, severity of crushing and scalping of oversize or undersize coke during sample preparation has great influence on particle size, shape, surface irregularities and VBD. The current study could not determine the influence of such surface irregularities on VBD, but it is observed that the low VBD coke has more porosity associated with surface irregularities.

Pore Size Distribution, 28x48 TM - - · - - Sample F, VBD = 0.893, Porosity = 25.09% · Sample D, VBD = 0.870, Porosity = 25.54%

Conclusions • • 134.790 89.750 32840 4.900 0.554 0.045 0.010 Pore Diameter (ìçôé)

0.004

•

0.003

Figure 13. Pore Size Distribution of a Sample with different VBD 28x48

•

Figure 13 shows the pore size distribution of two VBD 28x48 samples. The total porosity of these samples is very similar but the VBD results are very different. Pore Size Distribution, 28x48 and 8x14

•

- - · - - Sample 1, VBD 8 x 14 = 0.763 ■ Sample 1, VBD 28 x 48 = 0.840 Sample 2, VBD 8 x 14 = 0.820 Sample 2, VBD 28 x 48 = 0.893

References 1.

2.

3. 134.790 89.750 32.840 4.900 0.554 0.045 Pore Diameter (ìçç)

0.0t)

Vibrated Bulk Density does not universally correlate to total porosity. VBD correlates best with total porosity obtained by MIP, but it is not a good predictor of micro (less than 0.05 pm) or macro porosity (pore size distribution). Mercury Apparent Density at atmospheric pressure does not reveal sufficient data on pore volume distribution to be a reliable predictor of total porosity. Historical correlation between important anode properties and simple vibrated bulk density tests could be ineffective in the present calcined petroleum coke industry due to blending of different quality coke, degradation of coke quality, likelihood of misinterpretation of VBD test method(s) and other problems associated with correlating the test methodologies revealed in this study. Presently VBD results must be corroborated by MIP to provide all necessary information for predicting the pitch demand for a calcined coke and behavior of the sized aggregates for making higher quality anodes.

0.004 0.003

Figure 14. Pore Size Distribution of a Sample (Two Different Particle Size)

4.

Figure 14 shows a comparison of pore size distribution of two size fractions, 8x14 and 28x48, prepared from two different coke samples. For both samples for different particle size, there is a different pattern for porosity, but the porosity below 0.55 μηι pore diameter (above 400 psi) is very similar. This observation reveals the fact that the large pores are destroyed on reduction to smaller particle size (7). Also, the 28x48 material is showing high macro porosity especially up to 32 pm pore diameter (5.5 psi) compared to 8x14 material. This observation could be due to the fact that internal voids (large pores) are more easily accessible to mercury through cracks or paths in the smaller particles because they are closer to the surface of these smaller particles. This observation may also be a function of surface irregularities (jagged edges with pits larger than 32 pm) present on the surface of the smaller

5.

6. 7. 8.

930

P. Rhedey, "A review of Factors Affecting Carbon Anode Consumption in the Electrolytic Production of Aluminum," Light Metals 1971, TMS-AIME, New York, NY (1971) pp. 385-408. D. Belitskus, Optimization of Raw Materials and Formulations for Hall-Heroult Cell Electrodes," Third Yugoslav Symposium on Aluminum, (preprints) Edited by Prof. Andrej Paulin, University of Ljubjana (1978). R. E. Gehlbach, E. E. Hardin, L. I. Grindstaff, and M.P. Whittaker ; " Coke Density Determination and Its Relationship to Anode Quality," Paper Presented at the 109th AIME Annual Meeting, Las Vegas, NV, Feb 24-28 (1980). S. S. Jones, R. D. Hilberbrandt and M.C. Hedlund, "Variation of Anode Performance with Coke Quality," Paper No. A77-97, presented at the 106th AIME Annual Meeting, Atlanta, GA, 1977. D. Belitskus, D. J. Danka, "A comprehensive determination of effects of coke properties on aluminum reduction cell anode properties, Light Metals, 1989, 429439. D. Belitskus, "Evaluating Calcined coke for Aluminum Smelting by Bulk Density," Light Metals 1974, TMSAIME, New York, NY (1974) pp. 863-878. D. Belitskus, "Standardization of a calcined coke bulk density test," Light Metals, 1982, TMS-AIME, Dallas, Texas, (1982) pp. 673-689. R. E. Gehlbach and W. E. Walsh, "Influence of Sample Preparation on Petroleum Coke Properties", Light Metals 1995, pp. 539-543

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Prediction of Calcined Coke Bulk Density 1

Marie-Josιe Dion \ Hans Darmstadt1, Nigel Backhouse1, Mike Canada2, and Frank Cannova2 Rio Tinto Alcan, Arvida Research & Development Centre, 1955 Boulevard Mellon, Jonquiθre, QC, Canada 2 BP Coke, 5761 McFadden Avenue, Huntington Beach, CA, USA Keywords: Calcined coke, VBD, Calcination Abstract

• Coke A ■ Coke B o Coke C Ö Coke D

CO

E .o

The vibrated bulk density (VBD) is one of the most important calcined coke properties. In the context of changing green coke quality, a reliable forecast of the calcined coke VBD from small samples of green coke is required. To a certain extent, the VBD can be predicted from green coke properties, such as the Hardgrove Grindability Index (HGI) and the volatiles content. However, the precision of this forecast is not sufficient for procurement decisions which can reach millions of dollars. Thus, Rio Tinto Alcan (RTA) studied several laboratory calcination techniques, including the use of two different laboratory rotary kilns at BP Coke and COREM {Consortium de recherche minιrale). There was good agreement between the VBD and the degree of calcination (Lc) of calcined cokes produced in the kiln at COREM and in RTA's Arvida calciner. Adequate agreement for the VBD was reached with the BP kiln. Furthermore, it was possible to replicate phenomena in the BP kiln that cause problems on industrial scale, such as the formation of coke rings. The methods will therefore be used in the future to support RTA's calcining operations.

0.9

* ^ ί ^ — "

·

•

2

c m 0.8

m

0.

a m >

; o

" 1

· m

0.7

i

30

50

1

1

1

90

110

i

70

HGI [-] Fig. 1. VBD of calcined cokes as function of the HGI of corresponding green cokes calcined at Strathcona procurement decisions on such correlations alone. The three RTA calciners routinely perform single source calcination tests to assess the quality of new sources and verify the properties of regular supplies. Typically, these tests correspond to about 6 h of production. This corresponds to some 100 t of green coke feed. Thus, these tests cannot be used for exploring potential new coke supplies on a routine basis. For this, techniques based green coke samples of some kg are needed.

Introduction RTA operates three rotary kiln calciners in Canada that supply the majority of the calcined coke consumed by RTA's North American smelters. The land-locked Strathcona calciner mainly procures its green coke from local refineries, whereas the Arvida and Kitimat calciners, with near-by sea ports, have different and sometimes new suppliers. In the context of changing green coke quality, a reliable forecast of the calcined coke quality is required, ideally based on characterization of green coke samples of some kilograms. Calcined coke impurities and coke structure can be reasonably well predicted from green coke properties. However, this is not straightforward for one of the most important calcined coke properties, the VDB. The VBD is used to predict the optimum binder content [1] and it correlates with the baked anode density (BAD) [2,3]. In the context of deteriorating coke quality (such as lower VBD) there is strong pressure to increase BAD in order to support amperage increases. Consequently, selection of the best available green cokes is a priority for RTA.

For VBD prediction from small samples, the former Alcan developed a lab "flash" calcination method [4]. Historically, the VBD of lab-calcined and of plant-calcined cokes correlated well. However, in the recent years, the lab method considerably overestimated the VBD of some plant-calcined cokes, which is most probably due to the changing anode grade coke quality [5,6]. It was therefore decided to evaluate alternative methods. An important difference between the Alcan lab "flash" calcination and plant calcination is that in a plant rotary kiln the coke bed is strongly agitated, whereas during the "flash" calcination the coke bed is static. As it was believed that this has significant impact on the VBD, only lab calcination methods with agitated coke beds were considered. Experimental Details

To a certain degree, the VBD can be predicted from the green coke properties, such as the HGI and the volatiles content. The quality of such a prediction is illustrated in Fig. 1 for different cokes produced at the Strathcona calciner. There is a fair correlation between the green coke HGI and the VBD of the corresponding calcined cokes. The quality of the correlation is only slightly increased when other green coke properties are considered as well. For example, inclusion of the green volatile content in the correlation increases the r2 value only insignificantly (0.6059 vs. 0.6050). It can be concluded that although a rough prediction of calcined coke VBD from green coke properties is possible, it is somewhat risky to base

Coke calcination tests were performed in two labs: COREM (Quebec City, QC, Canada) and BP coke (Huntington Beach, CA, USA). Single source calcination tests of the same cokes were also performed at RTA's Arvida plant calciner (Jonquiθre, QC, Canada). COREM Lab Kiln The COREM lab kiln is usually used for the induration of iron pellets simulating the grate-kiln process. It is heated with propane,

931

which is mixed with primary air and, if so desired, with secondary air and oxygen (Fig. 2).

Typically, coke is sequentially treated in two kilns. Calcination is performed in an electrically heated rotary kiln with an inconel shell (length ~ 3 m, external diameter - 1 5 cm, Fig. 3). The kiln is equipped with a coke feeding system and a water-cooled coke outlet system. At the discharge end, the kiln is purged with nitrogen. However, air enters the kiln via leaks. The air inlet is controlled by the draft which is regulated by a valve close to the exhaust. As the maximum temperature of the rotary kiln is limited to 950 °C, a second static kiln (Fig. 4) is used to increase the calcination level of the coke to the desired level. For the calcination tests, the rotary kiln was heated to 950 °C. Once this temperature was reached, some 4 kg green coke (particles smaller than 1.25 cm and at least 50 % of the particles larger than 0.63 cm) were fed over a period of 45 min to the kiln. The coke residence time in the kiln was also approximately 45 min, which is comparable to the residence time in a plant kiln.

Fig. 2. COREM lab rotary kiln

After calcination in the rotary kiln, the samples were re-heated in the static oven at a temperature of 1225 °C for 30 minutes to obtain the desired Lc value.

For the coke calcination tests, the following procedure was used. The kiln was preheated to a temperature of 350-400°C. Once this temperature was reached, 13 kg of green coke were quickly introduced and the kiln was heated with a rate of 50 °C/min up to a final temperature of 1400 °C. Starting at 1100 °C, typically every 50 °C, 500 g coke samples were removed from the kiln with a ladle. Removal of larger amounts would disturb the thermal equilibrium of kiln. Once the final temperature was reached, the kiln was tilted to remove the remaining coke. Throughout the test, the kiln rotation was maintained at 3 RPM.

Arvida Plant Calciner Cokes calcined in the lab kilns were also calcined at the Arvida plant calciner. This plant is equipped with two kilns. The kiln gases of each kiln pass through a boiler for heat recovery and a bag house before they reach a common stack. During the single source tests, the kilns were run as during normal production. As the calciner moves towards under-calcination [7], the targeted coke calcination level presently corresponds to an Lc value of 24 A. Routine operation is without external fuel using burning coke volatiles as heat source instead. The calcination level is mainly controlled by the amount of tertiary air admitted [8].

BP Lab Kilns The BP lab kilns are exclusively used for studies on anode grade cokes such as:

Analytical Methods

• Impact of refinery operations on commercially calcined cokes • Verification if the structure of "opportunity" green cokes produced in the BP refining system is appropriate for the use in anodes • Under-calcined coke

The VBD was determined according to the ASTM method D7454 (Alcan VBD), whereas determination of the calcination level Lc is described in ref. [9]. Results and Discussion COREM Lab Kiln Properties of Lab-Calcined Cokes and Repeatability. As coke samples are collected over a range of calcination temperatures from the kiln, the dependence of calcined coke properties (such as the VBD and Lc) on the calcination temperature is obtained. The corresponding graphs show the known relationships for a lab setting (Fig. 5). The coke calcination level (Lc) steadily increases with the calcination temperature, whereas the VBD approaches a maximum which is associated with the start of coke desulfurization [10].

Fig. 3. BP lab rotary kiln used for calcination

The repeatability of the method was tested by calcining the same green coke three times. The corresponding VBDs are presented in Fig. 6. Up to a calcination temperature of 1200 °C, the VBDs were very similar (standard deviation < 0.006 g/cm3). However, for higher calcination temperatures small differences between the VBD values were observed. The standard deviation reached 0.013 g/cm3 after calcination at 1400 °C. Given that such high

Fig. 4. BP static kiln used for the adjustment of the calcination level

932

In another series of tests, three calcinations were performed up to a temperature of 1325 °C which corresponds to the temperature reached during plant calcination of the coke used. Once this temperature was reached, the kiln was emptied. In spite of the relative high calcination temperature, there were little differences between the VBD and Lc values of the corresponding calcined cokes (Fig. 7). The standard deviations were 0.004 g/cm3 for the VBD and 1.2 A for the Lc value. However, in this specific case, it has to be mentioned that no coke samples were collected during the tests, which may have contributed to a better thermal stability and hence, more consistent calcined properties.

0.85 5Γ

E .o

0.80

Q

0.75

m >

0.70 1100

1200

1300

Properties of Lab vs. Plant-Calcined Cokes. Comparison of the properties of the cokes calcined in the Arvida plant and in the COREM kiln at the same temperature showed good agreement for the VBD of four of the five cokes studied (Fig. 8) For coke Y, however, the plant VBD was considerably overestimated by the lab test. This coke had an extremely fine granulometry and a high volatile content. It is assumed that this changed the flow of the coke particles in the coke bed upon calcination, which is known to impact coke properties [11,12]. Furthermore, it is also well known that cokes with more volatiles have a tendency to release them faster, creating more porosity [10]. Given the very different dimensions of the plant and lab kilns, the lab method was most probably not capable to reproduce the plant heating rate for this "unusual" coke.

1400

Calcination T. [°C] Fig. 5. VBD and Lc values of coke "A" calcined in die COREM kiln as function of the calcination temperature ö Ject 1 m Toet 0

0.98

E S

0.96 0.94

û

0.92

>

0.90 ,

m

0.88 1000

n Toct Q

Ó

i

/o

i

1100

1

vj è

I

1200

1300

Not considering coke Y, the r2 value of the correlation was 0.96. For high VBDs, however, the lab method slightly overestimates the plant VBD. Nevertheless, the prediction is much better as compared to the method based on green coke properties (Fig. 1, r2 = 0.61)

I

14C

Ciilcinatl·3ΠΤ. [^C] Fig. 6. Repeatability of the COREM method. VBD of cokes obtained from the green coke "B" in three calcination tests

0.92

29

27 0.70

0.75

0.80

0.85

0.90 3

VBD COREM [g/cm ]

25

0.88

Fig. 8. Correlation between the VBD of cokes calcined in the COREM lab kiln and at the Arvida plant calciner

Test

The agreement between the calcination levels (Lc) of the cokes calcined in the two kilns was, with one exception, good as indicated by a r2 value of 0.84 (Fig. 9). Furthermore, the Lc value of coke Y calcined in the plant was reasonably well predicted by the lab test. Apparently, for a given calcination temperature, the coke granulometry and volatile content had less impact on the coke calcination level than on the VBD.

Fig. 7. Repeatability of the COREM method. Properties of cokes obtained from green coke "C" heated to 1325 °C temperatures are rarely reached during plant calcination, the VBD differences at high lab calcination temperatures were judged acceptable. It is also believed that these differences could be reduced by minimizing the number and the size of the coke samples collected during the calcination, which may have a significant impact on the thermal balance inside the kiln.

933

28

r2 == 0.84 /%

26 e J2

CL o

It should be mentioned that the correlation between the green coke properties and the calcination temperature is unique for every kiln and even for a given operation philosophy. However, once this relationship is known (as shown in Fig. 10 for the Arvida plant), the lab data (as shown in Fig. 5 for one coke) can be used to predict the properties of a calcined coke produced in a given plant.

ι

24

>>·'

22

Coke

Y

20

; Le

18

•

·

18

Furthermore, usually the plant calciners are fed with coke blends. These blends often contain green cokes with different volatiles levels. Thus, the calcination temperature reached and the properties of the corresponding calcined cokes differ between single source calcinations and calcination of coke blends. The data presented here allow predictions to be made even in these cases. Instead of predicting the plant calcination temperature from the green coke volatiles content of a single source, the volatiles content of the blend is used to estimate the plant calcination temperature (from graphs as shown in Fig. 10) and to estimate the properties of the corresponding calcined cokes (from graphs as shown in Fig. 5).

/ I

20

22

I

24

26

28

Le COREM [Β] Fig. 9. Correlation between the calcination level (Lc) of cokes calcined in the COREM lab kiln and at the Arvida plant calciner Prediction of Properties of Plant-Calcined Cokes from Lab Tests. After it was confirmed that for green cokes (without extremely fine granulometry and high volatiles content) the VBD of cokes calcined in the COREM lab kiln and in the Arvida plant kiln correlates well, a method for the prediction of the plant-calcined coke VBD from lab data was developed.

BP Lab Kiln Properties of Lab vs. Plant-Calcined Cokes. The VBDs of cokes calcined in the BP lab in the kiln are compared in Fig. 11 with the corresponding plant-calcined cokes. For the samples without subsequent heat-treatment in the static oven, there was good agreement (except for one coke discussed below). However, the lab calcination slightly underestimated the VBD. This was attributed to the fact that the lab calcination temperature (950 °C) was lower than the plant calcination temperature (Fig. 10). As the VBD decreases with decreasing calcination temperature (before significant desulfurization occurs), this lead to lower VBDs for the lab-calcined cokes. This assumption is supported by the observation that upon subsequent heating in the static oven, the VBD of the lab samples increased (Fig. 11).

The present control strategy of the Arvida kiln is based on a target coke calcination level, as opposed to a fixed calcination temperature. Under these conditions, the calcination temperature strongly depends on the green coke volatile content. The higher the volatile content the lower is calcination temperature (Fig. 10). Based on this correlation and the volatiles content of the green coke to be evaluated, the calcination temperature in the plant kiln is estimated. For example, upon calcination of a green coke with 10 % volatiles in the Arvida kiln a calcination temperature of about 1275 °C is expected. The properties of the corresponding calcined coke produced at this temperature are obtained from the lab tests. As shown in Fig. 5, calcination at 1275 °C is expected to yield a coke with a VBD of 0.82 g/cm3 and a Lc value of 24 A. This method allows therefore prediction of properties of plantcalcined cokes from reasonable small samples (-13 kg) of green coke.

,__,

1350

—

1300

9 H C

*

o

CO

E

c iS 0.75 a. a 0.70 m

1200

en O

1150

Ü

··- -, D

D

:

Coke Y

Φ

>

0.70

1250

c

J · Rotary Kiln : D Rotary Kiln & Oven

0.80

*

CO

0.85

11

0.80

0.85

VBD BP [g/cm 3 ]

# |~

10

0.75

12

Fig. 11. Correlation between the VBD of cokes calcined in the BP lab kiln (with and without additional heating in the static oven) and at the Arvida plant calciner

13

After the additional heat-treatment, for some samples the lab results considerably overestimated the plant VBD. This was in spite the fact that the lab heat-treatment temperature (1225 °C) was close to the plant calcination temperatures (Fig. 10). The structural changes in both kilns were apparently different. This is also suggested by the important differences between the

Volatiles [%] Fig. 10. Temperatures reached in the calcination zone of the Arvida plant calciner upon calcination of green cokes with different volatile contents

934

calcination levels (Lc) of cokes from the lab and the plant kiln (Fig. 12).

^ "^M Hfl·!

As for the COREM kiln, the BP lab data considerably overestimated the VBD of plant-calcined coke Y. Apparently the lab methods did not replicate the behavior of this high-fines, highvolatiles coke during plant calcination. However, there were clear indications during the BP lab calcination experiments that coke Y has unusual calcination behavior. A glass window at the coke outlet end of the kiln allows the observation of the burning coke volatile flames. The desired flame has a corkscrew-like shape and follows the kiln wall, occupying only a portion of the kiln crosssection (Fig. 13). With all cokes studied, with the exception of coke Y, such a flame shape could be easily obtained by manipulating the draft. However, upon calcination of coke Y, most of the time a turbulent flame, occupying the entire kiln cross-section of the kiln was observed (Fig. 14). Even by regulating the draft, a flame shape as shown in Fig. 13 could not be obtained. The turbulent flame was most probably due to the fast release of large quantities of volatiles. This usually causes a low VBD as the coke structure is blown-up (popcorn effect). It can be concluded that although the VBDs of lab and plantcalcined coke Y differed, there were clear indications during the lab test that the VBD of plant-calcined coke Y will be low.

Fig. 15. Coke ring formed upon calcination of coke Y in BP lab rotary kiln

coke volatiles) adhering to the kiln shell. Coke ring formation is also observed in plant kilns, where it has a significant negative commercial impact. The coke ring hinders the downhill flow of the coke bed towards the discharge end. Furthermore, the decreased free kiln cross section increases the velocity of the kiln gases, which increases entrainment of coke fines. Once the coke ring reaches a certain size, coke spills at the feed end of the kiln occur and the coke recovery rate decreases. The coke ring has therefore to be mechanically removed, which requires temporary shut-down of the calciner. It is known that the tendency for coke ring formation increases with the amount of coke volatiles and also depends on the nature of the volatiles [13]. Furthermore, it is especially observed in short kilns where the release of coke volatiles is concentrated in a relative small zone of the kiln. RTA adopted some operational changes which limit coke ring formation. However, until now, coke ring formation cannot totally be avoided at the Kitimat calciner (Fig. 16), which has the shortest kiln in the RTA system. More knowledge on coke ring formation is therefore required.

J · Rotary Kiln 28 ] G Rotary Kiln & Oven °<,

- · • 24

%

20

*T a

_i

D D

• Coke Y

m

0

DD

16

The BP lab kiln allows easy removal of coke ring samples. Characterization of these samples (Fig. 17) allows to gain more insight on early stages of coke ring formation. This is very difficult with samples removed from the plant kiln as they have been exposed for long times to very high temperatures and therefore underwent important structural changes (Fig. 18). Furthermore, the impact of kiln surface (such as its morphology and temperature) on coke ring formation can be studied.

I

16

20

24

28

L C BP[ΐ] Fig. 12. Correlation between the calcination level (Lc) of cokes calcined in the BP lab kiln (with and without additional heating in the static oven) and at the Arvida plant calciner

Fig. 13. Desired shape of flame upon calcination in BP lab kiln

Fig. 16. Coke ring formed in RTA's Kitimat plant calciner

Fig. 14. Undesired shape of flame obtained upon calcination of coke Y in BP lab kiln

Fig. 17. Coke ring sample removed from the BP lab kiln

Formation of Coke Ring. Upon calcination of coke Y, a so-called coke ring was formed in the coke pre-heating zone, close to the feed end of the BP lab kiln (Fig. 15). Such a ring consists of solid carbonaceous material (for example coke fines and condensed

Fig. 18. Coke ring sample removed from the RTA Kitimat plant kiln

Conclusions Based on green coke properties alone, reliable prediction of the VBD of plant-calcined coke is not possible. Due to the required

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logistics single source plant calcination tests cannot be used on a routine basis for the prospection of potential new coke sources. Prior lab calcining methods due to their static nature also had limited success in predicting industrial calcination results. On the other hand, the lab calcination methods that have been described in this paper, have been determined to provide good prediction of large scale rotary kiln calcined coke properties with reasonable effort. Lab calcination also allows the replication of phenomena such as coke ring formation that represent important problems on commercial scale. This makes it possible to test and develop strategies to improve the calcination process. Such lab calcination methods will therefore be used in the future to support the RTA and BP calcining operations. This includes the prospection of potential new coke sources as well as the optimization of calcination of coke blends. References 1. A. L. Proulx., "Optimum Binder Content for Prebaked Anodes", Light Metals, (1993), 657-661. 2. D. Belitskus., "Evaluating Calcined Coke for Aluminum Smelting by Bulk Density", Light Metals, (1974), 863-876. 3. D. Belitskus, D. J. Danka., "A Comprehensive Determination of Effects of Calcined Petroleum Coke Properties on Aluminum Reduction Cell Anode Properties", Light Metals, (1989), 429-442. 4. P. J. Rhedey., "Structural changes in petroleum coke during calcination", Transactions of the Metallurgical Society ofAIME, (1967), 239, 1084-1091. 5. F. R. Canno va and B. C. Vitchus "Controlling Calcined Petroleum Coke Quality in a Changing Oil Industry", 8th Australasian Aluminium Smelting Technology Conference, Yeppoon, Australia, 2004. 6. F. Vogt, R. Tonti, M. Hunt, L. Edwards., "A preview of anode coke quality in 2007", Light Metals, (2004), 489493. 7. J. Lhuissier, L. Bezamanifary, M. Gendre, M.-J. Chollier., "Use of Under-Calcined coke for the production of low reactivity anodes", Light Metals, (2009), 979-983. 8. F. J. Farago, D. G. Retallack, and R. R. Sood "Calcination of coke", US Patent 4,022,569,1977. 9. F. R. Feret, "Determination of the crystallinity of calcined and graphitic cokes by X-ray diffraction", Analyst, (1998), 123, 595-600. 10. L. Edwards, K. Neyrey, L. Lossius., "A Review of Coke and Anode Desulfurization", Light Metals, (2007), 895900. 11. M. F. Vogt, G. R. Jones, G. A. Tyler., "The Use of Refractory Lifters in Coke Calcination", Light Metals, (1984), 1697-1714. 12. L. P. Lossius, K. J. Neyrey, L. C. Edwards., "Coke and Anode Desulfurization Studies", Light Metals, (2008), 881-886. 13. D. DuTremblay, P. J. Rhedey, and H. Boden, "Agglomeration Tendency of Petroleum Coke During Calcination", AIME Annual Meeting, New Orleans, Louisiana, USA, 1979.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CALCINED COKE PARTICLE SIZE AND CRUSHING STEPS AFFECT ITS VBD RESULT Frank Cannova(1); Mike Canada(1); Bernie Vitchus; (2) *BP, 5761 McFadden Ave., Huntington Beach, CA 92649 2 Coke Advisor, [email protected] Keywords: Vibrated Bulk Density, VBD, Calcined Coke Size Distribution, Particle Shape Abstract The size of a calcined coke particle used in the Vibrated Bulk Density (VBD) test and the size of the particle before crushing affects its VBD analysis. That is, naturally occurring particles usually have a higher packing density and VBD compared with particles that are crushed to the same size. Consequently, calcined coke preparation crushing steps can dramatically affect the VBD result. Data, showing how calcined coke particle size and crushing steps affect the VBD result will be presented. These data help explain why the roll crushing steps need to be controlled to improve VBD repeatability and reproducibility. In addition, data will be presented showing how the roll crusher operation and maintenance affects the VBD result. Introduction The ASTM D4292 calcined coke Vibrated Bulk Density, (VBD) test defines sample preparation to control the process and improve the repeatability of the test. Although the VBD of a calcined coke is affected by many production variables, only the particle's size and crushing steps will be discussed in this paper. There have been many references in the literature as to how particle size affects the VBD test results (15).These papers show: •

The size of the coke particle affects VBD resultUsually Larger size particles have lower VBD, Crushing larger size particles to a given size, usually results in a lower VBD than particles that are naturally occurring. Since larger sized particles result in lower VBD, any change in the starting particle size distribution will affect the VBD result. For example samples that are subject to segregation will affect the VBD result. The following Figure 1 shows how the VBD varied as the %+4 mesh changed while a ship of calcined coke was being unloaded(1).

• •

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Figure 1 VBD of Calcined Coke vs %+4 mesh

1 80

1 90

Several papers (2' 4' 6) also suggest that particle shape has a significant affect on the packing density and therefore the Vibrated Bulk Density. ASTM Sample Preparation Procedure The sample preparation section of ASTM D4292 addresses the crushing steps and gives ranges for roll crusher settings. Even within the accepted range of at least 30% crushed and the ratio of the coarser to finer product 0.8 to 2.0, significantly different VBD results can be obtained. For most cokes analyzed for VBD (-28+48 Tyler mesh), greater % crushed and lower coarser to finer ratios result in lower VBD results. Outside of these specified ranges even larger VBD differences will be observed. Consequently, the crushing steps and roll crusher settings need to be monitored to get more repeatable results. The crushing steps consist of first a jaw crushing step with a recommended 5 mm at the closest setting. For our lab, the furthest gap is around 5 mm with the closest gap almost 2 mm. Using the jaw crusher closest gap at 5 mm generates a roll crusher feed that plugs the roll crusher. Even though the jaw crushing furthest gap is set at 5 or 6 mm, particles that are significantly larger in length or height get through with the width of the largest particles 5 to 6 mm. This VBD study was initiated because our carbon lab consistently obtained bias low VBD results compared to an outside lab. For Jan 2010 through early May, our in-house lab weekly VBD monitoring (40 samples) averaged 0.853 g/cc for -28+48 mesh calcined coke whereas an outside lab daily monitoring (104 samples) VBD averaged 0.872 g/cc. That is, our in-house lab averaged almost 0.02 g/cc lower VBD. When the same prepared VBD sample was analyzed at each lab, the results were almost identical. Even though each lab used the same roll crusher feeler gauge gap, it became apparent that the crushing steps were probably affecting the VBD result. This observation led to a close inspection of our roll crusher. By measuring the roll crusher gap at different positions before and after roll crushing, it was found that the rollers had worn unevenly such that the gap varied from the outside edge to the middle and at different positions of the rollers. Feeler gauges were needed to observe this variation. In addition, the gap changed from the beginning of roll crushing to after roll crushing. This change in gap from beginning of the roll crushing to after roll crushing had the largest impact for the closer roll crushing setting. Up to 0.1 mm reduction in gap was measured for the closest roller crusher gap setting of 0.48 mm. A complete rebuild of the roll crusher was performed before performing the tests presented in this paper.

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The type of feeler gauges used to set the gap also makes a difference in the actual gap between rollers. Using automotive type feeler gauges to set a roller gap will result in a gap that is slightly greater than using a custom made single feeler gauge. Using automotive feeler gauge often results in using one very thin gauge which can bend. Consequently, the gap is often slightly larger when the automotive feeler gauges are used to set the gap. The final roll crusher setting is the most critical gap in producing the actual size distribution of the coke for the ASTM VBD measurement.

Performing a 5th roll crushing on the particles 28 mesh and larger produced a calcined coke having a 0.847g/cc VBD. When this material was blended with the previously crushed calcined coke, its VBD was 0.862 g/cc with 68 % crushed and a coarser to finer ration of 0.12. This experiment shows how continuing to crush larger particles of Coke C results in calcined coke with a lower VBD (poorer packing density). The same trend was observed for most other cokes tested (See following Figure 3). These four calcined cokes demonstrate that the natural occurring particles have a higher VBD than crushing larger particles to the same size. However, it has also been observed that the natural occurring particles can have lower VBD.

Experiment and Results Experiments were run to quantify the effect of crushing the larger particles to the desired particle size (-28 +48 Tyler mesh) for the VBD analysis. Four different single source cokes were tested. Their properties are found in the following table: Table 1 Chemical properties of Calcined Cokes in pa per A Coke B C Sulfur, % 0.69 2.72 0.96 Vanadium, ppm 246 133 378

D 3.18 401

Initially the ASTM D4292 was strictly followed. However, it became apparent that as the calcined coke was crushed more, the VBD of the calcined coke decreased. This observation led to additional crushing steps at closer roll crusher gaps followed by additional VBD analysis of the newly crushed calcined coke and another VBD analysis of the newly formed crushed coke blended with the formerly crushed coke. The following Figure 2 demonstrates how this was done and the VBD results obtained for Coke C. 0.895

20%

30%

40%

50%

% of sample crushed to -28+48 * Coke A — L i n e a r (Coke A)

* CokeB Linear (Coke C)

CokeC Li near (Coke B)

Figure 3 - VBD vs % Crushed to -28+48 mesh for Coke A through Coke D Larger Symbols points meet ASTM Crushing Over/Under Ratio The Vibrated Bulk Density of cokes crushed in a roll crusher, converge as the cokes are crushed more. See the following Table 2. The natural size calcined cokes have a delta VBD difference of 0.140 g/cc whereas the crushed calcined cokes have a VBD delta of only 0.031 g/cc. Since crushing larger calcined coke particles with greater internal pores opens up these pores, the finer cokes would be expected to have more similar packing densities. In fact, one of the techniques of increasing anode density of calcined cokes with lower packing density is to use them in the fines fraction. However, cokes with higher packing density do produce anodes with higher Baked Apparent Density.

0%

10%

20%

30%

40%

50%

60%

70%

% of coke crushed to -28-1-48 mesh

Figure 2 VBD of Coke C as function of % Crushed This plot shows that 11% of Coke C was in the -28+48 mesh size range and has a VBD of 0.889 g/cc. After following the ASTM roll crushing step with the 3rd roll crusher set at 0.6 mm, 41.6 % of the coke was in the -28+48 mesh range and had a VBD of 0.870 g/cc and a coarser to finer sized ratio of 1.4. Additional crushing of coke larger than 28 mesh at a 0.55 mm roll crusher setting resulted in a coke with a VBD of 0.855 g/cc. When this coke was blended with the previous -28+48 crushed coke, the VBD was 0.862 g/cc with 50% crushed and a coarser to finer ratio of 0.89. Both of these crushed samples meet the VBD sample preparation sizing requirements but with different VBD results.

Table 2 Comparing the VBD of Natural Occurring -28+48 mesh calcined coke to VBD of ASTM prepared -28+48 after crushing Coke Natural VBD % Crushed to ASTM VBD (-28+48 mesh), -28+48 for (-28+48 mesh) g/cc VBD Test g/cc B 0.962 54 0.893 A 0.952 63 0.885 D 0.893 56 0.870 0.889 68 0.862 c The calcined coke's natural occurring -28+48 mesh particles were examined under 210 magnification to see if particle shape differences can be observed to help understand differences in VBD. These pictures are presented in Figure 4.

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Recommendations This study was initiated due to observing that in-house lab obtained VBD results were biased lower than an outside lab. Due to this, a quality control comparison of results between labs has been established that compares the % crushed to the desired particle size and the over/under ratio. In addition, a maintenance monitoring of the roll crusher has been established and includes: • Daily checking that the roll crusher gap via a feeler gauge is the same before and after crushing coke particles. • Daily observing the grease around the bearings. • Monthly adding grease to the bearing casing. • Installing a clutch so the roll crusher starts smoother instead with a jerk. We suspect the greatest wear is due to the jerking start of the rollers. We recommend other labs performing VBD tests consider a similar VBD monitoring and roll crusher maintenance program. Figure 4 - Pictures of calcined coke particles which have different VBD properties Comparing cokes A & B particle shapes to cokes C & D particles shows cokes C & D particles have many bumps and rounded appendages which could reduce particle packing density. Conclusions •

• • •

The sample preparation procedure for ASTM D4922 defines how to crush the calcined coke, but even in the range specified the VBD can vary. Outside of this range significantly greater variability is observed. Crushing larger particles or increased crushing usually leads to lower VBD results. It appears that particle shape has an affect on the VBD result with smoother surface of naturally occurring having a higher VBD due to closer packing properties. Although often overlooked, the operation of the roll crusher can have a dramatic affect on the VBD analysis of a given coke. Monitoring the consistency of roll crusher gaps can identify when the roll crusher needs to be rebuilt.

Acknowledgement A special thanks to Jim Baker for rebuilding our Roll Crusher and Bob Marquez for performing all the tests. References (1) Frank Canno va, Bernard Vitchus, Howard Childs, Yen Hoang, "Managing the Impacts of Coke Sizing and Segregation on Anode Performance", presented at Seventh Australasian Aluminium Smelting Workshop, November 2001. (2) R.E. Gehlbach,et al, "Influence of Sample Preparation on Petroleum Coke Properties", Light Metals (1995), pp 539-643. (3) Bernie Vitchus and Frank Cannova, "Understanding Calcined Coke Bulk Density- Inventory", Light Metals (2007) pp 1035-1038. (4) Bernie Vitchus, Frank Cannova, Randall Bowers, Shridas Ningileri, "Understanding the Calcined Coke VBD- Porosity Paradox", Light Metals (2008) pp. 871874. (5) D. Belitskus, "Standardization of a Calcined Coke Bulk Density Test", Light Metals (1982), pp. 673-689. (6) Randall Bowers, Shridas Ningileri, David C Palmlund, Bernie Vitchus, and Frank Cannova, "New Analytical Methods to Determine Calcined Coke Porosity, Shape, and Size", Light Metals (2008) pp. 875880.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

BULK DENSITY OVERVIEW OF ASTM AND ISO METHODS WITH EXAMPLES OF BETWEEN LABORATORY COMPARISONS Lorentz Petter Lossius1, Bill Spencer2 and Harald A. 0ye 3 1 Hydro PMT, Primary Metal Technology, Ardal, Norway 2 Oxbow Calcining LLC, Kremlin, Oklahoma, USA 3 Norwegian University of Science and Technology, NTNU, Trondheim, Norway Keywords: Petroleum coke, bulk density testing, VBD, TBD, ASTM, ISO cokes have to be considered and evaluated as anode raw materials. It is difficult for the anode manufacturer to • Monitor the producers' coke • Compare and evaluate different producers' cokes

Abstract The bulk density of petroleum coke is an important property when evaluating a coke for use in anodes in primary aluminum metal electrolysis. It is also an important property in petroleum coke trade. There are international standards for testing coke bulk density; ASTM has two vibrated bulk density (VBD) methods, D4292 and D7454, and ISO has a tapped bulk density (TBD) method, ISO 10236. There is a concern in anode production that it is difficult to obtain sufficiently good between-laboratory comparisons with any of these methods, both for use in comparisons, and when used for distinguishing coke qualities from different producers. The paper will present the methods, give results from several interlaboratory studies and discuss the between-laboratory comparison results.

Grain Fraction Size and Bulk Density Value The reported VBD or TBD value is dependent on the selected grain fraction. By measuring over a wide range of sieves the variation with grain size can be shown. Taking VBD as an example, a 2008 Hydro test with 13 sieves and a bottom fraction of -0.180 mm is shown in two charts in Figure 1. Different anode raw materials were analyzed including three sponge cokes/blends, a soft sponge coke and a good quality butts.

Introduction This paper is part of the 2011 TMS special session on petroleum coke bulk density. It is meant to give an overview of methods and some quantification of the precision of the methods as regards comparison between laboratories. Due to variation in practice, both Tyler mesh and metric indication of sieve ranges are included.(1) Bulk Density A good discussion of the relationship between bulk density and anode quality is found in [1], starting on page 93. Coke bulk density expresses a combination of grain size, grain packing and porosity. It is an indication of the potential a coke has to contribute to good anode density and is much in use as a petroleum coke specification by coke calciners and anode manufacturers. The principle of measurement is an ordered, systematic filling of a volumetric cylinder with a test portion of a coke sample with a defined grain size range. For a specified time, vibration or tapping is applied to achieve packing. The mass to volume ratio is determined and the bulk density is reported in g/cm3. It has been recognized for some time that the between-laboratory reproducibility, which determines the quality of comparisons between the coke producer and anode manufacturer, is less good than what is needed for monitoring coke quality at laboratories. The precision for within-laboratory comparison is acceptable, and makes it possible for the producer to monitor own production and fulfill specifications. But the poor between-laboratory comparisons are troublesome, especially as more and different 1

Figure 1: VBD for sieve fractions down to 0.180 mm. Both charts are linear, with details -1.0 mm in the lower chart. Vertical lines are the Tyler mesh 28x48 range (0.3-0.6 mm).

Tyler mesh 28x48 is 0.30 to 0.60 mm particle size.

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• • • • •

A majority of 12 laboratories reported as requested, while six reported results with alternative grain ranges or analysis methods. Assuming this represents the standard practice of these laboratories, Figure 2 well illustrates the current situation with spread in practices and analyzed grain size ranges yielding a scatter of results for the same property.

The VBD increases as porosity is crushed out of grains. For the three regular sponge cokes/blends, the difference in level was stable across fractions so several grain fractions could equally be used for monitoring or comparisons. The soft coke differed from the sponge blends, and got low reported VBD especially if measured on medium fine grains. For the high quality butts, VBD was high and less dependent on the grain size across fractions than the cokes. In the detailed plot the size range 0.71 to 0.25 mm might look near horizontal, but the average VBD difference from the 0.710-1.0 mm fraction to the 0.250-0.355 mm fraction for the three cokes is as high as 0.036 g/cm3.

Factors that cause variation in the bulk density results The bulk density testing has historic roots and traditions, with companies having long-standing and different standard practices. This has led to a several coke grain fractions being in use; especially ASTM D4292, but to some degree also ISO 10236, were written to allow variation in grain size analyzed. This has over time been a disadvantage to comparison of cokes.

Different reported values when describing the same property An example of how confusing the situation can be when comparing different laboratories' reported VBD was taken from a CII Carbon round robin, RR #17, run in 2004. It is a very large RR with 45 participants, 18 of which reported VBD.

Other factors causing variation: • Bulk density is a complex material property, with contribution from grain porosity, shape and packing. • The sample preparation can be quite complex. • The requirements of sample preparation are not adhered to. • The spread in petroleum coke properties is widening; the wider range of sponge cokes, shot cokes and other anode raw materials is a challenge to traditional bulk density methods. • Method development has been slow; anode quality has developed significantly to adjust to increased cell amperage and current density requirements, and the need for accurate analysis methods is now felt more keenly.

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Development of ASTM D4292. VBD ASTM D4292 is the most widely used method for determining petroleum coke bulk density. It has gone through two revisions: original publication in 1983, and revisions in 1992 and 2010. The original round robin data for ASTM D4292-83 used a 20x48 Tyler mesh fraction (0.30-0.85 mm). The 1992 revision made a couple of minor changes, which include the addition of air drying the lab sample (section 8.1.1) and adding note 6 (precision of other sample size ranges not determined). The recent 2010 revision is an attempt to better specify equipment and procedures to reduce the variation in results between laboratories. In the industry, several particle size fractions are in use and it is critical to identify the appropriate particle size with the corresponding results.

A

Figure 2: VBD (g/cm 3 ) classed by method or grain size range. The upper chart is results for finer grains, lower for coarser grains. The value is the average. (2) A calcined coke RR sample was distributed prepared to -4 Tyler mesh.(3) For VBD, each lab was to return results for analysis by ASTM D4292 for the two common grain ranges • The finer ASTM VBD Tyler mesh 28x48 (0.30-0.59 mm) • The coarser Kaiser VBD Tyler mesh 8x14 (1.17-2.36 mm)

4

ASTM D4292 is under the jurisdiction of ASTM Committee D02 on "Petroleum Products and Lubricants" and is the direct responsibility of Subcommittee D02.05 on "Properties of Fuels, Petroleum Coke and Carbon Material".

2

ASTM D2854 is "Standard Test Method for Apparent Density of Activated Carbon". 3 -4 Tyler mesh is -4.75 mm.

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The 2010 revision precision statement is still under development in an ongoing ILS and all examples of comparisons presented in this paper have used the 1992 revision (reapproved 2007).

• • •

Overview procedure pre-2010 D4292 VBD • Grains within 3-65 Tyler mesh (0.21 to 6.68 mm) • A sieved, crushed test portion from jaw- and roll-crusher retained between screens differing less than 2Λ/2 • Test portion 100.0+0.1 g • Transfer time 70 to 100 sec prior to vibration • Graduated cylinder foot loose inside retaining ring • Vibration for 5 min at amplitude 0.20-0.22 mm • The coke height measured at eight positions around cylinder to 0.5 mm, then averaged • VBD is the average weight/volume ratio from two determinations, reported to 0.001 g/cm3

Feeding is cut by an automated optical method when the cylinder is filled to 50 ml Test portion is weighed to nearest 0.01 g The VBD is the average weight to volume ratio from minimum three determinations, reported to 0.001 g/cm3

Users of the old D4292 had freedom of choice of grain range. The most common grain size ranges reported are • ThefinerASTM VBD Tyler mesh 28x48 (0.30-0.59 mm) • The coarser Kaiser VBD Tyler mesh 8x14 ( 1.17-2.36 mm). Precision D4292-92 VBD The precision statement is based on determinations on the Tyler mesh 20x48fraction(0.30-0.85 mm, USA mesh 20x50). Precision at 95% confidence level is r = 0.014 g/cm3 R = 0.046 g/cm3 Comparison between laboratories, using the R limit: Given a determination on two test portions of the same material at two different laboratories, the difference in equivalent temperature should be within 0.046 g/cm3 for 95 out of 100 such comparisons. The old between-lab precision of 0.046 g/cm3 is poor compared to commercial requirements. It is also not good when considering coke qualities from different producers, if these have been analyzed at different laboratories. Precision 2010 D4292 VBD The precision of the D4292-10 revision will be determined through an ILS to be run concurrent with the ILS for the D7454 method and a test of the Micromeritics' GeoPyc.

Figure 3: STAS' semi-automated VBD apparatus. Photo courtesy of Rain CII Carbon, LLC. Precision ASTM D7454 VBD A repeatability standard deviation was given of 0.0036 g/cm3; this is single laboratory determination of a single coke eight times and does not fulfill the ASTM requirements. The ASTM Committee has accepted this method provisionally without the precision statement and an interlaboratory study is being run concurrent with the ILS for the new D4292-10 revision and a test of the Micromeritics' GeoPyc.

Vibrated Bulk Density, ASTM D7454(5) ASTM D7454-08 - Standard Test Method for Determination of Vibrated Bulk Density of the 1.17 - 4.7 mm Calcined Petroleum Coke Fraction Crushed to 0.42-0.83 mm, using a Semi-Automated Apparatus Overview procedure D7454 VBD • Grains non-crushed (natural) within 1.17 to 4.7 mm (4x14 Tyler mesh) are crushed in multiple passes by a roll-crusher to within 0.425 to 0.85 mm (20x35 Tyler mesh) • Volume based, test portion is 50 ml • Feeding rate target is 150+15 s for the test portion Figure 4: GeoPyc apparatus (Micromeritics model 1360). Photo courtesy of Rain CII Carbon, LLC.

5

ASTM D7454 is under the jurisdiction of ASTM Committee D02 on "Petroleum Products and Lubricants" and is the direct responsibility of Subcommittee D02.05 on "Properties of Fuels, Petroleum Coke and Carbon Material".

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GeoPvc The GeoPyc measures bulk volume using a controlled vibration and rotation movement to agitate the test portion. Precision GeoPvc VBD A precision study for the GeoPyc is being run concurrent with the ILS for D4292-10 and D7454-08. It is uncertain if sufficient participants with GeoPyc will take part to qualify a precision statement. Tapped Bulk Density, ISO 10236 Method(6) ISO 10236 (1995) — Carbonaceous materials used in the production of aluminium — Green coke and calcined coke for electrodes — Determination of bulk density (tapped) Overview procedure ISO 10236 TBD • A sieved, non-crushed (natural) fraction, 4.0-8.0 mm, 2.0-4.0 mm, 1.0-2.0 mm, 0.5-1.0 mm or 0.25-0.5 mm • With 1.0-2.0 mm most in use for instance by Esso, BP Veba, BP Lingen, 0MV, Statoli, Copetro Oxbow. • Test portion 100±5 g weighed to 0.1 g • Transfer time 45 + 15 sec, tapping runs during feeding • Gravity driven taps using a cam shaft with a step • 1500 taps, each drop 3+0.1 mm • The coke surface is flattened, and the volume recorded to 1 ml • The TBD is the average weight to volume ratio from two determinations, reported to 0.01 g/cm3 The equipment description gives details of dimensions, requiring a 250 ml measuring cylinder of 190±15 g with ±1 ml gradation, a plunger with mass 450±5 g, a tapping device frequency of 250+15 Hz and the 3±0.1 mm gravity drop for the taps. Precision ISO 10236 TBD The test procedure requires two parallel determinations. The r&R statement gives acceptance criteria. The precision statement is based on determinations on the 1-2 mm fraction. Precision at 95% confidence level was determined according to ISO 5725. r = 0.01 g/cm3 R = 0.02 g/cm3 Issues with TBD Overall the TBD method seems simpler than the VBD method, especially the direct screening. But it has some issues. • The r&R values are quite strict limits. R&D Carbon Ltd. was involved with the development of this ISO method and has commented that a critical step in the analysis is sample filling time. The given r&R limits were obtained with strict adherence to the standard requirement of a filling time of 45±15 seconds, as close to 45 seconds as possible. [2] • Filling issues are coke bridging in hopper, delaying transfer, also TBD increases slightly withfillingtime. • Regarding read-out, the adjustment of the coke level with the spatula to an even level before read-out is operator dependent and can shift the result. For a 100 g test portion with TBD 0.83 g/cm3 a 1 ml decrease in read-out volume will increase the TBD 0.007 g/cm.

Figure 5: Example of TBD apparatus. VBD -Comparative Studies In these comparisons, the reported values have been input to ASTM E691 to determine the between-lab reproducibility at 95% confidence. VBD. Comparative Study 1999 Keith Neyrey presented results from four CII Carbon RRs at the 1999 ASTM D02.05D committee meeting.[3] There was one material in each RR, see Table 1. Eight to sixteen laboratories reported results. The observed R values were above the R reproducibility limit of 0.046 g/cm3.

6

The TBD method ISO 10236 is the responsibility of ISO Technical Committee 226, "Materials for the production of primary aluminium".

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Table 4. VBD (g/cm3) average and between-lab R at 95 % confidence level from four CII Carbon RRs.

Table 1: VBD (g/cm3) average and between-lab R at 95 % confidence level from four CII Carbon RRs.(7) RR R#09 R#ll R#12 R#13

Average 0.89 0.84 0.85 0.86

R 0.11 0.07 0.09 0.06

R#17 Fine-28x48 R#17 Coarse-8x14

TBD in the 2008 R&D Carbon Interlaboratory StudvKl Some results from this RR were shown in Table 3. For TBD, sixteen laboratories reported results. TBD was measured on several coke fraction sizes. For the 1-2 mmfraction,the expected TBD range of sponge petroleum coke is 0.78 to 0.86 g/cm3 so the two cokes A and C were at each end of the TBD-range. Using ASTM E691, the TBD between-lab reproducibility R came out as shown in Table 5. The observed R values were above the ISO 10236 reproducibility limit of 0.02 g/cm3. • The increase in TBD with decreasing grain size is clear. • The between-lab precision showed no trend with grain size, so the level of porosity was not a significant variation factor, and, somewhat surprisingly, neither was the number of grains in the test portion.

Table 2: VBD (g/cm3) average and between-lab R at 95 % confidence level in the 2008 RDC RR. Size range Tyler 28x48 Tyler 28x48

Average 0.91 0.86

R 0.09 0.12

Table 5: TBD (g/cm3) average and between-lab R at 95 % confidence level in the 2008 RDC RR.

In the R&D Carbon ILS, VBD was measured on several coke fraction sizes, see Table 3. Table 3: VBD (g/cm3) average and between-lab R at 95 % confidence level in the 2008 RDC RR. Material A Material A Material A Material A Materiale Materiale Materiale Materiale • •

Size range VBD 2-4 mm 0.79 1-2 mm 0.86 0.5-1 mm 0.90 0.25-0.5 mm 0.91 2-4 mm 0.74 1-2 mm 0.81 0.5-1 mm 0.81 0.25-0.5 mm 0.84

Material A Material A Material A Material A Material A Material C Material C Material C Materiale Materiale

R 0.07 0.04 0.05 0.04 0.10 0.10 0.06 0.09

TBD Size range 4-8 mm 0.73 2-4 mm 0.79 1-2 mm 0.86 0.89 0.5-1 mm 0.25-0.5 mm 0.89 4-8 mm 0.69 2-4 mm 0.74 1-2 mm 0.79 0.5-1 mm 0.82 0.25-0.5 mm 0.83

R 0.04 0.03 0.04 0.04 0.04 0.04 0.04 0.05 0.05 0.03

Method Development in ASTM - VBD

The increase in VBD with decreasing grain size is clear. The between-lab precision showed only a slight improvement with smaller grain size, so the level of porosity was not an important variation factor, and, somewhat surprisingly, neither was the number of grains in the test portion.

Method development has been slow partly due to uncertainty how to improve. However, in ASTM two major steps have been achieved in the last years. ASTM D7454-08 VBD A new method for vibrated bulk density has been published, ASTM D7454-2008. In this, the ASTM Committee and Rio Tinto Alcan has addressed what grain size is allowed, and also given a more strict procedural description for the test portion preparation. A new, semi-automated analysis instrument has been introduced. The measurement is mass of a fixed volume rather than the readout of the level of a volumetric cylinder as in D4292 VBD and the TBD methods.

VBD. 2004 CII Carbon RR #17 Γ5Ί This RR was mentioned above with Figure 2. A calcined coke RR sample was distributed prepared to -4 Tyler mesh. For VBD, each lab was to return results for analysis by ASTM D4292 for the two common grain ranges • The finer ASTM VBD Tyler mesh 28x48 (0.30-0.59 mm) • The coarser Kaiser VBD Tyler mesh 8x14 (1.17-2.36 mm) Using ASTM E691, the average VBD and the between-lab reproducibility R came out as shown in Table 4. The observed R values were comparable to the reproducibility limit of 0.046 g/cm3 for the fine material, and above for the coarse. 7

R 0.04 0.08

TBD -Comparative Study

VBD. 2008 R&D Carbon Interlaboratory Study R&D Carbon ran a large interlaboratory study in 2008.[4] For VBD, two materials were distributed, a denser coke A and a more porous coke C. Eleven laboratories reported results. For Tyler mesh 28x48, the VBD between-lab reproducibility came out as shown in Table 2. The observed R values were above the R reproducibility limit of 0.046 g/cm3.

Material A Materiale

Average 0.88 0.82

ASTM D4292-10 VBD A new, 2010 version of this method has been published. It was developed by the ASTM Committee, mostly through the efforts of Bill Spencer.

Data in lb/ft3 was converted to g/cm3 by dividing by 16.0.

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might be an effect of quality laboratories participating in the precision statement development, but also an effect of differences in the actual practice of the standards. For an anode producer, the one way to address and solve this dilemma is to establish the most relevant bulk density method available in their own laboratory, and do all required analyses.

Effect of Differences in Laboratory Practice Examples of non-uniform practice are • Feeding with or without starting vibration • Fixing the measuring cylinder to the vibration table or not • Speed of sample introduction into the measuring cylinder • Different types of crushers for sample preparation • Length of time for vibrating the test portion

In the comparative studies shown the TBD method yielded better between laboratory precision than the old D4292-92 VBD method. The results from the currently running ASTM interlab study for the new D7454-08 and the D4292-10 revision will add useful information to such a comparison. It will also shed light on the precision of the GeoPyc apparatus.

Results of a bulk density test are dependent on the average particle size and particle size distribution. Because of this dependency, sample preparation of the material directly impacts the results. One aspect of measuring bulk density is to examine only "as-calcined" particles or particles that are of a "natural" fraction. The as-calcined or natural particles are particles of calcined petroleum coke that have not been subject to a crushing step. It is appropriate to perform a sieving step to separate the as-calcined or natural particles. If measuring bulk density of ascalcined particles, identify the results with the particle size and the label of as-calcined. Example: VBD result = g/cm3 (20x48 Tyler mesh range (0.30 x 0.85 mm), As-Calcined).

In the short term, other methods are unlikely to appear. For any anode manufacturer using bulk density to evaluate coke potential, developing a better bulk density apparatus for internal use might be an option. A larger test portion could address aggregate packing issues such as • The wall effect and shape of grains • Small 120 cm3 test portion volume versus read-out accuracy • Large test portion surface to volume of the cylinder

Prepared bulk density samples are from particles of calcined petroleum coke that have been crushed in the laboratory. It is critical for good repeatability and reproducibility results to have appropriate sample crushing equipment and procedures. The recent year 2010 revisions to ASTM D4292 (D4292-10) contains additional specifications on equipment and crushing operations. The roll crusher specification in D4292-10 specifies that both rolls must rotate to crush the material. Disc mills, disc type grinders or disc pulverizers are not appropriate since these contain only one stationary roll. D4292-10 also specifies in the crushing operations through the jaw crusher and roll crusher that the entire gross sample should pass through the crushers.

Finally, there is the possibility of a reference material. The authors should like to suggest users of these methods to consider if a reference material could be established, e.g. through NIST, as an aid in tuning the bulk density analysis equipment. Authors' Background in Standard Development Dr.ing. Lorentz Petter Lossius is a voting member of ASTM Committee D02.05 and a Technical Expert in ISO/TC 226 Work Group 2 on Solid Carbon Materials. Dr. Bill Spencer is Chair of ASTM Committee D02.05. Dr. Tech. Harald A. 0ye is Chair of ISO/TC 226 and Convenor of Work Group 2 on Solid Carbon Materials.

Method Development in ISO -VBD The ISO VBD method study was based on screening a natural fraction and mixing it with the crushed and screened oversize fraction. The measurement procedure was similar to D4292 with 100 g coke in a 250 ml volumetric cylinder with even filling over 115 seconds with vibration, followed by vibration for 60 seconds. For read-out, the surface of the vibrated coke was flattened with a spatula and the volume read to the nearest 1 ml. A comparison was run versus the TBD method ISO 10236 using two single source and two blended cokes, and using TBD fractions.[6] The ISO Committee concluded that the precision with the new VBD method was not better than for the TBD method, and the work was discontinued. The laboratory testers observed that the main contribution to the standard deviation was • Leveling the test portion with the spatula • The coarse volumetric cylinder gradation of 1 ml.

References [1] Kirstine L. Hülse, "Anode Manufacture Raw Materials Formulation and Processing Parameters", R&D Carbon Ltd., ISBN 3-9521028-5-7 (2000). [2] Jean-Claude Fisher, private communication. [3] Keith Neyrey, CII Carbon, "Minutes of Meeting from the 1999 ASTM Committee D02.05D". Results reproduced courtesy of Rain CII Carbon, LLC. [4] 2008 R&D Carbon Interlaboratory Study RDC 8843, results reproduced courtesy of R&D Carbon Ltd.

Discussion

[5] Trey Neal, CII Carbon, Result Memo on RR # 17, 2004.

The bulk density has fairly good within-laboratory repeatability, and has a useful role as a coke property if it is based on the reported values from one laboratory. But examples of betweenlaboratory comparisons have shown that the bulk density has severe limitations when comparing results from different laboratories. This is a severe drawback for a commercial standard. Furthermore, from the studies shown it is reasonable to say that the reproducibility limits given in the standards, although they are high, still are optimistic compared to practical experience. This

[6] Turid Vidvei and Lester McCoy, "Memo to ISO/TC 226 Paris Meeting 2006; VBD (New ISO) vs. TBD".

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

IMPROVING THE REPEATABILITY OF COKE BULK DENSITY TESTING Î e s Edwards, 2Marvin Lubin, 2Jim Marino ^ain CII Carbon, 2627 Chestnut Ridge Rd, Kingwood, Texas, 77339, USA 2 Rain CII Carbon, Lake Charles Lab, 1920 Pak Tank Rd, Sulphur, Louisiana, 70665, USA Carbon, Petroleum Coke, Bulk Density one advantage of the test relative to VBD/TBD tests is its ability to measure coke density/porosity independent of particle shape.

Abstract The "Pechiney" mercury apparent density test was used by the aluminium industry for many years to measure the density of calcined petroleum coke. Over the last five years, the industry has moved away from this test for occupational health and safety reasons. The current alternative tests are based on vibrated or tapped bulk density. The value of measuring bulk density is reduced by poor reproducibility due to differences in equipment and sample preparation. This paper reviews different bulk density test methods and presents repeatability data on a new method for measuring bulk density. The method is based on automated equipment which uses transverse axial pressure to measure bulk volume. The new equipment shows improved repeatability compared to existing equipment and can also be used to measure envelope density which eliminates intra-particle porosity problems associated with bulk density tests.

The test uses the Archimedes principal where particle volume is measured by mercury displacement. The 1.70 to 0.85mm (10x20 Tyler mesh) particles used for the test are placed in a pycnometer and mercury is added at atmospheric pressure. The mercury is non-wetting to carbon and is not affected by pore capillary forces. The Hg surrounds the particles uniformly and enters open pores. Based on the pressure exerted by the weight of mercury, it is estimated that pores down to a diameter of ~13um [2] are penetrated by mercury. This is approximately the pore size pitch penetrates when making anodes so the Hg AD test provides a useful analog. VBD and TBD tests are capable of giving the same type of information but particle shape is a complicating factor. All VBD and TBD tests rely on measuring the bulk volume of a packed bed of coke particles. The bulk density of the coke bed is therefore always a function of the density of individual coke particles and the packing density of the particles.

Introduction The mercury apparent density (Hg AD) test was the test preferred by many anode producers for predicting the pitch demand and anode density potential of different calcined cokes. The procedure was developed by Aluminium Pechiney and uses relatively large volumes of mercury. This makes it hazardous from an occupational health and safety standpoint and the test has now been largely abandoned by the industry.

Vitchus et al [3] presented a good summary of the importance of particle packing and the impact of parameters such as particle size, shape and roughness on packing density and bulk density. Porosity or void space between particles is referred to as intraparticle porosity. For cokes crushed to a particle size of 0.6 to 0.3mm (28 x 48 Tyler mesh), intra-particle porosities of 42-50% are reported for two different cokes crushed in different ways. To complete the picture, 41-46% of the volume space is occupied by solid coke and 8-13% by porosity within the coke particles.

Although the Hg AD test has been widely used, many other tests are also commonly used for measuring coke density/porosity. These tests are based on vibrated bulk density (VBD) or tapped bulk density (TBD). Most have been around for a long time and many papers have been published on the value of these tests. At least one aluminum producer [1] routinely uses a VBD test for determining pitch demand during anode production

The above highlights one of the challenges of using VBD or TBD as a reliable indicator of coke porosity and anode pitch demand and density. Sample preparation and anything that has the potential to affect particle shape and packing density will always have a significant impact on the final bulk density result.

Rain CII has the capability to run four different bulk density tests: three VBD tests and one TBD test. The repeatability and reproducibility of these tests varies widely. This paper will present a brief review of the different test methods and will then focus on work done to quantify and improve the repeatability of the measurement portion of the tests. Results with an automated bulk density analyzer will be presented.

VBD/TBD Procedure Review There are three well known VBD and TBD procedures in common use: ASTM D4292 (VBD), ASTM D7454 (VBD) and ISO 10236 (TBD). Without the Hg AD test, most anode producers use one of these tests when setting coke bulk density specifications with suppliers and/or when trying to correlate coke properties with pitch demand and anode density.

This paper will not make recommendations about the merits of using any particular VBD or TBD test to monitor coke quality or predict pitch demand and anode density. It will however, look at correlations between the results of the various VBD tests, the TBD test and the Hg AD test across a range of different cokes.

Rain CII conducts all these tests and one additional test known as the "Kaiser" VBD test. It was developed many years ago by Kaiser Aluminum and is used to monitor daily calcined coke quality. The test is a little quicker and easier to run than the ASTM VBD tests but it is not a certified standard test procedure.

He AD Test versus VBD/TBD Tests There are different opinions about whether the Hg AD test is a reliable indicator of pitch demand and anode density but the reality is that the Hg AD test will not be around much longer. The

947

The following provides a brief summary of the sample preparation steps required for each of these tests. A more detailed description is provided in a parallel paper in these proceedings [4]. The published standards for each test also contain much more detail and should be read before any of the tests are undertaken.

The rest of this paper will focus on the measurement portion of each test. Differences in sample preparation will always have a significant effect on the final VBD result and changes in the measurement method will not overcome this problem. Unless the sample preparation steps are followed exactly as stated in the published standards, different results can be expected between labs analyzing the same samples. This goes back to the problem of particle shape, surface roughness, etc.

ASTM D4292 (0.6 - 0.3mm or 28x48 Tyler mesh) The sample is jaw-crushed first and then fed through a roll crusher three times at successively smaller gap settings to produce material in the 28x48 mesh size range. 100 grams of the 28x48 sample is used for the VBD measurement. The test can also be used for different size ranges such as 1.18-2.36mm or 3.3-6.7mm.

In contrast to the VBD and Hg AD tests, the TBD test involves no sample preparation other than screening. This eliminates a source of error that occurs in the all the other preparation methods.

ASTM D7454 (0.85 - 0.425mm or 20x35 Tyler mesh) The sample is hand screened to remove coke in the 4.75 - 1.18 mm (4x14 mesh) size range. The 4x14 fraction is then fed to a roll crusher three times at successively smaller gap settings to produce coke in the size range of 20x35 mesh. The 20x35 mesh sample is run in duplicate on semi-automated equipment to measure VBD.

VBD/TBD Measurement Procedures The three VBD tests all use the same principal where sample is added to a graduated cylinder and vibrated for a specified time. The TBD measurement is similar but instead of being vibrated, the graduated cylinder is tapped. Photographs of equipment used for the D4292 VBD and ISO TBD test are shown in Figure 1.

Kaiser VBD (2.36 -1.18 mm or 8x14 Tyler mesh) The sample is hand-screened at 8x14 mesh. The +8 mesh fraction is jaw crushed once and then roll-crushed repeatedly until it is all less than 8 mesh. All the 8x14 mesh fractions are re-combined for the VBD analysis which is run on a 100 gram sample.

For the ASTM VBD, Kaiser VBD and TBD tests, a fixed weight of sample is added to the graduated cylinder and the height of the coke bed is read (by eye) at the end of the test to estimate the volume. The VBD or TBD is calculated simply as mass/volume.

ISO 10236 TBD The ISO TBD test is the simplest and easiest of all the procedures since it involves no crushing. The sample is screened at the following screen sizes: 8.0 - 4.0mm, 4.0 - 2.0 mm, 2.0 - 1.0 mm (9x16 Tyler mesh), 1.0 - 0.5 mm and 0.5 - 0.25 mm. 100 grams of any one or all of thesefractionsis then measured for TBD. The purpose of the above is to illustrate that all of the preparation methods are different. The particle size ranges used for tests are different and this has a large impact on the bulk density result. For a given coke sample, the ASTM D4292 test at 28x48 mesh always gives a higher bulk density than the ASTM D7454 test at 20x35 mesh or the Kaiser VBD test at 8x14 mesh. Porosity is removed by crushing and the finer particles pack to a higher bulk density.

Figure 1: ASTM D4292 and ISO TBD tests ASTM D7454 is the most recently published VBD procedure (Oct 2008) and it requires the use of semi-automated equipment, Figure 2. The coke is added to the graduated cylinder with a specially designed vibrating bowl and feeder system. A photoelectric sensor detects the coke bed height when it reaches the pre-set, 50 ml level. The feeder shuts off and the coke in the graduated cylinder is then weighed and the bulk density calculated.

An interesting comparison between the tests is the yield in the size fraction measured. Typical yield data from Rain CITs lab are shown in Table I. For some tests, a relatively small portion of coke ends up being measured for VBD or TBD. For the TBD test, some labs measure TBD on all size fractions whereas others measure only one fraction, typically the 1-2 mm fraction. The downside of a lower yield, is that it may not reflect the average coke density or porosity as well. This is the primary reason Rain CII uses the Kaiser VBD test - it gives a relatively high yield. Test Method ASTM D4292 (28x48 mesh) ASTM D7454 (20x35 mesh) Kaiser VBD (8x14 mesh) TBD Overall (0.25-8mm) TBD (0.25-0.5mm) TBD (0.5-1.0mm) TBD(1.0-2.0mm) TBD (2.0-4.0mm) TBD (4.0-8.0mm) Mercury AD (10x20 mesh)

Typical Yield 1

(%) 30 15 55 75 12 12 15 15 20 45

Figure 2: ASTM 7454 semi-automated VBD equipment Not surprisingly, the precision of the measurement in ASTM D7454 is better than the other two VBD methods and the TBD method (3-sigma is about half). Weighing the sample instead of estimating the volume by eye is more precise. The coke feeding and table vibration are also controlled very precisely

Table I: Yield of coke in measuring size range after preparation

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first part, the repeatabilities of the various tests are evaluated and compared to the GeoPyc using quality control (QC) standards from Rain O F s lab. In the second part, a wide range of different calcined coke samples are evaluated by the various VBD/TBD methods and compared to Hg AD results. In the final part, the GeoPyc envelope density test is evaluated.

Another difference with the ASTM D7454 test is the requirement to use calibration standards. The test was developed by Alcan and the standards are now available through Rio Tinto Alcan (RTA). The equipment must be calibrated before first use and then checked routinely thereafter (typically once/week). There are several adjustments that can be made for recalibration and these are all well documented in the procedure.

Repeatability of Test Methods

Micromertics GeoPyc Equipment

The repeatability data generated in the following section was produced by running QC samples 25 times (once/day) on each piece of equipment. Each test has its own QC sample due to differences in particle size. The QC tests were run on the standard equipment and then on the GeoPyc. The GeoPyc is fully programmable and several test parameters can be varied including cylinder diameter and volume, force and number of consolidation cycles. A standard set of conditions was used for all tests.

The Micromeritics GeoPyc 1360 Envelope Density Analyzer has been available since 1995. It is used to measure particle and solids bulk densities. Rain CII first tested the equipment in 1996 as a potential replacement for the Hg AD test. Initially, the instrument was available only as an envelope density analyzer. For this test, a special dry flow media (DryFlo) is used which comprises small, spherical particles that are non-wetting and free-flowing. The DryFlo media behaves like a fluid and when mixed with solid particles, it surrounds the particles eliminating intra-particle voids. This negates particle shape problems and their impact on packing density. A photograph of the equipment is shown in Figure 3.

An example of the frequency histograms for repeatability of the ASTM D4292 test is shown in Figure 4. This shows the range of bulk densities measured with the QC standard on the standard equipment and the GeoPyc equipment.

Figure 3: GeoPyc 1360 Density analyzer

Figure 4: Frequency histograms for ASTM D4292 test

Rain CII did not find a good correlation between envelope density and Hg AD for samples prepared by the Pechiney method in 1996. The test repeatability was not particularly good so no further testing was done. Micromeritics later developed a bulk density option for the equipment referred to as the T.A.P option (Trans Axial Pressure). It measures the bulk density in a horizontal cylinder as shown in Figure 3. This option was not available when the equipment was tested by Rain CII in 1996.

Results for all tests are summarized in Table II. The first point to note is that, for the three VBD tests, the GeoPyc bulk densities were very similar to those measured on the standard equipment. This was a little surprising given the obvious differences in the mode of analysis (different vibration table amplitude settings, methods of clamping, feeding of samples, etc.). Test

With the abandonment of the Hg AD test, Rain CII decided to reevaluate the GeoPyc equipment, particularly the T.A.P option. In principal, it provides a very accurate way of measuring the bulk density by controlling the force and measuring the displacement of a Teflon plunger used to compact the bed of coke.

ASTM D4292 (28x48 mesh) ASTMD7454 (20x35 mesh) Kaiser VBD (8x14 mesh)

The plunger applies the compaction force in stages and the sample is agitated by rotating the cylinder during measurement. After each compaction cycle (referred to as a consolidation cycle), the plunger retracts a short distance and then moves forward again until the desired force is reached. Multiple consolidation cycles allow the particles in the coke bed to re-arrange to ensure maximum packing density without particle breakage. When the cycles are complete, the instrument uses stored data to calculate plunger displacement and hence bulk volume. The sample weight is input via the keypad and bulk density is calculated.

ISO 10236 (9x16 mesh)

Mean 3-Sigma Mean 3-Sigma Mean 3-Sigma Mean 3-Sigma

Standard Equipment 0.904 0.019 0.870 0.010 0.819 0.022 0.874 0.023

GeoPyc Equipment 0.899 0.007 0.875 0.009 0.825 0.006 0.882 0.016

Difference in Means 0.005 0.005 0.006 0.012

Table II: Bulk density results with QC standards (g/cc) The other point to note is that the repeatability of the bulk density measurement with the GeoPyc equipment is significantly better (lower 3-sigma) than the standard equipment for the ASTM D4292 and Kaiser VBD tests. This was not unexpected because both tests rely on manual reading of the coke bed height in a cylinder. For the ASTM D7454 test, the repeatability of the GeoPyc was about the same as the semi-automated equipment.

The rest of this paper reports on the results of bulk density testing with the various different types of VBD/TBD equipment. In the

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For the TBD test, the repeatability was better for the GeoPyc but the 3-sigma was higher than all the VBD tests.

completed yet. Rain CII typically blends cokes to meet specifications.

For the ASTM D7454 test, one of RTA's certified reference standards was used (0.870 g/cc standard). One of the benefits with the GeoPyc is that no standard is required and no form of calibration is required. The instrument does its own zero calibration with the Teflon plunger when the glass cylinder is empty. Another benefit of the GeoPyc is that the sample can simply be poured into the glass cylinder. For all the other bulk density tests, it is critical to feed the coke into the graduated cylinder at a constant, well-controlled rate.

Sulfur, vanadium, real density and Hg AD results for the straight run cokes are shown in Table IV. The Hg AD tests were run by an external laboratory since Rain CII no longer runs this test in the USA. Coke I is a highly isotropie coke and coke J is a shot coke.

Coke Coke A CokeB CokeC CokeD Coke E Coke F Coke G Coke H Coke I CokeJ CokeK CokeL

For the TBD test, repeatability testing was done only with a 9x16 QC standard. Rain CII does not have any customer specifications based on this test but the 9x16 fraction is used by some in the industry. The TBD test results differed from those of the GeoPyc by the greatest amount. The repeatability of the standard equipment was about the same as the other two VBD tests, which makes sense given that all rely on manual reading of a graduated cylinder. The reason for the poorer repeatability of this test with the GeoPyc relative to the other tests has not been fully established but is likely due to the lack of any form of sample preparation (crushing). Naturally occurring particles may be more irregular in shape and/or more susceptible to breakage during the measurement itself. Crushing removes some of this variation.

ASTM D4292

# Samples

62

Mean Difference in Means

0.868

Geo.

ASTM D7454 Std.

Geo.

0.864

0.839

Coke A B C D E F G H 1 J K L

Geo.

80

33

0.004

Kaiser VBD Std.

0.846

-0.007

0.772

0.763

0.008

Max

0.909

0.895

0.867

0.870

0.840

0.837

Min

0.806

0.812

0.810

0.820

0.685

0.699

Range

0.103

0.083

0.057

0.050

0.155

0.138

Hg AD 1 (g/cc) 1.73 1.71 1.75 1.74 1.75 1.79 1.75 1.74 1.76 1.79 1.82 1.83

The bulk densities of all the cokes above were measured by the four tests already described. Standard preparation procedures were followed and bulk densities were measured using both the standard equipment and the GeoPyc equipment. This generated a large amount of data, some of which is shown in Table V. All the VBD and TBD measurements were performed with both standard equipment and GeoPyc equipment, but only GeoPyc results are shown. Standard equipment results are shown in Table VI.

A wide range of samples with different bulk densities were analyzed using the standard equipment (Std.) and the GeoPyc (Geo.) equipment. The results are summarized in Table III. The TBD test was not included in this evaluation. Most of the samples were production or shipment samples containing multiple cokes.

Std.

Real Density (g/cc) 2.088 2.079 2.055 2.071 2.078 2.044 2.060 2.076 1.988 2.009 2.041 2.083

Table IV: Straight run coke samples

Comparison of Results Across a Wide Range of Samples

Test

Sulfur % 4.00 1.52 2.72 1.45 1.48 3.72 2.22 0.78 4.66 4.54 2.85 2.49

Vanadium ppm 740 95 320 110 115 490 290 270 620 1750 190 170

Table III: Bulk density results across wide range of samples (g/cc) There are more Kaiser VBD results in Table III because Rain CII runs this test on all daily production samples. The differences in the means are all statistically significant but are well below the stated repeatability and reproducibility limits for the various tests. For example, the published repeatability and reproducibility limits for the ASTM D4292 test are 0.014 g/cc and 0.046g/cc respectively. For all tests, the range between maximum and minimum values was lower for the GeoPyc equipment which is consistent with lower variability in the measurement method.

ASTM D4292 0.920 0.848 0.882 0.858 0.889 0.839 0.923 0.875 1.014 0.983 0.943 0.972

ASTM D7454 0.914 0.797 0.862 0.821 0.821 0.832 0.911 0.847 0.960 1.019 0.838 0.889

Kaiser 0.841 0.726 0.784 0.754 0.756 0.754 0.836 0.776 0.913 1.066 0.888 0.859

TBD 5x9 0.806 0.661 0.759 0.707 0.708 0.706 0.843 0.722 0.919 0.996 0.808 0.814

TBD 9x16 0.879 0.724 0.823 0.766 0.769 0.727 0.927 0.784 0.973 1.128 0.900 0.892

TBD 16x32 0.893 0.791 0.887 0.803 0.831 0.749 0.960 0.837 1.025 0.999 0.986 0.973

Table V: VBD and TBD results with GeoPyc equipment (g/cc) The data were analyzed by Minitab to check correlations. All correlations between the Hg AD results from Table IV (Hg AD 1) and the various VBD/TBD preparations were poor with R2 values <0.3. The correlations between most of the VBD/TBD preparations were much better. The best correlation (R2 = 0.95) was between the Kaiser GeoPyc bulk density and the GeoPyc bulk density on 9x16 particles prepared via the TBD test.

Next, a wide range of samples with different bulk densities and other properties were selected. All initial testing was done on straight-run cokes generated during full-scale, calcination trials. Testing is still underway on coke blends and has not been

The lack of correlation between any of the above bulk density measures and the Hg AD results was a concern, so a slightly smaller subset of the cokes (A-J) were sent out to several different

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laboratories for Hg AD analysis. A combination of Pechiney Hg AD equipment and mercury intrusion porosimeters (MIP) were used for this round of testing. MIP is a well established method that allows accurate characterization of pore size distributions with Hg up to pressures as high as -400 MPa.

methods and sizes from 28x48 mesh in the ASTM D4292 test to 5x9 mesh sizes (4.0 - 2.0 mm) in the ISO TBD test.

iS :

Most of the correlations improved, suggesting that the Hg AD 1 results reported in Table IV were unreliable. This highlights one of the problems with the Hg AD test - it has relatively poor reproducibility. In past industry round robins, Rain CII has reported reproducibilities of-0.030 g/cc. The new Hg AD results along with the standard equipment results for all the other tests (non-GeoPyc) are shown in Table VI.

0.859

Hg AD 2 1.73

Hg AD 3 1.73

1.75

Hg AD 5 1.64

Coke

D4292

D7454

Kaiser

TBD 9x16

A

0.909

0.908

0.844

Hg AD 4

B

0.847

0.792

0.687

0.704

1.70

1.71

1.71

1.53

C

0.901

0.868

0.781

0.800

1.73

1.75

1.77

1.64

E

0.887

0.819

0.771

0.741

1.70

1.70

1.76

1.62

F

0.847

0.823

0.761

0.714

1.67

1.66

1.73

1.59

G

0.926

0.907

0.826

0.944

1.77

1.78

1.82

1.72

H

0.885

0.869

0.769

0.788

1.75

1.74

1.74

1.63

1

1.031

0.954

0.926

0.962

1.76

1.76

1.77

1.70

J

0.990

1.032

1.048

1.106

1.79

1.79

1.79

1.74

I^LXpi:

Figure 6: SEM image of DryFlo media and mixture of DryFlo and calcined coke in GeoPyc instrument When using the DryFlo media, -25% of the sample (by volume) must be mixed with -75% of the DryFlo media. The DryFlo surrounds or "envelopes" each particle. When the test is complete, a screen is used to separate the DryFlo mediafromthe sample and the separation is very clean due to the non-wetting nature of the DryFlo. It can be reused multiple times. Results for envelope densities are shown in Table VII and they were significantly better than expected. Correlations with all tests were generally good. A correlation matrix for all the GeoPyc bulk density tests and the Hg AD results is shown in Table VIII.

The best correlations were observed with the Hg AD 5 results. These were Hg AD's run at a lower pressure with a MIP. At the pressure used, it was estimated that pores down to a size of -ΙΟΟμιη were penetrated instead of the 13 μιη pores in the regular Pechiney Hg AD procedure. Figure 5 shows correlations for the D7454 and 9x16 TBD tests with the Hg AD 5 results.

1.60 1.65 HgAD5

Real Density (g/cc)

A

1.418

2.088

32.1

B

1.256

2.079

39.6

C

1.386

2.055

32.6

D

1.332

2.071

35.7

E

1.344

2.078

35.3

F

1.279

2.044

37.4

G

1.461

2.060

29.1

R2=0.85

1.70

1.75

Porosity %

H

1.423

2.076

31.4

1

1.530

1.988

23.1

§0.80

J

1.610

2.009

19.9

K

1.437

2.041

29.6

L

1.462

2.083

29.8

Ë

1.55

Envelope Density (g/cc)

f?0.90

Q1.00

1.50

Coke

9x16TBDvsHgAD5 1.10

0.70

^

0.60 1.50

1.55

1.60

1.65

HgAD5

f-

MM mm

Table VI: VBD/TBD and Hg AD results (g/cc)

D7454vsHgAD5

:

*' νϊί

1.70

Table VII: Envelope densities and calculated porosities

Figure 5: Correlations for ASTM D7454 and TBD vs Hg AD 5 Kaiser OEO

These results make intuitive sense. Bulk density tests include all forms of open and closed porosity in the coke as well as intraparticle porosity. The Hg AD test, on the other hand, excludes open porosity >13 μηι and all intra-particle porosity. An MIP test run at lower pressure where Hg penetrates pores down to 100 μιη should show a better correlation to bulk density test results.

ASTM GEO D7464 GEO ISO 9X16 OEO

Envelope Density Testing In the final phase of this work, the GeoPyc equipment was tested in the envelope density mode. The recommended particle size of the sample when measuring envelope density is >2mm. This is due to the relatively large particle size of the DryFlo media (100150 μηι). A scanning electron microscope (SEM) image of the DryFlo media is shown in Figure 6 along with a photograph of the coke and DryFlo mixture in the GeoPyc. For this work, it was decided to compare results across all the various preparation

ASTM GEO

ALCAN GEO

9X16 Geo

Δ

AD2

A

"T5" AD4

ADS

0.65

0.95

0.76

HgAD1

0.25

0.24

0.06

0.16

HBAD2

0.53

0.50

0.55

0.65

0.02

HgAD3

0.46

0.45

0.48

0.61

0.00

0.91

HgAD4

0.34

0.38

0.44

0.51

0.14

0.54

HgAOS I Envelope Density

0.69

0.69

0.79

0.81

0.26

0.75

0.64

0.81

0.85

0.80

0.83

0.91

0.16

0.74

0.64

0.51

0.87

I

0.88

0.79

0.84

0.92

0.18

0.68

0.58

0.48

0.85

0.88

2

0.58

Table Vili: R correlations between different density measurements for cokes A - L

951

Den.

0.72 0.81

Porosity

Ënvel l

0-96 j

Among the Hg AD tests results, the envelope densities correlated very well with those of the Hg AD 5 test with an R2 of 0.87. The correlation with the Hg AD 2 and the Hg AD 3 results were 0.74 and 0.64 respectively. The correlation with the Hg AD 4 results was not so good at 0.5, and there was no correlation with the Hg AD 1 results (0.16), confirming the poor quality of this data.

blended cokes are used, and this is probably not surprising given the different densities and hardness of cokes used in blends. The GeoPyc envelope density analysis shows great promise. When the equipment is used in this mode, it eliminates the effects of particle shape and roughness that contribute to packing differences in traditional VBD and TBD tests. Preliminary repeatability testing on 8x14 mesh samples shows a 3-sigma of -0.030 g/cc. Controlling the volume ratio of coke to DryFlo media within a narrow range (25-30%) appears to be important for achieving good repeatability.

Among the VBD and TBD tests, R2 values with the envelope densities were generally excellent except the TBD test with 32x60 mesh particles. This is not unexpected given the fine particle size of thisfraction.For the Kaiser VBD and coarser particle size TBD results, the correlations were all >85% which makes sense given the coarser sizing relative to the DryFlo sizing.

The envelope density test is simple and avoids the occupational health problems of mercury. It may not give results fully comparable to the Hg AD test, but it could offer an improvement over traditional bulk density tests as a predictor of pitch demand and anode density. It will not give the sort of detailed pore size distribution data available with an MIP, but MIP equipment is expensive and costly to operate and is more of a research tool than a routine analysis tool. It still uses mercury, of course.

Real density data can be entered into the GeoPyc and it calculates particle porosity as shown in Table VII. This includes open porosity not penetrated by the DryFlo media which is likely to be most of the open porosity in coke. MIP data for 3 of the cokes (B, G and J) are shown below. The cokes represent high, medium and low porosity cokes, and Figure 7 shows pore size distribution curves at a pressure of ~400Mpa of pressure. The data show that most of the differences in density and porosity are due to differences in the macropores >1 μπι in size.

No attempt was made in this paper to quantify the effects of sample preparation differences on bulk density results. All samples were prepared as consistently as possible in one lab. The same samples prepared in other labs could give different bulk density results due to different preparation methods. The GeoPyc envelope density analysis has the potential to significantly mitigate sample preparation issues that contribute so much to differences in packing density. This is arguably the most compelling reason to investigate this analysis in more detail. The ISO TBD test avoids sample preparation differences between laboratories, but correlations between TBD and Hg AD results appear to decrease significantly for blended cokes. This is still being investigated but is likely due to the lack of crushing in the TBD method. All other methods, including the Hg AD test, require crushing during sample preparation. Anode plants also crush coke before use.

Figure 7: Hg porosimetry plots of cokes B, G and J Discussion and Conclusions

No recommendations are made on the merits of the various bulk and apparent density tests as predictors of pitch demand and anode density. The Micromeritics GeoPyc 1360 instrument can be used for either envelope or T.A.P density analysis and it offers some significant advantages over other types of equipment being used today. Rain CII will continue to evaluate the equipment and will look at comparisons between Hg AD, MIP and envelope densities for coke blends rather than straight-run cokes.

The Micromertics GeoPyc 1360 appears to be a versatile piece of equipment for measuring coke bulk densities when used with the T.A.P option. It does not require calibration standards and has a higher level of precision than the graduated cylinder and vibrating/tapping equipment used in the ASTM D4292, Kaiser VBD and ISO TBD tests. The precision is comparable to the semi-automated equipment used in the ASTM D7454 test. The results presented in the paper show that the mercury apparent density test does not correlate well with some of the common VBD tests. The Hg AD test itself is subject to sample preparation and measurement errors just like the other tests. The Pechiney procedure quotes a repeatability of 0.011 g/cc for samples run in the same lab. Reproducibilities between labs of 0.03 g/cc were more typical of Rain CII's experience with the test. At least one set of data analyzed by a lab in this study was unreliable (Hg AD 1). The other Hg AD results (AD 2 - AD 5) showed better correlations with a number of the VBD and TBD test results.

References 1. 2. 3. 4.

It is important to note that the correlations between Hg AD and bulk density reported in the paper were limited to straight run cokes. Some work has been completed with coke blends but further work is in progress. Correlations between several of the VBD and TBD tests appear to deteriorate significantly when

952

Andre Proulx, "Optimum Binder Content For Prebaked Anodes," Light Metals, 1993, 657-661 R. Barrai, "Coke Apparent Density Mercury Pycnometry," Aluminium Pechiney Standard Procedure, 1999,A.07.11.V06 Bernie Vitchus, Frank Cannova, Randall Bowers and Shridas Ningileri, "Understanding the Calcined Coke Bulk Density Paradox," Light Metals, 2008, 871-873 Lorentz Petter Lossius, Bill Spencer and Harald A. 0ye, "Bulk Density - Overview of ASTM and ISO Methods with Examples of Between Laboratory Comparisons," Light Metals, 2011

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ASTM D7454 VIBRATED BULK DENSITY METHOD - PRINCIPLES AND LIMITATIONS Francois Laplante and Luc Duchesneau Rio Tinto Alcan; 1955 Mellon Boulevard; Jonquiθre, Quιbec, Canada Keywords: Vibrated bulk density, VBD, Coke specification Abstract

Experimental

This method, proposed by Rio Tinto Alcan and approved by ASTM in 2008, differs from method D4292 by the introduction of a semi-automated equipment and also by referring to a tighter sample preparation procedure. The performance expressed in terms of repeatability is < 0.008 g/mL when the preparation variance is not included and 0.01 g/mL when the preparation variance is included. The intra-laboratory reproducibility over a two-year period came out to 0.02 g/mL. The inter-laboratory reproducibility has not yet been systematically determined but appears to be high, considering the large punctual differences observed between coke providers and coke purchasers. The underlying principles of D7454 will be presented, the factors causing differences between laboratories will be discussed and a mitigation strategy will be proposed.

Automated Apparatus for VBD Measurement The vibrated bulk density semi-automated apparatus (VBD) is presented in Figure 1. The equipment was jointly developed by Rio Tinto Alcan and STAS; a company specializing in the development, fabrication and commercialization of new equipment for aluminum industry. The instrument consists of a control panel, a vibrating bowl, and a graduated cylinder equipped with a photo detector fixed to an electromagnetic jogger. The control panel allows the adjustment of the vibration set point of the electromagnetic jogger and the control of the material flow passing through the vibrating bowl. The complete details surrounding the instrument set-up and calibration, are provided in references [4-6]

Introduction The vibrated bulk density (VBD) is known as an important technique for calcined coke characterization. The VBD value influences the binder control equation, [1] which itself influences the amount of pitch required during the production of prebaked anode [2]. Concurrently, the VBD is used to predict the green apparent density, and to some extent, the baked anode density [2, 3]. In order to achieve good anode quality, the cokes with the highest purity and the highest vibrated bulk densities are wanted. However, reported coke bulk density is sensitive to sample preparation and the results vary according to the laboratory where the measure is taken. In addition, the degradation of coke quality exacerbates the difficulty to compare results between the coke providers and the coke purchasers. In the past, Rio Tinto Alcan established the minimum specification for coke quality at 0.82 g/cc. Below this value, the coke requires particular attention. Figure 1: Automated Apparatus for VBD Measurements

To remedy the situation when coke is close to the lower specification limit, the VBD technique must be improved to optimize the reproducibility value. The apparatus was designed to achieve this objective and was presented for the first time in 1997 by Duchesneau [4]. The main differences of the Alcan method (ASTM D7454) with respect to ASTM D4292 is i) the use of a photo detector which stops the analysis at a given volume of coke and ii) a tighter sample preparation procedure. Such an approach minimizes the external sources of error and promotes the interlaboratory reproducibility.

Sample Preparation According to ASTM D7454 method, 1 kg of the natural coke sample was collected and sieved on a -4 x 14 Tyler mesh sieve. Then, a portion of 180-200 g of -4 x 14 was collected and further reduced with a roller crusher to -20 x 35 Tyler mesh. A fraction of 100 g was collected for the VBD measurement. Sample Analysis

The present paper outlines the performance and the limitation of ASTM method D7454 using information and data extracted from different round robins held at different times over the past ten years and emphazing the influence of the coke preparation on the VBD measurements.

The -20 x 35 fraction was analyzed with the automated apparatus for the VBD measurement presented in Figure 1. The coke sample was placed in the upper funnel and the VBD apparatus was started. When the coke level reached half of the cylinder level, the apparatus was stopped and the coke remaining in the cylinder

953

was returned in the upper funnel. The purpose of this procedure was to assure a constant bed in the vibrator bowl.

y

0.870

Once the coke level was uniform in the upper funnel, the chronometer was reset and the analysis was started. The cylinder filling had to be completed in 150 ± 15 sec. When the coke level reached 50 mL, the photo detector stopped the feeding. The cylinder was removed and then weighted to the nearest 0.01 g. For each sample, two additional readings were achieved. The vibrated bulk density was calculated based on average weight of coke over the calibrated volume of the cylinder.

u "δ 0.850

y

/y

ta 0.830

o

7454 4292 [7]

Reproducibility (g/cc)

0.012

0.020

0.014

0.046

m Labi A

1 *%U O

A

] / 0.810

0.830

~

Laoz Lab 3

0.850

0.870

Figure 2: Influence of the sample preparation method during VBD measurement (2007 Mini-Round Robin Results) In an attempt to establish the cause of this scattering, all the prepared portions were sent to the reference laboratory to be analyzed by the same instrument and thus allow comparison between the laboratories for prepared samples. Figure 3 shows that the results grew closer when the variance coming from the preparation was taken out, indicating at the same time that the difference between the results is mainly due to the preparation step. i

*- y

0.870 ι

s

\

'""^֮

Ë^

JO

J3 Γ

Table II: Performance of ASTM D7454 Method to Determine VBD Repeatability (g/cc)

s'

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VBD Reference Laboratory (g/cc)

The performance of the method, evaluated and expressed as the instrumental repeatability (a 10 x 35 mesh prepared portion analyzed 12 times in a very short period of time), the whole analytical repeatability (8 representative original portion of a same coke prepared and analyzed within one working day) and the intra-laboratory reproducibility (drift standard analyzed over a three-year period), are summarized in Table II.

Instrumental Repeatability (g/cc) 0.008

*jf

s'

s''

0.790 0.790

The ASTM committee had since tried to find a way to obtain a better conformance to the stated analytical parameters. RTA then proposed its method, which appeared to be a bit more directive and offered the advantage of a semi-automated measurement. Discussions and evaluation lasted for many years until, in 2008, RTA method was finally approved as ASTM method D7454.

Method

s

y

> Table I shows the results of a survey held in 2001 as part of a round robin organized by ASTM Subcommittee D02.05.0D on Petroleum Coke. This highlights the large diversity in the way the VBD methods were performed by the participating laboratories, despite the fact that many of them reported following the ASTM method D4292. It was obviously not surprising to observe differences between the results obtained.

s'

yy

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CΫ

Results and Discussion

s s£

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0.830

9 o·*«

yy.

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1 ah 3

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0.790 0.790

Φ.81Φ

0.830

0.850

0.870

VBD Reference Laboratory (g/cc) Figure 3: Samples analysis from three labs for a unique preparation made by the reference laboratory (2007 VBD Mini-Round Robin Results)

The inter-laboratory reproducibility has not yet been systematically determined, the ASTM method D7454 has not yet been implemented in many laboratories external to RTA. However, punctual inter-laboratory comparisons held in the recent past years permitted to appreciate the achievable agreement.

To investigate the impact of the preparation step , particle size distribution of all 20 x 35 portions, prepared by each laboratory and used to get the results showed in Figure 2, was determined using Ro-Tap1 technology. The results presented in Figure 4 show good reproducibility of the preparation process within the laboratories.

In 2007, an inter-laboratory comparison, involving four laboratories and three samples, showed (Figure 2) a large scattering in the results.

[

954

Ro-Tap is a registered trademark of Haver Tyler, Ine

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Original portion 1 Kg Extract 4X14
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progressiveley decreased (4,1-5, 1.0,0.5 mm

divided in 4 portions each sieved Ro-Tap 5 min. and recombined

grind 28 X 48 roller crusher; gap progressiveley decreased {0.25, 0.07,0.02 inch

750 g

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100 g

200 g

100 g

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150 g

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grind 28 X 48 roller crusher; 2 steps

grind 300-500 g Jaw crusher, to <0\25 inch

grind 750 g Chipmunk; gap 1.5 mm (4 mesh)

grind 28 X 48 roller crusher; gap 1.5 mm

grind 600 g Jaw crusher; gap 1.5 mm (4 mesh)

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original crushed to 20 X 48 mesh

original crushed to 28 X 48 mesh

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Original AST«

Preparation

Instrumental Setup

Table I: 2001 Survey of experimental conditions used for VBD Measurement

250 mL

250 mL

250 mL

250 mL

graduated cylinder

certified calib.

yes; water

yes; water

yes; v

yes; water

Syntron jogger calibrated 020-0.22 mm (peak) SO sec during filling

Laooretta 24; S

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vibrating spatula; 70-110 sec

vibrating bowl

modified 500 mL buret 85-100 sec

vibrating spatula; 10 rnJJ30 sec

funnel; >90 sec

manuaify during filling

Syntron jogger calibrated manually; 80-90 sec

with standards

150 sec during fiiSng

Syntron jog ger calib rated

Syntron jogger calibrated 0.20-0.22 mm 5 min

Syntron jogger calibrated 0.37-0.387? 5 min

Syntron jogger calibrated looking for plateau... 5 min

Syntron jogger calibrated 5.7 Mils

60 Hz, 5 min

graphite load screening machine 30 min manually

sec

Syntron jogger calibrated 0.20 mm (peak) 90 sec during filling

vibrating spatula; 90 +-15 sec

Sy nfcron Jo gger calibrated 0.20*0,22 m m (pea k) at satt?

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yes; water

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cylinder

yea ; water

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Sample Preparation

yes

Vibration monitoring

850

During the round robin in 2009, the coke samples were analyzed by 14 laboratories. Among those labs, they were RTA labs and other laboratories. The bulk density was measured according to the technology available in each participant laboratory. Among all the technique used, several laboratories were performing bulk density measurement using the ASTM D4292 method and others were using the ISO 10236 method. Inside RTA, the D7454 method was preferred but two laboratories are still using a modified protocol for the D7454 and one laboratory is using the old Alcan 883 method [4]. The overall results of the 2009 round robin are gathered in Figures 6 and 7. Each graph shows two bullets, indicating that each sample was analyzed in duplicate. The results obtained by each of the laboratories grouped as per the method being applied; thus permitting to appreciate within and between laboratories, as well as within and between methods performance. The error bar for each point represents the intralaboratory reproducibility standard deviation (Table II).

<425 600 425 Sieve (micron) Figure 4: Particle Size Distribution of all the 20 x 35 Mesh Portions of the 2007 Mini-Round Robin 710

Differences in particles size distribution patern from one laboratory to the other cannot be correlated to the difference among VBD results.. One hypothesis to explain the variations observed at figure 2 is that for a given particle size distribution, the particle shape is likely the principal source of variations. However, this aspect was not investigated in the present paper.

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Many technical exchanges and discussions took place between the laboratories involved. Among the improvement resulting from this inter-laboratory collaboration, the sample preparation was improved. The feeding speed through the roller crusher was modified, the gap between the rolls was adjusted more precisely and the sieving time required preparing the sample before the final analysis was framed. This ensured that each lab was performing the method the same way, thus permitting to improve agreement between laboratories and getting closer to the expected relationship (dotted line in Figure 5). Framing the preparation step had a beneficial effect on VBD measurement as confirmed by the results (Figure 5) obtained by each laboratory during a round robin in 2009. The round robin involved 14 laboratories, two different samples (low and high VBD) and different methods. The graph presents the 2007 results (bold bullets) and shows the improved results from the round robin performed in 2009 (clear bullets).

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The figures 6 and 7 are showing a low variation when the ASTM D7454 method is performed in accordance to the written procedure. However, a modification to the operational procedure leads to an increase of the results range. This trend was observed with the modified D7454 method. In such a case, a lack of crusher or a different coke fraction is responsible for results deviations. The variations observed by the laboratory which performed the D4292 are induced by the latitude offered by the method. In fact, there is no sieving range imposed with the D4292 and the ISO 10236 and since the method of sample preparation

Lab3-2007Î...

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956

can affect the packing characteristics due to differences in the particle shapes, the variation observed with the ASTM D4292 and the ISO 10236 were expected. Despite the fact that those round robin results cannot be compared on a statistical basis, they permit to state that ASTM method D7454 performance is at least comparable, if not better, to the other standard method. On the other hand, there is room and need for improvement in the new context where, because of the decreasing quality of coke quality, the VBD of the available coke becomes closer and closer to the minimum specification. Conclusion Performance improvement of the method is directly linked to the refinement and tighter control of preparation process. Further investigation has been initiated to better identify and understand the factors impacting the quality of the preparation such as the particle size distribution, particle shape, and roller design, including the spacing between the rollers, the roller feeding rate and others with the desire to preserve the results historic. Other mid-term, long-term options are also being considered such as elimination of grinding, exploring other types of grinding, or the use of 3D image analysis.

References 1. A.L.Proulx. Light Metals 657 (1993). 2. D.Belitskus. Light Metals 863 (1974). 3. D.Belitskus and DJ.Danka. Light Metals 429 (1989). 4. L.Duchesneau, R.Lessard, A.Gendron, and G.Brassard. Light Metals 497 (1997). 5. STAS Inc. Vibrated Bulk Density Device (VBDD). 2009. 6. ASTM. Standard test method for determination of Vibrated Bulk Density od Calcined Petroleum Coke using a Semi-Automated Apparatus. ASTM method D7454, 2008. 7. ASTM. Standard Test Method for Determination of Vibrated Bulk Density of Calcined Petroleum Coke, ASTM method D4292, 2010

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

VIBRATED BULK DENSITY (VBD) OF CALCINED PETROLEUM COKE AND IMPLICATIONS OF CHANGES IN THE ASTM METHOD D4292 B. Spencer \ L. Johnsen \ D. Kirkpatrick 2, D. Clark3, M. Baudino 4 !

Oxbow Calcining LLC, Kremlin, OK; 2Oxbow Calcining LLC, Port Arthur, TX; 3Oxbow Calcining LLC, Baton Rouge, LA; 4Oxbow Calcining LLC, Copetro, Buenos Aires, Argentina Keywords: Calcined Petroleum Coke, VBD, Vibrated Bulk Density, ASTM D4292 The reapproved method [6] made several revisions to improve testing results. Those included: • Deleted section 1.3 of procedure not specifying how the 2-kg sample is obtained and the method of subsequent breaking up of particles with diameters greater than 20 mm, and riffling. • Added reference to D2013 Method for Preparing Coal Samples for Analysis. • Added note that VBD is based on packing of sized particles and the method of sample preparation can affect results. • Added section on air drying lab sample • Corrected note 3 (1983 version) about reduction of vibration time from 5 minutes to 1 minute only lower results 0.022 g/cm3. The 1983 version states 0.002 g/cm3. The opinion of the authors is this 0.002 g/cm3 value is a typographical error since the repeatability of D4292 is 0.014 g/cm3. • Added note 6 - precision for VBD on other sample size ranges has not been determined.

Abstract Vibrated bulk density (VBD) is a quantitative measurement used in the aluminum industry to evaluate the density of calcined petroleum coke. In the calcining industry, the reproducibility of the current ASTM International (ASTM) method D4292 generates a wide range of VBD data. Therefore, Oxbow Calcining investigated the VBD procedure - D4292. Issues with D4292 include the use of the appropriate crushing equipment, crushing of the gross sample, sieving of the prepared sample, determination of the sample volume using appropriate graduated cylinder vibration time and apparatus setup. This investigation led to a revision in the ASTM D4292 method. Introduction Oxbow Calcining produces calcined petroleum coke (CPC) which is used to make anodes in aluminum smelters. Raw material from petroleum coker units is devolatilized and densified in kilns to meet customer specifications [1] such as sulfur content, trace metals, air reactivity and density. The smelter blends the CPC with pitch to form the carbon anode [2]. VBD is the most commonly-used procedure to determine bulk density and provides the smelter with the best overall estimate of coke suitability for anode manufacture. With lower bulk density, coke porosity increases, effective pitch level increases and the overall electrical resistance of the electrode increases [3].

In 2009, Oxbow Calcining investigated several sources of potential variability in the test method in an attempt to improve the method. The analytical testing to improve the method focused on sample preparation and on sample introduction to the test equipment. The testing evaluated: • Separation of the natural fraction and crushing the oversized vs. crushing the entire sample • Roll crusher gap sizes • Use of a Plate Mill instead of a Roll Crusher • Rate of Sample Introduction into the VBD Apparatus • Effect of vibration duringfillingof the cylinder

As originally written, ASTM D4292-83 VBD test method [4] contained several ambiguities in the procedure leading to a repeatability and reproducibility of 0.014 and 0.046 g/cm3, respectively. The ASTM research report [5], which discusses the method development, gives a few details about the original round robin used to develop the precision and bias statements for D4292. This original round robin consisted of 6 labs with 8 test samples. Samples were sized 20x48 mesh. Results on the 8 test samples were between 0.76 and 0.92 g/cm3. There was no documentation in the research report that gives the actual method in use during the round robin. The round robin calculations were dated January 11, 1982. The original D4292 method has approval and publication dates of October 28, 1983 and January 1984, respectively. The time difference of 21 months between the round robin calculations and the ASTM approval of D4292 causes concern regarding to the actual procedure in use during the original D4292 round robin.

Sample Preparation Separation of the Natural Fraction One proposed idea for improvement in the results is to separate the naturalfractionprior to jaw crushing (Figure 1). The premise was that the crushing may introduce variation in the pore distribution. By separating the natural fraction and crushing just the oversized fraction, this variation is minimized in the tested sample. An inter-laboratory study was conducted to evaluate this theory using seven different cokes.

959

Comparing the results of Table I to Table II reveals the difference in VBD reproducibility between crushing the entire sample vs. crushing just the oversized fraction. Only one sample of entirely crushed material (Sample B) produced a significantly higher reproducibility between the laboratories. The remaining six samples produced similar or lower reproducibility than the crushing-just-the-oversized-fraction results. This suggests that crushing the entire sample would improve reproducibility of the sample when compared to separating the natural fraction and crushing the oversized particles. For this reason, section 8.2 of the new procedure now requires that the entire sample is passed through the jaw crusher.

Roll Crusher Gap Distance The next series of tests evaluated the effect of gap distance between the rollers of a roll crusher (Figure 2) on the D4292-92 results.

Figure 2. Roll Crusher

Table III. Evaluation of a Roll Crusher in VBD Sample Preparation Runs 1 and 2 are 300 grams each; Spring adjusted prior to test Roller Opening 1 pass

Figure 1. Chipmunk Jaw Crusher used at Port Arthur

Table I. D4292 Crushing All Material with Jaw Crusher (g/cm3)

Material Sample A Sample B Sample C Sample D Sample E Sample F Sample G

Labi 0.857 0.881 0.850 0.865 0.794 0.850 0.857

Lab 2 0.874 0.901 0.834 0.855 0.811 0.851 n/a

Lab 3 0.837 0.870 0.827 0.857 0.802 0.839 0.848

Lab 4 0.864 0.887 0.849 0.867 0.834 0.857 0.860

Reproducibility 0.037 0.031 0.023 0.010 0.017 0.018 0.012

|

Table Π. D4292 Crushing Only +4 Mesh with Jaw Crusher (g/cm3) 1

Material Sample A Sample B Sample C Sample D Sample E Sample F Sample G

Labi 0.850 0.873 0.850 0.865 0.814 0.857 0.842

Lab 2 0.832 0.879 0.835 0.862 0.799 0.838 n/a

Lab 3 0.837 0.866 0.828 0.855 0.800 0.837 0.842

Lab 4 0.871 0.889 0.848 0.865 0.839 0.856 0.869

Reproducibility 0.034 0.023 0.022 0.010 0.040 0.020 0.027

mm

mm

2.5

1.5

| |

mm

1 pass

1.0

1.0

1.0

4 passes

0.3

0.3

0.3 1

% on + 28

n/a

7.4%

6.0%

% 28x48

59.0%

55.3%

57.3%

% on + 28

n/a

4.5%

2.3%

% 28x48

58.3%

56.6%

59.7%

Run 1-A

VBD, g/cm3

0.840

0.833

0.844

Run 1-B

VBD, g/cm3

0.833

0.826

0.830

Run2-A

VBD, g/cm3

0.833

0.837

0.840

Run2-B

VBD, g/cm3

0.833

0.833

0.830

Average VBD

0.835

0.832

0.836

Runl

Run 2

\

Three tests were performed in duplicate. The first test used a 2.5 mm gap for the first pass followed by a run through a 1.0 mm gap between the rollers. The run ended with four passes through a 0.3 mm gap. The second test used a 1.5 mm gap for the first pass followed by a run through a 1.0 mm gap between the rollers. The test ended with four passes through a 0.3 mm gap. The third test

by a plate mill/pulverizer and a roll crusher (two rotating cylinders). Four different grades of calcined petroleum coke were used. For each grade, a 28x48 Tyler mesh sample was prepared using both types of crushers and a two step grinding process. For all four grades, the samples prepared using the plate mill had higher VBD results than the samples prepared with a roll crusher. The average difference between the methods was 0.052 g/cm3. For example, a sample prepared with the roll crusher would have a 0.854 g/cm3 VBD versus 0.906 g/cm3 VBD when prepared with a plate mill.

made one pass through a 1.0 mm gap between the rollers followed by four passes through a 0.3 mm gap. Table III presents the results of varying the roll crusher opening. Three different parameters were tested. The 2.5 mm gap test yielded no oversized material. However, no real difference was seen between the average results of each test and thus the D4292 was not modified. Evaluation of Plate Mill/Pulverizer

A second test, comparing different plate mill gap sizes, was performed on Material 1. One sample was crushed using a 1.5 mm initial gap followed by a 0.5 mm gap. An initial 2.0 mm gap size followed by a 0.8 mm gap was used on the second sample. The ratio of oversize (+28 Tyler) to undersize (-48 Tyler) was 0.59 for the 1.5/0.5 mm gapped sample and 2.63 for the 2.0 mm/0.8 mm gapped sample. The VBD result of the 1.5 mm/0.5 mm gap sample was on average 0.033 g/cm3 higher than that of the 2.0 mm/0.8 mm gapped sample.

One concern with a roll crusher is that the rollers compress the sample, which may alter the pore shape and volume. Therefore, a plate mill/pulverizer was evaluated. A plate mill/pulverizer consists of serrated parallel plates with one stationary plate and one moving plate (Figure 3).

Ratios of oversize to undersize were calculated for all the samples. Section 8.3.7 of D4292-92 states that the crushing level is satisfactory when the ratio of the coarser to finer particles is between 0.8 and 2.0 with a sample yield of at least 30%. Gaps were set the same, but the ratio varied based on the grade of coke. The values ranged from 1.65 to 2.52. Compared to the roll crusher, the plate mill/pulverizer produced lower test material yield. Based on the lower yield and much higher VBD results, the change to plate mill/pulverizer was not recommended. However, these results, as well as the gap size results, led to a change in section 6.2 in the new VBD procedure to provide clear specifications on the grinding. The new procedure specifies that only a roll crusher must be used for crushing. Figure 3. Plate mill/pulverizer Table IV presents the results of experiments that were conducted to compare the difference in the VBD values of samples crushed

Table IV. Comparison of Plate Mill to Roll Crusher 1

Material Unit Initial/final gap size, mm

Roll Crusher 1.5/0.5

1

2 Plate Mill

4

3

|

Roll Crusher

Plate Mill

Roll Crusher

Plate Mill

Roll Crusher

Plate Mill

1.5/0.5

2.0/0.8

1.5/0.5

2.0/0.8

1.5/0.5

2.0/0.8

1.5/0.5

2.0/0.8 551.3

200.3

510.1

166.8

Sieve Content, g +28 Mesh

269.1

64.7

28 x48 Mesh

155.9

101.4

176.6

164.0

177.2

-48 Mesh

106.7

109.9

210.0

96.6

189.2

539.0 1

482.0

176.4

147.6

125.0

171.3

149.2

92.6

269.6

107.2

257.2

Ratio +28A48

2.5

0.6

2.6

2.1

2.7

1.8

1.8

1.6

2.1

Sample Yield (28x48), wt.%

29

37

19

36

20

36

14

38

16

VBD, g/cm3

0.817

0.889

0.857

0.824

0.869

0.850

0.928

0.861

0.932

961

Sample Addition

Table V. Average VBD (g/cm3) vs. Cylinder Filling Time

The D4292 VBD apparatus is shown in Figure 4. Filling the graduated cylinder at a different rate can affect the particle distribution in the graduated cylinder, yielding a different VBD result. Tests were performed to determine the best fill rate.

SAMPLE I.D. 1 2 3 4 5

30 0.798 0.843 0.862 0.824 0.824

Filling Time, sec 90 60 0.815 0.833 0.872 0.882 0.872 0.882 0.862 0.872 0.862 0.852

120 0.833 0.893 0.882 0.862 0.872

Additional Changes Since it was found that vibrations can affect the VBD results, Section 9.2 in the new procedure specifies that the graduated cylinder should not be attached to the vibrating table, but should be free-standing on the table. The D4292 methods approved in 1983 [4] and 1992 [6] have a note which states that vibrating the graduated cylinder for only one minute instead of the required five minutes results in only 0.022 g/cm3 difference and states that for routine use, the shorter time may be preferred. The implications of this time difference directly impacts the results. For greater consistency, the vibration time should not vary. Therefore, the note was removed in the most recent procedure. Due to the highly dependent nature of VBD on sample preparation, several other changes were made to the VBD procedure to improve consistency between various facilities. Section 8.1 clarifies that dedust oil should not be removed from the sample, since the process of removing dedust can change the sample. Section 8.3.6 specifies that no sieving of samples should occur between crushing steps (with a roll crusher), and personnel must use a Ro-Tap sieve shaker for 15 minutes. Both of these steps are designed to reduce variability between facilities.

Figure 4. D4292 VBD Apparatus An experiment was run comparing two different cylinder filling methods. For one method, a graduated cylinder was filled while the cylinder vibrated. Material flowed from a vibratory feeder into the funnel on top of the vibrating cylinder. The amount of time it took for 100 g of material to fill the cylinder was greater than 90 seconds. After the sample was in the cylinder, the vibration was stopped and the material height was recorded. For the second method, 100 g of material was added through a funnel into a stationary cylinder. Once all the material was in the cylinder, the cylinder was vibrated for 5 minutes; then the height was recorded. Samples from two different grades of coke were run 5 times with each method. The method of vibrating the cylinder during filling gave an average 0.04 g/cm3 higher than the other method. Section 11.3 in the new ASTM procedure now specifies that the cylinder should not be vibrated during filling since this action will change the results.

Conclusions VBD values are dependent on the type of crusher used and the gap settings used. Different cokes will behave differently at the same gap setting. VBD results have been shown to be dependent on sample preparation as well as filling and vibrating techniques. In order to ensure that the test method is consistent, clearer instructions were required for the ASTM D4292 method. Changes mentioned in this paper were included to reduce variability in VBD testing. Special care should be taken when performing this method to adhere to the steps as closely as possible.

Several different grades of coke were tested at various feed rates of filling the graduated cylinder. Data, presented in Table V, shows that the speed at which the material is added to the graduated cylinder can affect the final VBD results. As the time increases, the effect of the filling time decreases to a point where the VBD value is not affected. For this experiment, that point occurred around 90 seconds. Because of these results, Section 11.1 in the new procedure clarifies that the sample must be poured slowly and consistently into the graduated cylinder at a rate of 10 - 14 grams per 10 seconds. Since this is hard to accomplish by hand, using a vibratory feeder for sample introduction into the graduated cylinder is recommended.

ASTM has incorporated these changes in the latest test revision, D4292-10 [7]: a.

b.

962

Section 6.2: Roll crusher specifications, requirement that both rolls must rotate, requirement to check spring tension per manufacturer specifications. Section 8.1: Comment added to not remove dedust oilfromthe sample.

c. d. e. f.

g. h.

Section 8.2: Requirement for entire sample to pass through a jaw crusher Section 8.3.6: No sieving of sample between crushing steps (roll crusher). Also, the Ro-Tap sieve shaker must run for 15 minutes. Section 9.2: Do not attach the graduated cylinder to the vibrating table. Allow graduated cylinder to vibrate freely. Section 11.1: Pour sample slowly and consistently into graduated cylinder at rate of 1 0 - 1 4 grams per 10 seconds. Suggested using vibratory feeder for sample introduction into graduated cylinder. Section 11.3: Do not vibrate graduated cylinder while adding coke fraction Note 5: Deleted note 5 that suggested a lab could vibrate the graduated cylinder for only one minute instead of the required five minutes since the difference in results is only 0.022 g/cm3.

2

Paul J. Ellis and Christopher A. Paul. Tutorial: Petroleum Coke Calcining and Uses of Calcined Petroleum Coke. Prepared for the AIChE 2000 Spring National Meeting. Atlanta, GA., March 5-9,2000. Unpublished. 3

S. Chakravarty and Arun Kumar. "Petroleum Coke for the Anodes of Aluminum-Reduction Cells," Carbon Technology Seminar Proceedings (1979), p. 11-15. 4

ASTM International. "Standard Test Method for Vibrated Bulk Density of Calcined Petroleum Coke," ASTM Method D4292 83. Approved October 28, 1983. 5

ASTM research report No. RR: D-2-1166. July 25, 1983.

6

ASTM International. "Standard Test Method for Vibrated Bulk Density of Calcined Petroleum Coke," ASTM Method D4292 92. Original Approval in 1992. 7

ASTM International. "Standard Test Method for Vibrated Bulk Density of Calcined Petroleum Coke," ASTM Method D4292 10. Approved July 1, 2010.

References 1 Harald Onder and Edward A. Bagdoyan. Everything You Always wanted to Know about Petroleum Coke - A Handbook (Allis Mineral Systems Kennedy Van Saun, 1993).

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION

Anode Quality and Rodding Processes SESSION CHAIR

Nigel Backhouse Rio Tinto Alcan Saguenay, Quebec, Canada

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

MULTIVARIATE MONITORING OF THE PREBAKED ANODE MANUFACTURING PROCESS AND ANODE QUALITY 1

Julien Lauzon-Gauthier1, Carl Duchesne1, Jayson Tessier2, Katie Cantin3, Isabelle Petit3 Aluminium Research Centre-REGAL, Chemical Engineering Department, Universitι Laval, Quebec, City, QC, GIK 7P4, Canada 2 STAS, Alcoa-STAS R&D Team, Chicoutimi, QC, G7K 1H1, Canada 3 Alcoa inc., Aluminerie de Deschambault, Deschambault, QC, Canada, GOA ISO Keywords: Prebaked anode, anode quality, statistical analysis, PLS investigation of the causes of anode quality variations. Finally, the anode quality prediction will be discussed.

Abstract Prebaked anode quality is available through a weekly average of core sample laboratory measurements. Unfortunately, long delays between production and results (approximately 4-6 weeks) can lead to poor abnormal operation and faulty anode detection, and difficult process control. Extensive raw material and process data are available at Alcoa Deschambault smelter's carbon plant. Using projection to latent structures, a multivariate statistical method, it was possible to correlate raw material and process conditions to weekly lab results. The effect of different petroleum coke and coal tar pitch was analyzed and instant weekly prediction of anode properties was achieved.

Process and data The data used in this article were collected from the Alcoa Deschambault smelter (ADQ), located in Deschambault near Quebec City. The ADQ smelter operates 264 AP-30 reduction cells with 40 anodes each. The smelter's carbon plant produces more than 150 000 anodes annually in order to fulfill the potroom needs. The green anode plant has a capacity of 30 tons per hour with continuous mixers and vibrocompactors. There are two baking furnaces with 34 sections each. Weekly data was collected from: raw material lab analysis and supplier's certificate of analysis (COA), green mill process and baking furnace data historian, and core sample laboratory analysis. The time spanned by the dataset is from December 29, 2008 to July 26, 2010 for a total of 82 weeks. During this period, Deschambault used 6 different coke suppliers and 2 pitch suppliers. The anode recipe usually combines 2-3 types of cokes and a single type of coal tar pitch. The different raw material blends used throughout the analysis period are presented in Table I.

Introduction Baked anode properties are evaluated using core sample laboratory analysis. Usually, the number of samples taken during production (approximately 1%) is too small to accurately measure the variability of anode properties produced at the carbon plant. Weekly averaged data collected on a number of variables are used to monitor process performance and deviations. Furthermore, laboratory results used to monitor process performance are only available 4 to 6 weeks after the anodes have been produced. Typically, results are available after the anodes have been set in the pots. Detecting deterioration of anode properties and feedback correction of possible root-causes is slow due to these limitations. Critical process conditions are monitored at a much higher frequency at the green mill and baking furnace, but final product quality cannot be measured online.

Table I: Raw material blends processed during the analysis period.

1 A 2 A A 3 4 A D 5 6 D *Coke suppliers are identified by with numbers 1-2.

Alcoa Deschambault has an extensive database on raw material properties, green mill and baking furnace operating conditions. Industrial databases are often not exploited to their full potential due to the enormous quantity of information that needs to be analyzed to extract useful information. The goal of this analysis was to use Deschambault's process database to model and explain variability in baked anode properties. Using multivariate statistical techniques, it is possible to correlate variations in raw material properties and process operating conditions to variations in anode properties. Using this model, it is possible to investigate which combinations of parameters have the greatest influence on anode property variation and to predict anode properties on a weekly basis. This analysis is based on historical process data, with no special measurements other than data normally collected during plant operation. It will be demonstrated that traceability of the effects of raw material on anode quality can be achieved.

B D 1 C D 1 E 1 C C E 2 E 2 F 2 letters A-F and pitch suppliers

Deschambault has some blending capability that was discussed in the paper by Gendron et al. [1]. All the data presented in this paper have been mean-centered and scaled to unit variance (i.e. auto-scaled). The variables included in the analysis are presented in Table II. Multivariate analysis The dataset used in this analysis contains a large number of variables. Industrial databases are noisy and typically contain a certain percentage of missing data. The variables are also generally highly collinear. All of these situations can cause problems when using classical multilinear least-squares regression analysis. To handle these data issues, a multivariate latent variable called Projection to Latent Structure (PLS) was used.

This paper explains in detail how the analysis was performed. First, the dataset used for the study is described in detail. A brief description of the method used is presented, followed by an

967

Table II: Variables included in the analysis divided into a regressor block X and a response block Y VarlD

X-variable

VarlD

X-variable

VarlD

X-variable

VarlD

Y-variable

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

Coke real density Coke Na Coke Ca Coke S CokeV Coke app dens Ree butts Ca Ree butt Na Ree butt S Ree butt V Ree butt Na/Ca Pitch SP Pitch TI Pitch QI Pitch Coking val Pitch S

17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32

Pitch Dist VC bellows P % Pitch % Coarse % Inter % Fines % Butts recyc % Green recyc Coarse Rt4 Inter Rt50+RtlOO Fines Pt200 Butts ree Rt3/8 Butts ree Rt3/8+Rt4 Aggregate Rt3/8 Aggregate Pt200 Green app dens

33 34 35 36 37 38 39 40 41 42 43 44 45 46

Mixer lpower Mixer 1 delta power Mixer 2 power Mixer 2 load HX paste T Pitch Temp Pitch/paste Temp delta % Green scrap Fire weight loss Baked weight Fire cycle time Fire start Temp Fire final Temp % Baked scrap

47 48 49 50 51 52 53 54 55 56 57 58 59

% Air dust %AirRx %C02dust %C02Rx Flexu Strength Fracture Energy Thermal conduc Coeff Expansion Elee resistance Lc Young Mod App dens baked Real dens baked

compute the prediction error sum of squares (PRESS) of the group of data left out of the model. This procedure is repeated until each observation (i.e. row of X and Y) have been left out once. The overall prediction error sum of squares for a PLS model with a latent variables PRES S (a) is then computed. The selected number of LVs (a) is that one minimizing prediction errors (PRESS). A summary of the model is presented in Table III. It shows the cumulative variance of the X and the Y (R2X and R2Y) data blocks explained by the first 5 PCs, as well as the cumulative variance of Y predicted by the model (Q2) through the crossvalidation procedure.

This technique, and other multivariate methods, have been presented in recent TMS papers [2-5] for the analysis of primary aluminum smelter data and in other journals for different applications [6]. The PLS regression method uses all the available X and Y variables and projects them onto a lower dimensional subspace called the latent variable space. These new latent variables (LV) are a linear combination of the original variables and are computed so as to maximize the covariance between the X and Y variables. The latent variables can be viewed as a small number of lurking variables driving the process, and hence the X and Y data, in certain correlated directions. The structure of the PLS model is shown below. X = TP 1 +E Y = TQT + F T = XW*

(1) (2) (3) (4)

VIPM = >/jL;>*,(SSY a /SSY, 0 ,)

(5)

w* = w^wy 1

Table III: PLS regression model overview PC R2X(cum) R2Y(cum) Q2(cum) 1 2 3 4 5

In the equations above, the P and Q matrices contain the loading vectors that best represent the X and Y spaces respectively, whereas W* contains the loading vectors that define the relationship between the X and the Y spaces. Projection residuals of each space are stored in E and F. The variable importance in projection (VIP) is an indication of the importance of a variable in predicting the Y space. In the equation (5), waj is the loading weight of the j t h variable in the ath PLS latent variable, SS Ya is the sum of squares of Y explained by the a* LV of the PLS model and SS Ytot is the sum of squares of Y explained by the model.

0,225 0,316 0,394 0,449 0,501

0,263 0,395 0,451 0,498 0,527

0,237 0,346 0,372 0,381 0,379

A cumulative Q value of 0,379 appears low, but this value is an overall value computed from all thirteen Y variables. Individually,, most variables have good predictions. This is shown in Table IV. The same comment applies to the explained R2X and R2Y, some variables are well explained and some are not which lowers the overall variance explained. Table IV lists the variance explained for each individual Y variables. The fit or variance explained (R2Y) for this model ranges from 0,300 to 0,712. Some would consider such a fit as low, but these are good results considering the industrial nature and the level of noise in the data as well as the uncertainties related to the measurement of raw material properties and the residence time within each piece of equipment. The same comments apply to the variance predicted (Q2) which ranges from 0,126 to 0,610. For some variables, (i.e. Coeff Expansion and Young Mod) the prediction ability of the model is low, but it can be considered good for most variables.

Results PLS regression model A PLS regression analysis was performed on the dataset which contains 82 observations, 46 X-variables and 13 Y-variables. The model contains five latent variables (LVs). They were selected by a cross-validation procedure to maximize predictive ability and to minimize overfitting. This technique leaves a randomized group of observations out of the dataset. A PLS model with a latent variables is built on the remaining observations, and then used to

968

2

«* U

1

t_ ^

0

1

1 2

i n i

2

1 VV ! 3

i

au i

Weekly obs.

1 E -1] -2

J> Λ,-ν" « * - ? Λ . £ JVfe <η> >ø

Λ

*

^ ' $> >£ ^j£>' fi^ ^^ ' f> Weeklv ohs.

îrfW^^ir/Nni

•^_ÎI | ; 3

/vvvvVWv' Weekly *»bs.

Weekly obs.

Figure 1: Some raw material properties and process conditions variations. and pitch ratio) with the three different groups of similar process conditions identified.

Table IV: Variance explained for each Y variables Var ID Variable 47 % Air dust 48 % Air Rx 49 % C02 dust 50 % C02 Rx 51 Flexion Strength 52 Fracture Energy 53 Thermal Conduc 54 Coeff Expansion 55 Elee Resistance 56 Lc 57 Young Mod 58 App Baked Dens 59 Real Baked Dens

R2Y(cum)

Q2(cum)

0,505 0,583 0,495 0,559 0,712 0,501 0,509 0,300 0,675 0,613 0,362 0,664 0,450

0,341 0,450 0,318 0,381 0,610 0,253 0,354 0,126 0,520 0,514 0,186 0,549 0,341

Table V: VIP Rank Var ID 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

From the PLS regression model using 5 LVs a list of the 15 most important variables in the prediction (VIP) is presented in Table V, As a rule of thumb, as discussed in papers by Chong and Jun and Ericksson et al [7, 8], variables having a VIP over 1 are considered important. Most of the variables listed in Table V are raw material properties. This was expected since Deschambault dealt with six different coke-pitch blends (i.e. different coke/pitch supplier combinations) over the time spanned by this analysis. However, more operating conditions were expected to show a greater influence on the model. For example coke or dry blend size distribution for coarse, intermediate and fines fractions and their ratio in the dry blend mix all had a VIP of less than 1. This can be explained by their lack of variability in the dataset. Except for the pitch ratio in the paste, almost all other operating conditions in the green mill were kept constant. A design of experiments on the operating conditions would have enabled the capture of more information from the process variables. Figure 1 shows the variation in some coke properties, pitch QI and two process operating conditions (vibrocompactor bellows pressure

42 9 18 10 19 4 14 10 45 13 43 32 15 1 24

Variable

VIP

Baked weight 1,539 1,518 Ree butt S 1,458 VC bellows P 1,421 Ree butt V 1,413 % Pitch 1,369 Coke S Pitch QI 1,345 1,297 CokeV 1,252 FirefinalT 1,239 Pitch SP 1,216 Fire cycle t Green app density 1,203 Pitch Coking val 1,178 Coke real dens 1,173 1,070 % Green ree

Process variability investigation One of the goals of this analysis was to investigate process variability. The plot of the first two latent variables (ti and t2), shown in Figure 2, provides an overview of the information captured by the PLS model on the joint X-Y data blocks. The first two LVs are shown because they explain approximately 80% of the modeled variability of the Y variables (i.e. 39,5% out of the 52,7% of the 5 LVs). Some information can be contained in the three other latent variables (around 20%). The results presented in this paper focus on the first two LVs for brevity and also because most of the important information is carried in these two LVs. The 82 markers within this plot correspond to the projection of each multivariate observation onto the plane formed by the first two

969

principal components. Weekly observations projected within a similar region of the t r t 2 score plot show similar patterns in their data structure, hence similar in the latent variable space (i.e. similar combinations of raw material properties, recipes, process conditions and anode properties) whereas those falling in distinct regions are different.

continuously raised during that period to compensate for the increase in QI content and to meet green weight set-point. There was also a step change in the pressure applied on the anode during vibrocompaction and it has been kept constant since then. Green apparent density (GAD) went up in correlation with the pitch ratio. There was a major change in the baked weight. This might be due to the pitch supplier change. The new pitch has a much higher QI and coking value which will lead to more binder cokιfaction in the furnace, as discussed in [9, 10].

Blend 1 - B l e n d 2 * Blend 3 Blend 4 * Blend 5 * Blend 6

Figure 2: Scatter plot of the first two latent variables of the PLS model. Three distinct clusters are observed in Figure 2. Group 1 represents baked anodes produced from December 29, 2008 to July 13, 2009, group 2 represents anodes manufactured between September 14, 2009 and January 18, 2010, and group 3 between January 25, 2010 and July 26, 2010. The different causes leading to process movement in the latent variable space will be analyzed. One of the tools that can be used for interrogating the PLS model is the contribution plot, indicating which variables are associated with a movement in the t r t 2 latent variable space. Figure 3 represent changes from group 1 to group 2 and Figure 4 represents changes from group 2 to group 3.

Variables

Figure 4: Contributors to changes from group 2 to group 3 Differences between group 2 and group 3 are also due to raw material variations. Coke real density in the new coke blend was lower and these affected GAD, pitch demand and baked weight. It should be noted that process operating conditions except from pitch ratio and vibrocompactor bellows pressure, were kept constant during the time spanned by the analysis. Thus it is not possible to capture the effect of most process variables since they were not changed during the time period considered. Table VI: Average Y predicted for the three different operating groups (auto-scaled values) Ypred- variable Group 1 Group 2 Group 3 % Air dust %AirRx % C0 2 dust % C0 2 Rx Flexu Strength Fracture Energy Thermal conduc Coeff Expansion Elee resistance Lc Young Mod App dens baked Real dens baked

Figure 3: Contributors to changes from group 1 to group 2 The first transition between groups 1 and 2 is associated with several different factors. The first contribution comes from the sulfur and vanadium content of the coke and recycled butts. This change started with blend number 3 due to the change from coke D to coke E. The transition was accentuated by the change in pitch supplier which can be seen by the pitch coking value, quinoline and toluene insolubles. The pitch ratio in the anode was

0,183 -0,666 -0,584 0,670 -0,952 -0,689 -0,749 0,328 0,878 0,700 0,486 -0,439 0,725

-0,623 0,975 0,265 -0,289 0,725 0,280 0,575 0,056 -1,064 0,049 -0,371 1,054 -0,406

0,245 -0,032 0,635 -0,735 0,479 0,723 0,340 -0,584 -0,091 -0,992 -0,444 -0,450 -0,542

Averaged autoscaled predicted values for each group are listed in Table VI. Group 1 has the best C0 2 reactivity. Group 2 has the best air reactivity; this can be linked to the lower content of sulfur and vanadium during that period. The real baked density was higher in group 1 since coke real density has been gradually decreasing over time. Lc in baked anode seems to be highly

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correlated to coke real density. Baking final temperature was increased from group 1 to group 2, but the Lc has decreased. It is difficult to determine which group was better in terms of mechanical properties. Electrical resistance and baked apparent density were higher in group 2. This could be due to higher QI and coking value in the pitch used during that period. It is difficult to determine which of the 3 operational regions led to the best overall anode quality when looking at the quality variables one at a time. An important input would be to know which group had the best behavior in the potroom. Prediction of anode properties Weekly obs.

The second goal of this analysis is to predict anode quality variables. Figures 5 to 8 represent the predicted and measured values on time series graph for Lc, Electrical resistance, % CO2 Rx and % Air dust. The operational groups described earlier are marked on the charts. The variance explained by the model for each of these variables is listed in Table IV. The prediction model can be useful for a faster detection of large deviation in anode properties. Usually, complete lab results of anode core sample are available only 4 to 6 weeks after sampling. Being able to instantly estimate anode properties fabricated during the week can lead to earlier process deviation detection and faster problem solving. Therefore, such a tool would enable process engineers and technical staff to operate the carbon plant and baking furnace closer to the optimum production state, with respect to anode properties. However, the laboratory analysis of core samples will always be needed to validate model predictions.

Figure 7: Measured and predicted values for % C0 2 Rx

Figure 8: Measured and predicted values for % Air dust Conclusions The goals of this analysis were to investigate the causes of anode quality variability and the possibility to predict anode quality using raw material and process information. Using the projection to latent structure regression method, correlations between process data and anodes quality were modeled. Most of the variability affecting anode quality came from raw material variations (coke and pitch binder) mainly due to supplier changes. Since process conditions were kept constant, no variability could be captured. Variation within the same suppliers could be explored by performing separate analysis on each blend. It was also demonstrated that it is possible to predict a number of anode quality variables. Variance captured by the model (obtained by cross validation) ranged from 0,186 to 0,610. Variables with higher explained variance can be predicted on a weekly basis and at least 4 weeks before laboratory results. This can help in detecting problems, investigating them and taking corrective action much faster than when one has to wait for laboratory results.

Figure 5: Measured and predicted values for Lc

Acknowledgments This research was funded by the Fonds quιbιcois de la recherche sur la nature et les technologies (FQRNT), through the Aluminum Research Centre-REGAL, by the National Science and Engineering Research Council (NSERC) and by Alcoa Inc. The authors would like to thank the Alcoa Deschambault anode plant technical staff for their help and Alcoa Inc. for permission to publish this paper.

Figure 6: Measured and predicted values for Electrical resistance

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References 1. Gendron, M., Whelan, S., Cantin, K., 2008, Coke blendind and fines circuit targeting at the alcoa deschambault smelter, TMS Light Metals 2008, pp. 861-864 2. Tessier, J., Duchesne, C, Tarcy, G.P., Gauthier, C, Dufour, G., 2008, Analysis of a potroom performance drift, from a multivariate point of view, TMS Light Metals 2008, pp. 319324 3. Tessier, J., Zwirz, T.G., Tarcy, G.P., Manzini, R.A., 2009, Multivariate statistical process monitoring of reduction cells, TMS Light Metals 2009, pp. 305-310 4. Tessier, J., Doiron, P., 2010, Statistical investigation and modeling of bath level in hall-hιroult cells, TMS Light Metals 2010, pp. 553-558 5. Majid, N.A.A., Young, B.R., Taylor, M.P., Chen, J.J.J., 2009, Detecting abnormalities in aluminium reduction cells based on process events using multi-way principal component analysis (MPCA), TMS Light Metals 2009, pp. 589-593 6. Kourti, T., MacGregor, J.F., Process analysis, monitoring and diagnosis, using multivariate projection methods, Chemometrics and Intelligent Laboratory Systems, 1995, Vol. 28, pp. 3-21. 7. Chong, I.-G., Jun, C.H., 2005, Performance of some variable selection methods when multicollinearity is present, Chemometrics and Intelligent Laboratory Systems, Vol. 78, pp. 103-112. 8. Eriksson L., Johansson E., Kettaneh-Wold N., Trygg J., Wikstrom C, Wold S., 2006, Multivariate and Megavariate Data Analysis: Part I - Basic Principles and Applications. Umetrics, Umeβ, Sweden, 307p. 9. Wombles, R. H., Baron, J. T. McKinney, S., 2009, Evaluation of the necessary amount of QI in binder pitch, TMS Light Metals 2009, pp. 913-916 10. Krupinski, K. C, Windfelder, J.J., 1992, Effects of tar quinoline insolubles on manufactured carbon performance, 1992 Ironmaking conference proceedings, pp. 487-492

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CHARACTERIZATION OF A FULL-SCALE PREBAKED CARBON ANODE USING X-RAY COMPUTERIZED TOMOGRAPHY Donald Picard1'2, Houshang Alamdari1'2, Donald Ziegler3, Pierre-Olivier St-Arnaud2, Mario Fafard2 1 Department of Mining, Metallurgical and Materials Engineering, 1065 avenue de la Mιdecine Laval University, Quebec, QC, G1V 0A6, Canada 2 NSERC/Alcoa Industrial Research Chair MACE3 and Aluminum Research Center - REGAL Laval University, Quebec, QC, G1V 0A6, Canada 3 Alcoa Canada Primary Metals, Aluminerie de Deschambault, 1 Boulevard des Sources, Deschambault-Grondines, QC, GOA ISO, Canada Keywords: Carbon Anode, Computerized Tomography, Density CT scanners may be influenced by a number of parameters such as the heterogeneity of the material [4]. The present work therefore aims at the development of a CT-based method for estimation of apparent density of anode materials. Full-scale prebaked anode core samples were used for calibration and validation of this method.

Abstract In the conventional Hall-Hιroult electrolysis process, the carbon anode is formed either by pressing or by vibro-compaction. The final properties of an anode are influenced by many parameters such as raw materials properties and manufacturing process. Presently, the aluminium producers have to deal with continuous variation of raw materials properties. To minimize the effects of the raw materials variations on the final product quality, numerical modeling of the forming process is of great interest. However, it is imperative to collect data on real anodes in order to calibrate these models. Some of the most valuable data are the density and porosity distribution of a full-scale baked anode obtained with computed tomography (CT). To test the method, three cored samples of 300 mm in diameter were taken from an industrial anode and scanned with an X-Ray tomograph. Calibration standards were also used to fit the CT scan results with the experimental data.

Methodology Sample preparation Three core samples of 292.1 mm (11.5 in) in diameter were investigated in this study. The samples were taken from a fullscale prebaked anode, produced at Alcoa Deschambault Smelter and scanned using a Somatom Sensation 64 at INRS-ETE research centre in Quebec City (Figure 1). The locations of the three core samples are illustrated in Figure 2.

Introduction In the Hall-Hιroult electrolysis process, the cell is composed of various materials including steel potshell, insulators, refractory concrete, carbon lining and carbon anodes. Among them only the carbon anode could be considered as a consumable item requiring regular replacement. These anodes are consumed during electrolysis and replaced after approximately 28 days of operation. Depending on the cell technology, approximately one anode per cell is replaced each day. Hence a large number of anodes and consequently a large quantity of raw materials are required to operate a plant. The aluminium producers need to deal with continuous changing of raw materials properties resulting in a wide variation of physical properties of the pre-baked anodes. One solution to minimize the effect of the variation of raw materials properties is to use numerical simulation methods to model the manufacturing process. The objective is to predict the anode characteristics, to control the process parameters more efficiently, and to take corrective actions before the anode is produced. To achieve this goal, a series of experimental data must be first collected in order to validate the models. In the present case, the focus will be on the apparent density distribution of prebaked anodes which can be measured either by destructive sampling [1] or by nondestructive testing (NDT) methods [2, 3]. To obtain the apparent density distribution of a full-scale prebaked anode the NDT X-Ray computerized tomography (CT) method could be used. However, previous work [2] has shown that the apparent density obtained with the CT method needs to be calibrated with standard samples. In fact, the density estimated by

Figure 1. Somatom Sensation 64 located at the INRS-ETE (Courtesy of INRS-ETE). In order to reveal the size effect on the CT scan results, two core samples of 152.1 mm and 50.8 mm in diameter were also taken from the same 292.1 mm samples and scanned (Figure 3). The difference in length between the samples is due the coring apparatus limitations. The upper surface of the 152.1 mm and 50.8 mm diameter samples corresponds to the bottom surface of the stub holes as shown in Figure 3. Thus, each core sample was analyzed in three different sizes. The core #3 was broken during the coring process.

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respectively [2]. According to this calibration, the apparent density can be calculated using the following equation: p=0.00UCTnumber+l

(l)

The CT images were obtained by setting the X-Ray tube at 120 keV and 300 mA. X-Ray attenuation was measured using 0.6 mm progress steps. Each voxel (volumetric pixel) is therefore an average of the X-Ray attenuation of material with 0.6 mm of thickness. The volume of each voxel is related to the sample diameter presented in Table I. Table I. Voxel sizes. Core sample diameter (mm) Volume element size (mm) 1 292.1 0.7X0.7X0.6 152.1 0.3X0.3X0.6 50.8 0.13X0.13X0.6

Figure 2. Anode core samples locations.

Results and discussion X-Rav computed tomography Only two of the three initial core samples have been analyzed (core #1 and #2). The CT scan results obtained with the core #1 are shown in Figures 4 to 6. Due to the large number of images, only some of them are shown in these figures. Each image on Figure 4 and 5 is separated by a distance of 7.8 mm while those of Figure 5 are separated by a 11.4 mm distance. Similarly, CT scan results of core #2 are shown in Figures 7 to 9. Densities are calculated based on the entire volume of the samples rather than on one image.

Figure 3. Location of the cores samples of different sizes, (green: ö 50.8 mm, yellow: ö 152.1 mm, red: ö 292.1 mm). Anode apparent density measurement The apparent density was measured according to the ISO 129851:2000(E) standard method. The samples were weighed with a precision balance (Sartorius CPA16001S) and their dimensions were measured using a CMM DEA-0101. Due to the presence of the stub holes the geometry of the 292.1 mm core samples was not regular enough to calculate its volume with high precision. Therefore, the apparent densities were measured only on the 152.1 mm and 50.8 mm samples.

Figure 4. CT images of core #1 (ö 50.8 mm).

X-Rav computed tomography The X-Ray computed tomography is a method widely covered in the literature and thus will not be detailed here. In summary, this method gives the so-called CT number expressed in Hounsfield Units (HU) which is related to the X-Ray attenuation coefficient [2] and ranges from -1000 for air to +3000 for very dense materials such as metals. The density of materials with low atomic number (e.g. carbon) can be estimated by assuming a linear relation between the CT number and the density [5]. In that case, the calibration can be performed assuming that the density of air (CT=-1000) and that of water (CT = 0 ) are 0 g/cc and 1 g/cc,

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Figure 8. CT images of core #2 (ö 152.1 mm).

Figure 5. CT images of core #1 (ö 152.1 mm). Figure 9. CT images of core #2 ( ^ 292.1 mm).

Figure 6. CT images of core #1 (^292.1 mm).

By comparing the images obtained from samples with different sizes, the effect of size on the resolution of CT scan images can clearly be seen. Aggregates and porosities are clearly visible in the small sample (Figures 4 and 7) while they are not revealed in the larger one. This is due to the fact that the voxel volume increases by increasing the sample size resulting in lower resolution. To quantify this effect the CT number distributions were calculated for three measurements of the core #1 and shown in Figures 10 to 12. The average values and standard deviations of CT number obtained for core #1 and core #2 are summarized in Tables II and III, respectively. The high standard deviations could be related to the high porosity volume fraction of material (approximately 20-25%) and to the presence of the very dense impurities with much higher CT number than that of carbon. These impurities are represented by the white spots on the CT images (Figures 4 to 9). The standard deviation decreases by increasing the sample size from 50.8 to 152.1 mm. This is essentially due to the increase of the voxel size. The unexpected increase in standard deviation of the CT number for the 292.1 mm samples is however attributed to the presence of slots in these samples. In spite of the fact that a smaller standard deviation is expected for the large samples (larger voxels), the presence of the slots introduces a large number of low density voxels, leading to increased standard deviation of the whole data set. The average CT number decreases slightly (approximately 1%) by increasing the sample size suggesting that the estimated density decreases as the core size increases. Since the length of the largest diameter samples was greater than that of the smaller samples (Figure 3), the observed variation in density could also be attributed to the variation of anode density along the sample rather than only related to the sample diameter. In both cases, the density gradients within the anode and more particularly under the stub holes could explain the slight variations.

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From the CT numbers, the density was estimated according to equation (1) and the results were plotted in Figure 13 as a function of sample size. The estimated density was then compared to the measured apparent density as shown in Figure 14. It should be emphasized that the apparent density of the 292.1 mm samples was not measured. It can be seen that the estimated values differ significantly from the measured ones suggesting that the equation (1) does not result in accurate density estimation. The estimated apparent density is approximately-10% lower than the measured

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as great as 3000 HU. Thus, by using equation (3), a porosity volume fraction of 60% is obtained. This value is definitely inaccurate for the carbon anode. To overcome this problem one may consider limiting CTmax to 1000 HU, i.e. a value close to that of graphite (960 HU). The resulting porosity values with CTmax= 1000 HU are presented in Table IV and V. The estimated porosity correlates with the measured apparent density for all samples assuming that the theoretical density of carbon is 2.21 g/cc. Typical values are however closer to 2.1 g/cc for the anode. In the later case, the porosity would be thus overestimated by equation (3) and it is thus necessary to adjust it for the anode material. This still nevertheless suggests that limiting the CTmax to a maximum value corresponding to that of graphite eliminates the effect of non representative voxels and results in more accurate estimations.

A new calibration was therefore performed by using the air (CT=-1000) and a representative sample instead of water. The sample of 152.1 mm diameter (taken from core #1) was used as the representative sample. Recalibration of equation (1) using air and this sample resulted in the following equation: p = 0.0011 x CTnumber +1.1081

(2)

This equation is obtained using a sample taken from the anode being studied. Therefore, one could assume that it fairly represents the effect of raw materials, fabrication process and baking procedure by which all other samples have been influenced. The apparent density of the samples was calculated again using the equation 2. The new estimated values (called "calibrated densities") were compared with the measured densities in Figure 15. Regardless of the sample diameter, the gap between the measured density and the calibrated one has been reduced to less than 2%. This suggests that the apparent density could be accurately estimated using CT numbers when a similar material is used as calibration sample instead of water.

Table IV. Estimated total porosity of samples taken from core #1 % voxel > ì (HU) ö (mm) P(%) 1000 HU 25 50.8 485 0.03 26 152.1 468 1.43 292.1 460 0.63 27 1 Table V. Estimated total porosity of samples taken from core #2. % voxel > ì (HU) ö (mm) P(%) 1000 HU 0.03 25 50.8 488 2.76 25 152.1 476 0.63 26 292.1 467

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Three samples were initially cored from a full size prebaked carbon anode in order to calibrate the relation between the X-Ray tomography intensity and the apparent density. From each main core sample, two smaller samples were cored to reveal the size effect on the average CT numbers. It appears that no size effect is required be taken into account.

300

The apparent density has been determined as a function of the CT number. Instead of the water, the sample of ö 152.1 mm of core #1 was used for calibration. The gap between the measured density and the estimated one was reduced from 10% to less than 2%. Future CT scans on different prebaked carbon anodes have to be performed to expand the validity of the new relationship within the range of anode process variations. The porosity has also been estimated using a best fit relationship. The estimated porosity tendency correlates well with the measured density of the samples.

Sample Diameter (mm)

Figure 15. Comparison between the calibrated and the measured apparent densities. The CT technique could also be used to estimate the pore volume fraction in porous materials. According to Boespflug et al. [5] the porosity of a homogeneous material could be estimated by the following relation: CT+1024

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Acknowledgements Where p is the total porosity volume fraction in %, CT is the average CT number in HU, CTmax is the maximum CT number in HU and 1024 is the CT number of water in HU. The main issue with anode materials is to determine the CT^. The prebaked anode contains not only carbon, but also other impurities with very high voxel intensity. So, only one voxel related to the impurities would be sufficient to result in an overestimated porosity percentage of the sample. In the prebaked anode studied in this work less than 1% of the voxels have intensity greater than 1000 HU. Some voxels reach however very high intensity values

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The authors gratefully acknowledge the financial support provided by Alcoa Inc., by the Natural Sciences and Engineering Research Council of Canada and the technical support of the Aluminium Research Centre - REGAL.

References 1. 2.

3.

4.

5.

Frosta, O.E., Foosnaes, T., 0ye, H.A., and Linga, H., Modelling of anode thermal cracking behaviour. Light Metals 2008, p. 923-927. Adams, A.N., Karacan, O., Grader, A., Mathews, J.P., Halleck, P.M., and Schobert, H.H., The non-destructive 3-D characterization ofpre-baked carbon anodes using X-ray computerized tomography. Light Metals 2002, p. 535-539. Suriyapraphadilok, U., Halleck, P., Grader, A., and Andresen, J.M., Physical, chemical and X-Ray Computed Tomography characterization of anode butt cores. Light Metals 2005, p. 617-621. Duchesne, M.J., Moore, F., Long, B.F., and Labrie, J., A rapid method for converting medical Computed Tomography scanner topogram attenuation scale to Hounsfield Unit scale and to obtain relative density values. Engineering Geology, 2009. 103(3-4): p. 100105. Boespflug, X., Ross, N., Long, B., and Dumais, J.F., Axial Tomodensitometry - Relation between the Ct Intensity and the Density of the Sample. Canadian Journal of Earth Sciences, 1994. 31(2): p. 426-434.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

FEM ANALYSIS OF THE ANODE CONNECTION IN ALUMINIUM REDUCTION CELLS Susann Beier 1 , John J. J. Chen 1 , Hugues Fortin 2 Mario Fafard 3

4

1

Light Metals Research Centre, University of Auckland, Auckland 1142, New Zealand Institut de recherche Hydro-Quιbec - LTE, 600 avenue de la Montagne Shawinigan, Canada, G9N 7N5 3 NSERC/Alcoa Industrial Research Chair MACE Aluminium Research Centre - REGAL, Laval University, Sciences and Engineering Faculty, Quιbec, Canada, G1V 0A6 2

Keywords: anode connection, yoke bending, stub deterioration, stub diameter, stub replacement Abstract Achieving voltage savings over the anode assembly in an aluminium reduction cell, particularly at the anode connection, is a worthwhile approach within a wider programme of improvement in energy efficiency. Experiments carried out using operating cells are very difficult and expensive, however finite element method (FEM) simulations as used in this study are a cost efficient and accurate method to understand the behaviour of the anode connection, and to identify the constraints to voltage savings. This study investigates the impacts of stub deterioration and yoke stiffness on the anode connection and hence the performance of the anode assembly. An ideal stub diameter for the investigated configuration was found, and the increased voltage drops for various level of stub deterioration were identified. The results show that a yoke cross-bar with reduced height and hence reduced stiffness decreases the tensile stress developed in the carbon anode, which lowers the risk of anode cracks. A limit for stub service life is suggested, showing a potential saving of US$0.8m annually. Introduction The electrolysis process for aluminium production consumes the carbon anode block. After the anode is spent, the remaining 'butt' of the anode has to be replaced with a new anode. The anode assembly must be taken out of the cell, and carbon butt and cast iron thimble are stripped off the stubs. The steel yoke assembly is then inspected for damage, and re-used if the level of damage does not exceed specified smelter limits. New carbon anode blocks are connected to the anode assembly using molten cast iron, which is poured into the spatial gap between the stubs and anode stub holes. The hot cast iron heats up the steel stubs, causing thermal expansion, giving a shrink fit of the cast iron onto the

stubs. The cast iron shrinks when it solidifies, causing an air gap to develop at the cast iron - carbon interface. This air gap is critical to anode performance, as it increases the contact resistance of the anode connection. Research has shown that a voltage drop of 80 to 130 mV appears at the interface from stub to cast iron to carbon [1]. More than 90% of this is due to the cast iron to carbon interface [2]. The aggressive environment of high temperature and electro-chemical attack in the reduction cell can also cause deterioration and deformation of the anode assembly. The degree of stub deterioration and yoke bending are the major factors that determine the performance of the anode connection, influencing the stress, voltage and temperature distribution significantly. The cast iron thimble formed on a new, non-deteriorated stub is relatively thin, and hence the shrinkage of cast iron is small, giving a narrow air gap at the cast iron - carbon interface. The small air gap results in rapid tightening of the connection when the anode heats in the cell. This saves early-cycle voltage losses, but also causes higher stresses in the carbon anode and greater risk of harmful vertical anode cracking. Conversely, when deteriorated stubs are re-used, more cast iron is required at the connection to fill the anode stub hole, giving increased cast iron shrinkage and hence a larger air gap. This air gap creates high early-cycle voltage losses, as the air gap closes later in the anode cycle. The stress in the carbon anode is significantly lower however, and the risk of anode cracking is decreased. The high temperature environment in the aluminium reduction cell causes significant thermal expansion in the yoke. For a three-stub assembly geometry as studied in this work, the yoke expansion causes a bending moment and hence radial pressure increase at the outer stubs. This can have an significant impact on the stress development in the carbon anode block and needs to be further investigated [3].

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Research A i m The cast iron - carbon interface at the anode connection has a thermo-electro-mechanical interrelated behaviour which is strongly influenced by stub deterioration and yoke bending. The investigation of this behaviour by in situ plant measurements during cell operation is a very difficult and expensive approach. This study focuses on computational FEM simulation of an anode assembly, presenting a cost efficient and qualitative approach. The following key research questions were developed, which form the main objectives of this study:

FEM solver for Hall-Hιroult aluminium reduction cells developed by REGAL [6]. FEM Model The FEM model of an anode assembly consists of an aluminium rod, a steel yoke, three aligned steel stubs with cast iron thimbles, and the carbon anode block as shown in Figure 1. The models dimensions and design are widely used in industry and are based on Fortin et al. [9].

• What is the ideal ratio of stub diameter to cast iron thickness for the investigated configuration, to prevent anode cracks and give maximum voltage savings at the anode connection? • How significant is the impact of stub deterioration on the anode connection and hence on the anode assembly performance? • Does yoke bending increase tensile stresses in the carbon anode, causing increased risk of vertical anode cracks? • How long is the ideal stub service life before replacement for greatest economic benefit? Figure 1: Generated FEM anode assembly model

Previous Work This research is based on the work by Richard et al. [4] who developed the first weakly thermo-electricalmechanical (TEM) coupled FEM model, which contributed to the development of a constitutive cast iron carbon resistance model [5]. Goulet [6] further improved the TEM coupling, and implemented the FEM solver F E S h + + for Hall-Hιroult Aluminium reduction cells. This work was progressed by Fortin et al. [7] and Richard et al. [8]. Fortin investigated the impact of stub deterioration by varying the stub shape, showing that the outer stub holes of a three stub anode assembly have more impact on the anode performance than the middle stub. Richard simulated and compared anode assemblies with a different number of cast iron flutes, showing that the benefit of an increased number of flutes is limited. The anode assembly model presented and studied in this work is based on the simulation code used by Fortin and Richard, provided by the Aluminium Research Centre REGAL from Laval University. The anode assembly code was enhanced in APDL (ANSYS Parametric Design Language), and the simulation was solved with F E S h + + , the

The main model dimensions are: anode width 700 mm, anode length 1500 mm, anode height 650 mm, stub centre distance 500 mm, stub diameter 190-160 mm, stub hole diameter 209 mm, stub hole depth 130 mm, yoke crossbar width 120 mm, yoke cross-bar heights 93 and 100mm. Various boundary conditions were applied to represent the operational conditions in an aluminium reduction cell:

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• Immersion height of the anode into the bath of 400 mm. • Hot gas immersion from the bath of 150 mm height. • Height of cover and crust layer of 350 mm. • The bottom face of the anode block is defined as zero voltage. • The applied current density is dependent on rod diameter and assumed a total line current of 330 kA and 40 anodes per cell. • Gravity and the buoyancy force are applied to the anode block.

• Several convective and radiative heat fluxes were applied over the anode assembly taking into account the hot air, crust, gas and bath conditions as well as anode position in the cell [6, 9]. Air Gap Prediction and Plant Measurement Comparison

thickness, and Richard's method is only valid for a certain range of cast iron thickness. Simulations with various air gap sizes were developed and solved, and compared to in situ measurements from industry. It was found that a reduction of 13.5% of the initial calculated air gap, as given by Equation (4), best matched actual measured voltage drops from industry.

The air gap developed at the cast iron - carbon interface depends on various factors which are not yet fully understood. Richard [8] proposed a comprehensive method to predict the air gap. This validated simplified analytical method is given in Equations (1) to (3). *gap

= 75 + (* - 75) · dei · (Tei - T0) 75 = rs · as · (Ts ~ TQ) t=-7jT-rs

Where: Sgap

is t

=

Air gap size at To, mm

= Change in radius of the steel stub, mm

OLCI

=

Tei To rs as Ts

= = = = =

dsh

=

Stub hole gap, mm Thermal expansion coefficient of cast iron, Temperature of cast iron, °C Ambient temperature, °C Stub radius, mm Thermal expansion coefficient of steel, ^ Temperature of steel stub, °C Stub hole diameter, mm

Sgap flutes

Where: Sgap flutes =

= Sgap ' 0 . 8 6 5

(4)

Air gap size at cast iron flutes, mm

The air gap used in the FEM model at the cast iron body (1) was calculated by Richard's method, and the air gap at the cast iron flutes was reduced as given by Equation (4). (2) This approach gave the best agreement with the temperature and voltage measurements of 10 anode assemblies taken in Smelter A ' . Temperature measurements were (3) taken at 14 points over the anode assembly, with voltage drops taken at 9 different measurement points. All predicted model results fell within the range of both the temperature and voltage measurements. For reasons of confidentiality no further information can be published on the plant measurements. M o d e l Application

The initial air gap size sgap is calculated for the ambient temperature TQ as shown in Equation (1). The cast iron solidification temperature Tei is assumed to be 1200°C with the ambient temperature To of 20° C and the thermal expansion coefficient of cast iron aci of 9 · 1 0 - 6 - ^ . The change in radius of the steel stub is described by 75 in Equation (2). It is determined by the ambient temperature To, the effective stub temperature when cast iron solidifies Ts at 200°C [10], the stub radius rs and the thermal expansion coefficient of steel as of 1.5 · 10~ 6 ·^. The change in spatial gap, £, between stub hole and stub is given by Equation (3). Limitations of this method were found for cast iron configurations with flutes, where the prediction method results in a too large air gap for the large cast iron thickness at the flutes. Cast iron shrinkage appears non-linear to its

Two case studies were performed: firstly, the base model was modified by using different stub diameters of 190, 180, 170 and 160 mm with a constant stub hole diameter of 209 mm, representing different stages of stub deterioration. The different stub configuration models were solved and compared with respect to temperature, stress, voltage and current distributions. These models allow conclusions to be drawn on the influence of re-usage of deteriorated stubs. Secondly, each of the stub configuration models were solved using a reduced yoke cross-bar height of 93 instead of 100mm. The reduced height results in a decreased second moment of area, and reducing flexural stiffness by 20% as shown by Equations (5) and (7). w

Jx

1

=

Jx 2 =

981

•/i3

12

(5)

120 · 1003 — = 10,000,000 12

(6)

120 · 93 3 = 8,043,570 Ο2~

(7)

Where: Jx = Second moment of area, ram4 w = Width of the sectional area, mm h = Height of the sectional area, mm

Q = I2RE

The reduced flexural stiffness results in greater bending of the yoke, causing increased radial pressure at the outer stubs and stub holes. The comparison of various stub configuration models with 93 and 100mm yoke height clarified the effect of yoke stiffness on the anode connection. These predicted model results allowed development of a cost analysis, where voltage drop increase due to stub deterioration, anode cracking, material and replacement costs are taken into account. Potential voltage savings at the anode connection are identified, and suggestions for an ideal stub service life are made.

The cast iron temperature increased by 13% with stub deterioration by 16%. This temperature is conducted to the cooler upper parts of the anode assembly, and the connection, stubs, yoke and rod all show a slightly increased temperature as shown in Figure 3. The higher temperature of the materials leads to slightly higher electrical resistance causing a further increased voltage drop as shown in Figure 2. 700 -r 600 -E 0 500 {

·.· ';:

r

£. 400 -P g

j

T

"

w. 140

>

;

r

< 1 ~» Rod

!

f^**^^^T_^

åd 120 O 100

""" """j

o 80

J2

60

0

===4 -♦•Yoke

\ -*" Carbon around stub holes

250 ·§ 200 ■£

160

170 180 Stub diameter; mm

190

Figure 3: Temperature prediction for stub configurations The highest stresses in the carbon anode was found at the stub hole flutes. The local stresses were very high, and decrease rapidly from the carbon top to the bottom. For a stub diameter of 190mm with a cast iron thimble of 9.5 mm thickness, the local stress predicted is very high and likely to exceed the carbon anode stress limits. The local stresses decreased strongly by 33% with a 16% decrease in stub diameter as shown in Figure 4.

I ~*~-Cast iron

z'.'.'.'.'.'.'..i

§ 40 20

350 f:::

:::::::::::l ~&~ Stubs

"¼

60

\ ~*-Cast iron

300 -p

■'■'

"~~"~*-*~*-**».~_

160

j -*- Stubs

1 450 | E

The simulation of anode assemblies with various stub configurations showed that the normal stress at the interface is significantly lower for smaller stub sizes when the cast iron thickness and hence the air gap is increased. The increased air gap however causes a higher contact resistance and higher voltage drop as shown in Figure 2. A voltage drop increase of 68% was found at the anode connection for a stub diameter reduction of 16%. !

■+ Yoke

V 550 -j|

Stub Deterioration

180

-* Rod

650 f

Results and Discussion

200

(8)

t "'

WWV»-«...».«..«.,^.^ ,..„

360

».«,.^,.^^ö^,«

Carbon

■ ^ - 4 v ~ - - ^.v4

* "~ 165

170

175

11Ì5 180

190

Stub diameter, m m ^TLL

Figure 2: Voltage prediction for stub configurations The larger air gap causes low contact pressure, which increases the electrical resistance RE- This leads to high local current density over the actual contact surface. The high current density / increases the ohmic heat Q generation as given by Equation (8).

982

170

• j — | -*- Minimum HzzI stress at flute bottom

180

Stub diameter, mm

Figure 4: Predicted tensile stress in stub hole flutes

Additional material and operational costs are incurred due to earlier stub replacement. The material cost of a Simulation models were developed with various stub di180 mm steel stub was assumed to be US$47. With an ameters and yoke cross-bar height reduced from 100 to average stub life cycle of 15 anode cycles and a replace93mm, giving 20% reduction in yoke stiffness and hence ment cost of US$51.3, additional annual costs due to earincreased yoke bending. The simulation results for the lier stub replacement accrue as a function of the specified various stub diameters from 190 to 160mm were comminimum stub diameter before replacement. pared, and no significant differences in voltage or temperIn addition to these costs, the risk of anode cracks was ature distribution were found with reduced yoke height. also taken into account. The maximum stress developed The tensile stress strongly reduced in average in the carin the anode block was found from the simulation results, bon flutes however, by around 22% as shown in Figure while a stress limit for the carbon anode was assumed. 5. It was seen that the increased radial pressure due to The local stress was used to find a percentage increase increased yoke bending reduced the tensile stress at the in voltage drop, assuming the maximum voltage drop octop of the stub hole, and increased the compressive stress curs in the case of an anode breakage when the maximum at the bottom of the stub hole. The stress limit of carstress limitation is reached. The maximum voltage drop bon anode is significantly higher in compression than in due to an anode breakage was simulated, and the voltage tension [11]. This results in a lower risk of anode cracks drop for lower local stress values was found assuming a formation for the 180mm diameter stub configuration. It linear relationship due to partial anode cracking. These can be concluded that increased yoke bending and the radial pressure increase has a positive impact on the stress additional annual costs were added to the calculated cost due to stub replacement, and are also shown in Figure 6 development in the anode carbon block. as dashed line. The cost due to increased voltage drop outweighed the cost caused by stub replacement and risk of anode cracking below a stub size of 139mm, equivalent to 41mm or - - 100mm 28% deterioration in stub diameter. ; yoke 7~T~~j~~T~X" I ^*~«~*^~*'^ ~ „Ì~"""" ■ ~ \~\ ...i.j height A common smelter procedure found in an industry surS 6 vey is to replace a deteriorated stub after approximately -*- 93mm •HT" I l ""."'; ; I I '. '."V V I I'.' I I '. '.""] '.' I ", ' Γ Ú . 1 28% diameter reduction to 130mm [13, 14]. This study yoke height suggests replacement after approximately 23% reduction to 139 mm, giving US$800,000 annual saving in voltage drop for a smelter with 264 pots operating.

Yoke Stiffness

160

165

170

175

180

185

190

50,000

Stub diameter, mm

45,000

—Cost due to additional voltage drop caused by stub deterioration — Cost due to stub replacement and anode cracks

40,000

Figure 5: Predicted tensile stress in the carbon anode flutes

35,000 30,000 25,000 20,000

Cost Analysis

15,000

The lowest anode voltage drop found without exceeding carbon stress limits was found for the configuration with 180mm stub diameter and 93mm yoke height. This is taken as the initial configuration in the cost analysis, which assumes a operation at 330kA with US$0.044 costs per kWh and uses various assumption based on smelter cost statistics as discussed by Beier [12]. The simulation results showed that the voltage drop increases with decreasing stub diameter, giving additional annual operational costs as shown in Figure 6 as solid line.

10,000 5,000 0 130

140

150

160

170

180

Stub diameter, mm

Figure 6: Comparison of annual cost per pot by voltage drop against cost by stub replacement and anode cracks

983

[5] D. Richard, M. Fafard, R. Lacroix, P. Clery, and Y. Maltais, "Carbon to Cast Iron Electrical Contact Resistance Constitutive Model for Finitie Element Analysis," Journal of Materials Processing Technology, vol. 1, pp. 119-131, 2003.

Conclusions In this study a FEM model of the anode assembly was used with an adjustment of the air gap prediction for a fluted cast iron connection. The adjusted simulation model results compared well with in situ voltage and temperature measurements from industry. Various stub and yoke configurations were subsequently investigated regarding voltage drop, temperature, current and stress distributions. It was found that reduced yoke stiffness significantly decreased the tensile stresses in the carbon anode. For the anode assembly design investigated, the configuration with 180 mm stub diameter and 93 mm yoke crossbar height gave the best compromise between low voltage drop and stress development below the critical carbon anode tensile stress limit. A cost analysis based on the simulation results showed that the cost due to increased voltage drop caused by stub deterioration is very high. A common smelter practise is to replace stubs after approximately 28% deterioration, while this analysis however showed that earlier replacement after only 23% stub deterioration could lead to significant annual savings due to reduced voltage drop. Acknowledgement I would like to thank Olivier Trempe and the Aluminium Research Centre - REGAL, Laval University, Quιbec City, Stew Woodward (Rio Tinto Alcan, NZAS), Barry Sadler (Net Carbon Consulting Pty Ltd) and Michael Gasse (Alcoa, Deschambault) as well as Pascal Lavoie and the rest of the Light Metals Research Centre team for your collaboration, advice and support. References [1] W. T. Choate and J. A. S. Green, "U.S. Energy Requirements for Aluminium Production," 2003. [2] J. Wilkening and S. Cote, "Problems of the StubAnode Connection," TMS Light Metals, pp. 865-873, 2007. [3] N. A. Ambenne, "Vertical Anode Cracking - The VALCO Experience," TMS Light Metals, vol. 1, pp. 577-583, 1997.

[6] P. Goulet, Modιlisation du comportement thermoιlectro-mιcanique des interfaces de contact d'une cuve de Hall-Hιroult. PhD thesis, Laval University, Quιbec, Canada, 2004. [7] H. Fortin, M. Fafard, N. Kandev, and P. Goulet, "FEM Analysis of Voltage drop in the Anode Connector Assembly," TMS Light Metals, vol. 1, p. 6, 2009. [8] D. Richard, P. Goulet, O. Trempe, M. Dupuis, and M. Farad, "Challenges in Stub Hole Optimisation of Cast Iron rodded Anodes," TMS Light Metals, vol. 1T p. 6, 2009. [9] H. Fortin, "Modιlisation du comportement thermoιlectro-mιcanique de l'anode de carbone utilisιe dans la production primaire de l'aluminium * Master's thesis, Laval Univeristy, Quιbec, Canada, 2010. [10] O. Trempe, D. Larouche, D. Ziegler, M. Guillot, and M. Fafard, "Real time temperature distribution during sealing process and room temperature air gap measurements of a Hall-Hιroult cell anodeT" TMS Light Metals, 2011. to appear. [11] O. E. Frosta, T. Foosnaes, H. A. 0ye, and H. Linga, "Modelling of Anode Thermal Cracking Behaviour," TMS Light Metals, vol. 1, pp. 923-927, 2008. [12] S. Beier, "A Study of an Anode Assembly with Focus on the Anode Connection used in Aluminium Reduction Cells ," Master's thesis, Auckland Univeristy, July 2010. [13] S. Woodward, "Personal communication," from Rio Tinto Alcan, New Zealand Aluminium Smelters (NZAS), 2010. [14] M. Gasse, "Personal communication," from Aluminerie Deschambault Quιbec, Alcoa, through the Aluminium Research Centre - REGAL, Quιbec, Canada, 2010.

[4] D. Richard, M. Fafard, R. Lacroix, P. Clery, and Y. Maltais, "Aluminum Reduction Cell Anode Stub Hole Design using weakly coupled Thermo-ElectroMechanical Finite Element Models," Finite elements in analysis and design, vol. 37, pp. 287-304, 2001.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DEVELOPMENT OF INDUSTRIAL BENCHMARK FINITE ELEMENT ANALYSIS MODEL TO STUDY ENERGY EFFICIENT ELECTRICAL CONNECTIONS FOR PRIMARY ALUMINIUM SMELTERS D. Molenaar1, K. Ding2 and A. Kapoor2 1 CSIRO. Melbourne, Australia. 2 Swinburne University of Technology. Melbourne Australia Keywords: contact resistance, contact pressure, stub to carbon voltage drop, cast iron, anode carbon, anode assembly Abstract Process improvements of 5 MW per plant (50,000 t C02e pa for coal based electricity) are possible through optimisation of the complex cast iron to carbon contacts within aluminium smelter anode and cathode assemblies. Finite element analysis is considered the tool of choice within industry for assessing potential improvements; however there are limitations with existing models regarding handling of contact resistance and carbon stress state. A study has been undertaken using thermoelectrical-mechanical finite element analysis of the cast iron to carbon contact for an anode assembly. The contact pressure and electrical resistance and its dependence on temperature have been derived from data available in the public domain. This paper presents development of the benchmark model including results. The benchmark model will be used as the reference point for the development of more advanced models in ongoing studies to assist primary aluminium smelters achieve these substantial savings in energy efficiency and reduced greenhouse gas emissions.

Figure 1. Components of a typical anode assembly They require many months preparation and will generally mandate a very high number of sample repeats to form a statistically valid result due to the inherently high levels of variation within and between anode assemblies. The second method is the off-line experimental laboratory. Major issues with this approach are; (a) significantly reduced current density making it very difficult to detect small changes in contact resistance, (b) heating and protection of the anode carbon from combustion during testing and (c) it is common to employ smaller portions of an actual anode in the test rig resulting in non-representative overall geometry which will cause significant alterations to the constriction of current through the portion of carbon being tested. Also, in the latter case the smaller sized carbon may not withstand the stress generated from differential thermal expansion and will likely cause the carbon to crack, invalidating the test data. It is not practical or efficient to test entire anode assemblies in an offline laboratory. The third approach employed is finite element analysis (FEA) and this is considered the tool of choice within industry for assessing these potential improvements [3-6]; however there are numerous limitations with existing models, the two most important ones being the handling of contact resistance and carbon stress state. The purpose of this paper is to describe in detail the development of a benchmark finite element model using a proper analysis procedure to achieve a thermo-electrical-mechanical analysis of the cast iron to carbon contact for an anode assembly. This benchmark model will be used as the reference point for the development of more advanced models in ongoing studies towards assisting primary aluminium smelters achieve substantial savings in energy efficiency and reduced greenhouse gas emissions.

Introduction Aluminium smelters operate at currents in the range of 100-400 kA DC. Most reduction cells operate at approximately 4.5 volts DC each and, as the theoretical reduction of alumina requires about 1.8 volts DC, it is clear that there are significant energy losses in the process. Of the excess 2.7 volts, power losses associated with electrical conductors and connections represent approximately 0.2-0.4 volts and a significant portion of these losses are contained within the complex cast iron to carbon contacts within aluminium smelter anode and cathode assemblies [1]. Figure 1 shows the components of a typical anode assembly used within the aluminium smelting process comprising (1) aluminium rod, (2) aluminium-titanium-steel transition joint or clad, (3) steel crossbar or yoke, (4) steel stubs, (5) cast iron thimbles and (6) carbon anode. The anode assembly shown in Figure 1 has been specifically designed for this study. Of particular interest in this paper is the electrical connection system of the steel stub to cast iron thimble to anode carbon, referred to as the stub to carbon (STC) connection. There are generally three methods employed by industry to study potential energy efficiency savings that may be present in the anode assembly STC connection. The first has historically been to undertake extensive in-plant trials of suitably instrumented anode assemblies set into actual reduction cells and monitored over the entire period of operation [2]. The operating environment within a reduction cell offers many technical challenges to overcome in order to ensure that a measurement system is sufficiently robust and will generate valid data throughout the measurement campaign. In-plant trials are logistically demanding, expensive and labour intensive.

985

Development of Contact Resistance Relationship To be able to utilize fully coupled thermo-electrical analysis in the finite element model, it is first necessary to define the contact resistance as a function of both temperature and contact pressure. The basis for developing the required relationships draws upon the original work undertaken by Rhedey and Castonguay [7] in which a chart is presented showing the impact on contact resistance for a steel-carbon system as a function of contact pressure. In this study the original data set was divided into three distinct pressure regions being (A) low, (B) medium and (C) high. Basic curve fitting techniques were used to define a suitable set of equations that adequately calculated the effect of contact pressure on contact resistance for all temperatures and pressures expected within the model. The following overarching set of equations was found to describe the relationship within the temperature range of interest between 0-950°C: Region A (0.00029 MPa < P < 2.4525 MPa) CRa=ao*exp(a1*T)*(P*100/9.81)A(a2*T2+a3*T+a4)

(1)

Region B (2.4525 MPa < P < 9.81 MPa) CRb=(b0*Tz+b1*T+b2)exp(b3*T/+b4*T+b5)(P*100/9.81)

(2)

Region C(P> 9.81 MPa) CRc=(c0*T2+c1*T+C2)expl00*(c3*T2+C4*T+c5)

(3)

Detailed Model Geometry A generic three dimensional anode assembly model is constructed and run over several orders of magnitude scale, for example; the domain is approximately 3 m x 2 m as shown in Figure 3 and Figure 4, but the radii at the base of the stub hole and cast iron thimble are 3 mm as shown in Figure 5. This imposes considerable density of meshing issues at the radii to ensure that all elements maintain a proper aspect ratio and will not impact convergence. To more accurately model the electrical and thermal conduction through weldments, the components of assemblies being joined by each weldment were separated by 1 mm to ensure electrical isolation, forcing the current to flow through all welds as is observed in reality. ;

1"Γ*|==^ï|

If P < 0.00029 MPa, CR will be very high = 2000 ohm.mm2. Where: T - mean temperature at the contact surface (°C) CRj - contact resistance (ohm.mm2), i=a,b,c io, ii, Ì2> 13, U and i5 - material constants obtained through curve fitting Figure 2 shows the contact resistance relationship defined by equations (1), (2) and (3) which are presented as a function of both contact pressure and temperature between 0-950°C.

« Γ ~ ^ - 20 X 45.0" Chamfer

4 8 X 4 5 . 0 » Chamfer

20 X 45.0° Chamfer

Figure 3. Full anode assembly dimensions

Contact Pressure (MPa)

Figure 2. Chart of contact resistance relationships used in the subroutine with overlay of original data from [7]

Centre Channel

ade Channel

Figure 4. Key dimensions for half rota operation

986

each nodal temperature history in the anode assembly from step 1 as the thermal inputs; Step 3: Thermal-electrical analysis was performed using the ABAQUS user defined subroutine in terms of a function of temperature and contact pressure. The contact pressure was read from the previous thermo-mechanical analysis (step 2). TWmbte Cross Section

Software and Hardware All analyses were run on a personal computer. This was done in order to assess the performance statistics of a PC available as a simple facility in the industry work environment in handling such computationally intensive tasks. The details of the computer configuration and the computational software used for this project are shown in Table I.

C < TWmbfe Vertical Cross Section

Figure 5. Cast iron thimble dimensions

Table I. Computational configuration and environments Hardware

Modelling Technique The STC electrical contact resistance is defined as a set of functions of temperature and contact pressure using equations (1), (2) and (3). The functions are programmed into a user defined subroutine for the ABAQUS modelling software, which specifies the gap conductance for the STC contact surfaces. The thermomechanical modelling process deals with the air gap change in expansion and shrinkage and contact pressure in the STC interfaces for both the casting operation and the cell operation. In a coupled thermal-electrical modelling process, the nodal temperature becomes a known variable while the node contact pressure can be read in from an external file generated from the thermo-mechanical modelling process. This approach allows the electrical contact resistance for the STC interfaces to be automatically determined within the ABAQUS subroutine and thus no calibration or manual adjustment is required.

Software

Model Setup The FEA model used 4 or 8 node brick continuous elements. The FEA mesh was kept the same, but the element type was in accordance with each analytical scheme. For heat transfer analyses, DC3D8 and DC3D4 (linear bricks) were used. For thermo-mechanical analyses, C3D8R and C3D4 (linear brick) were used. For thermal-electrical analyses, DC3D8E and DC3D4E (linear bricks) were used. The details of element and node numbers of the FEA model for each part are shown in Table II. The actual models are shown in Figure 7. Table II. Details of FEA mesh model

Modelling Flow Chart There are three main steps in the modelling procedure, shown as the flow chart of Figure 6.

Anode assembly parts Hanger (Al rod, steel, weld) Anode Thimble (total 4) Cover Bath Total

Single Anode Assembly

m

Node numbers 15646 15965 5856 9126 2815 49408

L^:

Cell Operation Modelling

(uncoupled heat transfer analysis)

Resultant temperature profiles of anode assembly for step 1

Element numbers 25730 50271 17388 9065 1584 104038

constant value

FEA Models Casting Modelling

HP Z400 6-DIMM Workstation. Intel Xeon W3580 3.33GHz CPU-8MB of RAM ABAQUS 6.9 -EF, Microsoft Visual Studio 2008, and Intel® Visual Fortran Compiler Professional Edition 11.1 .054 for Window XP 64bits

(fully coupled thermal-electrical analysis)

Sequentially Coupled Thermal-Stress Modelling Step 1 - Casting Process Step 2 - Cell Operation Process

Resultant temperature profiles of anode assembly for step 2

New gap electrical conductance (resultant contact pressure-electrical resistance relationship subroutine)

X

Cell Operation Modelling (fully coupled thermal-electrical analysis)

I

Final Results

Figure 6. Flow chart of modelling procedures Step 1: Transient thermal analysis and thermal-electrical analysis were performed to determine the temperature profiles and historic data at each node of anode assembly in the casting operation and the cell operation (at half rota), respectively; Step 2: Thermo-mechanical analysis was undertaken to obtain the main output of contact pressures on the interfaces between cast iron thimble and carbon in the stub holes, reading the results of

(a)

(b)

Figure 7. (a) Sequentially coupled thermo-mechanical and uncoupled heat transfer models; (b) fully coupled thermalelectrical model

987

Material Properties The FEA model includes 8 kinds of materials specified as follows: • Stub, crossbar and steel weld: SAE1020 • Aluminium rod: similar to A16061 • Aluminium weld: similar to A14043 • Transition joint includes three parts of material: SAE1020, A14043 and CP-Ti • Thimble: cast iron (grey cast iron) • Anode carbon: baked carbon material - special C • Cover: 50% alumina + 50% solid cryolite • Bath: molten cryolite.

>1

Process Parameters and Boundary Conditions The key operational parameters are: • Anode rod current of 7500 Amps • Anode Cathode Distance (ACD) of 35 mm Bath height of 205 mm (35 + 170 mm) • Cover height profiled at 100 ± 30 mm • Half rota anode carbon dimensions of 1670 x 625 x 515 mm Bath temperature of 960 °C • Temperature of cast iron at casting of 1450 °C The boundary conditions for the thermo-mechanical analysis are addressed below: • Fixed boundary conditions were applied to the side surface of Al rod where clamped to the anode beam. • The bottom surface of the anode carbon during the thermomechanical analysis of the casting operation was supported by the ground, but during the cell operation, the vertical constraints on the carbon were released. • All contacts were applied for the 'hard' contact relationship in normal direction and a small sliding relationship in tangent direction with friction coefficient of 0.2. The contact between thimble and stub was assumed as a tied contact.

'Ère'"

Figure 8. Simulated voltage and temperature probe locations Modelling Results Model Run Time Duration Running of the model includes a very time consuming debugging process which may consume up to about 4 days to execute. This is quite a normal part of this modelling process owing to the very complex interactions between cast iron, air gap and anode carbon. Once the model has been completely debugged and convergence is possible then the final computational effort required is still quite significant in that it requires several hours to complete. To assess the impact of using multiple CPU cores on the same machine the final computational effort was performed using 1, 2, 3 and 4 CPU cores. The results are shown in Table III.

Simulated Voltage and Temperature Probe Locations Validation of the modelling will require wiring of an actual anode assembly and collecting both voltage and temperature data. This has been anticipated and so predefined locations have been established in the modelling geometry to allow reporting of key voltage and temperature data for later comparison of the actual anode assembly performance. The predefined locations for monitoring and reporting voltage and temperature data are as follows and shown in Figure 8: • STC voltage drop from the stub 50 mm above the thimble top surface to in-line with the base of the stub hole at 70 mm out from the flute base widest point. Rotated 45 degrees away from the centre of the anode. Each of the 4 stubs from side channel (S/C) to centre channel (C/C). All voltage drops are recorded from the right hand side (RHS) of the anode when looking into the cell from the side channel. • Middle anode temperature 100 mm deep from anode top surface. STC temperature exactly as per voltage probes but on the left hand side (LHS) of anode with the stub thermocouple in 25 mmfromthe surface.

Table III. Details of computational speed for each analysis

FEA analysis Heat transfer analysis Thermal-electrical analysis (first pass) Thermo-mechanical analysis Thermal-electrical analysis with subroutine

# CPUs 1,2,3,4 1,2,3,4 1,2,3,4 1,2,3,4

Clock time (hrs) 0.45,0.22T 0.17,0.14 0.20, 0.10,0.06,0.05 2.40,0.85,0.66,0.54 0.40, 0.20,0.15,0.12

Temperature Profile The resultant temperature profile for the anode assembly model is shown in Figure 9. The specific temperature probe locations data is contained in Table IV. It can be seen that the mid anode temperature at 100 mm below the carbon top surface is in the order of 560 °C by mid rota. The stubs are operating closer to 600 °C at the lower portions and around 350 °C at 50 mm above the carbon top surface. The temperature of the transition joint region is determined to be approximately 220 °C. A parameter study was undertaken for the anode cover material thermal conductivity to establish the influence on the overall temperature profile of the STC region. A node in the middle of the anode at approximately 84 mm below the carbon top surface was utilized for this parameter study. Data are presented in Table V.

988

Voltage Profile The resultant voltage profile for the anode assembly model is shown in Figure 13. It can be seen that the lines of constant voltage distort as they approach the STC connection and that there is a very significant distortion of potential around the lower corners of the stub holes.

HTIX Ö H -

+9.6004.+02 -»-9.200«+02 +«.800*+02 +8.4G0*+02 +8.000·+02 +?.600*+02 +7.200·+02 , „ +fi.800*+02 U4. +S.400»+02 +6.000«+02 , . +5.*00·+02 f-4 +5.200*+02 ' +4.800*+02 Uf+4<400*+02 +4,000*+02 , . +3.600*+02 H-+3-2û0*+M M - +2.SÛO«+02 ' ' +2.400*+02 +2.000«+02 +1.600*+02 +l.200«+02 +8.0G0«+01 +4.000«+01 +0,000«+00

Voltage and Temperature Probe Location Data The information presented in Table VI shows the voltage potential for each of the predefined locations in the anode carbon and stubs. The STC voltage drop can then be determined and is summarized in Table VII.

Midi/Cat 100 deep

ί

Figure 9. Temperature distribution in the anode assembly. Tioo=557.5 °C at mid-anode and 100 mm deep Table IV. Measured temperature points of anode carbon and stubs at mid rota Location Tioo Tel Tc2 Tc3 Tc4 Tsl

Ts2 Ts3 Ts4

Temperature [°C] 557.5 578.5 594 591 577.5 335.0 321.0 352.0 358.0

Location Descriptions Mid-anode 100mm deep Near stubl, 45°, out 70mm from flute (LHS) Near stub2, 45°, out 70mmfromflute(LHS) Near stub3, 45°, out 70mmfromflute(LHS) Near stub4, 45°, out 70mmfromflute(LHS) Stubl, 45°, up 50mmfromanode surface Stub2, 45°, up 50mmfromanode surface Stub3, 45°, up 50mm from anode surface Stub4, 45°, up 50mmfromanode surface

1. MM, *m#pι M4NH«

**> !*?*+<*&

Table V. Sensitivity study on anode cover thermal conductivity Anode cover thermal conductivity (W/m.K) 13 6 3 0.5

+&333#*βΦ

Ts4 Temperature [°C] 514 549 591 705

**>

>«+*»

Figure 10. Maximum principal stress in the carbon anode at mid rota (Maximum value at the base of stub hole = 73.7 MPa)

Carbon Stress The highest tensile maximum principal stress (73.7 MPa) is located at the bottom of the stub holes, specifically at the small curve of the C/C stub hole as shown in Figure 10. Parameter studies, altering only the Young's modulus (E) of cast iron and carbon materials, found a significant improvement in maximum principle stress in this region when using a low value of E. In the thermo-mechanical analysis, the highest tensile maximum principal stress predicted in the carbon anode would indicate a potential cracking problem; however, as a linear mechanical property is assumed in the model, the stress was deemed to be overestimated.

CPRESS +8.580«+01 +S.000*+01 +4.383#+01 +4ι16?·+01 +3. ?5Q*+01 +3.333«+01 +2.9Î7«+01 +2.500«+0l +2.0β3·+01 +1,ί.*7«+01 +1.250·+01 +Ô.333*+0O + 4 . 1 6 7* +00 +O.O00*+0O M«MI + « . 5 β 0 * + 0 1 £l«mt MtOMi 3 6 9 8 0 «*d«! 3 5 0 7

Figure 11. Contact pressure distribution in the mid stub holes at mid rota (Max. value at the base of stub hole = 85.8 MPa)

Stub Hole Contact Pressure Distribution The resultant stub hole wall contact pressure for a middle stub is extremely low as shown in Figure 11. The values here are close to zero with only a few very discrete locations of higher stress shown around the bottom edge of the stub hole wall. However, the maximum contact pressure generated is significantly larger at 85.8 MPa for a middle stub hole and over 90 MPa for an outer stub hole.

tCt>, M«9nitu4« (Av9: 75%)

Ú . .

}·#·

Current Density The current density distribution for the lower portion of the anode assembly model is shown in Figure 12. It clearly shows that the current flow is concentrated through the walls of the stub holes and not through the base.

, , »41 -

Ú

+1.337·*01 $.5β3·-0Ú +9.16 7«- 02 +S.750«-0δ 4«.333«·02 ·*7.«7·-0Ú +7.500«-02 +7.083*02 +*.δ«7·-02 +*.ÎS0«-02 +5.833·-02 +5.00«·-02 +4.593«·β2 +4Λ*7··02 +3.?5β«·02 +3,333«-»ί +2.917·¼2 +2.500«-02 +2.033·-02 41.6δ7*-02 +1.250·-02 +«.333*03 +4.1*7·-03 +Ο.0ΟΟ«·Κ)Ο

Figure 12. Current density distribution at mid rota

989

temperature are causing the very low contact pressures throughout the bulk of the stub hole wall, but it is presently unknown what is causing such extremely high contact pressures at the very bottom of the stub hole walls. Further investigation is required to resolve this issue.

EPOT

Λ

+1.241«+03 +1.086*+O3 +1.034«+03 +9.823«+0£ 4-9.307«+02 +8.790·+02 +8.2?3*+©2 +7.750*402 +7.239«+02 +6.722*+02 +6.205«+02 +5.i,89*+02 +S.l7i*+02 +4.ι54«+02 +4.l37«+02 +3.620«+02 +3.103*+02 +2.585*402 +2.Q68«+02 +1.55l«+02 +1.034«+O2 +5.17X*+01 +0.000*4-00

Voltage Profile The observed distortion of the voltage profile around the base of the stub holes is expected as it is a result of non-contact through the base of the stub holes forcing the current to flow through the walls of the stub holes. It is encouraging to see that for a four stub anode assembly that there is a relatively uniform voltage profile through the bulk of the anode. It would be expected that with fewer stubs to feed the current into the anode there would be substantially more disruption to the voltage profile in the anode. In short, a disrupted voltage profile equates to electrical inefficiency within the system.

Figure 13. Voltage potential distribution at mid rota Table VI. Measured voltage points of anode carbon and stubs at mid rota Location

Voltage [mV] 909.0 901.5 908.0 913.5 1181.0 1200.0 1197.0 1179.0

Vcl

vc2 vc3 Vo4 V.i

vs2 vs3 Vs4

Location Descriptions Near stubl, 45°, out 70mmfromflute(RHS) Near stub2, 45°, out 70mmfromflute(RHS) Near stub3, 45°, out 70mmfromflute(RHS) Near stub4, 45°, out 70mmfromflute(RHS) Stubl, 45°, up 50mm from anode surface Stub2, 45°, up 50mmfromanode surface Stub3, 45°, up 50mm from anode surface Stub4, 45°, up 50mm from anode surface

Conclusions This paper has discussed the successful development of a fully coupled thermo-electrical-mechanical finite element model and presented the results of this development. Further work to improve the materials properties data set relevant for this analysis is required. The benchmark model will be used as the reference point for the development of more advanced models in ongoing studies towards assisting primary aluminium smelters achieve these substantial savings in energy efficiency and reduced greenhouse gas emissions.

Table VII. STC voltage drops at measured points at mid rota Location

VSTCI(VS1-

Vc l )

VsTC2(Vsr V c 2 ) VsTC3(V s3 - V C3)

^^^

VsTC4(Vs4- V

Voltage [mV] 272.0 298.5 289.0 265.5

Location Descriptions Stubl to anode near stub hole 1 at 45° Stub2 to anode near stub hole 2 at 45° Stub3 to anode near stub hole 3 at 45° Stub4 to anode near stub hole 4 at 45°

Acknowledgments The authors would like to thank Dr Neal Wai Poi and CSIRO Light Metals Flagship for supporting this work and to Dr Dayalan Gunasegaram of CSIRO for trawling the public domain to provide all of the required materials properties input data. The assistance of Mr Russell Bartels of Veldor Advanced Pty Ltd is greatly appreciated in building the initial model geometry for FEA input.

Discussion Contact Resistance Relationship The form of equations was confirmed as suitable for the required analyses. The slight discontinuity between the pressure regions A and B is deemed to be acceptable.

References [1] D. Molenaar, "High Amperage Electrical Connections for the Aluminium Smelting Industry" Masters thesis, Monash University, 2003. [2] R. Peterson, "Studies of Stub to Carbon Voltage" Light Metals 1978, pp 367-378. [3] M. Dupuis, "Development and Application of an ANSYS based thermo-electro-mechanical anode stub hole design tool", Light Metals 2010, pp 433-438. [4] S. T.X. Hou, Q. Jiao, E. Chin, W. Crowell and C. Celik, "A numerical model for improving anode stub design in aluminium smelting process", Light Metals 1995, pp 755-761. [5] D. Richard, P. Goulet, O. Trempe, M. Dupuis and M. Fafard, "Challenges in stub hole optimisation of cast iron rodded anodes", Light Metals 2009, pp 1067-1072. [6] D. Richard, M. Fafard, R. Lacroix, P. Clery, Y. Maltais, "Aluminium reduction cell stub hole design using weakly coupled thermo-electro-mechanical finite element models", Finite elements in analysis and design 37, 2001, pp 287-304. [7] P. Rhedey and L. Castonguay, "Effects of Carbonaceous Rodding Mix Formulation on Steel-Carbon Contact Resistance" Light Metals 1985, pp 1089-1105.

Temperature Profile Whilst the distribution of temperature throughout the anode assembly model seems quite reasonable, it does appear that the values in and around the stub holes are low. The sensitivity study conducted on the value of thermal conductivity obtained for the cover material indicates it may be too high, allowing too much heat loss to be calculated for the top of the anode assembly. Anode Stress The anode stress is reported as very high and would certainly exceed what could be tolerated from real anode material. The sensitivity study conducted into the Young's Modulus for both the cast iron and the anode carbon show that the resultant anode stress is sensitive to this parameter and it may be that subsequent materials testing is required to confirm these values for smelter grade materials. The over prediction of anode stress is common place for FEA modelling of this assembly and this error represents one of the two major limitations of the present modelling technique and is the topic of future study in this area. Stub Hole Contact Pressure It is surprising that for the bulk of the stub hole wall area the contact pressure is almost zero, furthermore that the intensity of the contact pressure at the bottom of the stub hole wall is in fact extremely large. It could well be that the low values of

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

REAL TIME TEMPERATURE DISTRIBUTION DURING SEALING PROCESS AND ROOM TEMPERATURE AIR GAP MEASUREMENTS OF A HALL-HÉROULT CELL ANODE Olivier Trempe1, Daniel Larouche2, Donald Ziegler3, Michel Guillot2, Mario Fafard1 ^SERC/Alcoa Industrial Research Chair MACE3 and Aluminium Research Centre-REGAL, Laval University, Sciences and Engineering Faculty, Adrien-Pouliot Building, Quιbec City, Canada, Gl V 0A6 2 Aluminium Research Centre-REGAL, Laval University, Sciences and Engineering Faculty, Adrien-Pouliot Building, Quιbec City, Canada, G1V 0A6 3 Alcoa Canada, 1 blvd. des Sources, Deschambault-Grondines, Qc, GO A ISO, Canada Keywords: anode, cast iron, temperature, thermocouple, air gap, metallography, Hall-Hιroult Abstract

Measurements of the air gap at the carbon/cast iron interface; Measurements of the temperature field during the sealing process; Metallographic analysis of the cast iron; Evaluation of the heat transfer coefficients at the interfaces;

An experimental investigation of the sealing process of an anode assembly used in electrolysis cells has been performed to better define the thermal and mechanical aspects of the cast iron thimble solidification between the steel stub and carbon hole. Three holes of a baked anode have been thoroughly measured with a Coordinate Measuring Machine (CMM) to obtain a high precision three-dimensional map of the carbon interface. These measurements were then compared to the outer surface of the frozen thimbles to obtain a room temperature air gap dimension. Thirty-nine thermocouples, placed in a strategic configuration, allowed the reconstruction of the temperature field in the steel stub, carbon block and solidifying cast iron thimble from the pouring to room temperature. Hence, heat transfer coefficients can be evaluated at the carbon/cast iron and steel/cast iron interfaces with a thermal model. Metallographic analysis is matched with the cooling curves.

Air gap measurement methodology Two different methodologies, both with their advantages and disadvantages have been considered to measure the air gap. Realtime measurements with Linear Variable Displacement Transducers (LVDT) have been extensively used in experimental setups [6-12] and allows tracking the opening of the air gap as the molten metal solidifies and cools down. It is therefore an easy task to estimate the Heat Transfer Coefficient (HTC) at the interface as a function of gap opening and temperature along with thermocouple measurements so the method serves well for numerical model validation. The second method, differential measurements, doesn't yield a real time air gap magnitude but rather a final value at room temperature. Three-dimensional measurements of the stub holes and cast iron thimbles are compared to obtain the air gap. The latter method has been chosen for reasons of logistics and simplicity.

Introduction Prediction of the initial air gap in an anode is of high interest for people who aim at modeling the thermo-electro-mechanical behavior not only of an anode but also the whole aluminium electrolysis cell. It is a key parameter that defines the quality of thermal and electrical contact at the connection, thus influencing the voltage drop and current distribution in the anode and cell.

The stub holes are measured prior to casting with a coordinatemeasuring machine (CMM) by taking over 2000 points spread over equally spaced heights along the interior surface of each hole. The spatial coordinates of each point are stored in the CAD drawing file of the stub hole, indicating their position with respect to the theoretical surface. The same process is used with the frozen iron after is it withdrawn from the stub hole. The two sets of measurements are aligned with a best fit method and the air gap is computed locally from the sum of the deviation of the measured coordinates versus the CAD coordinates.

In the recent years, many attempts have been made to model the anode with generally satisfactory results [1-5] using a simple relation proposed by Richard [4] to predict the air gap : Y — rstub

x

a

steel

x

(Ja ~ ?o)

gap = y + (/ - γ) x airon x (Ts - T0)

(1)

Where, Ta is the steel temperature at the moment of the cast iron solidification, Ts is the solidification temperature of cast iron, I is the cast iron thickness and a are the thermal expansion coefficients.

CMM stub hole and thimble measurements Because of its bulky size and weight, the anode was mounted sideways next to the CMM as shown in Figure 1. An extension was used so the probe could reach the stub holes.

However, no actual measurements have been made in an industrial environment to assess the quality of the predictions made by this relation.

A parametric Computer Numerical Control (CNC) program controls the CMM and the work is done with an automated procedure. The probe has a target measuring point corresponding to the CAD surface and an approach vector which is perpendicular to the surface. It records the actual measured point and its

This current work is an effort to provide crucial information to validate a 3D thermo-mechanical solidification model of an anode sealing process. Within this scope, the following results are presented :

991

deviation from the stored drawing. Figure 2 shows a 3D representation of the measured points. The green ones are within an arbitrary tolerance compared to the CAD.

Figure 1 : Set-Up of the stub holes measurements with the CMM

analysis because of their obvious incongruity, that is, negative or excessive magnitude.

Figure 3 : Measurements of the cast iron thimble A clear trend appears as the air gap is smaller in the bottom section of the stub hole. This is in agreement with the fact that the hole has a tapered profile over the height, thus resulting in the cast iron section being thicker in the upper part of the thimble. It is therefore expected to observe a larger gap in the upper part because of thermal shrinkage. The results were compared with the predicted gap given by the Richard's relation, Equation (1). Confidentiality does not allow quantifying the difference but it can be said that the relation underestimates the magnitude of the air gap. 1r

T

1

ι

2

3

1

1

1

1

1

1

1

1

10

11

0.9 0.8 0.7 0.6 0.5

Figure 2 : 3D representation of a stub hole measurements

0.4

A complete three stub assembly was not used for the experiments but rather three separated stubs allowing for good alignment in the holes and a clean surface. The cast iron usually cracks when cooling due to stress arising as the thimble shrinks around the stub. During the withdrawal of the thimbles and stubs from the carbon, the first thimble fell apart from its stub because it was cracked. As a precaution, the two other ones were tack welded to ensure that the assembly remained in one piece. Figure 3 shows the measurement of one thimble following the same procedure as the stub holes except that the approach vectors are inverted.

0.3 0.2 0.1 4 5 6 7 Thimble Height Position

Figure 4 : Scaled air gap magnitude as a function of height in the cylindrical part of stub hole number 2 Choice of thermocouples

Air gap dimension

The molten cast iron is poured in the anode at temperatures rising up to approximately 1350°C depending on the circumstances of the process. As a first approximation, it is fair to expect the cast iron will solidify within two minutes, considering that the anodes are suspended from the transport conveyor in the plant within two minutes of pouring.

Figure 4 contains the results computed locally in the cylindrical part of the thimble number 2. The magnitude is scaled for reasons of confidentiality. The position refers to equally spaced heights in the stub hole, position 0 being the top of the hole and position 11 being the bottom. All the locally computed gaps around the stub hole are represented as data points and the average gap is plotted as a line. Many measured points have been excluded from the

992

It therefore important to use thermocouples (TC) well adapted to the process to be investigated, that is, resistant to high temperature, fast response time and small thermal inertia in order not to disturb the temperature field. Unsheathed tip, 24 gauge, 3.175mm bead, type K and type S TCs were chosen for the experiments based on preliminary trials conducted for qualification purpose.

1300

1100

Thirteen TCs have been placed in each stub hole, giving a total of thirty-nine measurement points. Holes drilled in the baked anode and steel stub allow insertion of the TCs at the desired location in the steel, carbon and cast iron. Figure 5 shows the configuration for the whole anode with the following notation: A letter referring to the material in which the TC is (Cast Iron, Steel, Carbon), a number to identify the position of the TC and a second number identifying the stub hole.

| 2

900

VM :

g

800

I

1

I

1

I

1

1

1

1

1

B_l 141

:

V \ \ \

5 700

1

600

M-MMM

\;

500 400

I

ni

\M

Q 1000

Thermocouple configuration

1

\

1200

0

50

r

Mv^

100

150 200 Time (s)

250

300

350

100

150 200 Time (s)

250

300

350

Sets of TCs are placed on a radial pattern at two different heights, that is, at one and two third of the depth of the stub holes. Two TCs are placed in the cast iron halfway between the two interfaces in the cylindrical part. Two TCs are in one fin of the thimble. Two TCs are placed 1 cm inside the steel surface and four TCs are in the carbon 2 and 8.5 cm away from the interface. In addition to that, there are three TCs placed in the middle of the steel stub distributed over the height. A silver compound has been used to ensure a good thermal contact between the thermocouple and the surfaces. The data acquisition is performed with a DataTaker® DT85 for a period of 24 hours after the time of the anode casting. The data logger records one data point per TC per second.

Figure 5 : Thermocouple configuration in a stub hole Thermocouples results The casting was carried out with a single crucible, thus there was a time lag between the corresponding curves for each hole as the first stub to be cast was the right one (3) and the casting ended with the center one (2). The superimposition of the curves was adjusted with the moment the molten iron touched the TCs.

Figure 6 : First 350 seconds of the temperature recorded in the three cast iron thimbles during sealing There is an appreciable difference between the corresponding curves of the three holes. Some hypotheses are formulated to explain this observation. The casting was made in an industrial environment so some parameters could not be controlled as would have been possible with an entirely experimental setup. The stubs and TCs were placed in the holes with great care to alignment and centering but imprecision in the order of a millimeter was likely to occur. In addition to that, having three holes in an anode with random fins orientation makes it impossible to have the same

Cast Iron The first 350 seconds of the cooling curves of the cast iron are presented in Figure 6. From the numbering shown in Figure 5, the solid lines are the TCs in the cylindrical part and the dashed lines are the TCs in the fins. Thick and thins lines represent the high and low locations of the TCs.

993

boundary conditions for every hole. Also, the filling rate and mass were not controlled. It resulted the center hole (2) being overfilled. The heterogeneous state of the carbon material can also have an influence on the thermal field. The casting sequence (3,1,2) can be deduced by the maximum temperature recorded. The stub hole 3 has a maximum temperature of almost 1300°C and the stub hole 2 doesn't exceed 1200°C. The cast iron was heated up to 1485°C in an induction furnace, transferred into a crucible and brought to the anode to be cast. In each of these steps, heat was lost and it explains the maximum temperature difference. The slight time lag between the thick and thin lines at Os is the time taken to fill the cavity to reach the two highest TCs. It is obvious that the cooling is slower in the fins and in the higher part of the stub holes as the 13 TC is generally the slowest to cool down and the 12 TC is the fastest.

0

200

400

600

800 Time (s)

1000

1200

1400

1600

Figure 8 : First half hour of the temperatures recorded in steel for stub holel

The latent heat released by the phase changes during cooling is easily observable with two major plateaus in all the curves: one at around 1150°C associated with the formation of the austenite and one at around 750°C being the eutectoid reaction. However, two different behaviors can be observed depending on the TC locations. The curves representing the cooling in the fins, in which there is more mass, have a steadier and slower cooling rate and there is a subtle slope variation between 1050 and 1000°C suggesting there is another phase change there. The same behavior is not recorded by the TCs in the cylindrical part of the thimble. It also appears in thimble 1 and 3 that the thinner section (TC 12) experienced undercooling at the austenite formation temperature. Metallographic analyses presented in the next section confirm that a different solidification path was followed by the iron at the different locations.

Cast Iron Metallography Cast iron samples have been cut off the thimbles to perform metallographic analysis. The samples were cut to observe the microstructure at the corresponding location of TCs II, 12,13 and 14 in the stub hole number 1. The results presented here come from optical microscopy without etching. Further analysis with microprobe will be performed later to observe the matrix composition and alloying elements. It will then be possible to make more definitive statements on the phase changes observed in the cooling curves. Figure 9 to Figure 12 represent the microstructure from the thicker to the thinner section, corresponding to TCs 13, 14, II and 12 respectively. There is an obvious evolution in the microstructure as a function of the cooling rate. The slowest one resulted in a dendritic structure and the fastest one gave graphite flakes. The two intermediate cooling rates yielded a hybrid structure made of dendrites andfinerflakes.

Carbon and Steel The curves recorded in the carbon (Figure 7) and steel (Figure 8) show a similar pattern to the three holes and for reasons of conciseness, only the curves for stub hole 1 are presented here. The differences between holes can be attributed to the same hypotheses stated in the previous section.

■

■

■

■

'

.

.

■ "

·

■

El=

Time (s)

Figure 9 : Microstructure of the cast iron at TC D location

Figure 7 : First half hour of the temperatures recorded in carbon for stub holel

994

through a GapCond subroutine in which a particular relation was applied to each contact interfaces. Based on the work of Trovant and Argyropoulos[10, 11] with a casting in a sand mould, the HTC at the carbon/cast iron interface is expected to drop rapidly in the first seconds of the casting and then slowly decrease because the air gap opens. For this reason, the chosen form for the HTC relation is:

È

(2) The relation depends on time. A, B and C are parameters identified with a trial and error method by running the model several times until the cooling curves closely match the 39 experimental curves. Table 1 shows the final parameters obtained with the model and Figure 13 shows the curve fitting at four TCs locations.

20 pm Figure 10 : Microstructure of the cast iron at TC14 location

A

%i /...

* L·

■ i

Table 1 : Parameters identified in equation (2) for the heat transfer coefficients at the casting interfaces Interface Carbon/Cast Iron Steel/Cast Iron Carbon/Steel

?

iJt * '

'

,

400 350 325

A

0.0005 0.00001 0.0005

B

Because of its very small parameter B, the HTC at the steel/cast iron interface decreases slower (almost linearly) than the carbon/cast iron one. The cast iron shrinks around the stub as it cools down, hence the interface pressure rises. It helps maintain a good contact and heat transfer.

■

* *

r*t · ^

%r- ?

One drawback of this method is that the HTCs can not be related to the air gap opening at the carbon/cast iron interface or to the pressure at the steel/cast iron interface. A thermo-mechanical solidification model that is currently developed to predict the air gap would have been of great use to better evaluate HTCs with rather physical relations to gap opening or surface pressure than time, but the model is not advanced enough to accomplish this task due to convergence problems with viscoplastic and damage issues. However, the current relation still gives valuable information.

*

20 ìm

Figure 11 : Microstructure of the cast iron at TC II location

3600

&J 4 ίιΐ **(

C 400 600 400

r

-*

'

*

4

/ "A-' v :

· · * *

- <

'^

» *

%

"

L.

* ** · '

20ì/ç

Tîme(s)

Figure 12 : Microstructure of the cast iron at TC 12 location

Time (s)

3600

Figure 13 : First hour of the thermal model cooling curves (dashed) against TCs cooling curves from the three stub holes

Interface Heat Transfer Coefficient In order to evaluate the HTCs at the casting interfaces, a 3D thermal model was built in Abaqus. The HTCs were imposed

995

3. Fortin, H., M. Fafard, N. Kandev, and P. Goulet, FEM Analysis of Voltage Drop in the Anode Connector Assembly. Light Metals, 2009, p. 1055-1060.

Conclusion The results of an experimental investigation of the sealing of an anode assembly have been presented in order to provide information for the validation of a thermo-mechanical solidification model to predict the air gap at the carbon/cast iron interface.

4. Richard, D., Conception des tourillons d'anode en usage dans une cuve de Hall-Hιroult à l'aide de la mιthode des ιlιments finis, Master Thesis, Universitι Laval, Quιbec, Canada, 2000.

It was found that the relation previously proposed by Richard (Equation (1)) underestimates the magnitude of the air gap at the carbon/cast iron interface. Given that many works have used this relation to evaluate an initial air gap condition, it is suggested that the results of these works be reevaluated with the new available data.

5. Richard, D., P. Goulet, O. Trempe, M. Dupuis, and M. Fafard, Challenges in Stub Hole Optimisation of Cast Iron Rodded Anodes. Light Metals, 2009, p. 1067-1072. 6. Celentano, D., D. Gunasegaram, and T. Nguyen, A thermomechanical model for the analysis of light alloy solidification in a composite mould. International Journal of Solids and Structures, 1999. 36(16), p. 2341-2378.

Cooling curves recorded with thermocouples during the sealing of an anode assembly allowed observation of the thermal field as a function of time. Those curves will be useful for the validation of a thermo-mechanical solidification model to predict the air gap.

7. Coates, B. and S.A. Argyropoulos, The effects of surface roughness and metal temperature on the heat-transfer coefficient at the metal mold interface. Metallurgical and Materials Transactions B, 2007. 38(2), p. 243-255.

From the cooling curves, it was possible to evaluate the heat transfer coefficient at the interfaces with a 3D thermal model. The coefficients found here are functions of time and are therefore only good as a guideline for modeling. It should be considered to carry out real time air gap measurements with LVDTs in conjunction with thermocouples in order not only to better evaluate the heat transfer coefficients, but also to have a better understanding of the strains as the cast iron cools down with the help of the forthcoming thermo-mechanical model.

8. Grandfield, J., D. Mortensen, H. Fjaer, P. Rohan, V. Nguyen, H. Sund, and T. Nguyen, Remelt ingot mold heat flow and deformation. Light Metals, 2006, p. 869-876. 9. Gunasegaram, D.R. and T.T. Nguyen, Effect of cooling rate on air gap formation in aluminium alloy permanent mould casting. International Journal of Cast Metals Research, 2006. 19(2), p. 116-122.

It was also possible to observe the importance of the effect of cooling rate on the final microstructure of the solidified cast iron through metallographic analysis going from dendritic structure to large flakes of graphite. Still to be determined is the need to take this phenomenon into account in an air gap prediction model.

10. Trovant, M. and S. Argyropoulos, Finding boundary conditions: A coupling strategy for the modeling of metal casting processes: Part I. Experimental study and correlation development. Metallurgical and Materials Transactions B, 2000. 31(1), p. 75-86.

Reference [13] provides more details about the work presented in this paper.

11. Trovant, M. and S. Argyropoulos, Finding boundary conditions: A coupling strategy for the modeling of metal casting processes: Part II. Numerical study and analysis. Metallurgical and Materials Transactions B, 2000. 31(1), p. 87-96.

Acknowledgments The authors gratefully acknowledge the financial support provided by Alcoa Inc., the Natural Sciences and Engineering Research Council of Canada, the Fonds quιbιcois de la recherche sur la nature et les technologies and the Aluminium Research Centre - REGAL. We also thank the following people for their indispensable help: Hugues Ferland, Edmond Rousseau (REGAL), Maude Larouche(Universitι Laval), Michel Gasse, Jean-Pierre Rapin, Daniel Lapointe (Alcoa Deschambault).

12. Vicente-Hernandez, P., F. Decultieux, P. Schmidt, I.L. Svensson, and C. Levaillant, Mushy State Behavior - Rheological Characterization and Influence on Air-Gap Formation. Isij International, 1995. 35(6), p. 805-812. 13. Trempe, O., Ιtude expιrimentale et modιlisation du scellement d'un ensemble anodique d'une cuve Hall-Hιroult, Master Thesis, Universitι Laval, Quιbec, Canada, 2011.

References 1. Beier, S., J. J. J. Chen, and M. Fafard, FEM Analysis of the Anode Connection in Aluminium Reduction Cells. Light Metals, 2011. 2. Fortin, H., Modιlisation du comportement thermo-ιlectromιcanique de l'anode de carbone utilisιe dans la production primaire de l'aluminium, Master Thesis, Universitι Laval, Quιbec, Canada, 2010.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECTS OF HIGH TEMPERATURES AND PRESSURES ON CATHODE AND ANODE INTERFACES IN HALL-HEROULT ELECTROLYTIC CELLS 1

Lyne St-Georges1, Lβszlo Istvβn Kiss1, Mathieu Rouleau1, Jens Bouchard1'Daniel Marceau1 Universitι du Quιbec à Chicoutimi; 555 boul. de l'Universitι; Chicoutimi, Quιbec, G7H 2B1, Canada Keywords: cathode interfaces, anode interfaces, electrolytic cell of the contact surface reduces the contact resistance. Among all the investigations presented in the literature on the subject, some authors have reported higher values than those predicted by theoretical models and irregularities at high temperature [5-6].

Abstract This paper deals with the physical modifications occurring at high temperatures and pressures at the interfaces found in the anode and cathode of a Hall-Heroult electrolytic cell. The anode and the cathode are fabricated from carbon blocks, with steel bars inserted and sealed with cast iron. Consequently, two different types of interface are found in the anode and the cathode assembly: castiron with steel and cast-iron with carbon. For the investigation presented here, an experimental setup was built to heat and load anode and cathode samples. Specific attention was put on the sample preparation, to reproduce real cathode/anode sealing conditions. During the heating and the loading of the samples, fluctuations of electrical and thermal contact resistances were observed and related to physical transformation at the interfaces. These transformations could potentially explain the nonhomogeneities of voltage and current distribution occurring in a Hall-Heroult electrolytic cell.

In this paper, the effect of high temperatures and pressures at the anode and cathode interfaces on the electrical and thermal contact resistances is investigated. The specific interfaces considered are those in the anode and the cathode of an electrolytic cell: • anodic cast iron / steel stub rod; • anodic cast iron / carbon; • cathodic cast iron / steel of collector bar; • cathodic cast iron / carbon. Material properties The anode and the cathode carbon used are different materials with specific composition and properties. The main differences between these two materials are listed in the Table 1. For the sealing of anodes and cathodes, different cast iron is also used. The main difference between the two types of iron used stands in the material composition. Both the anode and the cathode castiron are of the grey type.

Introduction The anodes and the cathodes of a Hall-Heroult electrolytic cell are composed of carbon blocks in which steel rods or bars are inserted. To provide a good electrical and thermal contact and to give good mechanical strength to the steel and carbon assembly, liquid iron is cast between the steel rods/bars and the carbon block to seal the anode or the cathode. The quality of the contact between the components used in the anode and cathode is of primary importance because a great amount of energy is transmitted through these interfaces to the electrolytic process. During the aluminum electrolysis, a bad contact between the castiron and the carbon block or between the steel and the cast-iron can modify the current and the heat distribution in the electrodes and decrease the effectiveness of the electrolytic cell.

Table 1 : Differences between anode and cathode carbon Composition Graphitization

Anodic carbon Coke+Pitch Non-graphitized

Porosity

Relatively porous

Hardness

Low

Cathodic carbon Anthracite+Pitch Partially graphitized Relatively less porous than anodic carbon Relatively hard, function of the graphitization

Sample preparation

For 30 years, many efforts have been made to quantify experimentally or in-situ the voltage and the temperature drops at the cast iron / carbon interface on the cathodes and the anodes of the electrolytic cells [1-6]. These voltage and thermal drops are expressed in terms of electrical and thermal contact resistance respectively.

To produce test samples, the steel, the cast-iron and the carbon samples (both anode and cathode) were fabricated to reproduce the surface characteristics obtained in industrial anodes and cathodes. The steel samples were obtained directly from the steel supplier. Cylindrical samples with a diameter of 50.8 mm and a length of 330.2 mm were used for the experiments. The cathode and the anode carbon samples were cut respectively from an industrial cathode and anode block. Cylindrical samples with a diameter of 50.8 mm and a length of 330.2 mm were extracted. The surfaces of the samples were not modified from the carbon block to conserve their original characteristics.

The contact resistances normally decrease with the applied pressure. This phenomenon can be explained by the augmentation of the real area of contact (or the increase of the number of contact points) with the pressure. This increase of the contact surface decreases the throttling of the current lines at the interface, responsible for the contact resistance, and consequently decreases the contact resistance. Furthermore, based on the literature review, the contact resistances are also expected to be reduced by an increase of the temperature. As the temperature increases, the mechanical strength of the material decreases, asperities begin to collapse and the real contact surface increases. This augmentation

The carbon and the steel samples were placed separately in hollow graphite tubes (508 mm long with an internal diameter of 50.8 mm and an outside diameter of 76.2 mm). These tubes, with either the carbon or the steel samples, were inserted into an anode carbon block to provide the same cast iron cooling rate as that

997

used in industrial anode sealing. The carbon block was then heated with the samples up to 100°C (the normal cell pre-heating temperature used in anode/cathode sealing). Once the temperature was reached, the melted iron was cast at the top of the carbon and the steel samples, using an industrial anode sealing procedure to form the cast-iron/carbon and cast-iron/steel pairs of samples. Once cooled, the pairs of samples were delicately extracted from the carbon block using a hydraulic press. For the fabrication of the cathode samples, based on numerical simulations, two side-byside anode blocks were used instead of a cathode block to provide a cooling rate similar to that obtained in industrial cathode sealing. Experimental tests

For the experimental determination of the contact resistances, the samples were then heated, using the experimental setup, up to 1000 °C, using steps of 150°C (except for the last step, from 800°C to 1000°C). For each step of temperature, steady-state contact resistance was measured with applied pressures between 0.5 MPa and 2.5 MPa. These temperatures and pressures were used to reproduce the conditions found in an electrolytic cell. For each step of temperature, the conditions were maintained for 24 hours to reach a thermal steady-state condition. Each level of pressure was then applied for 2 hours (10 hours total) after which the steady-state measurements were taken. After the heat and pressure cycles experiment, electron and optical microscopes were used for microstructure and macroscopic observations.

The apparatus used to heat and load the samples was designed to determine both electrical and thermal contact resistance [7]. Figure 1 shows the experimental set-up used. In the apparatus, the samples are placed vertically between two electrodes. A furnace is used to heat the pair of samples and the geometry of the electrode outside the furnace is designed to act as a heat sink to create a heat flux in the samples through their interface. The samples are connected into an electrical circuit by the electrodes and consequently, a voltage and a temperature gradient are created through the test section. A measurement of the electrical and thermal contact resistances at the interface is then possible.

Results and discussion Figure 2 shows the typical variation of the contact resistance obtained. Globally and as expected, the resistance decreases as the temperature increases. However, at about 550 °C, an augmentation of the resistance is observed This behavior occurs in all the samples tested, but is more visible in the results obtained for the electrical contact resistances. For the material tested, various metallurgical and mechanical transformations may occur during the experiment process: • desorption of oxygen by the carbon; • carbon combustion; • hardening of the cast iron surface (during the iron casting); • annealing; • chemical reactions in the cast iron; • cast iron creep; • welding between the steel and the cast-iron; These transformations could explain the fluctuations of the resistance observed and their effects are discussed in the following sections. 0.5 MPa A

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300

In the apparatus presented, the samples were protected from oxidation by various means: (1) the oxygen was purged by injection of argon in the furnace cavity. A semi-hermetic sheet metal retained the argon around the sample and a pressure gage was used to maintain the argon pressure higher than the atmospheric pressure, to avoid oxygen inlet in the system; (2) an anti-oxidation coating was applied on the samples. This antioxidant is very stable at high temperature and non-conductive thermally and electrically; (3) a small carbon piece was placed in the furnace, between the samples and the thermal shell and acts as a sacrificial sample to rapidly consume the oxygen before a reaction takes place with the samples.

350

400

450 500 550 Temperature f Q

600

650

700

Figure 2: Variation of the electrical contact resistance between cathodic cast iron and carbon Oxygen desorption and carbon combustion The observations made on the carbon samples have shown that the surface of the carbon samples was consumed during the test procedure. Considering the porosity of the carbon sample and the presence of several barriers to the oxygen inlet in the experimental device used, it appears that the oxygen does not come from outside the system butfromthe sample themselves.

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measurement were analyzed with an scanning electron microscope (SEM) equipped with an Energy Dispersive Spectrometer (EDS). The microstructures of the as-cast sample (non-heated) and of the heated sample after an experiment are presented in Figure 4. The as cast samples exhibit graphite plates near the surface. The graphite plates are generally produced by a very fast cooling rate [10] and are representative of the anode/cathode sealing process, where a fast solidification is obtained.

The oxygen adsorption by the carbon at low and high temperature is a known phenomenon. The carbon being relatively porous, a non negligible quantity of oxygen is adsorbed by the carbon surface before the sample fabrication. When heated at elevated temperatures, the oxygen reacts with the carbon. The carbon combustion has been frequently observed at temperature about 500-600°C. This range of temperature also coincides with the temperature of carbon desorption (550°C, [8]). Heat treatment of cast iron

For the heated sample, the heating during the experiment increases the size of the graphite plates. The large graphite plates, which are more numerous in the heated sample than in the as-cast sample, reduce the cast iron hardness and its electrical resistivity [11].

The microstructure of the iron is highly dependent on its thermal history. To evaluate the impact of the heating procedure on the microstructure of the iron used, micro-hardness measurements were made at several points along the depth of a non-heated (as fabricated) cast-iron sample and compared to a cast-iron sample after a normal heating procedure used during the experiments. For the hardness measurements, the samples were cut along their longitudinal axis and the measurements were made from the surface (depth ~ 0) to their opposite ends. Figure 3 shows the effect of the heating procedure on the micro-hardness of the cast iron, along the depth axis. The micro-hardness is higher for the as-cast sample than for the heated samples. Moreover, in contrast to the heated sample, the micro-hardness is significantly higher near the contact zone (depth ~ 0) for the as-cast sample. For the heated sample, the hardness is nearly constant over the depth. The surface hardening observed for the as-cast sample is produced by the fabrication procedure used: the liquid iron, initially heated at a temperature of 1400°C, is poured on a carbon block at 100°C. The instantaneous contact between the liquid iron and the carbon block causes a rapid cooling of the iron surface and produces the surface hardening. 600

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After the heating procedure used to determine the contact resistance at elevated temperature, an extremely friable, thin, grey-blue layer, not seen before, was observed between the heat treated cast-iron and the carbon samples (both anode and cathode). This layer, observed in the test samples, is related to the relatively long exposure to high temperatures. For similar environmental conditions, three specific phenomena are normally observed in iron: surface decarburization, surface oxidation and surface oxidation with the presence of silicon.

Figure 3: Vickers micro-hardness of (a) a non heat-treated and (b) a heat treated cast iron On the other hand, when heated at elevated temperature, the grey cast iron is subject to the following heat treatment 9]: (1) annealing (which eliminates the residual stresses generated by heterogeneous cooling); (2) ferritic annealing (which transforms the perlite into graphite and ferrite, resulting in a decreased hardness). These transformations also occur in the iron samples during the experiments. They are responsible for the hardness reduction and homogenization of the heated samples.

To determine the nature of the transformations observed, the surface of the cast iron was analyzed with an electron microscope. Pictures of the iron surface, after heating, are presented in Figure 5 for various magnifications (a to c), and compared with a cast iron sample, not heated (d). For the heated cast iron surface, two

To analyze the effect of the heating procedure on the iron, the microstructures of the samples used for the micro-hardness

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dark gray layers are observed. To determine the chemical composition of these layers, a Wavelength Dispersion Spectrometer (WDS) was used. The composition maps obtained are presented in Figure 6. The first layer observed is rich in oxygen and iron and the second layer is mainly composed of oxygen, silicon and iron. The composition of the first layer is representative of the presence of an iron oxide (such as FeO). The FeO, also called Wustite, is a ptype semi-conductor and is formed from 570°C [12], which is about the temperature of sudden jump of contact resistance observed in Figure 2. For the second layer, a silicon dioxide (silica) containing iron is suspected (Si02 + Iron oxide for example). In the iron, the silicon can be oxidized to form silica [13]. The oxygen needed for these chemical reactions may come from the carbon desorption, as mentioned before. Figure 5: Cast iron surface, heated sample at (a) 25X, (b) 100X, (e) 500X and non-heated sample at (d) 500 X

The decarburization at the surface of the cast iron, clearly visible in Figure 5, is also observable in Figure 4 (at a depth < 0.18 mm). This decarburization seems to be due to chemical reactions between the carbon contained in the cast iron and the oxygen adsorbed by the carbon sample in contact with the iron, to form gaseous CO and CO2. This phenomenon of cast iron decarburization is well known and reported in the literature [11, 13]. In addition to the carbon impoverishment of the cast iron produced by the decarburization, other oxidation or reduction mechanisms may occur: (1) oxidation of the iron contained in the cast iron; (2) removal of the exposed graphite by oxidation to carbon monoxide [83]. The friability of the grey blue thin layer observed after heating can be explained by the oxidation of the iron. It is known that over 700°C, the carbon monoxide and dioxide removal can create cracks and swelling in the iron. In the lamellate graphite, the graphite plates facilitate the oxygen penetration inside the cast iron and oxidized layers can then be produced inside the iron. These metallic oxide layers have a volume greater than the volume previously occupied by the metal, and may produce cracks and swelling of the iron, expose new surfaces to the oxygen and accelerate the oxidation process [11]. The decarburization as well as the oxidation of the iron has an important effect on the contact resistance since the cast iron conductivity is greatly dependent on the carbon content.

Figure 6: Chemical analysis of cast iron surface. Clear pixels represent high concentration of (b) carbon; (c) oxygen; (d) iron; (e) silicon; (f) sulphur; (a) is a 200X optical micrograph Λ

! ~° 1

The decarburization of the cast iron can be a priori difficult to accept, considering the presence of the carbon in contact with the cast iron in the pair sample. The diffusion or the migration of an element follows the direction opposed to the gradient of concentration of the element itself and, consequently, a carbon enrichment of the iron instead of an impoverishment would be expected.

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To explain this behavior and to visualize the real surface of contact between the carbon and the iron in the samples produced, a sample was cut, epoxy mounted and polished for microscopic observations. In the sample preparation, a mechanical pressure of 1 MPa was applied during the epoxy solidification to reproduce test conditions. A typical micrograph of the carbon / cast iron interface obtained under this condition of preparation is shown on Figure 7.

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Figure 7: Evidence of gap at the cast-iron/carbon interface under a mechanical pressure of 1 MPa A gap is observed between the carbon and the cast iron. In fact, a good contact is obtained only at a limited number of points. The gap between the cast iron and the carbon samples (approximately equal to 75 μιη) is greater than the gases mean free path, even at elevated temperature. Consequently, CO and C0 2 displacements

1000

are possible. This gap between the two materials is a potential obstacle to the diffusion and probably stops the diffusion of the carbon towards the cast iron. However, this spacing is large enough to allow oxygen, CO and C0 2 to circulate at the interface and to escape.

Conclusions During the heating of cathodes and anodes, several physical transformations may occur at the cast iron/steel and cast iron/carbon interfaces. The decarburization and the oxidation of the carbon/iron interface, the welding of the iron/steel interface and the material creep are the only potential transformations noticed in this investigation. More complex or combined transformations may occur in industrial anode/cathode interfaces and interface transformations being susceptible to variation with time, the contact resistances should also vary with time. These transformations have important effects on the contact resistances and could explain the electrical and thermal distributions heterogeneity observed in the electrolytic cells.

During the sample fabrication and the anode / cathode sealing, the liquid iron is poured directly on the carbon. The presence of the gap observed may have different sources (gas trapping, withdrawal.) but can be partially explained by the thermal contraction of the cast iron during its cooling. In reality, cast iron shrinkage is about 1%, which is sufficient to produce the gap illustrated in Figure 7. Cast iron creep During the heating and the loading cycles, under steady-state conditions of loading, it was observed that, for a constant displacement imposed, the pressure applied on the samples goes down with time, especially at elevated temperatures. This phenomenon could be explained by the creep in compression of the cast iron. It is known that the creep of the cast iron can becomes important when the material is heated at elevated temperatures [11]. The cast iron creep could increase the real surface of contact between the surfaces and thus reduce the contact resistance.

Acknowledgements This work is part of the research program of CRDA (Centre de Recherche et de Dιveloppement Arvida) of Rio Tinto Alcan and was also supported by the CURAL (Centre Universitaire de Recherche sur l'Aluminium), part of the Universitι du Quιbec à Chicoutimi. References [1] R.W. Peterson "Temperature and voltage measurements in hall cell anodes", Light Metals, Vol 1, 1976, p. 365.

Steel and cast-iron welding

[2] D.G. Brooks, V.L. Bullough, "Factor in the design of reduction cell anodes", Light Metals 1984, p. 961.

For the cast-iron/steel samples (anode and cathode pairs), an unexpected welding phenomenon was observed: welded spots with surfaces of about 150 mm2 were observed between the castiron and the steel samples. Figure 8 shows a typical weld broken for observation purposes. On each sample, only one of these spots was observed. This welded spot could be related to a local melting of the steel sample during the sample fabrication and is typically located where the liquid iron falls directly on the steel, i.e. where the heat is sufficient to melt locally the steel sample.

[3] PJ. Rheday, L. Castonguay, "Effects of carboneous rodding mix formation on steel-carbon contact resistance", Light Metals 1985, p. 1089. [4] M. Sorlie, H. Gran, "Cathode collector bar to carbon contact resistance", Light Metals 1992, p. 779. [5] F. Hiltmann, H. A. 0ye, "Influence of temperature and contact pressure between cast iron and cathode carbon on contact resistance", Light Metals 1996. p. 277. [6] L.I. Kiss, M. Rouleau, L. St-Georges, "Determination of the thermal and electrical contact resistances at elevated temperatures ", Proceeding of the twenty-eight international thermal conductivity conference, 2006, p. 224.

aίm b)a**«a· iiiiiM . Figure 8: Broken welded spots between (a) cast-iron and (b) steel

[7] St-Georges, L., Kiss, K.I., Rouleau, K., « Evaluation of contact resistance in electrodes of Hall-Heroult process», Light Metals 2009, p. 1103.

The metallurgical links created by the welding between the steel and the cast-iron have an extremely important impact on the current and heat transfer. For the cast-iron/steel samples, the assumption of uniform voltage distribution fails.

[8] Le groupe franηais d'ιtude des carbones, « Les carbones », Tome 2, Masson et Cie, 1965.

This first observation gives more insight on the cast-iron/steel interfaces but needs to be validated for real electrolytic cells. The procedure used to produce the cast-iron/sample is representative of the one used for the anode/cathode sealing and should provide at least a similar cooling rate. However, for large samples or during the sealing of an industrial cathode or anode, only a limited number of welding points are expected, where the heat provided by the liquid iron is large enough to melt the steel. To confirm this expectation and to validate the presence of these welded spots in industrial electrolytic cells, further work is needed.

[9] J-M. Dorlot, J-P. Bβillon, J. Masounave, «Des matιriaux», 2iθrae ιdition, Ιdition de l'ιcole Polytechnique de Montrιal, 1986. [10] F. Barralis, G. Maider, «Prιcis de Mιtallurgie», AFNOR, p. 58.

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[11] Centre d'information des fontes moulιes, «Manuel des fontes moulιes », Ιditions Techniques des industries de la fonte, 1963, p. 705. [7] A.S. Khanna, «Introduction to high temperature oxidation and corrosion », ASM International, 2002, p. 82. [83] W.F. Charles, «Gray and Ductile iron casting handbook», Editions Cleveland Ohio, 1971, p. 679. [94] S.C. Saxena, R.KJoshi «Thermal accommodation and adsorption coefficients of gases», Hemisphere publishing corporation, United-States, 1989.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

NEW APPARATUS FOR CHARACTERIZING ELECTRICAL CONTACT RESISTANCE AND THERMAL CONTACT CONDUCTANCE Nedeltcho Kandev1, Hugues Fortin1, Sylvain Chιnard1, Guillaume Gauvin2, Marie-Hιlθne Martin2 and Mario Fafard2

2

Institut de Recherche d'Hydro-Quιbec (IREQ), 600 avenue de la Montagne, Shawinigan (Quιbec), Canada, G9N 7N5 NSERC/Alcoa Industrial Research Chair MACE3 and Aluminium Research Centre-REGAL, Laval University, Sciences and Engineering Faculty, Adrien-Pouliot Building, Quebec City (Quιbec), Canada, Gl V 0A6

Keywords: Aluminium reduction cell, anode connection, anode stub contact, anode power losses, electrical contact resistance, thermal contact conductance Abstract A new apparatus for characterizing electrical contact resistance and thermal contact conductance has been developed and tested. The heating of the samples inside the apparatus was achieved via induction heating of a stainless steel billet to create a powerful heat generator instead of the commonly used convection furnace. With this equipment, the thermo-electro-mechanical (TEM) behavior of metal-carbon and metal-metal interfaces can be reproduced using a controlled inert gas environment, temperature up to 1000°C and mechanical pressure up to 2 MPa. A major advantage of this new concept is that the temperature equilibrium of the samples can be reached quickly (within two or three hours) while precisely controlling the heat flux. Recent experimental results performed on steel-carbon samples show that this concept is feasible and very efficient. This apparatus will be used to establish the constitutive laws of interfaces for electrical contact resistance and thermal contact conductance in the electrode connections to support numerical modeling of the aluminium reduction cell. Introduction Theoretical and experimental investigation of the thermoelectro-mechanical (TEM) behavior of the stub-anode connection is not an easy task, as it involves complex interactions between non-linear thermal, electrical, mechanical and surface phenomena. Nevertheless, computer modelling is exceptionally useful for understanding these complex phenomena and optimizing the anode design and the fabrication process. However, to build a representative numerical model of the aluminium reduction cell, it is important to correctly predict the thermal contact conductance and the electrical contact resistance (ECR) as a function of both pressure and temperature. The development of constitutive laws based on laboratory measurements is, therefore, a crucial task. Carbon-cast iron-steel interface phenomena are regulated by the powerful fully-coupled thermo-electro-mechanical behavior of the assembly. While increasing the temperature, the materials are expanding, creating a higher contact pressure at the interfaces and decreasing the ECR. With decreased ECR, the local Joule dissipation is lowered, affecting the thermal field and so on until the thermodynamic equilibrium is reached. This stabilization of the TEM fields can take some time.

The objective of this work is to design a new apparatus employing induction heating, rather than a conventional convection furnace, as a heat generator to achieve better control of heat flux and to reach thermal equilibrium more rapidly. With this equipment, it will be possible to identify parameters of constitutive laws for the ECR and the thermal contact conductance for metal-carbon and metal-metal interfaces for implementation into a Finite Element Method (FEM) code. These laws may lead to further promising research, investigating the effect of new cast iron recipes on the ECR as well as clad to transversal bar contact resistance or the steel-cast iron interfaces degradation that has been pointed out by Kandev and Fortin [1]. Previous work The ECR establishes the current distribution across the different interfaces and is responsible for the high voltage losses in the anode connection. To characterize the ECR, a lab-scale reproduction of the operating condition in the cell demonstrating the thermal, electrical and mechanical fields is needed. In 1976, Peterson [2] instrumented an anode with voltage probes and thermocouples and found that the ECR was responsible for 25% of the global voltage drop in the anode. Two years later, Peterson [3] studied the effect of changing cast iron volume in the stub hole relative to temperature distribution. In 1984, Brooks and Bullough [4] attempted to optimize cast iron thickness in the stub hole to minimize anode cracking. They found that ECR was a function of both temperature and contact pressure. In 1992, S0rlie and Gran [5] built an apparatus to measure the ECR of steel-to-carbon interfaces. A 60-mm diameter cylindrical piece of cathode block was linked with a cylindrical layer of mild steel at each end to form a sandwich sample of two different materials. Subsequently, the sample was put in a tube furnace having a controlled environment, using nitrogen gas to avoid corrosion. The same setup was later used by Hiltmann et al. [6] for the measurements of the cathode cast iron interfaces. From these experiments, unfortunately, no constitutive laws were developed for ECR that could be used in a FEM code. By 2000, Richard et al. [7, 8] had reproduced the S0rlie and Gran [5] experiment at room temperature, using anode carbon

1003

material. With a calibration based on experimental data from S0rlie and Gran [5], they were able to develop a constitutive law for ECR as a function of both temperature and pressure in a form of the Weibull law. Richard et al. [7, 8] incorporated this law into an FEM code and built a pie-shaped 3D model of a stub hole to study the impact of new flute design on the voltage drop. In 2003, Laberge et al. [9] built a new test bench to characterize the thermo-electrical properties of the anode-coke bed-cathode interfaces during the preheating of the cell. Based on the same principle as the S0rlie and Gran [5] experiment, the sample was placed under a controlled environment with argon gas and resistant insulating material was placed around the sample to avoid corrosion of the carbon material. By changing the temperature and the mechanical pressure, they were able to characterize the thermal contact conductance and the electrical contact resistance of the anode-bed coke-cathode interfaces which are critical for modeling the electrical preheating of cells. In 2009, Richard et al. [10] published the parameter values of the constitutive law at operating temperature which had been developed back in 2001. Using this constitutive law, recent publications [10, 11, 12] have since demonstrated that with FEM software, it is now possible to construct a coupled thermoelectro-mechanical (TEM) model of a full 3D anode. Such models are very useful for experimenting with new designs and will help the aluminium industry to achieve its energy efficiency objectives. Basic concept A new test bench was designed, constructed and tested, using induction heating of a stainless steel billet as a thermal generator to heat the sample. By precisely controlling the power injected in the sample, thermal equilibrium of the heat flux can be achieved very rapidly within two or three hours. A simplified schema of the concept design of the apparatus is shown in Figure 1. The stainless steel billet is heated by induction using cylindrical copper coil wrapped around the billet. To prevent overheating of the copper coil at high induction heating power, the coil is watercooled. The heat flux is transferred from the billet to the carbon-castiron-steel "sandwich" samples using a metal disk with good thermal and electrical conductivity (Ni for example). The disk is added to ensure that the thermal and the electrical flux lines will be perpendicular to the surface.

Thermal Insulation-1 Stainless Steei Inductor

Thermal Insulation-2

Figure 1 : Simplified schema of the concept design To minimize the heat losses, a high-quality thermal insulation (thermal insulation 1) is wrapped around the sample and around the stainless steel billet to limit the outgoing radial heat flux. Also, several layers of a special thermal insulation (thermal insulation 2) resistant to high mechanical pressure and high temperature are placed under the billet for thermal insulation and mechanical support. TEM behavior of the interfaces inside the apparatus can be reproduced using temperatures ranging from 20°C up to 1000°C and mechanical pressure up to 2 MPa. In addition, a DC current is applied between the stainless steel billet and the top of the samples, representing the same current density that is applied at the contact surfaces in an actual operating cell. Also, to limit the air oxidation of the billet and the samples during the experiments at high temperature, a controlled inert gas environment is anticipated. Numerical modelling To validate this new concept, a transient thermo-electromagnetic 2D axisymmetric model was built using the computer package COMSOL Multiphysics 3.4. This model is based on the geometry of the apparatus shown in Figure 1. The physical model chosen to solve this problem is the coupling of electromagnetic induction current model with the classical heat transfer by conduction model. The problem is considered as symmetrical and the thermal and electrical characteristics of the involved materials (steel, carbon, cast-iron and thermal insulations) are considered homogenous and temperaturedependent. To simplify the numerical solution, all interfaces between materials are considered ideal. Finally, the geometry of the coil is represented by a slab of copper having an equivalent current. The electromagnetic domain is delimited by an air cylinder. On the external boundaries of this air domain the magnetic and

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electric conditions are fixed to: n x A = 0 and n.J = 0 , where n is normal vector to the boundary, A is the magnetic vector potential and J is the current density vector. The interior boundaries conditions assume continuity of the medium, corresponding to a homogenous Neumann condition.

Temperature [°C]

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Several induction heating cases were simulated at different power supplies and frequencies to validate the concept and to optimize the system design. An example of temperature distribution solution at 7000 seconds time step is shown in Figure 2 for an induction heating power of 4050 W operating at a frequency of 2800 Hz. The results demonstrate that with such an induction heating generator, the temperature in the upper portion of the sample can reach 500°C very quickly.

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0.7

Figure 4: Temperature distribution on the cylindrical axis of the system at 1820 W and 2800 Hz Min: -5.684C-M

Figure 2: Temperature distribution in the model at 4050 W induction heating power and 2800Hz Figure 3 illustrates the temperature evolution close to the axis of the system (red line in Figure 2) during induction heating up to 7000 seconds and the temperature distribution of the system at different time intervals between 3000 and 7000 seconds. It can be seen that 3000 seconds after the start of the induction heating system, the maximum temperature in the stainless steel billet reaches 435°C and at 7000 seconds of heating time, the billet reaches 925°C. The software allows the calculation of the outgoing heat fluxes and both thermal and electrical losses in the system. For the case considered and illustrated in Figure 2 and 3, the outgoing thermal flux represents 1828 W of power. 7000 seconds after heating began the induction power was decreased to approximately 1820 W in order to compensate for the outgoing thermal flux and losses.

The simulation results have demonstrated that the concept of a water cooled induction coil provides a very efficient solution. At 2800 Hz, the total current in the coil is about 5800 Amps. By using a coil with 20 spires, for example, the supplying current will be:

/

=IISL = ^

n

0

20

=

290 Amps

(i)

This induction coil design was used to construct the apparatus. Design and construction of the apparatus The apparatus was constructed with a variety of components. For the electrical portion, a 35-kW induction generator was used, having the possibility of varying power and frequency with a capacitor bank for the frequency adaptation.

1005

From the numerical simulation results, the inductor was designed using 20 spires of W diameter copper pipe. To determine the optimal frequency for the generator, the inductor characteristics were measured with an HP 4194A Impedance/Gain-Phase Analyzer. Principal characteristics are listed in Table 1, where/is the frequency, L the inductance, R the resistivity with the electrical load (stainless steel billet), R0 resistivity without the load, Qmeas the quality factor measured, Cres the capacitance at resonance frequency and finally, the Eff elee is the electrical efficiency of the induction that is calculated from the equation (R-Ro)/R. Table 1 : Calculated and measured electrical characteristics of the induction coil

I

f Hz 700 800 900 1000 2000 2800 3000

L H 4.44E-05 4.37E-05 4.24E-05 4.13E-05 3.70E-05 3,68E-05 3.60E-05

R mOhm 60,00 66,00 69,70 76,92 109,00 126,00 130,00

Qmeas 2,80 2,90 3,00 3,05 4,00 5,67 5,68

Ro m Oh m 21,2 23,0 28,0 30,0 45,0 48,1 48,7

Cres MF 1165,33 905,68 737,20 613,47 171,15 87,80 78,18

Eff elee I

%

64,7 65,2 59,8 61,0 58,7 61,8 62,5

Figure 6: Sample under mechanical pressure

I

From those results, the frequency of 2800 Hz was chosen, respecting the capacitance limitation of the generator while demonstrating good electrical efficiency.

The temperature and the voltage drop monitoring were logged with a computer and an Agilent 34970A acquisition system. The thermocouples used were K-type and the voltage drops were measured directly from the thermocouples. The sample and the billet were instrumented with, respectively, 7 and 4 thermocouples measuring the temperature and the voltage drop. Probe arrangement in the billet (left) and in both carbon and steel (right) can be seen in Figure 7.

As for mechanical loading, initially force was applied on the sample using a lever arm system having different weight at the end, as seen in Figure 5. The inductor, the different thermal insulation and the water cooling system are also shown here.

50 , *

w

«I . CO

Figure 7: Probe arrangement in the stainless steel billet (SSx) and the carbon (Cx) (mm) Experimental study

Figure 5: Apparatus in operation In the future, the lever arm will be replaced by a manual press of 1-ton capacity which will apply a constant pressure up to 2 MPa on the sample. Also, to measure the pressure applied on the sample, a Tovey Engineering pancake type load cell will be used. In addition, all heating systems will be protected with a Nylon and G-9 shell. A stainless steel tube, with a half-open top, will contain the sample and the inert gas during the test, to avoid the oxidation of the material. For the initial test, only the stainless steel and the carbon were used. The samples of 50.8 mm diameter and 150 mm height were placed between the nickel disk and the mechanical pressure system seen in Figure 6.

The first part of the experimental study was conducted only on the induction heating generator to verify the plausibility of the concept and the efficiency of the coupling at a frequency of 2800 Hz. Using only the lower part of the apparatus in Figure 1, i.e. without the sample and the mechanical pressure, the stainless steel billet was heated for about half an hour at a power of 10 kW. After that time, the power generator was turned off, but the data acquisition system continued to monitor the temperature for another fifty minutes. Figure 8 presents the temperature for the thermocouple in different locations on the stainless steel billet as a function of time. The first test results showed that the electromagnetic coupling was excellent and that within half an hour, the temperature on the top of the billet could reach about 450°C. After the power generator was shut off, (B on Figure 8), the temperature slowly decreased to 400°C. It is interesting to note that two different incidents occurred during this test. At A, there was a power failure because the operating parameters of the generator were

too close to the limits. To overcome this problem, the power was decreased to obtain stability. Finally, at C, the thermal insulation on the top of the stainless steel billet was removed quickly to see the stainless steel surface at this temperature. This explains the temperature drop on the top thermocouple in the stainless steel.

500 9 400

SSI side bottom SS2side middle 553 side top

300

554 top at 70 mm

200

2000

3000

4000 Time (s)

5000

Figure 9: Temperature as a function of time in the stainless steel with the sample and the electrical current

100

20CX)

, ,3000

Time (s)

4000

5000

Figure 8: Temperature as a function of time in the stainless steel billet without the sample An interesting feature of this curve is the temperature evolution in the middle section of the billet, where the induction heating is greater. Calculating the slope of about 0.3°C/s, the estimated time to reach 925°C is about fifty minutes. This result compares well with the simulation results shown in Figures 3 and 4.

Figure 10 portrays the temperature in the carbon sample as a function of time. The results show that within the hour, the sample temperature is over 500°C in the bottom of the sample. The temperature distribution in the carbon showed that the thermal equilibrium could not be achieved within the hour due to the thermal properties of the carbon and the carbon-steel interface behavior and the controlled argon gas environment. After the test, no trace of oxidation was observed, validating the effectiveness of the controlled argon gas environment.

A second test, employing the same operating parameters as the first test, 10 kW and 2800 Hz, was performed to validate the sample heating time and procedure. This involved mounting the carbon and steel samples on the nickel slab on the billet surface as seen in Figure 6. All the materials were instrumented, respecting the thermocouple location detailed in Figure 7. A current of 20 Amps and mechanical pressure of 0.3 MPa were applied to the sample. Figure 9 shows the temperature in the stainless steel billet as a function of time. The billet was heated up for one hour. After a quick power adjustment (A), the sample was heated up until it reached 550°C (B). At point B, the power was decreased to achieve thermal equilibrium until the power was shut down at 700°C (C). The test shows that it is very easy to achieve temperatures up to 700°C within the hour. After the power was shut down, the temperature slowly decreased, following an exponential behavior. These results also demonstrate that the heat flux is controlled precisely, modifying the power input thus reducing the temperature in the sample rapidly. Efficiently attaining thermal equilibrium in this manner will be very useful.

3000

4000

Figure 10: Temperature as a function of time in the carbon sample with current applied In Figure 11, the temperature of all thermocouples is presented as a function of their positions, at a given time, in both carbon and steel. Results show that thermal contact conductance is weak between carbon and steel, having a drop of about 20°C at the interface (green line). Also, temperature distribution in carbon seems to follow more of a polynomial than a linear behavior as opposed to the linear behavior of the steel. For the characterization of the interfaces, a minimum gradient in the sample is required, especially in the carbon.

1007

Acknowledgements The authors would like to thank Donald Picard from Laval University for support for the apparatus design, especially for the mechanical part and Claude Belzile from IREQ for the experimental test measurements.

5 300

♦ Carbon # Steel

!

References

£ 200

0

50

100 150 200 Thermocouple position (mm)

250

300

Figure 11: Temperature of the different thermocouples as a function of thermocouple position in carbon and steel at 5000s Finally, differences between the experimental data and the numerical modelling stem from the initial condition that all interfaces between the materials are considered ideal. In the actual test, ECR and thermal contact conductance play a crucial role in electrical and thermal distribution like shown in Figure 11. Special care in the experimental setup should be given to interface cleanliness to avoid the presence of unwanted particles between the interfaces. Conclusion Using induction heating technology, a stainless steel billet has been designed as a powerful heat generator for heating the carbon-cast iron-steel material used in the aluminium industry. With FEM modelling, several induction heating cases were simulated to validate the concept and to optimize the design and procedure. The simulation results showed the importance of water cooling of the inductor for frequency of 2800 Hz, with a total current of 5800 Amps. The model could be reused for different applications involving induction heating technology such as the preheating of the stub before cast iron sealing. The main advantage of this new apparatus is that temperature equilibrium of the samples can be reached very quickly (within two or three hours) while precisely controlling the heat flux. In situ operating conditions in the anode stub hole connection could be reproduced by putting a carbon-steel sample in the heat generator and applying mechanical pressure and electrical current on the sample. With the same TEM behavior of the materials, this new apparatus will be used initially to characterize the metal-carbon and metal-metal interfaces to determine new constitutive laws for these interfaces. These laws will be included in FEM software such as FESh++ to test new anode concepts or optimize the anode connection.

1. N. Kandev and H. Fortin: Electrical losses in the stub-anode connection: Computer modeling and laboratory characterisation. In Proc. TMS Light Metals, pp. 1061-1066, 2009. 2. R.W. Peterson: Temperature and voltage measurements in Hall cell anodes. In Proc. TMS Light Metals, pp. 365-382, 1976. 3. R.W. Peterson: Studies of stub to carbon voltage. In Proc. TMS Light Metals, pp. 367-378,1978. 4. D. G. Brooks and V. L. Bullough: Factors in the design of reduction cell anodes. In Proc. TMS Light Metals, pp. 961976, 1984. 5. M. Sorlie and H. Gran: Cathode collector bar to carbon contact resistance. In Proc. TMS Light Metals, pp. 779-787, 1992. 6. F. Hiltmann, J. Mittag, A. St0re and H. A. 0ye: Influence of temperature and contact pressure between cast iron and cathode carbon on contact resistance. In Proc. TMS Light Metals, pp. 277-283,1996. 7. D. Richard, M. Fafard, R. Larcroix, P. Clιry and Y. Maltais.: Aluminum reduction cell anode stub hole design using weakly coupled thermo-electro-mechanical finite element models. Finite elements in analysis and design 37, pp. 287-304, 2001. 8. D. Richard, M. Fafard, R. Larcroix, P. Clιry and Y. Maltais: Carbon to cast iron electrical contact resistance constitutive model for finite element analysis. Journal of Materials Processing technology 132, pp. 119-131, 2003. 9. C. Laberge, L. Kiss and M. Desilets: The influence of the thermo-electrical characteristics of the coke bed on the preheating of an aluminum reduction cell, In Proc. TMS Light Metals, pp. 207-211, 2004. 10. D. Richard, P. Goulet, O. Trempe, M. Dupuis and M. Fafard: Challenges in stub hole optimization of cast iron rodded anodes. In Proc. TMS Light Metals, p. 1067-1071, 2009. ll.H.Fortin, M. Fafard, N. Kandev and P. Goulet: FEM analysis of voltage drop in the anode connector assembly. In Proc. TMS Light Metals, pp. 1055-1060, 2009. 12. Marc Dupuis. Development and application of an ANYS based thermo-electro-mechanical anode stub hole design tool. In Proc. TMS Light Metals, pp.433-438, 2010

Finally, the versatility of this new apparatus would permit the characterization of the steel stub deterioration in the anode or in the bus bar/aluminium rod connection. It could also be very useful for the optimization of the cast iron composition for the anode or cathode connection.

1008

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CARBON ANODE MODELING FOR ELECTRIC ENERGY SAVINGS IN THE ALUMINIUM REDUCTION CELL D. H. Andersen1,2, Z. L. Zhang2 Hydro Primary Metal Technology, Ardal, Norway 2 Dept. of Structural Engineering, Norwegian University of Norway 1

Keywords: FEM, Anode - to cathode distance, carbon anode,cast iron, non homogeneous, aluminium reduction cell, contact, rough surface, asperity, Hertz, Greenwood-Williamson Abstract The carbon anode geometrical design influences the energy consumption in aluminium production. A 2D finite element model (FE-model) of an anode immersed in an aluminium reduction cell has been developed to study how the anode geometry affects the variation in the anode to cathode distance (ACD). Large variation in ACD will prevent a systematic reduction of the average ACD and thereby hindering a reduction of electrical power loss in the bath. Another modeling example focuses on the large amount of energy loss occurring at the anode-cast iron interface due to the roughness-induced contact resistance. An analytical equation for the real contact area has been established to link the electrical power loss from the contact resistance with the pressuredependent interface properties. The proposed contact model can be implemented in a full scale electrical FE-analysis of an anode, and used to optimize energy savings. Introduction Over the last 10 years there has been focus on current increase actions and also on reducing the noise in the ACD resistance in the aluminium industry. One of the major improvements of reducing the noise was the implementation of slots in the wear surface as illustrated in Figure 1 [1]. The produced carbon dioxide creates bubbles in the bath and the slots function as an escape route for the bubbles. The lowfrequencynoise components in the liquid metal have also been studied in a magneto hydro dynamic (MHD) aspect that has influenced design optimizations of the cell and also the bus bar system. There is also a focus on how the non homogeneous density of the bath creates non homogeneous ACD's [2], This gives motivation for alumina distribution to the cell through individual feeder control in order to reduce the density differences in the bath. Still there exist variations in the ACD resistance. In normal situations today it is reported that the standard deviation in the current load on individual anodes in the same cell often are more than 10% of the average current [3]. Actions for reducing the noise in the ACD resistance by slots in the anode has also lead to negative effects on other parameters, like current efficiency and increased non homogeneous anode consumption. It is emphasized that slot implementation must be treated individually for each plant in order to avoid pitfalls [4]. The numerical results of the initial state of the anode electrical current was used to analytically describe how this will affect the variation in the anode- to cathode distance (ACD) in a steady state scenario after several hours in the electrolysis bath.

parameterized with three new roughness parameters that control the root mean square height, the slope and the curvature of the rough surface, respectively. Unlike the roughness parameters used in the well known Greenwood-Williamson (GW) model [6,7], the three parameters introduced in this paper are independent of each other, which is an advantage as this facilitates the characterization of the isotropie random surface of the material [8,9,10,11]. However, a disadvantage with this parameterization is that the roughness parameters are not linked to a contact area equation any more, like the Hertz equations that express contact with linear materials. Different asperity profiles defined by the new roughness parameters were analyzed by a 2D axial symmetric FEmodel. The numerical results suggest a procedure for developing a contact area function that are parameterized with the three new roughness parameters, which is also suitable for non linear brittle materials. The real contact area function can be used to compute the contact resistance in the carbon-cast iron interface in different ways; either by implementing the equations in new finite contact elements, or in a boundary condition in a FE-model. The paper will first focus on the effect from electric current distribution in the anode and secondly the contact resistance in the cast ironanode interface. Parameterization of non homogeneous properties in the cell We can introduce a parameter, k, describing a relation of ideal electrical current density, / 2 , in the inner leg of the anode and a current density, J3, in the outer leg of the anode as shown in Figure 1. The current density, J3, can be taken as an abnormal current density and J2 as a reference current density. For the ideal initial state, k=ko, (assuming no non homogeneous bath effects change the current distribution) we have

*-**-£-*£

(1)

where I2 I3, A2, A3 are currents and cross sections of leg 2 and 3. For the ideal steady state, after several hours in the bath, different processes like surface overvoltage on anode wear surface, production of C0 2 and anode consumption will result in a homogeneously distributed current near the anode wear surface. This means that k - 1 and h - ^h = h, and the initial degree of non homogeneous current distribution, k0, has been transferred over to the steady state by equation 2.

Holms equation of contact resistance [5] is an example of how the contact area between two rough surfaces is used. The contact resistance is present due to asperities ("hills") on the rough surface which makes up the real contact in the interface. An asperity profile function has been proposed that describes the isotropie rough surface of a material. The function is

The components in the equation are illustrated in Figure 2 which shows the equivalent DC circuits of the anode in the bath. The high resistances in the cell and the resistances that are highly affected by the current density are included in the equivalent DC

1009

circuit (the Nernst potential, anode concentration overvoltage and the cathode overvoltage are therefore neglected). .L

m <*

Stφbbale

Jt

h

\

nap

Bottons of stuhhoW

//

%i

t/η ;

1

appears. Equation 4 can be used in anode design to reduce the ACD when the bath effects on the ACD are isolated ( m and n are set to constants). The bath conductivity relation, n, is defined with virtual conductivities. We can express this relation with real physical properties. It is mainly given by phenomena such as surface anode overvoltage and the presence of bubbles in the bath caused by formation of C0 2 . From Figure 2, a real physical expression for n can be derived from the voltage drops over each resistance in the bath.

Λ

d2

4

S i

Uba2

^ <#« 1

1«β 1

:

Ì Le$2

\

......μ^.

U'b«3 ~ Uov3 0 . + "Ubu3 ' + Uba3 1

Lea 3

Equation 5 can further be expressed by resistances and current densities [6] shown in equation 6,

; ULL

-h)h a-°baz

Figure. 1. The anode immersed in the electrolysis bath (the longitudinal direction of the anode is into the paper). The slots divide the lower part of the anode into three longitudinal "legs" with height d4. Dimension dj is the radius of the stub hole bottom.

tt--h)h 0ba3

(3)

l-U

L-L· ó

áçÆ

áç2

+ kQ-

RT 1.08F

{L-i2~db2y2

^i/;

CbtizO- - Ö2)

* ©

(6)

VbaZ

where R is the universal gas constant, F the Faraday constant, J0 the limiting current density in the surface overvoltage term, and db the bubble layer thickness which has a tendency to decrease with current in an electrolysis cell [12]. The bubble coverage, Ö, increases with the current density and isolates the anodes wear surface with bubbles in a higher degree [13]. We rearrange

When the resistances in the equation are expressed with its conductivities and dimensions (·* = 'óΑ) we get

ó

(5)

u ov2 + T Ubu2 Vbu2 + t ^ba2 Uba2 UQV2

13

'/

la equation 6 to find ba2. The expression for n can also be modified and include extra effects like anodic- and cathodic concentration overvoltage in the bath. The voltage drops from these effects must be added in the right side of equation 5 and a modified expression for n is derived from the extended equation (an extra resistance in each branch in the left DC circuit in Figure 2 will appear). We can also simplify equation 6 if the surface overvoltage term is linarized. The current densities, J2 and 73, in equation 6 will be omitted. The anode-bath conductivity relation, m, has also a virtual component in its expression. This can also be represented with physical parameters by Ohms law. From u=zRI>R= ΙóΑ and using Figure 2, the physical expression for m is shown in equation 7.

We assume homogenous conductivity in the anode and introduce the conductivity relations Equation 3 can then be expressed as

7a

ba2

n = <W . and

kQnl2 + k0mn(L — l2) — ml

< Ran

Anode

I Anode wear surface

(4)

<***!

X è Ri

^αη^αη^

I ™U

Bath

J

ba2

& bmsW

Metal pad

Figure. 2. Equivalent DC circuit of an anode in the bath. The current density, /, is affected by anode resistance, Ran, surface overvoltage resistance, Rov, bubble resistance, Rbu, and the resistance in the bath, Rba. The left circuit can be simplified to the right DC circuit where Rba is the resistance that includes Rov, Rbu and Rba. For the case of n = k0 = 1, no difference in the ACD, due to anode design, will appear since l3 = l2 in equation 4, ACDref = L-l2- ACDabnorm = L- l3^ an( j n o n o n homogeneous anode consumption, due to anode design, will occur. If the reduction cell and the anode are designed and will run with k0-, mand n-values that result in *3 ö h in equation 4, non homogeneous consumption of the anode takes place and a difference in the ACD

1010

bajneas

(7)

ACDrefIan

The area, Aan, is the anode wear surface, óαη the anode conductivity, AC®ref = L — l2 m e reference ACD, Ian the electrical current through one anode, and Ubameas is the measured/calculated voltage drop between the anode and the metal. The real conductivity of the bath is normally around 200 300 S/m. The total virtual conductivity of the bath, ?ba (Figure 2), which also includes surface overvoltage and bubble overvoltage, will be much less. For example, a conservative cell with ACDref=0.04m and Ian=S 000 A is about m » 195. There is a different m for each aluminium plant and it depends on how the manager decides to run the cell. For a current increase project in a plant the reference ACD will be decreased. Typical values [14] can be ACDref= 0.03m and a chosen Im= 10 000 A. This results in m a* 220 from equation 7. In the numerical study k0 will be determined and the results will be linked to the variation in the ACD in equation 4 with specific values on m and n. It is not the

intention of this paper to calculate m and n for many different aluminium reduction cells. The purpose of this paper is to define the parameters and illustrate them with one cell solution (m = 200, ACDref= 0.04m).

distribution along the boundary 2 cm above the bath level for the best- and worst &-case is shown in Figure 4.

r;

Results and Discussions of the FE-Analyses on Anode Current Distribution

<

£·

In the numerical 2D analyses the electric current has been parameterized to enter in four different ways; i0: upper half of stub hole wall, ij: whole stub hole wall, i2: the whole stub hole (wall and bottom) and i3: bottom of stub hole.

=X

7 -0

Ö

1 0,95

" —"Best k-case

Worst k-case

v N-

J\

0,9 0,85

^^^^^

0,8

= {1-4,1.75}

and The slots got two different depths, Ύ^ΙΑ -{1.00,1.176,1357} , . _ _ a positioned at * (Figure 1). Data for estimating k0 was found on the boundary 2 cm above the bath shown in Figure 3 (or 17 cm above the anode wear surface). The current density, J3, is found by integrating the current along the boundary crossing the outer leg of the anode. The current density, J2, is found by integrating the current along the boundary crossing the inner leg of the anode. The conductive Media DC module in the FE-software, COMSOL 3.4, was used with 2D quadratic Lagrange elements. A mesh was defined for the whole domain with maximum element size scaling factor = 0.08, element grow rate= 1.2, mesh curvature factor = 0.25, mesh curvature cutoff = 0.0003, resolution of narrow regions = 1. This gave an element number of 75 000 for the whole domain shown in Figure 3.

Anode width

Figure. 4. Electrical normal current density distribution over the anode width along the boundary 2 cm above the bath for the best k-case (feo = 1.06 forγ=Λ,â = 1.4, ί = ί 3) md w o r s t k . c a s e (fc0 = 1.20/or r = 1.36,/? = 1.75, ί = î 0 ). The boundary line is shown in Figure 3. With k0 found above, and with a simulated cell setting of m = 200 we can set these values into equation 4 and study how an initial current distribution in the anode can affect the ACD. Figure 5 describes the critical n (nc) we must have for keeping l2 = l3 for different ko when m = 200, using equation 4. The worst case of anode design, k0= 1.20, demands a bath conductivity relation, n = 0.83, to avoid a variation in the ACD by ko. If n = 0.9, it is seen from Figure 6 that '3 " . With l2 = 600 mm, we have l3 = 596.5 mm, which means a difference in the ACD of 3.5 mm. Values of k0, found from realistic anode designs, can easily induce a difference in the ACD around 10% of its value if the critical n is not reached. We should keep in mind that the k0 found from the numerical analyses are based on average values of current densities. In real cases, if we take into account non homogeneous frozen bath on the anode wear surface, the "real k" can be much higher than the model shows.

ø Current input: /a„=7,000 A

Boundary 2 cm above the bath

1,1 1,05

Anode

Bath Ground on metal pad

Variations in the ACD

ACD ref=0.04m

Figure. 3. The 2D domain consists of the sub domains yoke, cast iron, anode and bath. The current enters from the top of the yoke and penetrate down to the metal pad defined as the boundary of ground. All other outer boundaries got an electrical insulated property. Here the current distribution is shown with current setting, i0.

0.5 1:1

1:2

IA

1.5

lt6

1.7

k0

Figure. 5. The critical n for different ko to keep l2 = 13 so that no difference in the ACD will occur. The anode-bath conductivity relation is set to m = 200, ACDref = 0.04m and l2 = 600 mm in equation 4.

The best case, &o=1.06, has a slot position of γ= 1, slot depth of â = 1.4 and electrical current input setting, i3. This is an anode with deeper slots closer to each other, and with an electrical current entering the anode in the bottom of the anode stub hole. The worst case found from the numerical analyses, k0 =1.20, has a slot position of γ= 1.36, slot depth of â = 1.75 and electrical current input setting, i0. This is an anode with shorter slots further away from each other, and with electrical current entering in the upper half region of the conic stub hole wall. The current density

It can also increase by asymmetric electrical coupling of the yoke to the anode with cast iron. The electrical coupling between the stub holes in one anode can also differ and increase k0. A real k0 of 1.4 demands a bath conductivity relation, n < 0.71, if a variation in the ACD is to be avoided (Figure 5). Another point to remember is that the bubble coverage under the anode wear

1011

surface as function of anode current density has a lower slope for current densities above 1 A/cm2 than for current densities beneath 1 A/cm2 [13]. 1,05

-B-0,9 -A-0,8 -tt-0,7 1

1,1

1,2

1,3

1,4

1,5

1,6

Figure. 6. The relation ljl3 as function of initial anode current distribution, k0, and the bath conductivity relation, n, for a reference anode height, l2 = 600 mm, m = 200 and ACDref= 0.04 m. This means that the bath conductivity relation, n, is larger for high amperage cells. The risk is higher that the initial electrical current distribution in the anode creates variations in the ACD for a high amperage cell. Parameterized Contact area function The contact resistance in the cast iron-yoke interface is present due to rough surfaces, where the "hills" (asperities) on the surface makes contact and decides the real contact area. The real contact area is therefore a sum of contact "islands" within the nominal contact area. ac : Contact pressure z(r) Initial asperity profile Deformed asperity profile

compressed for different initial profiles decided by the developed profile function, z(r). From each of the numerical analyses the contact radius, rc, contact stress, °e, and contact strain, f c , was found. The profile function, z(r), was developed and parameterized with three independent roughness parameters. A roughness parameter with similar properties as the known rms roughness is chosen to be the dimensionless crest factor, Ccn (peak- to rms height ratio defined in equation 11). If the peak height is constant, the crest factor has a linear relation to the rms height of the asperity. The crest factor does not need to change if the territorial width of the profile changes. The territorial width can be linear to the rms slope of the profile if the peak height is 1. The second roughness parameter is therefore defined as the widthto peak height ratio, rvv,=rv//, and illustrated in Figure 7. It should be noted that rw and Ccr are dimensionless since they are ratios of distances. These two roughness parameters can also be made independent of each other; they do not need to correlate. The roughness parameter, rw , has similar properties to the well known Q _ #summits/

density roughness parameter, / Apparent Area^ s j n c e they both contain information of the territory of an asperity and the frequency of asperities along the surface. The third roughness parameter is a shape parameter, ί , describing a shape of the profile in the same way as the circular shape parameter, /?, in the GW-model. When both the rms slope and rms height for a profile is set (and peak height is unity), the profile is still free to vary its curvature within these limits. This means that the shape parameter, ί , is defined to take those shapes that do not alter the rms height and the rms slope of the profile. The shape parameter can therefore generate different shapes instead of only one specific shape as the circular shape described by ί in the GWmodel. The proposed profile function defines it in more detail below. The shape parameter has to change the curvature within a peak height of unity and a constant finite width. This can be fulfilled by moving the function downward along the negative zaxis, but at the same time keep the maximum value of the function equal to 1. The first version of the profile function can now be given in equation 8: 1 + βτ

*M--

0.5

1+

—

)

-Êt

(8)

2 q

'3dB )

Figure. 7. Schematic drawing of an asperity profile compressed by a contact pressure down to a new height, zc — *- ~ Ec, and contact radius, rc. The asperity is axial symmetric about the z-axis. The initial asperity profile is given by the developed function, z(r).

The shape parameter, ί , is defined in the range, ° < £ f < 1, due to the unity of the peak height (and 0 < z < 1 )t Now, we want to find an expression for the order, 2q. It should be expressed by the width of the asperity at z=0. We solve for 2q by setting z(rw=0). The order is then given by 1 + 2/3(t1 Ú lnJ

(9)

2q =

For the cast iron - anode interface, the contact resistance has earlier been found experimentally [15], and laboratory data have been curve fitted [16] to find expressions for the contact resistance, mainly as a function of pressure and temperature. In this paper it has been focused to find an expression of the contact resistance by parameters describing the roughness of the interface, and how the roughness changes by the contact pressure. This can improve the accuracy of calculating the contact resistance for different anode stub hole designs A single asperity (one single contact "island") on the rough isotropie surface was analyzed in a 2D axial symmetric model in COMSOL 3.5. The asperity was

In

' 3 dB

We set equation 9 into equation 8 and get the asperity profile function in equation 10.

1012

ι + £*

φ)=-

-ίJ

L±2£i t i

(10)

3dB )

1+

l U*J

Now we have two of the defined roughness parameters in the equation. We still miss the crest factor. Instead of expressing z(r) with r3dB, we rather exchange it with the crest factor. The crest factor, Ccn of z(r) from equation 10 is given as the peak- to rms height ratio in equation 11. 2(0)

Lrr

(11)

J£j>(r)Pdr

Equation 11 was solved for different values on rw*r3dB>ί , since the whole symbolic equation is not solvable. It was found that the r

ZdBf

ratio,

ίf.

*rw, has a power relation,

r*dB=(a(ί^ίf)

+

r

b .

aL,

cr "+"c, for each value of

c(ίi))rw

r3dB= [{1.034·0f3 - 0.7977-ίf2 + 0.6285/3f + l.On)· c(-0J371./

+ 1.701.flt-iJ76)

+

(.j^./

+

better than R2=0.9995. If we set r3dB in equation 12 into equation 10, we have the complete asperity profile function with three independent roughness parameters, rw> ^crS . The function, z(r), shown in Figure 8 and expressed in equations 10 and 12, can appear complicated. But r3dB is only a single value when the roughness parameters are set. It contains no argument, r, of the function. The argument of z(r) occurs only once in equation 10, and the order (2q) of r is also a single value when the roughness parameters are set. This simplifies mathematical handling considerably. In the numerical analyses a 2D axial model was used due to the symmetry of the asperity. Special attention has been paid to the mesh of the asperity, with a maximum element length of le-4 in the boundary of contact, element growth rate of 1.2 and mesh curvature factor of 0.1. Triangular elements were chosen for the asperity and quadrilateral elements for the smoothed stiff material which compressed the asperity downward. The number of elements varied with the roughness parameters since the area of the material changes. The yield stress, ^ 5 , of the anode material was set to 20MPa. The stress-strain curve of the material followed the Ramberg-Osgood stress-strain relation. The hardening parameter, n, represents a linear material for n=l and ideal plastic material forîl -* 00. The material was analyzed for n - {1,2,3,4}9 w i m a degree of plasticity far greater than a carbon material, which in many cases are regarded as a linear material. The general yield offset value was set to €py* ~ a ' E ~ · The parameter, φr, was set to 1 for all the numerical analyses. All the asperity profiles were analyzed with contact strains, €c ^ 1%·. Asperities with low crest factors have been analyzed for contact strains greater than 1%, so that all asperities have reached the limit, ó° ~ >r*. One specific asperity was compressed to 3% contact strain. For the territorial radius, two values were chosen; one minimum value of 0.5, and a maximum value of 1.3 (rw = 10.5,1.3]). p o r m e c r e s t factor, five values were tested, £cr = {1.2,1.4,1.6,1.8,2,0} ^ hardening value was tested with four values of n ~ i 1*2,3,4] p o r m e shape parameter, only the value of ί = ®^ was used. The variation in the shape parameter creates only a small variation in the contact area compared to the other roughness parameters, rw and Ccr. We therefore focused on the parameters that created large variations in the contact area. This resulted in a set of 2x5x4x1=40 parametric numerical models, where each model was displacement controlled.

(12)

Qj^ff2

-0.6236-^-0.01087)]·^ 0 < z < 1,0.1 < /?f < 0.9 0 < 7-3dB < rw.l < Ccr< 2.0,rw > 0

Results and Discussions of the parameterized contact function From the numerical analyses it was found that the contact radius

Figure. 8. Plots of the asperity profile function z(r) given in equation 10 and 12 with rw=0.5 for all the plots. Each of the three figures is plotted with five different shapes, . The five different shapes in each plot have a common crest factor; left plot with Ccr=2.0, middle plot with Ccr=1.6 and the right plot with Ccr=1.2.

had to be scaled by the contact strain, as /rc which is the same as the mean slope of the compressed "vanished" part of the asperity profile (Figure 7). When this ratio became a function of the ratio,

The power fit was better than R2=0.9998, where R2 is the

the parametric settings in the numerical study, as shown in Figure 9. The proposed form of the function for expressing the contact radius is given in equation 13, where the argument is

coefficient of determination (on 50 samples within ' rw 0 for each P in the range, Each coefficient, a, b, c, in the power relation is a function of that r3dB c a n b e expressed as in equation 12. The coefficients of equation 12 were found for nine values on ί from 0.1 stepped with 0.1 up to 0.9. The polynomial fits described in equation 12 of each coefficient was

€c

/'rd, it was observed a high degree of linearity regardless of all

' r «, and the parameters are **cr* rwf ί *çº and rci=r(zc)=r(l-ec) from equation 10 and 12 and Figure 7.

-(I)"

1013

= OliCc,

^,«)(|)'

«j(«OH»«·/» ·¼

(13)

Condition: {^yVE = const.,sPys = const.] or [(«Pjfi) ^<(0,β}

distribution in the roughness parameters. As long as the distributions are known, the total real contact area will be a sum of each distributed asperity.

<

Conclusions

All the curves are above the bold line, y=x, in Figure 9. This means that rc< rci (except for the zero displacement case where r < c rci). The four types of effects in Figure 9 are forcing the curves to a lower slope, which means they increase the contact area of an asperity. For an ideal plastic material (?i —> OO) the curves will follow the line, y=x, since rc ~* rci (the deformation is following the original profile). A decreasing crest factor, Ccr-*1 (solid cylinder) will also force the curves close to y=x (if we neglect the Poisson effect). The linear properties of the function in equation 13 holds also for contact strains larger than 1 %. The asperity, fa = a 5 ' c « - = 1 · 6 ' ίf = a i » " = 3 ), was numerically compressed to 3% contact strain, and the linear property shown in Figure 9 was maintained.

For realistic anode designs the ko, describing the degree of initial non homogeneous electrical current distribution in the anode, can range from 1.06 to 1.20 (from 2D modeling). If the anode is designed with a k0= 1.20, the bath conductivity relation, n, should be less than 0.83 for the analyzed cell if the variation in the ACD is to be avoided. The best fc-case (low k0) describes an anode with deeper slots closer to each other, and with an electrical current entering the anode in the lower parts of the anode stub hole. For an isotropie non linear brittle material the contact area of a rough surface can be found from a function that is parameterized with three new roughness parameters and the hardening value of the material. The independence of the parameters facilitates the characterization of random rough surface and increases the probability to develop contact area functions from regression and curve fitting of data from numerical studies. References [I] [2] [3]

[4] Figure. 9. The function, results.

' r c versus ' r « , for all the numerical

[5] [6]

There are many reasons for the function, in equation 13, to be solvable by curve fitting: 1) A wide range in each of the

[7]

r

roughness parameters in equation 13 give values of * c in almost the same range. 2) The high degree of linearity of the function in equation 13 makes it solvable by curve fitting. 3) All the parameters have a property of converging the slope of the function

[8] [9]

to 1 ( frc ~ 'ra\ as shown in Figure 9: i) A lower crest factor, Ccr, will move the profile towards a cylindrical shape (Figure 8). Without any Poisson effect, rc is approximated to rci for a crest factor equal to 1 (solid cylindrical asperity), ii) A larger rw will increase the average of rc for the same variation in the other parameters in equation 13. iii) A smaller ί will transfer more mass of the asperity to the peak (figure 8) and reduce the local slope around the peak, and rc has to increase, iiii) A higher degree of plasticity (increasing n) will reduce the difference between rc and rci. The contact radius (contact area) of one asperity on the rough surface can therefore be found if the contact strain, £c, is known and used in equation 10 with the right values on the defined ~ I , /rlra.t The argument, r o roughness parameters to find the argument,

[10] [II] [12] [13]

[14] [15] [16]

r

ci and the values of the parameters is set into equation 13 €c / which computes the ratio, /rc, to determine the contact radius. The asperities can be distributed along a surface width a

1014

M. W. Meier, R. C. Perruchoud, W. K. Fischer. "Production and performance of slotted anodes". Light Metals 2007, 277-282. B. Moxnes, A. Solheim, M. Liane, E. Svinsβs, A. Halkjelsvik. "Improved cell operation by redistribution of the alumina feeding". Light Metals 2009, 461-466. A. Solheim, B. Moxnes. "Anodic Current Distribution in Aluminium Electrolysis Cells". Paper presented at the XIII International conference "Aluminum of Siberia-2007", Krasnoyarsk, Russia, September 11-13,2007, 21-27. X. Wang, G. Tarcy, S. Whelan, S. Porto, C. Ritter, B. Quellet, G. Homley. "Development and Deployment of slotted anode technology at ALCOA". Light Metals 2007, 299-304. R. Holm. "Electric Contacts". Springer 1967. M. Braunovic, V. V. Konchits, N. K. Myshkin. "Electrical Contacts - Fundamentals, Applications and Technology". CRC Press, 2007 J. A. Greenwood, J. B. R. Williamson. "Contact of nominally flat surfaces". Proceedings of the Royal Society, A295: pp 300-319, 1966 D. J. Whitehouse, J. F. Archard. 'The properties of random surface of significance of contact". Proc. Roy. Soc. Lond., A316:pp 97-121, 1970 Myshkin, N. K., "Tribology of electrical contacts", In Tribology in the USA and Former Soviet Union, Allerton Press, New York, pp. 341-364,1994 B. Bhushan. "Contact Mechanics of rough surfaces in tribology: multiple asperity contact", Tribology Letters, 4: pp 1-35, 1998 R. A. Onions, J. F. Archard. "The contact of surfaces having a random structure". J. Phys. D, Appi. Phys., 6: pp 289-304, 1973 W. Haupin. "Interpreting the components of the cell voltage". Light Metals 1998, 531-537. N. Richards, H. Gulbrandsen, S. Rolseth, J. Thonstad. "Characterization of thefluctuationin anode current density and "bubble events" in industrial reduction cells". Ligth Metals 2003, 315.322. A. Solheim. "Anode-Cathode Distance in ΐI and ΑΠ". Internal Report 2008, Sintef Materials and Chemistry. M. S0rlie, H. Gran, Cathode collector bar-to-carbon contact resistance, Light Metals, 1992,779-787 D. Richard, M. Fafard, R. Lacroix, P. Clery, Y. Maltais. "ThermoElectro-Mechanical Modeling of the Contact between Steel and Carbon Cylinders using the Finite Element Method", Light Metals 2000, 523-528.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION

Cathode Design and Operation SESSION CHAIR

Richard Jeltsch Consultant Spokane, Washington, USA

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PREHEATING COLLECTOR BARS AND CATHODE BLOCKS PRIOR TO RODDING WITH CAST IRON BY PASSING AN AC CURRENT THROUGH THE COLLECTOR BARS 1

Erik A. Jensen1, Hans Petter Bjornstad2' Jan D. Hansen2, EAJ Consulting, 11020 Crossdale Lane, Mechanicsville, Virginia 23116, USA, 2 ALMEQ Norway AS, P.O. Box 50, N-1405, Langhus, Norway

Keywords: Cathodes, Cathode cast iron rodding, Collector bar and cathode block preheating, Collector bar as resistance heater Abstract Three basic methods for heating collector bars and cathode blocks, prior to pouring cast iron, are in use today: gas burners directly impinging on the collector bars, ovens for heating bars and blocks separately, and third, passing an alternating electrical current through the collector bars to heat bars and blocks simultaneously. This paper examines electrical heating using the collector bar as the heating element. Passing an alternating current through the collector bar produces an easily regulated and uniform temperature throughout the bar. Radiant energy from the bar heats the slot area of the cathode block. Temperature levels are adjusted by time and voltage selection. Electrically heating collector bar/cathode block assemblies uses less than 15% of the energy required for propane gas burner heating. The method is quiet, requires little or no supervision, has no products of combustion to exhaust, and temperatures are highly repeatable.

procedures used during rodding (sealing) with cast iron may be the cause of these types of cracks.

Introduction The voltage drop between collector bars and carbon cathode blocks depends on the contact pressure between them. If the pressure is too low, the voltage drop will be high. If the pressure is too high, the block may develop a longitudinal wing crack, eventually allowing metal pad penetration to the collector bar and ultimately cathode failure.

Figure 1. Types of potential cathode block cracks from casting (WC = wing crack, TC = transverse crack) (redrawn from reference [2]). When molten iron at 1,300°C - 1,400°C is poured into the gap between collector bars and cathode blocks, the carbon experiences a thermal shock. Collector bars initially have a higher temperature at the surface closest to the slot bottom, due to the freezing iron. The temperature differential between bar top and bottom causes the bars to bend with ends higher than the middle. Because the bar slot is keyed and the iron freezes rapidly, this initial curvature of the bar creates bending stresses in the block wings. The stresses, thus created in the slot corners, may be near or exceed the elastic limit of the cathode block and create wing cracks (WC in Figure 1). As the temperature in the bar equalizes, the bar straightens, and the forces on the block wings diminish [3]·

Cast iron rodding or sealing of collector bars to cathode blocks is the most common method of connecting the two components in use today. When done properly, the method provides a good mechanical and electrical joint between bars and blocks. When done improperly, longitudinal wing cracks and transverse cracks may develop. Cracks, created by improper preheating and casting procedures, may appear after iron pouring and then seem to disappear as the blockfàar/iron assembly cools. These cracks are still there, even though they are not visible [1] and may be the cause for premature cathode failure.

As the iron cools and solidifies, the block develops a higher temperature in the slot than at the bottom of the block, causing the block to bend. The portion near the middle of the bar becomes higher compared to the block end or portion near the bar ends. This bending, created by block differential expansion, may create transverse cracks in the tops of the block wings, if block tensile strength is exceeded (TC in Figure 1) [3].

The purpose of heating collector bars and cathode blocks, prior to sealing with molten iron, is to minimize potential cracking due to thermal shock, differential thermal expansion [1], and rapid expansion of water vapor in the blocks [2]. Another purpose of heating bars and blocks is to prevent steam blow-back of molten iron.

Proper preheating of bars and blocks and other procedures help reduce the thermal shock and stresses generated by the hot iron.

Cathodes experience a complex set of stresses during their lifetime, starting with the molten iron sealing of steel collector bars in prebaked cathode blocks. Blocks are susceptible to wing cracks (WC) that start at the slot corners, and transverse cracks (TC) located at or near the mid-points of the bar length. Improper

Most of the current between collector bar and cathode block flows through the sides of the cathode block slot [4]. The electrical resistance of this connection goes down with increasing contact

1017

pressure between the cast iron and carbon. However, too much pressure can exceed the bending (tensile) and shear stress limits of the block, resulting in wing cracks [5]. Small wing cracks that appear immediately after pouring iron, and then disappear, may result later in separation of the wing during pot operations. If wing separation occurs, electrical resistance between bar and block increases and the block integrity is compromised. An increased potential for wing separation may occur during operation, if bars are not preheated prior to pouring cast iron.

Table I lists bar and block temperatures prior to pouring cast iron specified by different block suppliers and users. Recommendations for block temperatures vary from 200 °C to 350 °C. Recommendations for bar temperatures are in the range of 500 °C to 700 °C. Heating time for cathode blocks should be sufficiently long to minimize moisture, especially in the slot area. In practice, these temperature ranges may be wider depending on the individual cathode sealing operators.

Table I. Collector Bar and Cathode Block Preheating Requirements Prior to Pouring Iron.

Collector Bars Bar Temperature Time bars heated Cathode blocks Block Temperature Minimum temperature at cast iron pouring Time blocks heated 1 Iron temperature at pouring A B

Units

Cathode BlockMfg.A

Cathode Block Mfg.B

Pianti

Plant 2

Plant 3

°C

650-700lAJ

600-700lBJ

650-670

500

500

120

180

180

320-340

300

300

300

300

300

120 1430-1460

180 1450

180

Minutes 200-300

°C °C Minutes °C

1300-1370

Assumes bars and blocks are heated together by gasflameon collector bars. Cathode block slot edges protected by steel angles. Recommended minimum of 550°C at bar ends. commercial installation of the conductive cathode bar and block heater. The system has been further refined to heat collector bars and cathode blocks in a variety of sizes, and the heating time has been shortened. A schematic diagram of the method is shown in Figure 2.

Discussion There are three primary methods for preheating bars and blocks: 1. Gas or oil burners withflamesimpinging directly on collector bars mounted in the slots of cathode blocks. This method heats the top of the collector bar and the slotted surface of the block. 2. Furnace heating of the bars and blocks separately. Bars are placed into the block slots hot, prior to pouring cast iron. 3. Electrically heating the bars using the bars as the heating element. In this method the bars are mounted in the blocks and provide radiant and convective heating to the block. This method is the subject of this paper.

Transformer to Bar electrical

The third method consists of passing an alternating current (AC) through the collector bars at a controlled power load and duration, to generate the heat necessary to raise the bars and blocks to their specified temperatures. The collector bars are placed in the slots of the cathode block and then connected in a series circuit with the alternating current (AC) electrical supply.

Transformer & Alternating Current Electrical Supply

Collector Bar

Insulated chamber

Cathode Block

Electrical Jumper Cable

Figure 2. Schematic diagram of collector bar heating method. Note the jumpers that connect the collector bars into a series electrical circuit.

The concept was initiated by ALMEQ together with the Norwegian inventor, Asbjorn Moen. The concept was tested at The Electrical Research Institute of the Norwegian SINTEF Group of the University of Trondheim. The initial tests used a two slot cathode block and two collector bars. The test results became the foundation for the development of the first

Electrical contact is made to the collector bars by conductive shoes clamped to the bars by pneumatically operated cylinders.

1018

The two drawings shown in Figures 3 and 4 demonstrate how this works.

The controller also has the capability to accept the input of individual cathode block serial numbers and to match these to their respective temperature curves for eventual use as pot lining historical data. The heat generated by the bars is radiated to the bottom and sides of the slot, and to the inside of the insulated chamber, where it adds heat to the block by convection. The latest system is equipped with a multi-winding transformer with a tap changer to optimize the heating cycle and to allow the system to accommodate different types and sizes of cathode blocks and bars that may be in use at the same plant. The transformer primary is wound to accommodate the plant distribution voltage. The secondary windings are designed to produce less than 40 volts. 12

Figure 3. Split bar jumpers.

|*«S

Ö

The collector bar ends are clamped between contacts in a series arrangement (creating a series electrical circuit as shown in Figure 2). For block/bar assemblies with split bars, contact blocks in the middle of the unit bridge the current from one bar to the next as shown in Figure 3.

m II r"~ - ir m I ^

1

ÎWMm

*_

^W

1 86 1 M iSf >w

c

1

%

* *

.

400 |

Collector Bar ^Temperature

1300 I

Cathode Block Temperature

2

15

30

45

60

75

m

1100

H

106 120

Time (Minutes) Figure 5. Bar and block temperature and amperage development over time at Plant 1.

I

pi

S

Jj

4

0

Mjj

1

10

s

A typical graph of current, collector bar temperature and cathode block temperature is shown in Figure 5. The steps in the current curve represent transformer tap changes to compensate for the increased resistance of the circuit, caused by rising collector bar temperature.

m f

Comparison of Bar/Block Heating Methods

}

Direct flame heating of bars in blocks usually involves a series of burners mounted in a line above the collector bars with flames directed to the top of the collector bars. The corners of the slots are usually protected by steel angles to prevent excessive air burn of the slot corners. Control is sometimes based only on length of time under the burners. The uniformity and repeatability of temperature from bar to bar, or within a single bar, are not assured. The uniformity and repeatability of temperatures from batch to batch are also not assured. The process is frequently noisy and creates products of combustion.

Figure 4. Contact shoes and bar to bar jumpers at nontransformer end of the assembly. The heating elements of this system are the collector bars. They in turn heat the cathode blocks. An insulated cover is lowered over the assembled bars and blocks, prior to heating, creating an insulated electric oven with collector bar heating elements. A programmable logic controller (PLC) records the temperatures and amperage during the heating cycle. When the set-point temperatures are reached, the system maintains the temperatures and alerts the operator.

Oven heating bars and blocks provides controlled temperatures to bars and blocks separately, but requires assembling hot bars to hot blocks prior to pouring cast iron. One would expect temperatures in each of the two furnaces to be highly repeatable. Energy consumption is relatively high, because heating requires either burners or electrical heating elements in each of the two furnaces (one for bars and one for blocks).

Block and bar temperatures are measured by thermocouples and continuously recorded by the system controller.

1019

Collector bar resistance heating by alternating current is efficient, quiet and produces no products of combustion. The bars are mounted in the blocks cold, prior to heating. After reaching the required temperature, the powered trolley holding the bar/block assemblies brings them to the pouring position for immediate sealing with cast iron. The system is efficient, because the electrical resistance of the collector bar is used to heat itself and the hot bar then heats the cathode block. Little or no oxidation of the slot edges has been detected. Therefore, steel angles are not needed to protect the slot edges from airburn. The process is very quiet.

The maximum temperature difference of steel collector bars was 24°C. The maximum temperature difference of cathode block measuring points was 14° C. Table II. Block and Bar Temperatures after 180 minutes. Recording Time: 180 Minutes

Energy Consumption During a recent expansion, Plant 2 converted from propane heating of cathode bar/block assemblies to alternating current heating of the collector bars. Plant 2 reported that the average propane energy needed was 946 kwh/cathode bar-block assembly. The energy consumption using the AC electrical heating system is 101 kwh/assembly. This is more than an 89% energy savings over the previously used propane heater. Another plant, heating with oil burners, required 1,335 kwh/cathode bar-block assembly. The AC collector bar heating system now in use consumes 142 kwh/assembly. This is also more than an 89% energy savings from the previous oil burner energy consumption. Temperature Studies

Measuring Point

Cathode Block Temperature °C

IR01 IR02 IR03 IR04 IR05 IR06 IR07 IR08 IR09 IR10 IR11 IR12

216

Steel Collector Bar Temperature °C 478

230 462 218 486 222 474 230 462 223 484

A temperature study using infra-red (IR) instruments was conducted at Plant 2 in 2009. Unfortunately, there is no corresponding temperature study of the propane burner heater used previously at this plant for heating collector bars and cathode blocks, prior to sealing with cast iron. Six cathode blocks, each with one slot and split collector bars, were heated for 180 minutes. Infra-red (IR) temperature measurements were taken at the twelve points shown in Figure 6.

-A X X r

Recordings on (op:

?

;FLIR Figure 7. Infra-red photograph of bars and blocks at Plant 2. Another indication of temperature distribution is the start-up data recorded at Plant 1. Bar and block thermocouple temperature readings were recorded for twelve batches of blocks and bars. Each batch consisted of 6 blocks, 3,250 mm long and each block had one collector bar.

Figure 6. Locations of infra-red temperature measuring points. Measurements were taken when power was off and with the cover open. The cycle was interrupted several times to move the cathode bar/block assemblies out from under the heater for IR photographs. Table II lists the temperatures measured by the infra-red devices and Figure 7 is an infra-red photograph illustrating the relative uniformity of temperatures.

The cathode block temperatures for the start-up batches are shown in Figure 8. There were three temperature measuring points for each batch and 12 batches are presented. The maximum cathode block temperature differential batch to batch was 50° C, and the maximum temperature differential within a batch was 16°C.

1020

400

ΤΤ09

•«•ττïβ

380

-^~ΤΤ10

>3β0 »340

-»—«h

1320

Once activated, the system automatically positions the trolley containing the bar/block assemblies in the heating position, then applies pneumatically activated electrical contact shoes and switches on the alternating current electrical power. The system is designed to accommodate blocks and bars of different lengths at the same plant, by entering the appropriate lengths in the control panel.

300 280 1

2

3

4

6

6

7

8

9

•

10 11 12 13 14 1S

Batch Number

* : m**

Figure 8. Cathode block temperatures for 12 batches of blocks.

-. », :::

Figure 9 shows the collector bar temperatures for the 12 batches. There were three bar temperature measurements per batch. The maximum collector bar temperature differential, batch to batch, was 24° C, and the maximum temperature differential within a batch was 16° C.

1,

«""$

^-..-fffh *ft 1. Ü

"é

*·**■*
ft. 1 Tray& Trolley #1 1 Môvfog into héâtftf

700

{

*«Wk

1... Ë

I m &t3 «wa i m-

ÎI"W i Λ^

\

Ü" \xm i r i ^

MoSi&n Iron Being I Poured 1 Tray & Troltòy «2 1 1 Control Panel ■

1 H«rter

■

I

. . .

Figure 11. Heating System at Plant 1. Note: one tray of block/bar assemblies is entering the heater and one tray of assemblies is being sealed with molten iron. 640

When the required bar and block temperatures are reached, the power is shut off, the contact shoes are released, and the trolley is moved out to the pouring position. However, bars and blocks can remain in the heater with the cover on, to retain heat until the iron crucible is ready. Movement of the block trolley from release of the conductor clamps, until it is located in the pouring position, takes about 90 seconds. Pouring iron from a single crucible for six, 3,225 mm long blocks, each with one bar, takes about 10 to 12 minutes.

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 Batch Number

Figure 9. Collector bar temperatures for 12 batches of bar/block assemblies at Plant 1. Equipment Arrangement and Operation The equipment layout for Plant 1 is shown in Figure 10. A single heater serves two trolleys. With this arrangement, one batch of bars and blocks can be heated, while another batch is either sealed with cast iron, or bars and blocks are assembled prior to heating. Figure 11 is a photograph of this arrangement.

There appears to be no impediment in using the AC resistance heating method to heat composite collector bars (a copper rod inserted in a longitudinal hole in the steel bar). Summary: • • • •

• Figure 10. Layout of Heating and Rodding (Sealing) System for Collector Bars and Cathode Blocks at Plant 1.

1021

Collector Bars and cathode blocks can be preheated successfully by connecting the bars to an appropriate alternating current electrical supply. The block bar assemblies are heated inside a covered and insulated box, effectively forming a furnace with collector bars acting as the heating elements. No products of combustion are emitted from the process. Steel angles are not required to protect the top corners of the collector bar slots. No, or extremely small, evidence of block oxidation has been observed in this area. The process makes efficient use of energy to heat the bars and blocks. Two users of the electrical heating system have reported energy savings of 89%, compared to their previously used propane and oil fueled heaters.

Block and bar temperatures are measured by thermocouples and controlled and recorded by the system PLC. Noise levels of the process are very low. The heater can accommodate bars and blocks of different sizes and lengths at the same plant. The system works with either continuous or split collector bars. References: Morten Sorlie and Harald 0ye, "Cathodes in Aluminium Electrolysis T** Edition ", Aluminium-Verlag GmbH, Düsseldorf, (1994), Page 52. D. Dumas and J. Vallon, "Improvement of the Casting with Cast Iron of Collector Bars in Large Length Cathodic Carbon Blocks", Light Metals (1973), 641-646. Bιnιdicte Allard, et al., "Modelling of Collector Bar Sealing in Cathode Blocks with Cast -Iron", Light Metals (2009), 1097-1102. Morten Sorlie and Harald 0ye, "Cathodes in Aluminium Electrolysis 2nd Edition ", Aluminium-Verlag GmbH, Düsseldorf (1994), Page 301. Morten Sorlie and Hermann Gran, "Cathode Collector Bar-to-Carbon Contact Resistance", Light Metals, (1992), 779-787.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DEVELOPMENT AND APPLICATION OF AN ENERGY SAVING TECHNOLOGY FOR ALUMINUM REDUCTION CELLS Peng Jianping1, Feng Naixiang1, Feng Shaofeng2, Liu Jun3, Qi Xiquan4 School of Materials and Metallurgy, Northeastern University, Shenyang, 110819, China 2 Zhejiang Huadong Aluminum Corporation Ltd., Lanxi, 321103, China 3 East Hope Baotou Xitu Aluminum Ltd., Baotou 014030, China 4 Northeastern University Engineering & Research Institute Co. Ltd; No.73, Xiaoxi Road, Shenhe District, Shenyang, 110013, China Keywords: aluminum electrolysis, novel structure cathodes cells, energy savings In the near future, the NSC with crisscross ridges as shown in figure 2 will also be applied.

Abstract An energy saving technology based on novel structure cathodes in aluminum reduction cells has got wide application and development in many smelters in recent years. Structural and operating characteristics of the cells are described in this paper. Some details, such as lining structure, heat balance, current efficiency, effect of bath chemistry and wear of cathodes, are discussed according to the present applications. The results of six months' testing in a 200kA potline showed energy savings of 1030 kW»h per tonne of aluminum. Erosion of the ridges is projected to be less than 10% per year. Introduction Though lots of work has been done, much energy is consumed during the aluminum reduction process. About 14100 kW»h was required to produce one tonne primary aluminum ingot in China last year. At present 1 kg of standard coal generates 3 kWeh energy and produces 2.5 kg of C0 2 from coal-fired power plants which supply over 80% electric energy output in China. Therefore energy saving in the aluminum reduction process decreases both production cost and green house gas emissions. For a smelter producing 200 000 ton per year, if 100 kW»h energy for each ton of aluminum produced is saved, it can win about ¥8,000,000 due to 20,000,000 kW*h electric power saved every year.

Figure 1 NSC pots with longitudinal ridges

So an energy saving technology with novel structural cathodes (NSC), invented by Prof. Feng in 2007, attracts people's attention [1,2]. Three 168 kA test pots with NSC technology in Chongqing Tiantai Aluminum Industry CO., Ltd., which started-up in March 2008, have been working ever since. NSC technology has also been applied in 170 kA, 200 kA, 240 kA, 300 kA, 330 kA, 350 kA, and 400 kA pots in recent years. These NSC pots can work steadily at 3.7 V to 3.8 V. Following the initial success of NSC technology, Feng puts forward some other structures of carbon cathodes with various shape ridges, such as rectangle or trapezoid in vertical section or column, and different ridges arrangements, such as end to end, side by side, or crisscross, in his patents [2,3]. There are mainly three kinds of cathode structures as shown in figures 1-3. From Figure 1 it is shown that there is a gap, with width 150 to 200 mm, between two ridges of the carbon cathode, and there are two rows of ridges for a wider carbon. These two types of cathode are widely applied in smelters in China at present.

Figure 2 NSC pots with crisscross ridges

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Another NSC in Feng's patent that is made up of graphitized carbon block and anthracite cylinders, as shown in figure 3, should have more advantage in energy saving, because of lower resistance in the graphitized carbon block.

aluminum baking with strong thermal shock could damage the cathode. So in the tests both NSC pots and Ref pots have been baked by aflame-liquidaluminum preheating method, including 24 hours flame preheating from room temperature up to about 700C, and then 72 hours baking with liquid aluminum as a resistance. These pots were divided into three batches to bake in order. The first batch, including 29 NSC pots and 29 Ref pots, was baked in June 2009. The second batch, including 34 NSC pots and 2 Ref pots, was baked in July 2009. And the last batch, including 31 NSC pots and one Ref pot, was baked in August 2009. Technology of 200 kA NSC potline NSC pots use process technology conditions shown in table 1. Table 1 The cell operation conditions of NSC pots 206±lkA Current 3.700-3.750V Working voltage 19~21cm Bath level Aluminum level 18~20cm Electrolysis temperature 950~960°C Initial crystallization temperature 945-955 °C Molar ratio 2.3-2.5

Figure 3 NSC pots with graphitized carbon block The application and development of NSC technology and its related technology are discussed in the present paper. Application and Development of energy saving technology The energy saving technology with NSC invented by us has already been applied in many smelters after successful testing in Chongqing Tiantai Aluminum Ltd. Remarkable energy saving efficiency of 800 to 1000 kW-h/t has been achieved in most smelters using this technology; however, there are some differences in the energy saving efficiency due to technological conditions, operation methods and structural design of cells. A typical case[4] is the 200 kA NSC potline in Zhejiang Huadong Aluminum Corporation Ltd..

Results Working voltages from Sept 2009 to July 2010 are listed in table 2. In order to work normally at lower voltage, the potline amperage has been increased gradually from 200 kA to 206 kA since Sept 7, 2009. And now most NSC pots can be work steadily at 3.72 V. Table 2 Comparison in working voltage (in Volts) between 94 NSC pots and 32 Ref pots ^_^_^__ Mar Oct2009 Nov Dec Jan 2010 Feb 94 NSC pots 3.742 3.715 3.726 3.714 3.716 3.715 4.082 32 Refpots 4.069 4.026 4.028 4.029 4.028 Comparison 0.340 0.343 0.312 0.313 0.313 0.313

NSC pots baking and startup An old 200 kA potline, with average cathode life of 2300 days, was shut down in September 2008. After the 200 kA potline was overhauled, NSC carbons were used for 94 pots with improved heat insulation in the sidewall, and traditional cathode carbons were used for another 32 pots (as reference pots, named Ref pots). Fig 4 shows a NSC cell.

Current efficiency (CE) of the NSC pots and 32 Ref pots is listed in table 3. The values are calculated according to weight of aluminum tapped from the pots. From that, the NSC pots can work steadily at lower voltage without any CE loss. Moreover, if the difference of aluminum level between before and after the six month is taken account, CE in 94 NSC pots and 32 Ref pots is 93.105% and 93.001%, respectively. Table 4 shows a comparison of DC power consumption between the NSC pots and the reference pots. Table 5 shows a difference of overall alternating current electric power consumption between the two type pots. From tables 4 and 5, it is shown that the DC power consumption in NSC pots is only 12043 kW#h, which is lower than that in Ref pots, and that the overall alternating current electric power consumption of NSC pots is only 12791 kW#h, which is obviously less than 14171 kW#h, the average value of traditional pots in China last year. So NSC pots can be applied rapidly due to their advantage over traditional pots in saving energy.

Fig.4 the photo of NSC pot For traditional pots, the coke baking method and liquid aluminum baking method are popular. The coke preheating method not only consumes coke but also brings about trouble in clearing residue of the coke from pots afterward.. However, because of these special structural ridges, like walls, on cathode carbon in NSC pots, liquid

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Table 3 Comparison in CE between 94 NSC pots and 32 Ref pots Nov Dec Oct2009 93.43% 92.15% 94 NSC pots 93.15% 93.21% 93.47% 32 Refpots 92.16% Comparison -0.04% -0.01% -0.06%

Jan 2010 92.74% 92.55% 0.19%

Feb 92.87% 92.96% -0.09%

Table 4 Comparison in DC Energy Consumption (in kWh/t) between 94 NSC pots and 32 Ref pots Dec Oct2009 Nov Jan 2010 Feb 12052 94 NSC pots 12109 12023 12070 11993 13144 13108 13080 13009 32 Ref pots 13060 Comparison 1074 1001 1028 986 1067

Mar 92.97% 93.03% -0.06% Mar 12013 13006 993

Average 92.88% 92.90% -0.02% Average 12043 13068 1025

Table 5 Comparison in overall alternating current electric power consumption (in kW-h/t) between 94 NSC pots and 32 Ref pots Mar Nov Dec Jan 2010 Feb Average Oct2009 12794 94 NSC pots 12746 12896 12837 12808 12791 12663 13834 32 Refpots 13937 13908 13828 13865 13775 13858 Comparison 1041 1034 1112 1071 1026 1119 1067 stability has an important influence on CE. The NSC design can improve the stability of the pad because of its special structural cathodes and thus increase CE. However, CE may be lowered if the NSC pot works under an anode-cathode distance (ACD) that is too low. So CE of the NSC pots depends on the two aspects.

Heat balance technology of NSC pots NSC pots usually work at cell voltage of 3.70 to 3.75 V, which is about 0.3 V lower than traditional cells. As is known to all, the resistance drop of the heat balance system from anode rods to cathode bars is only 2.3 to 2.4 V for traditional pots at cell voltage of 4.1 to 4.2 V. Input heat will reduce about 13% due to 0.3 V decrease in cell voltage for NSC pots. It means that 13% of the total heat loss should be decreased to keep the heat balance for stable production at normal electrolytic condition.

The three 168 kA NSC pots in Chongqing at about 3.76 V of cell voltage in the past two years achieved 93% CE, which is 1% more than other traditional pots in the same potline. The CE for 200 kA NSC pots in Huadong, operated at 3.72 V, is no less than the 32 traditional pots at 4.05 V in the last year. However, not all NSC pots at the same lower cell voltage achieve higher CE. Technological parameters, operation methods and structural design of pots can cause different CE. In order to achieve higher CE in NSC pots, three effective methods should be used as following,

An excellent design is crucial for steady and effective operation. The design concept for traditional pots is insulation in bottom and heat emission by sidewall. However, the concept is not fit for NSC pots due to a lower cell voltage. According to the working state of pots applying the NSC technology, a new heat balance technology should include several important aspects as follows: 1. 2. 3. 4. 5.

1.

Lowering properly the aluminum level to reduce heat losses by pots sidewall under good technological conditions; Adding the thickness of insulating layer of A1203 cover to reduce heat losses on the top, equivalent to the space left by the lower aluminum level; Radiating properly by sidewall as in traditional pots is of advantage to form thicker ridge which reduces surface area of metal pad and the dissolving loss of aluminum; Insulation amplified in all corners and in a waist zone between cathode bars and pot bottom to reduce heat losses; Insulating properly in bottom to meet two requirements: one is that the temperature in the narrow zone near carbon bottom is about 880C, and the other is to form a decreasing temperature gradient from pad, cathode carbon, refractory under carbon, to shell, which lets resistance heat of cathode carbon be delivered out effectively.

2.

In spite of all methods above used, there is still a possible shortage of heat input for NSC pots, so an amperage increase should be another more effective method to make up for it. Current efficiency of NSC pots It is well known that the loss mechanism of current efficiency (CE) is that aluminum produced dissolves into the bath by chemical or physical way and then is oxidized by anode gas. So the metal pad

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3.

To remain operating at proper temperature. Too-low temperature usually causes higher bath viscosity, lower dissolution rate and dissolvability of alumina, poor bath conductibility, and higher overvoltage. Under too-low superheat of the bath with low molar ratio, cells will be unsteady and alumina sludge will form easily on the cell bottom. For both NSC pots and traditional pots, the operating temperature should be controlled in the range of 940 to 960 °C for more efficient production, such as AP39 at about 960°C [5]. But the superheat should be controlled in the range of 8 to 10°C. However, the operating temperature may be decreased if the bath contains higher content LiF or/and KF under the condition of 10°C or less superheat. To reduce non-anode-projected area. Two opposite processes, aluminum producing and dissolving, occur on the pad surface which can be divided into two parts, anode-projected zone and non-anode-projected zone. In the non-anodeprojected zone, the aluminum dissolving process dominates and the aluminum production is less due to poor current density. So CE in the zone is low or even negative. The area of pad depends on the superheat and insulation in sidewalls. Either higher superheat with lower molar ratio or strong insulation by sidewalls leads to bigger area of non-anodeprojected zone, which causes higher dissolution of aluminum and lower CE. To avoid too much insulation on the pots bottom as mentioned above.

Table 6 Aluminum height (in cm) andridgesheight (in cm) in NSC Pot 125 First Measurement (December 2009) Aluminum height Aluminum Height of the Position of Cathode above the ridge level ridge A8 6 17 11 A14 7 18 11 Up-stream A15 8 19 11 A16 9 19 10 A17 7 17 10 B2 8 18 10 DownB8 7 18 11 stream B13 8 10 18 7.5 Average 18 10.50 Erosion and wear of ridges of the cathode in NSC pots NSC pots can decrease the energy consumption of aluminum producing with lower cell voltage due to lowering anode-cathode distance (ACD) with the fluctuation of pads weakened by these ridges, which are 11 cm of original height on the cathode carbon. So the life of the ridges is a key to energy saving. In order to minimize the ridges' erosion by bath, the height of liquid aluminum above the ridges should be kept in the range of 5 cm to 8 cm. It is important to learn any change in the height of ridges, which can be calculated by the height of liquid aluminum above the ridge and the aluminum level, which are measured during the anode change. One example on NSC Pot 125 is shown in table 6.

The reason for the terrible consumption rate of 168 kA NSC pots lies in the lower aluminum level. On one hand, it causes stronger horizontal current in the metal pad above the ridge. In addition, if the surface of the metal pad goes lower than the ridge, there will be occurrence of discharge of Na+ and Al3+ on the surface of the ridge in molten bath, and production of carbides, which dissolve easily into the bath. Therefore proper metal pad control to protect the ridge from immersing in the bath is critical to lower the erosion and consumption of the ridge. Contribution of the bath component on lowing cell voltage

The Pot 125 was started-up on June 26, 2009, and 0.5 cm of the ridges had been broken down or worn off during the first six months, and 0.25 cm had been worn during the next three months. In order to know the changes of the ridges in detail, the erosion of the ridges in all 94 NSC pots were measured and calculated on December 2009 and March 2010, respectively. The results are listed on table 7.

As known, the cell voltage is mainly related to ACD, temperature, and conductivity of anodes, cathodes and bath. The conductivity of bath can be improved with higher electrolysis temperature, higher molar ratio, or adding LiF and KF. For an industrial pot, the voltage is about 4.1 V, with operation temperature at 950 to 965°C, molar ratio at 2.3 to 2.6, CaF2 at 3 to 5%, A1203 at 2 to 3%, ACD at 4.5 to 5.0 cm. If LiF is added into the bath, the voltage will be decreased due to improved conductivity of the bath. Though KF can also improve the conductivity, it may damage heavily the cathode carbon due to stronger expansion of K than Na.

Table 7 Rate of erosion of ridges To Dec 2009 To Mar 2010 Total Rate Total Rate (mm) (mm/mo) (mm) (mm/mo)

Batch 1 (29 pots) (Start June 6-27) Batch 2 (34 pots) (StartJull4-Aug3) Batch 3 (31 pots) (StartAugl7-Sep6)

6.30

1.1

7.55

0.4

6.17

1.3

7.05

0.3

5.35

1.5

6.36

0.3

Second Measurement (March 2010) Aluminum height Aluminum Height of above the ridge level the ridge 7 18 11 8 18 10 7 18 11 8 17 9 8 18 10 9 19 10 6 17 11 8 18 10 7.625 17.875 10.25

However, some alumina produced from bauxite in Henan province, China, contains usually rich K 2 0 and Li20. As this alumina is added in the bath, LiF and KF will be generated by reactions as: 3Li20 + 2AlF2- >6LiF + Al203 3K20+2AIF3- >6KF+ALO,

From Table 7, erosion and wear of the ridges of all three batches during the first three months is more than 5 mm and it exceeds 7.5 mm for 9 months for batch 1. However, erosion and wear of ridges during the later three months, from Dec 2009 to Mar 2010, is only about 1.05 mm. So the wear of the ridge mainly occurs during the pot baking and start-up, and it is forecast that the erosion and wear ofridgeswill be less than 1 cm per year.

Most of the KF and LiF enriches into the bath except for that permeating into the cathode lining. It's possible that the total content of KF and LiF enriched would reach 10% to 14%, and bath conductivity would be over 20% higher than one without KF or LiF. The voltage drop of the bath with 10% or more KF and LiF would be less 0.3 V than one of the bath without KF or LiF, with the same ACD, from 4.5 to 5.0 cm. If this type of alumina were used in NSC pots with ACD reduced below 4.0 cm, with lining design to further improve heat balance as mentioned above, the voltage might be further decreased about 0.2 V .

Theridgeswere consumed at 2.5 cm to 3.0 cm per year according to the measurement of the three 168 kA NSC pots, the first NSC pots in the world, in Chongqing Tiantai Aluminum Ltd., since startup in March 2008. Now it is proved that the cell needs only 4 to 5 cm highridgesto achieve about 0.3 V reduction.

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Future Prospects NSC pot technology has been accepted by most smelters in China, because it can reduce cell voltage and energy consumption. However, the energy saving capacity in NSC pots depends not only on the structural design but also on technical operation and management. NSC pots shown in figures 2 and 3 are predicted to be steadier than the pot shown in figure 1 and have an obvious advantage in decreasing cell voltage and in increasing CE. These improved NSC pots are under testing just now, and test results will be given in the near future. Acknowledgement The authors wish to thank the financial support provided by the National Natural Science Foundation of China (No. 50934005) and a grant from the National High Technology Research and Development Program of China (No. 2009AA063701). References 1. 2. 3. 4. 5.

Feng Naixiang, Tian Yingfu, Peng Jianping, et al. New Cathodes in Aluminum Reduction Cells, Edited by John A. Johnson, Light Metals 2010, TMS: 405-408. Feng Naixiang. A novel cathode structural aluminum reduction cell, CN 200710010523.4 Feng Naixiang. A novel cathode structural aluminum reduction cell, CN 200810228017.7 Feng Shaofeng. Technology Study In 200ka Pots Using Novel Cathodes With Ridges. Journal of Materials and Metallurgy, 2010, Vol. 9, SI: 23-29. Martin O., et al., Development of the AP39: The New Flagship of AP Technology, Edited by John A. Johnson, Light Metals 2010, TMS: 333-338.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

STUDY OF ELECTROMAGNETIC FIELD IN 300KA ALUMINUM REDUCTION CELLS WITH INNOVATION CATHODE STRUCTURE Baokuan Li^Xiaobo Zhang1, Sui-rui Zhang1, Fang Wang1, Nai-xiang Feng1 1

2

School of Materials & Metallurgy, Northeastern University, Shenyang 110819, China School of Materials & Metallurgy, Wuhan University of Science and Technology, Hubei 430081, China

Keywords: aluminum reduction cells; innovation cathode structure; electromagnetic field; numerical simulation Abstract

Equations

Three dimensional steady finite element models have been developed to study the effect of the novel structure cathode (NSC) on the electromagnetic field in aluminum reduction cells. In addition to reduction cells, the finite element system includes the busbar, ferromagnetic shell and air. Numerical results show the current density, magnetic flux density, electric potential and electromagnetic force is different between NSC and traditional cells. The NSC with the convex block can divide the aluminum melt into a number of parts, where the electromagnetic force decreases. There are different small circulations formed around the convex blocks in aluminum melt, which not only reduce the fluctuation and the anode effect caused by lower alumina concentration, but also promote the dissolution of alumina. Compared to the traditional aluminum reduction cells, the NSC cells have the advantage of reasonable current density distribution and relatively uniform electromagnetic force distribution, so that the aluminum melt has smaller fluctuations and higher efficiency of current.

The Maxwell equation, Lorentz law and Joule law describing the magnetic field is:

Introduction

electric scalar potential equation for electrostatic analysis is derived from governing Equation (2) and Equation (4), and constitutive Equation (3):

Ampere law: Faraday law:

Vx// = J +

^

~

Vx£ =

dt

dB

(1) (2)

dt Gauss law: V D =p (3) Constitutive equation of magnetic flux is V · B = 0 (4) Since the frequency is less than 50Hz, the displacement current is ignored, as

dt

= 0-

As follows from Equation (2), the electric field is irrotational, and can be derived from:

È = -VV

Aluminum production consumes large amounts of energy, requiring 14000 ~ 14500kWh/T-AL for lt m . On the other hand, the current efficiency and thus energy efficiency in the aluminum reduction cell may be decreased due to the melt accelerating cyclic aluminum surface uplift, deflection and fluctuation12'31. The Novel Structural Cathode (NSC)[[?1 technology, shown in Fig. 1, has a pattern of raised ridges where the flow field of molten aluminum shall be divided[4]. The ridges decrease the flow velocity and weaken significantly the effect of electromagnetic force and gravity waves of molten aluminum. Coupled action of electric field, magnetic field, flow field and joule-heat field in aluminum reduction cells has a significant influence on the current efficiency, energy consumption and the cells life[5"6]. So, deeper understanding of the coupled relation of electric field, magnetic field-three field, choosing the appropriate mathematical model, and improving precision of computed results are very important in theoretic and practical guideline for improving of the optimization, design, engineering analysis and development of new cells171.

-V{ê{VV)) = p

(5)

(6)

Equation (6) is solved in an electrostatic field analysis of dielectrics with the conductor of anode, electrolyte, molten aluminum and the cathode using elements SOLID231. The equivalent resistance network model is used to solve the busbar system, according to KirchhofTs law. The current distribution is calculated by Equation (7) to solve the potential of each node and the current of the busbar segment by choosing SOURC36 unit. :

Ó^=° Ó7/=°

Where J = ó(Ε + í×Â) B=VxÂ, Lorentz law:

(7)

,

(8)

Λ = -—-VV dt F = JxB

(9) (10)

where H is Associated magnetic flux density; J is total current density vector; D is electric flux density vector; F is Lorentz force; B is magnetic flux density vector; E is electric field intensity vector; A is magnetic vector potential; V is magnetic vector potential; Q is Joule heating; p electric charge density; σ is electrical conductivity matrix; ε is permittivity matrix; μ is magnetic permeability; v is velocity vector; V is electric scalar potential; CO is joule-heat power density; t is time.

In this paper, the model of the structure of NSC systems is set up using PRO/E. Based on ANSYS platform, the finite element models of traditional and NSC cells are established to simulate the electromagnetic coupling field, using the equivalent resistance network method and edge element method.

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The mesh system is shown in Figure 2, where the cathode and metal pad are meshed by structural grid and electrolyte and air are meshed by nonstructural grid. The total number of finite elements is 245542. Table 1 Geometry and operating parameters of the aluminum reduction cell Chamber Anode carbon Cathode Cathode Bus

Geometry /m 14.85X4.20X1.34

Parameter Current /KA

300

1.64X0.66X0.55

Voltage /V

4

3.44X0.51X0.45

Number of bar

26

1.57X0.18X0.13

Number of bus

26

Table 2 Physical parameters for electromagnetic field calculation Parameters of electrical( Ω · m ) Resistance of molten 2.4 X10-7 aluminum Resistance of electrolyte 4.5 X10"3

SïdeB

Resistance of carbon anode

3.5 X10-5

Resistance of steel yokes

2.34 X10"7

Resistance of cathodes

3.8 X10 5

Resistance of the cathode bar

1. Aluminum conduction rod 2. Cathode 3. Cryolite 4. Aluminum melt 5. Anode rod 6. Anode bus bar 7. Anode 8. Cathode collector bar 9. Riser bus bar 10. Ramming paste 11. Steel shell 12. Cathode bus bar Fig. 1 Structure of 300kA cell with NSC technology

7.78 X10-7

Result and discussion The current density distribution

Method and gridding The commercial package ANSYS 10.0, which is based on the finite element method, is used to analyze the electromagnetic field and the current field system. Nodal-based method (Solidi 17 element type) is used to solve the three-dimensional static electromagnetic field. The current amplitude loaded on the upper surface of the anode rod is 300KA with coupled VOLT freedom, and the electric potential on the top surface of the cathode collector bar is zero. The relative permeability of the electric conductor is set to 1. The physical properties, geometrical and operating conditions are shown in Tables 1 and 2. Fig 1 shows a schematic of the cell that was modeled.

The total current density distributions of (a) traditional aluminum cell (b) NSC cell are shown in Figure 3.The current density distribution of the traditional cell is more uniform than that of the NSC cell, due to the flatness of the traditional cell compared to the projections from the surface of the NSC cell.

m

B sm m

p% mix U jut*

■ me?

J mn inn m im

$

If0?If f i ! f f I I f f f i l !?ff f i l

m

I

ì0243

S é«â«

m

S mi

3 »*

y ma

Fig. 2 Thefiniteelement mesh model of electromagnetic field with a quarter of the air

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m * ·of Fig. 3 The total current distribution on the bottom surface molten aluminum (a) Traditional cell (b) NSC cell

Figure 4 is a cross sectional view o f the current densities (x direction on figure 1). Both distributions are symmetrical about the centerline o f the cell, but the traditional cell has a more uniform distribution with lower maximum values. 1.5

-

— C o n v e n t i o n a l model - • I n n o v a t i o n mo d e l ^ - x

3

1.2

-

0.9

-

(

/ /\\ s i

i 1

0.6 o.

x

1

v

'/

\s

\ '

\

<J -

1

I

.

I

.

I

Fig. 4 The current density distribution on the cross section o f molten aluminum in the traditional and NSC cell The potential distribution The potential distributions of molten aluminum for the two types of cathode are shown in Figures 5 and 6. The potential distributions are symmetrical along both the long and short axes. The maximums are at the short ends; the minimums are at the long sides. The voltage drop of the traditional cell is 0.009V; compared to 0.006V for the new cell from table 3. .2936

: .,';νΐ5.ί:;

'tMM&ίy-M^

Fig.6 The anode system unit potential distribution (a) Traditional cell (b) NSC cell The voltage drops in the two cells are shown in Table 3. The total voltage drop o f the N S C cell is 0.066V lower than the traditional one, so the new cathode is more energy-efficient. A little more than half o f the gain comes from the 0.035V difference in the electrolyte.

3905 .4874 5843

6812 .7782 .8751

Table 3 voltage drop of traditional and NSC cell tradition innovation 1 parameter aluminum cell aluminum cell voltage drop of 0.248 0.219 cathode voltage drop of 0.638 0.638 anode voltage drop of 2.959 2.924 electrolyte voltage drop of molten 0.009 0.006 aluminum voltage drop of 0.336 0.337 cathode bar

.972 0689

1658 2483 3205 3926 4648 537 6092 6814 7536 8258

Fig.5 The molten aluminum potential distribution (a) tradition aluminum cell (b) innovation aluminum cell

The magnetic flux density distribution The magnetic flux density B x , B y and B z distribution o f molten aluminum in the two types o f cell is shown in Figures 7-10. Both cathode structures show the same type o f symmetry: B x and B z have inverse symmetry along the short axis while B y has symmetry along the long axis as well as the inverse symmetry along the short axis. The fluctuation of molten aluminum in the traditional cell is more intense. Magnetic induction B x , B y have little difference in value, while the magnetic induction B z has slightly larger difference in value. That means the molten

The potential distribution of aluminum rod, steel yoke and anode block is the same in the new and traditional cathodes, as their structure, materials, and electrical load are the same. In both cells the anode assembly voltage drop is 0.638V from the top o f the rod to the bottom of the anode block.

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aluminum is predicted to have better flow stability in the new cathode.

Convent i onal rrodel

(a) Side A

- - - H nnovat i on frode!

«s

0.60 Conventional

(b) SideB

0.55

model

1 nnovat i on model

0.50

ì\

// \\

0.45

Fig. 7 The magnetic density distribution of molten aluminum in the traditional cell

0.40

'

0.35

!

030

L

1

v

1

1

s

< \

/

- ' ^ ~ > "* - ,

*" /

v

1

1

1

1 _

u )

Fig. 9 The magnetic density distribution of molten aluminum

(a)

·

·,,.

.vW

4

> v

#

Fig. 8The magnetic density distribution of molten aluminum in the NSC cell Table 4 Magnetic density on different plane of molten aluminum in the NSC cell parameter Top of molten Bottom of molten I aluminum | aluminum 31.23 23.79 1 Βχ 1 max 4.13 3.87 1 Dx 1 ave 70.82 62.65 1 Dy 1 max 24.11 22.18 1 Dy 1 ave 37.77 30.24 1 Dz 1 max 8.86 7.83 1 Dz 1 ave 70.82 62.65 iBlmax 28.19 25.68 1 B1 ave

Fig. 10 The magnetic density vector distribution of molten aluminum on different planes in the NSC cell (a) top surface, (b) bottom surface

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The electromagnetic force distribution Acknowledgment

The electromagnetic force distribution of molten aluminum along the horizontal direction is shown in Fig. 11. From Fig. 11(a), the electromagnetic force along the short axis x direction is dominant and the electromagnetic force along the long axis y direction is smaller. The reason is that level current of molten aluminum in the x direction and vertical current in the z direction are dominant. From Fig. 11(b) for the NSC cell, we can see the electromagnetic forces in the x-axis and y axis have been significantly reduced because there is an opposing level electromagnetic force. The net effect reduces the fluctuation of molten aluminum in the NSC cell.

The project is supported by the National Natural Science Foundation of China ( No.50934005 ) and National High Technology R & D program "863" project(2009AA063701). Reference [1] J. I. Buiza, "Electromagnetic Optimization of the V-350 Cell," Light Metals, (1989), 211-214/ [2] M. Dupuis and V. Bojarevics, "Weakly Coupled Thermoelectric and MHD Mathematical Models of an Aluminum Electrolysis Cell, "Light Metals, (2005), 449-454.

-+ 0.27N

[3] V. Bojarevics and K. Pericleous, "Comparison of MHD Models for Aluminum Reduction Cells," .Light Metals, (2006), 347-352. [4] G. V. Arkhipov and A. V. Rozin, "The Aluminum Reduction Cell Closed System of 3D Mathematical Models," Light Metals, (2005), 589-592.

•

6

-

4

-

2

0

2

4

[5] M. V. Romerio and M. A. Secretan, "Magne-tohydynamic equilibrium in aluminum electrolytic cells." Computer Physics Reports, (1986), 3: 327-360.June II.

6

y/h»

[6] M. Dupuis, V. Bojarevics and D. Richard, "Impact of the Vertical Potshell Deformation on the MHD Cell Stability Behavior of a 500 kA Aluminum Electrolysis Cell," Light Metals, (2008), 409-412. [7] Feng Naixiang, "New Cathodes in Aluminum Reduction Cells," Light Metals 2010, (2010), 405-408.

Fig. 11 The Fxy distribution at the bottom plane of molten aluminum in the traditional(a) and NSC(b) cell Conclusions (l)The voltage drop of the NSC is 0.066V lower than the traditional one, which means it is more energy efficient. (2) The current density distribution in the NSC cathode is symmetric along the center line, there is a large level current at the short axis, (3) Electromagnetic force in the NSC is smaller, since there is a force in the opposite direction which can reduce the electromagnetic force.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Evaluation of thermophysical properties of silicon carbide, graphitic and graphitized carbon sidewall lining materials used in aluminium reduction cell in function of temperature Ayesha Khatun, Martin Dιsuets LAPSUS (Laboratoire des Procιdιs industriels et Simulations numιriques de l'Universitι de Sherbrooke) Universitι de Sherbrooke 2500 Boul. de l'Universitι, Sherbrooke, Quιbec, J1K 2R1, Canada Keywords: Thermal conductivity, Thermal diffusivity, Heat capacity, Temperature varying properties, Sidewall lining material. Abstract The thermal properties of the sidewall lining materials are required to ensure good predictions of the dynamic thermal behavior of Hall-Heroult cells. A precise estimation of energy losses and location of the side freeze are made possible when these materials are well characterized as a function of temperature. The present work uses transient characterization techniques to measure the thermal diffusivity, heat capacity and thermal conductivity of silicon carbide, graphitic and graphitized carbon materials. The thermal diffusivity and the heat capacity are measured using a state-of-the-art transient laser flash analyzer and a differential scanning calorimeter respectively. The thermal conductivity is calculated by assuming a constant density. Finally, based on the calculations conducted with a 2-D numerical model, the effect of the temperature varying thermal properties of the sidewall materials on the dynamic behavior of a laboratory scale phase change reactor is presented.

determine the thermal properties of materials like insulation, refractories, ceramics, etc., used in industrial aluminium cells. They described the limitations of the steady state method and focused on the use of transient methods to measure the thermal properties. In steady state methods, a massive sample cylinder is typically required which introduces significant temperature variations during the measurements. Also, the heat losses through the thermocouples are source of errors during the measurement. Finally, the development of a steady state experimental setup and the measurement operation are lengthy and time consuming [6]. To overcome these limitations, the present work aims at using transient methods for measuring the thermal properties of three different sidewall materials. In combination with computerized data processing, such techniques can produce accurate and reliable data from room temperature to high temperatures for all types of materials [6]. The methods used lead to the measurement of heat capacity and thermal diffusivity in order to calculate the thermal conductivity. In the current work, the density of each sidewall material has been taken from the literature and assumed to be temperature independent. Two cutting edge measurement techniques, Laser Flash Analyzer (LFA) and Differential Scanning Calorimeter (DSC), have been used to measure respectively the thermal diffusivity and heat capacity.

Introduction Generally, the carbon materials used in cathodes are also used for the sidelining of aluminium pots [1]. Sidewalls made of graphitic carbon are losing interest because of their lower resistance to air oxidation and chemical corrosion. The molten metal and corrosive electrolytic bath deteriorate graphitic carbon easily and rapidly which leads to higher risk of sidewall failure and shorter cell life [2]. Graphite has better thermal properties than graphitic material. However, at cell operating temperature, graphite can also be oxidized to CO2. Nowadays, ceramic sidewall lining material like silicon carbide is typically used for the construction of the sidewall of the aluminium cell due to its high chemical resistivity and its excellent thermal conductivity. It can withstand the extremely corrosive molten electrolyte for long periods [3].

An empirical correlation for each property has also been derived by fitting the measured data with polynomial equations. These equations can be directly used by modelers to calculate the sidewall properties during the estimation of the heat losses and side ledge formation. This paper also discusses the effect of the sidewall materials on the prediction of the static and dynamic behaviour of the ledge in a laboratory scale phase change reactor. This physical model of an aluminium electrolysis cell represents the cooling of liquid electrolytic bath inside a crucible made out of industrial sidewall materials.

The thermal properties of sidewall lining material have a significant impact on the thermal equilibrium inside the pot. For instance, a high thermal conductivity helps to maintain the desired amount of ledge protection on the sidewall. In consequence, the thermophysical properties of such materials have a significant effect on the heat losses through the sidewall.

Measurement of thermal properties Thermal diffusivity

The thermal conductivity of different carbon materials used as cathodes has been measured by Dumas et al. [4], using a direct steady state method named KOLHRAUSCH. This same method has been adapted by Allard et al. [5] using a radial heat flow mode to measure the thermal conductivity of graphitized carbon. Llavona et al. [6] described the different measurement methods to

The principle of the laser flash method (Figure 1) is based on the heating of a specimen by a short laser pulse on its front side and the detection of the temperature increase at its rear side. The thermal diffusivity is determined based on the relative temperature change as a function of time only.

1035

# # Specimen

P-

crucible surface and to use a sample of cylindrical shape. It is also essential to have plane and smooth surfaces.

Signal -^ Laser

Figure 1 : Laser applied at the front side of the specimen. The thermal diffusivity a is calculated according to equation (1), where d is the thickness of the sample and ty2 is the time taken to reach half of the temperature increase due to the laser pulse [7] : a = 0.1388— (1)

A standard diameter of 5+0.2 mm is required to avoid the direct contact of the sample with the inner sidewall of the crucible while measuring the heat capacity. The standard thickness of the sample is kept to approximately 1 mm. This value however depends on the density and mechanical strength of the sample material at such low thicknesses. Thermal conductivity One of the main key parameters in heat conduction is the thermal conductivity. In this paper, the thermal conductivity is measured indirectly. If the thermal diffusivity (φr), density (p) and heat capacity (Cp) are known, the thermal conductivity (Λ), can be estimated from: ë= apCp (2)

2

This equation is based on the following assumptions: the laser energy is applied instantaneously i.e. the heat pulse duration is negligible, there is no heat loss from the side of the sample, heat conduction in the sample is one dimensional and unsteady.

The calculation of the relative error of the thermal conductivity is essentially the sum of the relative errors on a, p and Cp:

The sample specimen was prepared by machining the material into a long cylinder by using a lathe machine. The desired sample thickness is then obtained by using a diamond saw. A special care has been given on the sample dimensions, sample surface finishing and surface blackness. These parameters play a vital role for obtaining good thermal diffusivities. The sample surface should be smooth enough and plane parallel. Its color should be as black as possible. Smooth and black surfaces give good signal responses and good absorption of the laser energy.

ë

Cp

p

In theory, the density of a material can have a great impact on the thermal conductivity. In practice however, the density of a solid material does not vary much as a function of temperature. For example, sintered bonded silicon carbide has a density of 3160 kg/nr at room temperature and 3110 kg/m3 at 1000 °C [8]. This represents a 1.5 % variation over a fairly large temperature range. Obviously, neglecting such a variation is introducing an error of 1.5 % in our estimation of the thermal conductivity value at high temperature. The analyses conducted in this paper being comparative, the impact of this density variation on the measurement error is the same for all materials. It has thus been neglected.

The sample diameter should be within 12.66 to 12.68 mm to fit the sample holder of the Netzsch LFA model 457 used for the measurement of the thermal diffusivities. This range of diameters prevents the direct penetration of laser at the rear surface of the sample. The sample thickness has been chosen according to the rules of thumb suggested by the manufacturer of the equipment [8].

Results and discussion Evaluation of the heat capacity in function of temperature

Thermal diffusivity values have been measured in a 25 to 1000 °C temperature range for three different thicknesses at the same time. The optimum thickness has been chosen based on the following criteria: • Good signal versus time curve provided by the detector. •

a

The heat capacity for each material was measured from room temperature to approximately 1000°C. For each material, an increase in heat capacity as a function of temperature is observed. Heat capacity of silicon carbide is gradually increasing by following two linear curves (Figure 2). The heat capacity of silicon carbide is lower than that of graphitic and graphitized carbon. This difference becomes stronger as temperature increases. At high temperature, the heat capacity of silicon carbide is approximately 1.2 times less than the heat capacity of graphitic and graphitized carbon.

Good matching of the thermal conductivity values with the literature values already reported for similar materials, as the thermal diffusivity values of the specific materials studied here are not available in the open literature.

Graphitic and graphitized carbons have almost similar heat capacity temperature functions except at high temperature where the two functions behave differently. At high temperature, the heat capacity of graphitic carbon is lower and more stable than the heat capacity of graphitized carbon.

Heat capacity The heat capacity of each sample has been measured by using a transient relative method called Differential Scanning Calorimetry (DSC). This cutting edge technology can measure two quantities at the same time, the heat flow rate and the corresponding AT between a sample and a reference material. This measurement is usually accomplished in three steps - baseline measurement, sample measurement and reference measurement. The most important factors to get precise measurements with this method are to maintain a good contact between the sample and the

Polynomial expressions have been fitted on the experimental data. Table 1 presents the regression analysis as applied to the heat capacity (Cp) values, previously adimensionalized by dividing each data by the maximum Cp value of the data set.

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Table 1 : Empirical correlations for the heat capacity of sidewall lining materials. Name of materials

<~*~<SilkoB carbide

Regression 1 coefficient

Empirical correlations Cp = - 3 x 10" 7 7 2 + 0.00087 + 0.2329 Cp = -10~ 6 7 2 + 0.00267 - 0.2413 Cp = - 1 0 - 6 r 2 + 0.00257 -0.2272

Silicon carbide 1 Graphitic carbon Graphitized carbon

1 i

rlv = 0.987 rlv = 0.996 r?p = 0.999

300

400

500

60Φ

700

800

900

1000

1100

Temperature,! fC)

Figure 3: Thermal diffusivity of the sidewall lining materials. Table 2: Empirical correlations for the thermal diffusivity of sidewall lining materials.

300

400

500

600

?00

800

900

1000

1100

Tempera ture ,T(C)

Figure 2: Heat capacity of the sidewall lining materials. Evaluation of the thermal diffusivity in function of temperature For each sidewall material, an important decrease in the heat diffusion rate or thermal diffusivity is observed between room and high temperatures (Figure 3). The change of the heat diffusion rate in graphitized carbon is faster than the change observed for graphitic carbon and silicon carbide. At low temperature, the thermal diffusivity of graphitized carbon is 4.5 times higher than that of silicon carbide (SiC) and 3 times higher than that of graphitic carbon. This difference gradually decreases as temperature increases. At high temperatures, the ratio of thermal diffusivities becomes 2.45 for (XgrapWtized/ίgraphitic and 2 . 1 7 for OCgraphitized/^SiO

The thermal diffusivity of SiC is higher than the thermal diffusivity of graphitic carbon. However, at high temperature, the thermal diffusivity values of these two materials are very close to each other. As for the heat capacity, polynomial expressions have been developed (see Table 2) to represent the temperature variation of oAXmax» otmax being the maximum value of the thermal diffusivity, found here with the graphitized carbon at high temperature.

Name of materials

Empirical correlations

Silicon carbide

a = 3 x 10" 7 7 2 - 0.00077 + 0.4629

ra2 = 0.992

Graphitic carbon

a = 3 x 10" 7 7 2 - 0.00057 + 0.3676

ra2 = 0.992

Graphitize d carbon

a = 2 X 10" 6 7 2 - 0.00377 + 1.996

ra2 = 0.980

Regression 1 coefficient

Evaluation of the thermal conductivity in function of temperature While the thermal conductivity of graphitized carbon and SiC gradually decrease with temperature, the thermal conductivity of graphitic carbon is nearly constant (Figure 4). Initially, the thermal conductivity of graphitized carbon is 5 times higher than that of graphitic carbon and 2.4 times higher than that of SiC. At high temperature, the thermal conductivity of each material is becoming very close to each other, a similar conclusion drawn by Dumas et al. [4] for the thermal conductivity of graphitic and graphitized carbon. Finally, Table 3 presents the empirical correlations obtained by conducting a polynomial regression on the adimensionalized values of the thermal conductivities.

300

400

500

600

700

800

900

1000

1100

Temperature ,T(*C)

Figure 4: Thermal conductivity of sidewall lining materials. Table 3: Empirical correlations for the thermal conductivity of sidewall lining materials. Name of materials

Empirical correlations

Silicon carbide Graphitic carbon

λ = 3 x 10~7TZ - 0.0006Γ + 0.6132 λ = - 1 x 1(Γ 7 Γ 2 + 1 x 1(Γ4Γ + 0.1971 λ = 7 x 10" 7 Γ 2 - 0.0018Γ + 1.5413

Graphitized carbon

Figure 5: Schematic of the phase change reactor.

Regression I coefficient

To simulate the operation of the oven, a constant heat flux of 20 kW/m2 is applied at the top of the crucible and of the liquid electrolytic bath, used here as our Phase Change Material (PCM). The bath is numerically solidified in the crucible by extracting the heat through a forced convection air stream applied at the exposed bottom surface of the crucible. The constant forced convection heat transfer coefficient at the bottom of the crucible was calculated as 53 W/m2/K, based on empirical correlations [10] obtained for similar configuration. The heat loss through the inner exposed surface of the firebrick and castable concrete is also due to this forced convection. The calculated free convection heat transfer coefficient was estimated at 2.6 W/m2/K for the south west side of the reactor and at 5.86 W/m2/K for the west side exposed surface of the insulation package. These boundary conditions were kept constant in the evaluation of each crucible material. Starting from an initial condition where the electrolytic bath is liquid, the solidification gradually occurs due to the heat losses from the reactor to the ambient environment through these convective boundaries.

ri = 0.999 r/ = 0.975 ri = 0.999

Effect of the sidewall materials' thermal conductivity on the formation of the ledge A 2D transient mathematical model in axisymmetric coordinate, similar to the one derived by Marois et al. [9], has been developed to study the effect of the crucible material on the formation of ledge and on the distribution of heat losses inside an experimental phase change reactor. The phase change phenomena has been characterized by simulating the 2D Stefan phase change problem taking into account the latent heat evolution by using an enthalpic formulation. The impact of the thermal properties on the prediction of the ledge profile and on the thermal balance is analyzed based on the numerical results obtained by the mathematical model.

In the numerical simulations, the temperature varying thermal properties (heat capacity and thermal conductivity) of the each sidewall lining material has been taken into account by using the empirical correlations described in tables 1 to 3. The thermal properties of the other materials like firebrick or castable concrete, were determined based on manufacturers' technical data.

The schematic of the cylindrical shape reactor has been described in figure 5. The sidewall of the reactor is made out of an insulation package. The bottom of the experimental setup is built up from firebricks and castable concrete. The simulation was carried out to predict the dynamic behaviour of the ledge when SiC, graphitic and graphitized carbon are respectively used as the crucible material. The dimensions of the crucible were kept constant for all simulations.

The effect of the sidewall materials' thermal properties on the thickness of the ledge is very interesting. As expected, the ledge thickness directly depends on the thermal conductivity of the sidewall materials. The global effect can be easily explained by representing the one-dimensional heat flow between liquid electrolytic bath and ambient environment (Figure 6), using the electrical network analogy. The thermal resistance due to the convection heat transfer outside of the reactor and the global temperature difference (ΔΓ) between the reactor and the ambient air are considered constant in our analysis. The higher is the thermal conductivity of the sidewall, the lower is the thermal resistance in the crucible region. The constant heat flux entering at

1038

the top of the crucible and liquid electrolytic bath can be expressed as:

o =—

1

* ' » Ledge profile for graphitic carbon -Ledge profile for silicon carbide -·«-·♦ Ledge profile for graphitized carbon

0.9

(4)

0.8 I-

1"0.7

Where

Rfnt — Rh+Rr + Rn

i

(5)

5 0.6 g

Rb= thermal resistance of the bath zone, essentially due to solidified bath,

*

Rc= thermal resistance of the crucible,

0.5

g 0.4

Ramb- thermal resistance due to convection to the environment.

^

0.3 0.2 0.1

0.86

1

0.7 0.6 0.42 0.28 Radius of the crucible (r'LpcMjt)

0.1

0

Figure 7: Effect of side wall materials' thermal conductivity on the prediction of the steady state profile of ledge. Figure 6: Thermal resistance network representing one dimensional heat flow between the liquid electrolytic bath and the ambient environment.

Table 4: Power losses in the reactor for different crucible materials.

The only way to keep Qn constant when ΔΓ is constant, is to keep Rtot constant. When a crucible material of higher thermal conductivity is used, Rc is expected to be lower. Thus Rb will have to compensate the reduction of Rc by forming more ledge on the inside bottom surface of the crucible. When the thermal conductivity of crucible material is low, the reverse phenomenon will be observed. The ledge formed in the graphitized carbon crucible is the thickest one as this material is introducing the lowest resistance in the thermal resistance network. For the same reason, the ledge is thicker in the SiC crucible than the one formed in graphitic carbon. At steady state conditions, the thickness of the ledge formed at the center of the graphitic carbon and silicon carbide crucible are almost identical, whereas the ledge formed at the center of the graphitized carbon crucible is about 0.5 cm thicker (Figure 7). The difference between the behaviour of these materials is best seen at the peripheral regions of the crucible. The system goes to steady state conditions more slowly when graphitized carbon is used; the additional ledge formed causing an inertial effect on the thermal system. At steady state conditions, the power losses through the exposed bottom surface is boosted up by 4.2 % for the graphitized crucible and by 1.3 % for the SiC crucible when compared to graphitic carbon (Table 4).

Crucible

South side exposed bottom surface of the crucible, (W)

Graphitic carbon Silicon carbide Graphitized carbon

1325.6 1342.4 1381.3

v%$v ax \ \\\

ÌI a

\

^

V**H

Percentage 1 of variation with respect to graphitic carbon 1.3 % 4.2 %

1

— * — Ledge profile at 2 h —»· — Ledge profile at 2.5 h — —- Ledge profile at 3.3 h — 0 — Ledge profile at 5 h —·· — Ledge profile at 10 h : * Ledge profile at 20 h ""P" -~ Ledge profile at 30 h — i»--- Ledge profile at 40 h — ♦"' — Ledge profile at 50 h Ledge profile at 60 h

. f t * » » « » » » » » » * » * * » * « » »

m m mm m m 0.7 mm m ~m 0.6 WΙ

0.42 0.28 Radius of the crucible (r/L μ

The dynamic behavior of the reactor has also been evaluated for each of the crucibles. The simulations have been conducted up to the steady state condition of the system. As explained above, the movement of the ledge in the upward direction is slower in the graphitized crucible (Figure 8). The moving rates of the solidification front in silicon carbide and graphitic carbon crucible are very close to each other (Figure 9 and 10). However, at steady state conditions, more ledge is formed in the SiC crucible than in the graphitic carbon crucible.

Figure 8: Dynamic behaviour of the ledge formed after starting the cooling process of liquid electrolytic bath in the graphitized crucible.

1039

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formed at the steady state conditions is the thickest one, followed by the silicon carbide and by the graphitic carbon. The thermal conductivity of the material used for the sidewall construction significantly affects the side ledge thickness over it, especially during the transient conditions. These variations in ledge thickness also lead to variations of the sidewall heat losses.

» Ledge profile at 2 h ~~#> -Ledge profile at 2.5 h Φ Ledge profile at 3.4 h -~~*— Ledge profile at 5 h ~~ *~ r Ledge profile at 10 h —»· -Ledge profile at 20 h - $ - -Ledge profile at 30 h ~~Φ~ -Ledge profile at 40 h "" """Ledgeprofile at 50h |

The authors are grateful to Rio Tinto Alcan for their financial support.

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References

V " » " ' » „φ,,,ιφ,,η^,,,^,,,^,,.,φ,,,,^,,,,φ,,,,^,,,,φ,,,,ρ,,,φ.

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,

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»

,

»

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0.6

.

/

.1

1

0.42

[1] S0rlie, Morten, and 0ye, Harald A (1994). Cathodes in aluminium electrolysis, 2nd edition, Aluminium -Verlag, 1994, p. 12-23.

1

0.28

0.1

Radius of the crucible (r/L FC]Ä )

[2] Boily, P., Kiss, L.I., Bui, R.T. and Desclaux, P. (2001). Sensitivity analysis of the thermal detection of the Ledge profile in an aluminium reduction cell. Light Metals 2001, p. 1209.

Figure 9: Dynamic behaviour of the ledge formed after starting the cooling process of liquid electrolytic bath in the silicon carbide crucible.

[3] Pan, Yuhua., Wright, Steven, and Sun, Shouyi (2009). Review and applications of thermal conductivity models to aluminium cell sidewall refractories. International Journal of Modern Physics B. Vol. 23, Nos. 6 & 7 (2009), p.790.

— * · — Ledge profile at 2 h —#- — Ledge profile at 2.5 h — * — Ledge profile at 3.3 h Ledge profile at 5 h — <*■" — Ledge profile at 10 h — e- — Ledge profile at 20 h — · " — Ledge profile at 30 h — «— Ledge profile at 40 h — ·- — Ledge profile at 50 h

[4] Dumas, D. and Lacroix, P. (1994). High temperature measurement of electrical resistivity and thermal conductivity on carbon materials used in aluminium smelters, Light Metals, 1994, p. 751. [5] Allard, Benedict. Dreyfus, Jean-Michel, and Lenclud, Michel (2000). Evolution of thermal, electrical and mechanical properties of graphitized cathode blocks for aluminium electrolysis cells with temperatures, Light Metals, 2000, p. 515.

~ * ~ · ~ % *~ -*v

O- O- ·<> «-<* H

K

^

» » » fe \ .

|Í»»#«Ö*#

»

♦

»

»

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»■■■»■■■» »■■■»■■»

[6] Llavona, M.A., MaAyala, J., Garcia, MaP., Zapico, R. and Sancho, J.P. Thermal conductivity measurement of the materials in the aluminium industry, Light Metals 1994, p.467.

< < < ^ ^ 4 ^ ^ ^ . < 0 m <

0.7 0.6 0.42 0.28 Radius of the crucible (r/L_CMk)

[7] Laser flash analyzer instrument Manual provided by Netzsch. [8] http://www.ceramics.nist.gov/srd/summarv/scdscs.htm, NIST, Property Data Summaries, page consulted at 13 September 2010.

Figure 10: Dynamic behaviour of the ledge formed after starting the cooling process of liquid electrolytic bath in the graphitic crucible.

[9] Marois, M., Bertrand, C, Desilets, M., Coulombe, M and Lacroix, M (2009). Comparison of two different numerical methods for predicting the formation of the side ledge in aluminium electrolysis cell, Light Metals, 2009, p. 563.

Conclusions The presented research work aims at the precise measurement of the thermal properties of graphitic and graphitized carbon and silicon carbide sidewall lining materials used in the aluminium reduction cell. The heat capacity and the thermal diffusivity of each material have been measured and used to estimate the thermal conductivity at constant density. The effect of the thermal conductivity of the each sidewall material on the prediction of the bottom ledge inside a laboratory scale phase change reactor has also been estimated by using the empirical correlation derived by using polynomial fitting of the estimated thermal conductivity.

[10] Incropera, Frank P., Dewitt, David P., Bergman, Theodore and Lavine, Adrienne S. Introduction to Heat Transfer, 6th edition, Chapter 7, John Wiley & Sons, Ine, P. 423-428,431.

As expected, graphitized carbon has the highest thermal conductivity in comparison to the other sidewall lining material evaluated. When graphitized carbon sidewall is used, the ledge

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ADVANCED NUMERICAL SIMULATION OF THE THERMO-ELECTRO-CHEMO-MECHANICAL BEHAVIOUR OF HALL-HÉROULT CELLS UNDER ELECTRICAL PREHEATING Daniel Marceau1, Simon Pilote1'2, Martin Dιsuets3, Jean-Franηois Bilodeau2, Lyθs Hacini2, Yves Caratini4 university Research Centre on Aluminium (CURAL) - Aluminium Research Centre (REGAL) - University of Quιbec at Chicoutimi; 555, Boul. de ΓUniversitι, Chicoutimi (Quιbec), Canada, G7H 2B1 2 Rio Tinto Alcan (Arvida Research and Development Center), P.O. Box 1250, Jonquiθre (Quιbec), Canada, G7S 4K8 3 Aluminium Research Centre (REGAL) - Sherbrooke University; 2500, Boul. de l'Universitι; Sherbrooke (Quιbec), Canada, J1K 2R1 4 Rio Tinto Alcan (Laboratoire de recherche des fabrications); BP 114, 73303 Saint-Jean-de-Maurienne Cedex - France Keywords: Hall-Hιroult Cell, Electrical Preheating, Energy Consumption, Thermo-Electro-Mechanical Simulation, Finite Element Method In particular, recent technological advances have still demonstrated that the lifespan of a reduction cell is strongly conditioned by the start-up stage. However, it remains very difficult to evaluate the quality of a good start-up because of many factors influencing this stage mainly dedicated to ensure smooth transition to the production stage. Figure 1 shows one published opinion that the start-up stage represents 25% of the total influence on its lifespan.

Abstract In today's context, aluminum producers strive to improve their position regarding energy consumption and production costs. To do so, mathematical modeling offers a good way to study the behavior of the cell during its life. This paper deals with the numerical simulation of the electrical preheating of a Hall-Hιroult cell using a quarter model of the cell. The fully coupled thermoelectro-mechanical model includes material non linearities and multiphysical behavior at interfaces allowing accurate evaluation of the stress distribution in the cathode blocks and surrounding components. The baking of the ramming paste as well as the evolution of its thermo-electro-mechanical properties are updated via the baking index computed using a kinetic of reaction. The model is initially calibrated with in situ measurements and then used to estimate the effect of preheating on the behavior of the cell including temperature, current, deformations as well as the contact conditions at critical interfaces. Introduction The Hall-Hιroult process is still and will remain for a long time, the most popular process used to produce aluminum (Al). This high-temperature electrolytic process starts with the dissolution of alumina in a molten cryolite. The electrical voltage used in a typical reduction cell is in the range of 4 to 5 V, but requires high amperage in the range of 150 to 500 kA depending on the technology.

Figure 1 - Relative importance of impact factors on the lifespan of a typical reduction cell [1]. According to [2], a good start-up should include a preheating phase to smoothly increase the temperature in the cell in order to ensure sealing of the cathode plane and avoid thermal shock in the cathode blocks during the bath/molten metal additions. The sealing of the cathode plane is ensured by the ramming paste located between the carbon blocks and between the carbon blocks and the side wall. The quality of the sealing is not only associated to the recipes of ramming paste but also, to the global stiffness of the reduction cell as well as the intensity of the compressive stress distribution developed in the cathode blocks during the preheating stage.

Considering today's available data, the best technology uses approximately 12.5 to 13.0 MW/t of Al which is quite far from the theoretical value of 6.0 MW/t. Applied to the Quιbec smelters, this actual consumption corresponds to 14% of the installed capacity of Hydro-Quιbec. A value of 11 MW/t of Al is targeted by the industry before 2020. In the same way, the typical reduction cell has an average lifespan of about 2,500 days which corresponds to approximately 25 kg of spent pot lining per ton of Al. Canadian smelters produce approximately 65 kt of spent pot lining per year. Moreover, considering that the replacement cost of a typical cell is roughly ranging from 100 to 350 k$ CND, an increase of its lifespan is therefore an efficient way to reduce the production cost as well as the middle- to long-term impacts for the environment.

The preheating rate should also be slow enough to avoid a non uniform thermal distribution allowing a large thermal gradient in the cathode blocks. Excessive thermal gradient may result in a loss of stiffness due to crack initiation in the carbon blocks and therefore, in a loss of sealing. On the other hand, a too slow preheating rate may also result in inadequate sealing conditions due to insufficient baking of the ramming paste to produce sufficient compressive stress in the cathode blocks. At the end of the preheating, the temperature in the cathode blocks should be high enough to avoid bath freezing as well as the flash pyrolysis of the ramming paste during bath addition. At this time, the

Considering the large number of operating reduction cells in the world, it is natural to investigate the ways to increase the amperage of these cells without any significant negative impacts on its operating cost, lifespan and for the environment. To do this, a solid understanding of the behavior of the cell during its start-up stage as well as during its production stage is a crucial issue.

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cathode blocks and the ramming paste should provide a sufficient mechanical resistance to allow the bath/molten metal additions without any damage. All these concerns show that the investigation of reduction cell under preheating is important but not an easy task. In this paper, the electrical preheating of a P155 reduction cell is investigated using a thermo-chemo-electro-mechanical quarter cell model. The model takes into account the realistic behavior of the most significant components in the cell as well as the interfacial behavior between them. The fully coupled problem is solved using the in-house FE code FESh++ [3] which is one of the most efficient tools to solve this transient highly non linear multiphysical problem. Previous work During the last 35 years, several numerical models have been proposed to simulate the behavior of the electrolytic cell or parts of it such as anode assembly, shell, cradles, coke bed, etc. During the first 25 years, most of these models were dedicated to the thermo-electrical behavior of the cell. Despite the large interest of these models, none of them consider the impact of the thermal field distribution on the electrical, mechanical and chemical response of the cell as well as the impact of the mechanical stress distribution on the thermo-electrical equilibrium. Couplet thermoelectro-mechanical modeling of electrolytic cells was first introduced in 2000 by Hiltmann & Meulemann [4], Richard et al. [5], and Dιsilets et al. [3]. More recently, Richard et al. [6] presented a thermo-chemo-mechanical slice model used to predict the impact of the preheating scenario on the stress distribution in the cathode block and the baking level of the ramming paste. In this last paper, the cathode block, the side block, the castable and the ramming paste were modeled using more realistic constitutive laws that contribute to a more accurate evaluation of the stress distribution in these components during the simulation. In 2010, Dupuis [7] presented three approaches for the mathematical modeling of the deformations inside the potshell of the reduction cell including sodium swelling in the cathode block. However, as underlined by the author, these models neglected the impact of the chemical swelling/shrinkage deformation mechanisms of the ramming paste in the early age of the cell on its stiffness changes as well as that caused by a more realistic mechanical response of the carbon components as used by Richard et al. [6].

additions. With the exception of the classical thermal expansion and elastic strain which take place within the cell during the preheating, other mechanisms such as plastic strain of the carbonlike components and chemical swelling/shrinkage of the ramming paste should be considered. Over this preheating phase, other mechanisms such creep of steel and cast iron components as well as the sodium swelling of the cathode block should be considered but are not treated in this paper. At the interfaces, all these changes may have an impact on the electrical and thermal contact resistances which are function of the temperature and the mechanical pressure [5]. Some of these interfaces are critical regarding the energy consumption of the cell such as the cast iron to carbon interfaces where an initial gap distribution takes place before the start-up. Considering the possibility of failure in the carbon block, the contact forces at these interfaces should be computed using the appropriate frictional coefficient. Indeed, the difference of thermal expansion coefficient for the cast iron and the carbon allows moderate relative sliding between both components. If sliding is not allowed, the resulting contact force (normal and frictional) may induce failure in the carbon block. Also, the evolution of the thermo-mechanical contact condition between the shell and the cradle must be considered to allow possible separation. Otherwise, the global stiffness of these components will be overestimated leading to an unrealistic stress state in the components of the cell. The solution of such a problem is obtained by solving a transient coupled non linear system of equations representing the global thermo-electro-chemo-mechanical equilibrium at anytime. Finite element model The finite element model presented in this paper is based on the PI55 technology such as that used at Usine Grande-Baie in Saguenay (Quιbec). This cell consists of 24 anodes and 16 cathode blocks and operates at approximately 185 kA. The model is built using ANSYS® version 11.0 and solved using the finite element toolbox FESh++ which is widely used for the solution of complex multiphysical problems. Description of the geometrical model Figures 2 to 4 show a step-by-step construction of the P155 quarter cell model. More specifically, Figure 2 shows the shell/cradle assembly. The shell is supported by the cradles via thermo mechanical contact at the interfaces, and the cradles are simply supported by the concrete beam located below the cradles via mechanical contact. Regarding the components inside the cell, Figure 3 shows the side wall and pier components as well as the carbon blocks and the ramming paste. The side wall is connected to the shell, pier and ramming paste via thermo-mechanical contact. Also, the cathode blocks are linked to the ramming paste using a thermo-mechanical contact considering that the ramming paste is not a good electrical conductor during the first hours of baking. The link between the small joint and the big joint is ensured via thermo-mechanical contact to avoid excessive tensile stress during the preheating. At the cast iron to carbon interface, a thermo-electro-mechanical contact law is used with an initial air gap distribution and the interface between cast iron, the collector bar being assumed perfectly linked. It should be noted that the end and corner walls are not included in the model in order to decrease the number of mechanical DOFs. These components are replaced by equivalent thermo-mechanical contact properties using the thermal conductivity, the Young's modulus and the

Physical aspects Mathematical modeling of the behavior of electrolytic cells during the electrical preheating phase involves complex physical phenomena and their interactions in the continuous media but also at the interfaces between solids. During the electrical preheating, a high amperage electrical current is applied to the cell during a specific time allowing an important increase of the temperature in the components of the cell due to the Joule's law. These temperature changes have an impact on the electrical, thermal and mechanical properties of the cell components including the effect of the baking level of the ramming paste as well as the curing level of the refractory concrete. These complex phenomena (baking, curing) are related to the evolution of a kinetic model assuming a first order reaction rate such as described by the Arrhenius law. From the mechanical point of view, all these changes allow various strain mechanisms which play a major role in the conditioning of the cell before the molten bath and metal

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located in the side wall and the pier, their corresponding materials are considered thermo-elastic during the preheating.

thickness of each component. Also, it is assumed that the components below the cathode blocks do not contribute to the bending stiffness of the cathode assembly. In this way, the components are replaced by equivalent mechanical stiffness and thermal conductivity using the approach described above. As shown on Figure 3, a virtual support is used to ensure the mechanical stability of the cathode blocks, the side wall and the pier. To avoid an increasing of the shell stiffness, this virtual support is linked to the shell floor but in contact with the inclined wall of the shell. Figure 4 shows the full model including the anode assembles which are connected to a virtual anode beam mainly used to apply the current into the cell allowing a more realistic current distribution in each anode. A thermo-electrical contact interface is used to simulate the coke bed using thermo-electrical contact resistance based on the work of Laberge et al. [10]. Finally, the flexibles are perfectly connected to the collector bars as well as to the busbars which are thermo-electrical components.

Figure 4 - PI55 quarter cell geometry: anode assembly, anode beam,flexibleand busbar. Regarding the contact interfaces, the thermo-electrical contact resistance used for the coke bed is such as described by Laberge et al. [10]. A pressure dependant thermo-electrical contact resistance is defined using the approach proposed by Richard et al. [5] at the cast iron to carbon interface. For mechanical interfaces where frictional contact occurs, a frictional coefficient should be used. Considering the contact interfaces with carbon-based materials, the frictional coefficient is set to 0.08 which represents a moderately lubricated interface.

Figure 2 - PI55 quarter cell geometry: shell and cradles.

Most of the material properties used in this model are confidential and therefore, cannot be published in this paper. Boundary conditions The essential electrical boundary condition consists of a zero voltage applied at the exit of the busbars. The essential mechanical boundary conditions are mainly those needed to describe the double symmetry of the quarter cell model. Also, the three displacements on the surface of the concrete beam are prescribed. Regarding the natural electrical boundary conditions, amperage of 185kA/4 is applied to the quarter cell on a small area located at mid-length on the virtual anode beam. For the thermal part, equivalent natural convection/radiation is applied on external surfaces of the shell and cradles using appropriate coefficients. Inside the cell, the presence of crushed bath around the anode blocks is modeled using a small natural convection coefficient on the sides of the anodes and side wall as well as on the surface of the ramming paste. Considering that the tops of the anodes are insulated with mineral wool, natural convection is also applied on these surfaces using appropriate coefficients. The mechanical part consists of the effect of the weight of the anode assemblies on the cathode plane which is modeled using uniform pressure applied on the total surface of the cathode plane.

Figure 3 - PI55 quarter cell geometry: side wall, pier, ramming paste, cathode blocks, collector bars, cast iron and plug. Material properties As described previously, the good representation of the confinement level in the cell during the preheating involves the utilization of realistic constitutive laws. To do so, the carbonbased materials such as cathode block and the ramming paste are modeled using the quasi-brittle with softening constitutive law developed by D'Amours et al. [8]. To adequately describe the evolution of the thermal and mechanical parameters during baking of the ramming paste, the above-mentioned law is coupled with a kinetic model described by first order reaction rate as proposed by Richard et al. [9]. Considering the large stiffness of components

To ensure the mechanical stability of the numerical model during the first time steps, the mass of each component (except the anode assemblies) is included as a body load. Otherwise, it remains very

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difficult to stabilize the contact conditions on the unloaded mechanical system. Mesh The mesh presented on Figure 5 is obtained after some mesh refinements required to obtain a converged solution in the regions of elevated electrical, thermal and mechanical gradients. The global mesh includes 350080 elements distributed as shown on Table 1 regarding the type and physic of the elements. Considering the number of degrees of freedom (DOFs) per node, the thermo-electric and mechanical sub-problems include 464855 and 595595 DOFs respectively. Table 1 - Number of elements for each part of the model Physic Total E T M TE TM TEM Solid 1248 1776 60776 37794 106000 207594 Shell 8358 - 8358 Contact 5032 3867 7692 51218 24000 91809 BC 68 38751 3500 42319 Total: 68 45031 9143 68468 97370 130000 350080 E: electrical, T: thermal, M: mechanical, TE: thermo-electrical, TM: thermo-mechanical, TEM: thermo-electro-mechanical, BC: boundary condition. Type

Newton-Raphson solution technique until convergence. The global convergence is reached when the correction of each field is less than a specified tolerance. The third approach is to solve the fully coupled non linear problem at each time step using the Newton-Raphson solution technique until convergence. The advantage of this approach is that the linearized system of equations gives a correction for each field simultaneously. This approach allows generally a faster convergence, especially during highly non linear sequences. However, the memory requirement to allow the factorization phase is bigger than that required to solve both sub-problems separately since new couplings appear in the linearized system. For the time stepping aspect, the problem is solved using a backward Euler time integration scheme with an initial time step of one hour. An adaptive time stepping algorithm is used with a maximum time step of three hours. A bisection technique is used to avoid divergence of the Newton-Raphson solution technique during highly non linear sequences. The problem is solved using 20 dual core Opteron™ 64 bits compute nodes with 2 Go of RAM per compute node. As a simple demonstration of the second and third approaches, the Figure 6 shows the CPU time required to solve each time step of a typical preheating of 24 hours using the fully coupled method and the fixed point method. The speed-up of the fully coupled method is a consequence of a better rate of convergence coupled with the adaptive time stepping algorithm. More specifically, the overall CPU time needed to solve this problem by the fully coupled method is 13 hours compared to 24 hours for the fixed point method. The additional memory needed for the factorization of the global problem is simply spread over the 20 cluster nodes used for the simulation.

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Solution technique Such kind of problem can be solved using different approaches based on the coupling level. A basic approach is to initially solve the thermo-electrical sub-problem for all the preheating phase without any adjustment regarding the possible effect of the mechanical conditions on the thermo-electrical equilibrium. Later, the mechanical sub-problem is solved for each time step considering the corresponding temperature and voltage distribution for each specific time. Unfortunately, this simple approach cannot be retained in our context considering that the thermo-electrical contact resistance is pressure dependant.

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Figure 6 - Fixed point -vs- fully coupled method: Comparison of the CPU time. Results and Discussion Sensitivity at the contact interfaces Before any serious numerical study, the model should be validated regarding the ability of the non physical parameters to give a unique solution. Regarding our problem, the penalty numbers used in the mechanical unilateral contact law are such parameters that need to be calibrated. To do so, six contact interfaces are

The second approach, which could be designated as the fixed point method, is to solve for each time step the thermo-electrical sub-problem using the last mechanical conditions followed by the solution of the mechanical sub-problem using the last temperature and voltage distributions. Considering that both sub-problems include non-linearities, each of them should be solved using a

1044

the horizontal displacement of the deck plate at point B (see Figure 2) during the preheating is shown on Figure 10. The in situ measurement is obtained using a survey device. Considering the accuracy of this device, the numerical result extracted at 18 hours is in that range which is ± l m m .

investigated. These interfaces are those perpendiculars to or in the vicinity of the critical path where forces go from the cathodes to the shell. Table 2 shows the interfaces under investigation and the four cases which define the combination of penalty number. Table 2 - Contact interfaces used for the convergence test Case

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The results are obtained for a short preheating of 10 hours. As shown on Figure 2, the horizontal (Z) displacement of point A is monitored for each case. Figure 7 shows that the displacement seems to be converged for the case C or D . However, the case C is a good compromise between speed and accuracy since the computation time generally increases with an increase of the penalty number.

Figure 8 - Comparison with in situ results: temperature at three positions on the shell.

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Figure 9 - Comparison with in situ results: average voltage drop. Impact of contact loss To demonstrate the ability of the model to evaluate the thermoelectro-mechanical contact condition, Figure 11 presents the contact status at the cast iron to carbon interface for each cathode block after 24 hours of preheating. These contact conditions show that sliding and separation occurs in the last four cathode blocks and become more significant for the last cathode blocks. This situation may be explained by the lack of confinement of these cathode blocks combined with the lower temperature in this region of the cell. This separation and/or decrease of contact pressure will adversely affect the current distribution which will be partially redirected to the cathodes located in the center of the cell increasing the current density, and hence, the temperature of these cathode blocks. This situation can be fixed by increasing the stiffness of the shell in this region and by improving the insulation of the end walls which are less exposed to high temperature.

Preheating time (h)

Figure 7 - Effect of the penalty number on the horizontal displacement of point A. Comparison with in situ measurements Some results are now compared with in situ measurements obtained on a P I 5 5 electrolytic cell. Considering a preheating of 24 hours, Figure 8 shows the heat-up of three points located on the external face of the shell. At the end of the preheating, the computed temperatures are consistent with those measured with the thermocouples. Considering the electrical results, the average voltage drop between the clads and the end of the collector bars is computed using numerical voltage and compared with the in situ average voltage drop. Figure 9 shows that the model can adequately predict the early voltage drop in the cell without any fudging factors to control the total resistivity of the cell. Finally,

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To ensure a better understanding of the cell over the preheating phase, the actual works are dedicated to the identification of performance indicators for the cell start-up including the bath/metal addition period. To do so, the development of phase change of molten bath, the strain mechanism associated with the sodium migration as well as the creep of cast iron and steel are in progress.

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Acknowledgements The authors would like to thank Dr Guillaume D'Amours from the Aluminum Technology Center for his guidance regarding the thermo-mechanical behavior of the ramming paste and Dr Patrice Goulet from Laval University for his help during the first debugging stage of the quarter cell model. Finally, we thank the Natural Sciences and Engineering Research Council of Canada for its financial contribution via the Industrial Postgraduate Scholarships Program.

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Figure 10 - Comparison with in situ results: horizontal displacement of the deck plate (point B).

References [1] W.R. Hale, "Improving the useful life of aluminum industry cathodes", JOM, 11 (1989), 20-25. [2] C. Zangiacomi, V. Pandolfelli, L. Paulino, S.J. Lindsay & H. Kvande, "Preheating study of smelting cells", TMS Light Metals, 2005, pages 333-336. [3] M. Dιsuets, D. Marceau & M. Fafard, "START-CUVE: Thermo-electro-mechanical transient simulation applied to electrical preheating of a Hall-Hιroult cell", TMS Light Metals, 2003, pages 247-254. [4] F. Hiltmann, K.H. Meulemann, "Ramming paste properties and cell performance", TMS Light Metals, 2000, pages 405-411.

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[5] D. Richard., M. Fafard, R. Lacroix, P. Clιry, & Y. Maltais, "Thermo-electro-mechanical modeling of the contact between steel and carbon cylinders using the finite element method*. TMS Light Metals, 2000, pages 523-528.

1.183

Figure 11 - Contact status at the cast iron to carbon interface (0: non contact, 1 : stick contact, 2: sliding contact). Conclusion Important works made over the past 10 years have resulted in the development of efficient numerical tools and solution techniques for the simulation of complex multiphysical problems. Combined with the high performance computing facilities, it is now possible to solve very large complex industrial problems in just a few hours. This paper deals with the development of a quarter cell model for the prediction of the thermo-electro-mechanical behavior of Hall-Hιroult cell under electrical preheating. More specifically, this P155 model takes into account the quasi-brittle with softening behavior of the carbon-based materials, the change of thermo-mechanical properties of the ramming paste with the baking index as well as the presence of the most critical contact interfaces.

[6] D. Richard, P. Goulet, M. Dupuis & M. Fafard, "Thermochemo-mechanical modeling of a Hall-Hιroult cell thermal bakeout", TMS Light Metals, 2006, pages 669-672. [7] M. Dupuis, "Using Mathematical modeling of aluminium reduction cell potshell deformation", TMS Light Metals, 2010, page 417-422. [8] G. D'Amours, M. Fafard, A. Gakwaya & A. Mirchi, "MultiAxial Mechanical Behavior of the Carbon Cathode: Understanding, Modeling and Identification", TMS Light Metals, 2003, page 633-639. [9] D. Richard, G. D'Amours, M. Fafard, A. Gakwaya & M. Dιsuets, "Development and validation of a thermo-chemomechanical model of the baking of ramming paste", TMS Light Metals, 2005, page 733-738.

It was demonstrated that the use of a fully coupled solution technique could drastically decrease the computation time. The model has been validated against temperatures, voltages and displacements obtained from in situ measurements at Usine Grande-Baie in Saguenay. Finally, the model has been used to evaluate the contact conditions at the cast iron to carbon interfaces in the carbon blocks after a preheating of 24 hours. It shows that the confinement of the last cathode blocks and the insulation of the end wall should be revised.

[10] C. Laberge, L. Kiss & M. Dιsuets, "The influence of the thermo-electrical characteristics of the coke bed on the preheating of an aluminum reduction cell", TMS Light Metals, 2004, page 207-211.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INFLUENCE OF TECHNOLOGICAL AND CONSTRUCTIVE PARAMETERS ON THE INTEGRITY OF THE BOTTOM OF ALUMINUM REDUCTION CELLS DURING FLAME PREHEATING Alexander Arkhipov, Gennadiy Arkhipov, Vitaliy Pingin RUSAL Engineering & Technological Centre LLC, Pogranichnikov str. 37 bild. 1, Krasnoyarsk, 660111, Russia Keywords: preheating, mathematical modeling, ANSYS, CFD, cell life. Abstract

•

The effects of preheating technology, cathode block materials and other factors on the integrity of the cell during flame preheating of aluminum reduction cells were investigated. The temperature fields, gas dynamics, and stress-strained state of the reduction cell during flame preheating were studied using mathematical modeling in an ANSYS finite-element package and Star-CD CFD-package. The best way to cover the peripheral seams during preheating has been identified.

•

For the problem of heat conductivity, these are the coefficients of heat exchange (allowing for emission) on free surfaces of the shell, the temperature of the surrounding medium, and the initial temperature of the cell. For the SSS these are the loads from the anode assembly and the limitation of displacements of the cathode assembly depending on the conditions of its mounting in the electrolysis shop.

(iii) The specification of the properties of materials and the gas-air medium, namely: density, heat conductivity, heat capacity, dynamic viscosity, elasticity modulus, the Poisson coefficient, yield point, and secant elasticity modulus (the temperature range under consideration for materials isfrom-40 to +1000°C).

Introduction The main cause of premature failure of an aluminum reduction cell is the violation of the integrity of the bottom and, as a consequence, the leakage of the melt to the collector bar and into the base of the cathode. One of the main factors affecting the integrity of the bottom (25%, according to the literature [1]) is the technology of preheating and startup of the reduction cell. The goal of this work was to analyze the effect of technology of flame preheating on the integrity of the bottom and to select the optimal parameters for its performance on the basis of mathematical modeling. To achieve this goal, it is necessary to solve a set of equations of heat and mass transfer and fuel combustion, thermal conductivity, and the mechanics of a deformed solid body [2, 3]. Their analytical solution is inapplicable to reduction cell preheating; therefore, to calculate the temperature fields of the latter, we used a STAR_CD program based on the control volume method; and to calculate the stress-strained state (SSS) of the cathode we used an ANSYSfiniteelement program.

(iv) Computations of the transient temperature field of the cell using a Star-CD program from the beginning to the end of preheating while allowing for fuel combustion and heat-and-mass transfer in a gas-air medium. (v) An analysis of the temperature field for an evaluation of the correspondence to the calculated plot of the increasing temperature to the plot designed according to the process flowsheet of the preheating cell.

Computational procedure The procedure for calculating the temperature field of the reduction cell and the stressed-strained state of the cathode, as well as the results of an analysis and the results of their correspondence to the criteria of high quality preheating, involves the following: (i) The development of a three-dimensional computer model by the drawings and production flow sheet of preheating, which repeats the geometry of the operating or designed reduction cell and gas-air medium (Fig. 1). (ii) The specification of necessary initial and boundary conditions. •

For the gas-air medium, these are the reaction of fuel combustion C7H16 + 1102 -> 7C0 2 + 8H20

Places for mounting the burners

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the velocity of output of the fuel and air from the burner and zero pressure in the orifices for the output of the gases through the cover material.

1047

Figure 1. Finite volume model of the cell for a calculation of the flame preheating, (a) Solid state part and (b) gas-air part with places of mounting burners and outlets.

(ix) If necessary, a change in the process parameters of preheating, lining design, and covering of the cell, as well as repeated calculations of the temperature field and SSS.

(vi) If necessary, a variation in the process parameters of preheating, lining design, and cover of the cell, as well as a repeated computation of the temperature field.

(x) The development of recommendations for a change in the process parameters of preheating and/or design of the cathode assembly to provide the integrity of the bottom in the preheating process of the cell.

(vii) Transfer of the thermal field of the cell obtained in STAR_CD into the ANSYS program and the fulfillment of calculations of the stress-strained state of the cathode facility for different durations of preheating.

The computational procedure and preheating model of the cell were verified by a comparison of calculated and measured (embedding thermocouples into the lining and mounting displacement sensors) data on the temperature of the bottom surface, as well as the temperature and deformations of the cathode shell and lining. Air-gases media

(viii) An analysis of the strains and mechanical stresses and evaluation of integrity of the bottom with the use of the strength criterion of its materials.

Figure 2. Variants 1-4 for a calculation of the flame preheating of the cell. the boundary "end periphery seam-cathode block" (Fig. 2a). The preheating time is 64 h, and the fuel consumption is 6970 kg. (2) Variant 7, but the EAS is charged with CRC from silicon carbide plates to end anodes (Fig. 2b). (3) Variant 7, only the EAS is not charged with CRC but is covered with a heat-insulating plate from above (Fig. 2c). (4) Variant 3, only the SAS is not charged with CRC, but is covered by a heat-insulating platefromabove (Fig. 2d). (5) Variant 7, but the CB contains 100% graphite. (6) Variant 1, but the CBs are graphitized.

Results of calculations The flame preheating of the cell for a current of 300 kA with prebaked anodes and a cradle cathode case was modeled. Cell was preheated by Hotwork diesel equipment with 4 burners pointed out on Figure lb. The calculations of the following types of preheating and cell design were performed. (1) The starting variant: the cell design has a 30% graphite content in cathode blocks (CB), the side-anode space (SAS) is completely charged with crushed recycled cryolite (CRC), and the end-anode space (EAS) is charged with CRC from silicon carbide plates to

1048

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(7) Variant 1, but the preheating time is reduced to 50 h due to an increase in process intensity after 35 h. The rate of the temperature rise is increased from a factor of 1.5 at τ = 35 h to a factor of 1.67 to the end of preheating compared with variant 1. The fuel consumption is 5715 kg.

Variant 4. The absence of charging into the EAS and SAS leads to a substantial increase in the temperature of the longitudinal and end shell walls, as well as the end and longitudinal peripheral seams with a decrease in the bottom temperature in general. An increase in the temperature of the longitudinal shell walls determines the larger extension of its upper part, which is indicated by considerable displacements of the ends (32.4 mm), a substantial lift of the bottom (86 mm), and the displacement of the edge cathode block by a factor of more than 2 compared to variant 1. Compression of the bottom considerably worsens, and the danger that the interblock seams will open appears.

The results of calculation of the field of velocities of the gas-air medium, temperature field, and the SSS of the cathode facility for variant 1 of preheating cells are presented in Figure 3. The results of calculation of the temperature field and SSS for other variants of preheating and cell designs are presented in Tables I and II. Their analysis showed the following. Variant 1. The average bottom temperature in the end of preheating is 892°C. For the interblock seams over the entire height, t > 600°C; i.e., for them, as well as the periphery seams near the cathode units, the temperature exceeds the temperature of the onset of coking the cathode paste (400°C). At distance above approximately 100 mm, the temperature is below 400 °C in the periphery seams.

Variant 5. When using cathode blocks with graphite content of 100% instead of 30%, an increase in the average temperature of the bottom surface from 892 to 911°C is observed. The temperatures of the longitudinal and end peripheral seams and CBs on the boundary with them considerably increase because of the higher thermal conductivity of the CBs. The maximum temperatures of the longitudinal and end shell walls at the level of the bottom surface are elevated by 40°C.

Compression of the bottom in the longitudinal direction is sufficient to conserve its integrity, excluding the peripheral seam, in which tensile stresses up to 2 MPa appear. Compression of the bottom in the transverse direction is also sufficient to conserve its integrity, since tensile stresses are observed only in the interblock seams as well as in the edge cathode blocks (up to 1.2 MPa). The causes of the appearance of the latter are the high temperature gradient and the insufficient rigidity of the paste in the end peripheral seam. However, in this variant of flame preheating, tensile stresses to not reach values dangerous for bottom integrity.

Compared with variant 1, there is no noticeable improvement or worsening of the bottom SSS. Preheating according to variant 2 with CBs containing 100% graphite leads to a noticeable improvement of bottom compression due to an increase in the temperature of edge blocks; end periphery seams; and, consequently, to the weakening of the tensile stresses in them. Variant 6. The use of graphitized cathode blocks makes it possible to substantially increase the temperature of the cathode assembly, and it weakly affects the improvement of its SSS during preheating by variant 1 ; however, it promotes a decrease in or the exclusion of negative consequences of incorrect preheating technology, for example, by variant 2.

Variant 2. The complete charge of the EAS to the bottom surface of the end anodes leads to a considerable decrease in the temperature of the end periphery seams and end cathode blocks (by 140300°C). As a result of an increase in temperature gradients and a decrease in the rigidity of end seams, tensile stresses at the edges of the cathode blocks substantially increase and the danger of violation of the integrity of the bottom appears.

Variant 7. Shortening the preheating time from 64 to 50 h due to an increase in the process intensity leads to a decrease in the average temperature of the bottom surface from 892 to 827°C. Bottom compression in the longitudinal and transverse sections is almost indistinguishable from variant 1. The periphery seams, in which the tensile stresses are practically absent, and edge CBs, where the tensile stresses decreased from 1.2 to 0.5 MPa (which is associated with the lower preheating temperature), are exceptions.

Variant 3. The complete absence of charging in the EAS (the EAS is only covered from above by an MKRKG_400 heat_insulating board or a similar material) leads to a decrease in the average temperature of the bottom surface when compared with variant 1 from 892 to 817°C because of a large heat loss through the side walls, which is indicated by a substantial increase in the temperature of end walls of the shell from 143 to 557°C. The temperature of the longitudinal seams and CBs at the boundary with them decreased by 20-40°C. The average temperature of end seams increased more than by 100°C.

Conclusions (i) Of all the preheating technologies we considered, the best is variant 1 ; in this case, the whole surface of the cathode blocks contacts with the gas_air medium and periphery seams are charged with recycled cryolite. (ii) Charging the EAS to the anode edge leads to the partial insulation of end CBs from the gas_air medium, a larger temperature gradient in the edges of the blocks in the direction of the longitudinal axis of the bath, and a lower temperature of end peripheral seams, which causes large tensile stresses in the end blocks and can cause their destruction during preheating or after the startup of the cell.

An analysis of the SSS showed that, in this variant of preheating, the possibility of formation of the cracks in end CBs decreases, but the deformation of end walls of the shell increases and a large danger of appearance of longitudinal and transverse cracks in the end seams and the detachment of the top of end seams from the silicon carbide plates appears.

1050

(iii) The complete exclusion of charging into the EAS causes strong heating (up to 557°C) of the end wall of the cathode case and its considerable deformation. In connection with this, the danger of appearance of longitudinal cracks in end peripheral seams or their detachment from silicon carbide plates emerges.

2. Mikheev M.A. and Mikheeva I.M., Osnovy Teploperedachi (Principles of Heat Transfer), Moscow: Energiya, 1977. 3. Arkhipov G.V., Barantsev A.G., and Pingin V.V., Abstracts of Papers, V Mezhdunarodnaya konferentsiya "Alyuminii Sibiri_99 (V Int. Conf. "Aluminum of Siberia_99"), Krasnoyarsk, KGU, 2000, p. 147.

(iv) With the complete exclusion of charging into the EAS and SAS, the strong overheating of end and longitudinal walls of the cathode case is observed, which is accompanied by their substantial bending and a lift of the bottom. This in turn leads to the danger that the interblock and peripheral seams will open.

4. Arkhipov A.G. and Polyakov P.V., in Alyuminii Sibiri 2004: Sbornik nauchykh trudov (Aluminum of Siberia 2004: Collected Articles), Krasnoyarsk: Bona Kompani, 2004, p. 149.

(v) The use of graphitized and graphite CBs gives no valuable improvements in the conservation of integrity of the bottom with the correct preheating technology, but decreases the probability of negative consequences in the case of nonoptimal preheating technology.

5. Arkhipov A.G. and Polyakov P.V., Abstracts of Papers, 4_ya konferentsiya poVzovatelei programmnogo obespecheniya CAD_FEM GmbH (4th Conf. of Users of CAD_FEM GmbH Software), Shadskii A.S., Ed., Moscow, Poligon, 2004, p. 323.

(vi) Shortening the preheating time from 64 to 50 h with an increase in the rate that the temperature rises after 35 h does not lead to an increase in the danger that the integrity of the bottom will be violated at the end of preheating; however, it could be accompanied by a larger electrical voltage during startup because of the weaker heating of the bottom.

6. Arkhipov A.G., Polyakov P.V., Dekterev A.A., and Litvintsev K.Yu., in Alyuminii Sibiri_2004: Sbornik nauchykh trudov (Aluminum of Siberia_2004: Collected Articles), Krasnoyarsk: Bona Kompani, 2005, p. 9.

REFERENCES 1. Sortie M. and Oye H., Katody v alyuminievom elektrolizere (Cathodes in Aluminum Cell), Krasnoyarsk: KGU, 1997.

7. Gol'denblat I.I. and Kopnov V.A., Kriterìi prochnosti iplastichnosti konstruktsionnykh materialov (Criteria of Strength and Plasticity of Construction Materials), Moscow: Mashinostroenie, 1968.

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Table L Results of temperature calculations during the preheating of the cell i, XI, for variants Zone of cathode facility 1 1 2 | 3 4 5 Bottom 932 969 1061 1050 1011 Bottom surface, imax 757 911 892 817 846 Bottom surface, iav 795 639 595 586 606 Lower surface of cathode blocks, iav 891 792 679 736 767 Average cathode blocks values Middle of the longitudinal wall (input/output side) 820/ Top of the cathode block at the boundary with the 630/ 757/ 626/ 633/ 624 597 794 peripheral seam 613 656 Bottom of the cathode block at the boundary with the 453/ 465/ 609/ 690/ 467/ 605 peripheral seam 472 453 430 707 540/ 508/ 483/ 746/ Center of the peripheral seam 529/ 496 553 475 767 522 1 Average bulk values of longitudinal peripheral seams 344 332 331 586 465 1 Middle of the end wall (duct/tap end) 639/ 701/ 813/ Top of the cathode block at the boundary with the peripheral 568/ 300/ 599 712 seam 294 706 576 550/ 407/ 1 Bottom of the cathode block at the boundary with the 458/ 331/ 670/ 411 464 460 328 679 peripheral seam 644/ 497/ 708/ 1 Center of the peripheral seam 505/ 331/ 488 326 550 718 513 490 587 495 1 Average bulk temperature of the end peripheral seams 358 209 1 Maximum of the cathode shell walls 77 557 437 182 1 Ends 143 564 160 202 | Longitudinal 160 160

6

7

1055 918 831 912

998 827 527 720

739/ 725 729/ 715 764/ 750 499

581/ 599 376/ 386 467/ 480 306

823/ 797 779/ 755 830/ 806 525

575/ 576 400/ 400 477/ 477 366

187 198

165 165

Table IL Results of the calculation of the SSS

Variant no. 1 | 2 | 3 | 4 | 5 | Displacement of the shell, mm 8,2 13,4 32,4 9,9 Top of the end wall in the longitudinal direction 9,3 8,32 13,7 End wall at the level of the middle of the CB height in the 8,7 7,7 9,7 longitudinal direction 9,14 12 9,7 Top of the longitudinal wall in the transverse direction 10,5 12,3 Bottom center in the direction Z 2,2 2,5 82,9 2,5 1,6 Bottom displacement, mm 9 18,1 10,5 Top of the edge CB in the longitudinal direction 10,7 11,7 2 -0,3 Longitudinal deflection of the edge CB 1,7 1,3 1,5 4,9 End of the middle CB in the transverse direction 5,8 5,5 5,1 6,1 4,9 86 Center of the top of the bottom in the vertical direction 5,2 5 5,1 Normal stress, MPa Edge* of the CB in the transverse direction 1,2 6,3 0,7 -1,0 -1,5 2,9 RP of the end seam in the transverse direction 0,8 1,2 1,0 0,1 Maximal value of the Gol'denblat-Kopnov strength criterion [7] 3,4 Ramming paste 2,9 3,3 6,1 5,1 0,94 0,99 1,53 1,05 Carbon blocks 1,3 0,2 0,41 Edge of last carbon block 0,4 1,07 0,51 * Boundary of edge CBs and end peripheral seams. Zone of the cathode facility

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6

|

1

7

12,4 12,2

8,3 7

14,7 3,1

8,8

13,4 1,9 8,3 7,2

9 1,3 5,1 5

-2 1,8

0,5 0,63

2,7 0,84 0,43

2,63 1,3 0,34

|

2,8 1 |

|

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CREEP BEHAVIORS OF INDUSTRIAL GRAPHITIC AND GRAPHITIZED CATHODES DURING MODIFIED RAPOPORT TESTS Wei Wang1, Jilai Xue1, Jianqing Feng2, Qingsheng Liul1,2, Lei Zhan2, Jun Zhu1, Hua He2 department of Nonferrous Metallurgy, University of Science and Technology Beijing Xue yuan Road 30, 100083 Beijing, China 2 Ningxia Qingtongxia Energy Aluminium Group, China Power Investment Corporation, Xinchang west road 168, Yinchuan, Ningxia Keywords: Creep deformation, Penetration, Carbon cathodes, Aluminum electrolysis Abstract Creep is of importance for evaluating the materials deterioration and deformation of the cathodes in aluminum reduction cells. The purpose of this work is to obtain the creep data for various industrial cathode products. A modified Rapoport equipment was used for measuring the creep behaviors during aluminum electrolysis with CR=4.0 and at temperature of 965°C. Testing samples were taken from three typical industrial cathode blocks: semi-graphitic, full graphitic and graphitized carbon products and were characterized for their graphitized degree using XRD method. The values of doo2 for all cathode samples were lowered after aluminum electrolysis, and the graphitized cathode showed a smaller creep deformation than those of semi-graphitic and full graphitic cathode samples. The results have demonstrated a correlation between the graphitization degree and the creep deformation using a consistent testing procedure, and the obtained data will be useful for better quality control in cathode manufacture and improvement in cell structure design. Introduction Numerous efforts have been made for several decades in improving the cathode properties for a long service life of aluminum reduction cells. The major deformations of the cathode during aluminum electrolysis are found in close relationship with thermal and sodium expansions [1-7]. Recently, creep deformation has received research attention due to its effect on materials deterioration of the cathodes. The creep phenomenon in various cathode carbons was observed using uniaxial compression tests with free lateral strains in laboratory aluminum electrolysis [8,9]. Similar creep behaviors with TiB2/carbon composite cathode materials were investigated in a modified Rapoport apparatus, where the addition of TiB2 in the carbons showed an improved resistance to the creep deformation [10].

Experimental Creep Measurement in a Modified Rapoport System Figure 1 shows the modified Rapoport system used for the measurements of creep strain during aluminum electrolysis. The testing sample rested on a machined Alsint support and acted as the cathode during electrolysis, while the graphite crucible was the anode. The graphite crucible was placed in a vertical resistance furnace and was fixed onto an anode rod. In the electrolysis process, a direct current was passed through the anode rod, the sidewall of the graphite crucible, molten electrolyte and the testing cathode sample to a negative end. The creep strain of the sample during tests was measured by a LVD transducer (range 10 mm, resolution 1 urn) located on the top of the furnace. The signals of the creep strain were logged once a minute into a computer connected to the LVD transducer. The experiments were conducted in this furnace that was flushed with argon gas (99.99%) through the gas inlet and the gas outlet, and the temperature was controlled with a thermocouple in the crucible. The external load was provided by a constant pressure system, which could maintain the pressure at a given value for a period of time. During the testing, the applied pressure was increased each time in steps of 2 MPa.

The creep data can not only be useful in cell structure design and cathode construction due to its effect on the cathode deformation, but can also be applied as a quantitative indication for evaluating the materials deterioration and their service life under operating environment. In this work, the creep behaviors of the cathode carbons during aluminum electrolysis have been investigated against varying degrees of graphitization. All cathode samples were uniformly taken from various industrial cathode blocks that were produced using different technologies; and the electrolysis performances were carried out under operating conditions similar to the industrial environment. The aim of this work is to obtain technical data using a consistent procedure to be used for better quality control in cathode manufacture and improvement in cell structure design.

Figure 1. A modified Rapoport system for creep measurement: 1Load; 2-LVD transducer; 3-Loading frame; 4-Measuring extension pin; 5-Gas outlet; 6-Loading extension rod; 7-Graphite crucible; 8-Cryolitic melt; 9-Testing cathode sample; 10-Alsint support; 11-Resistance furnace; 12-Anode rod; 13- Gas inlet

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The experiments of aluminum electrolysis were performed with a current density of 0.5 A/cm2 at the cathode and a cryolite ratio of 4.0. The cathode sample was 40 mm immersed in the melt and was subject to creep measurement after 2 h electrolysis. Materials and Chemicals Figure 2. Schematic drawing of a sample holder for X-ray diffraction measurements of graphite carbon materials. A: Carbon powder sample at the gluefilm;B: Paper mask.

All testing samples were machined to a cylindrical form for use as a cathode sample in laboratory aluminum electrolysis. Table I shows the density and the porosity of the testing samples taken from three different types of industrial cathode blocks of HC35, HC100 and SMH. As the heating treatment temperature for the carbons increased, the apparent density of the cathode products rose while the porosity lowered.

Results and Discussion Creep Deformation Measurement of Various Cathode Materials During aluminum electrolysis , metallic sodium is generated by the chemical reactions below

Table I. Density, Porosity and Heat Treatment Temperature of Various Industrial Cathode Samples

Materials

Sample

Apparent Density (g.m-3)

Apparent porosity (%)

Heat treatment temperatura (°C)

HC35

Anthracitic+ graphite (35%)

1.57

14.04

1200

HC100

Graphite (100%)

1.64

18.12

1350

SMH

Fully graphitized

1.58

22.27

2850

3NaF(l) + Al(l) = A1F3(1) + 3Na(in C)

(3)

Na(in Al) = Na(in C)

(4)

Previous studies showed that the creep of the carbons was related to the diffusion of this metallic sodium in the cathodes [9,11]. Figure 3 show the creep strain curves vs. testing time for a graphitized cathode sample of SMH with the loading pressures of 2 MPa, 4 MPa and 6 MPa, respectively, at the temperature of 965°C and the cryolite ratio of 4.0. Here, the reason for using the high ratio of the cryolitic melt is to accelerate the rate process of the sodium diffusion in order to reduce the testing time. It is found that no matter which cathode material used, most of the creep curves can be divided into two stages. In the first stage or in the period of 10 min to 15 min from the starting time, the curve exhibits the feature of a typical transient creep that the creep rate (the curve slope) is high at first but soon decreases. This is followed by the secondary stage of a steady-state creep (relatively flat section of the curves), where the creep rate is small and the strain increases slowly with increased testing time, especially when the pressure is 6 MPa. No third stage, i.e., accelerating creep, is recorded which might result in a fracture in the testing samples.

The chemicals used were cryolite (industrial grade with a cryolite ratio of 2.0), and the additives of A1203, CaF2, NaF (analytical reagent). All chemicals were dried at 400°C for 4 hours before testing in aluminum electrolysis. For each run, the total mass of the electrolyte was 160 g containing 5 % CaF2 and 8 % A1203. XRD Measurement for Interlaver Spacing in Carbons For XRD measurement of interlayer spacing, the carbon specimen was cut from the cathode samples and ground to fine powders. The interlayer spacing (doo2) in cathode carbons could be calculated through the Bragg equation: 2d sin È = À

(1)

d=-

(2)

2 sin 6

where ë is the wavelength of X-ray, and Θ is the incident beam angle that can be acquired by XRD data. Figure 2 is a sample holder for XRD analysis, where the sample powders of the sieve fraction < 75 pm were spread on the central plane area using spray adhesive and sieving the sample powder regularly on the glue film to reach a random particle distribution. This treatment can ensure the operating error of XRD measurements less than 5% for various carbon materials [11].

—T—

10

- i —

20

1

30 t (min)

—i— 40

— I 50

—I 60

Figure 3. Creep strain curves with testing time for cathode sample SMH (loading pressure: 2 MPa, 4 MPa, and 6 MPafromtop to bottom) during aluminum electrolysis with CR=4.0 and at 965°C

1054

In Figures 4 5, there are creep strain curves vs. testing time for the semi-graphitic (HC35) and the fully graphitic (HC100) cathodes, respectively, with the loading pressures and electrolysis parameters the same as those used for the graphitized cathode sample of SMH. Among them, HC35 had the highest specific creep strain amplitude, followed by HC100; and both are higher than that of SMH. This phenomenon has a close relationship with the heat treatment temperature of the cathode samples. SMH had the highest heat treatment temperature, followed by HC100 and HC35. Therefore, the SMH sample showed the lowest specific creep strain amplitude. This is because that the higher heat treatment temperature will make a higher Femi level with less Fl layer structure within the graphite materials, resulting in fewer amounts of intercalation compounds between the graphite layers [12].

~ -0.08

I

<° -0.10

(5)

OSeL.

l

ref

where t is time required to reach the steady state, which is the total creep strain, and tQ^otal is t n e time elapsed when 50% of the total creep strain is reached. Table II gives the estimated normalized reference times for the carbon samples of HC35, HC100 and SMH under the external pressure of 2 MPa, 4 MPa, and 6 MPa. Here the creep deformation of carbon cathodes can be considered as time dependent and permanent at high temperatures. The axial strains are used to calculate these values. The estimations are based on the creep strains curves presented in Figures 3-5. For any pressure of the loading, the reference time for creep deformation of the HC35 is the longest, followed by HC100 and SMH. This order is quite similar to the one observed on those cathodes vs. the sodium penetration [2, 10], suggesting phenomenologically the same cause, that is, C-Na intercalation that results in the carbon structure deterioration including the creep. HC35 has the lowest degree in graphitization among the tested cathode carbons, leading to cleavage and slip of the basal planes (Van der Waals bonds) in the materials. Therefore, it requires the biggest J with the same pressure during aluminum electrolysis. More quantitative details about the creep deformation vs. service time for the cathode carbons under pressure are still under investigation.

30

Table II . Normalized Reference Time F Imin/iTiin)(CR=4.0)

t (min)

Figure 4. Creep strain curves vs. testing time for cathode sample HC100 (loading pressure: 2 MPa, 4 MPa, and 6 MPa from top to bottom) during aluminum electrolysis with CR=4.0 and at 965°C

Pressure (MPa)

SMH

HC100

HC35

2

0.2143

0.4286

0.4583

4

0.1892

0.1942

0.2642

6

0.0926

0.1235

0.1702

XRD and Interlaver Spacing of Cathode Materials

— -0.08 H

After the experiment, a small piece of specimen with 4 mm thickness, cut from the bottom 10 mm of the cathode sample, was analyzed by XRD. This interlayer spacing can be an important parameter for characterization of cathode creep. For the purpose of comparison, a sample of high purity graphite (HPG) with consistent quality could serve as a reference standard material in our laboratory.

t (min)

Figure 5. Creep strain curves vs. testing time for cathode sample HC35 (loading pressure: 2 MPa, 4 MPa, and 6 MPa from top to bottom) during aluminum electrolysis with CR=4.0 and at 965°C

Figures 6 and 7 show the XRD analysis of the carbon cathodes before and after aluminum electrolysis, respectively. Whether before electrolysis or after electrolysis, the 2Θ value of SMH at the characteristic peaks is the largest of the tested samples, followed by HPG and HC100, and finally HC35. The interlayer spacing (d) of SMH is the lowest according to Equation 2, which means smaller change in its graphite structure and stronger stability against the creep deformation. This implies that the interlayer

The creep strain is highly dependent on the applied stress, temperature and testing time. The existing creep process may be expressed by defining a normalized reference time, 7 , as shown in Equation (5):

1055

spacing values of cathode carbons obtained from the XRD data can serve as a characteristic parameter for evaluation of the cathode quality and its resistance to creep deformation.

sliding, mass penetration in micro-pores, migration and recrystallization, etc. It could be a time-consuming and laborintensive work to make quantitative measurements on all those effects in practice. The summary of the obtained results in this investigation demonstrate a consistent method applicable in creeprelated materials evaluation and structure design for reduction cathodes. WM Before electrolysis &23 After electrolysis

27

28 2è(°)

Figure 6. XRD analysis of the cathode samples without aluminum electrolysis at 25 °C 200000 A SMH 150000

HC100

HC35

Figure 8. Chart summarizing changes in the interlayer spacing of the tested cathode carbons before and after aluminum electrolysis

100000 H 50000

Conclusions

. HC35 HPG HC100 SMH

o-r 25

HPG

26

27

28

29

1.

Creep deformation of industrial cathode carbons with varying degrees of graphitization can be quantitatively differentiated using a modified Rapoport system under identical loading pressure, operating temperature and testing time, etc.

2.

The creep strain of the tested cathode carbons is decreased with increased graphitization: SMH (fully graphitized) < HC100 (100% graphitic) < HC35 (35% graphitic), while the time span in the creep process is increased with lowered degree in graphitization.

3.

A reduction in the interlayer spacing (doo2) of the cathode carbons after aluminum electrolysis is found with all types of tested cathodes that have exhibited a creep behavior, and the change in doo2 values before/after aluminum electrolysis is closely correlated to the creep resistance of the cathode materials.

30

Figure 7. XRD analysis of the cathode samples after aluminum electrolysis at 965°C and CR=4.0 Table 111 The Interlayer Spacing doo2 (A) in Cathode Carbons Time

SMH

HC100

HC35

HPG

Before electrolysis

3.3730

3.3856

3.4010

3.3735

After electrolysis

3.3595

3.3683

3.3656

3.3570

For comparisons of micro-structural change related to the degree in graphitization for the cathode materials, the values of the interlayer spacing (d) for the tested samples are calculated, as presented in Table III.

Acknowledgement Financial support from National Natural Science Foundation of China (NSFC), and Funding for Doctor Degree Education from Ministry of Education of China are gratefully acknowledged.

For clarification and comparison, a chart of various changes of the interlayer spacing in the tested cathode carbons before and after aluminum electrolysis is presented in Figure 8. The creep behavior of the carbon or graphite materials is the result of a number of chemical and micro-mechanical processes, such as the motion of defects, diffusion in grain boundary, grain boundary

References

1056

1. Morten Sorlie and Harald A. Oye, Cathodes in Aluminum Electrolysis (Dusseldorf, FRG: Aluminium-Verlag GmbH, 1994), 282-361. 2. J. Xue, Q. Liu and W. Ou, "Sodium Expansion in Carbon/TiB2 Cathodes during Aluminum Electrolysis," Light Metals, 2007, 1061-1066 3. P . Y . Brisson, et al., "X-ray Photoelectron Spectroscopy Study of Sodium Reactions in Carbon Cathode Blocks of Aluminum Oxide Reduction Cells," Carbon, 44(2006), 14381447 4. M. B. Rapoport, V.N. Samoilenko, "Deformation of Cathode Blocks in Aluminum Baths during Process of Electrolysis," Tsvet Me, 30(2) (1957), 44-51. 5. D. S. Newman, et al., "Technique for Measuring In Situ Cathode Expansion (Rapoport Test) during Aluminum Electrolysis," Light Metals, 1986, 685-688. 6. R . C. Asher, "A Lamellar Compound of Sodium and Graphite," Inorganic Nuclear Chemistry, 10(1959), 409-410 7. Nafaa Adhoum, Jacques Bouteillon and Daniel Dumas, "Electrochemical Insertion of Sodium into Graphite in Molten Sodium Fluoride at 1025 °C," Electrochemical Acta, 51(2006), 5402-5406 8. A. Zolochevsky, et al., "Creep and Sodium Expansion in a Semi-Graphitic Cathodes Carbon," Light Metals, 2003, 595-602. 9. D. Picarda, et al., "Room Temperature Long-term Creep/Relaxation Behavior of Carbon Cathode Materials," Materials Science and Engineering A, 496(2008), 366-375 10. J. Xue, Q. Liu and B. Li, "Creep Deformation in TiB2/C Composite Cathode Materials for Aluminum Electrolysis," Light Metals, 2008, 1023-1027 11. F. Aune, W. Brockner, and H. A. Oye, "X-ray Characterization of Carbon Cathode Materials," Carbon, 30(7) (1992), 100-1005 12. M. C. Robert, et al., "Reaction of Sodium with Graphite," Chemistry and Physics of Carbon, 10(1978), 83-106

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION

Cathode Materials and Wear SESSION CHAIR

Frank Hiltmann SGL Carbon GmbH Meitingen, Germany

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

MEASUREMENT OF CATHODE SURFACE WEAR PROFILES BY LASER SCANNING Egil Skvbakmoen, Stein R0rvik, Asbj0rn Solheim, Knut Ragnar Holm*, Priska Tiefenbach* and 0yvind 0strem* SINTEF Materials and Chemistry, NO-7465 Trondheim, Norway *Norwegian University of Science and Technology, NO-7491 NTNU, Trondheim, Norway Keywords: Cathode wear, laser scanning, wear profiles

Abstract The service life time for high amperage aluminium reduction cells with graphitized cathodes is limited by cathode wear. The wear is normally very non-uniform, and it is commonly documented by photography and/or manual point measurements. In an attempt to record the wear pattern in a much more detailed way, a laser scanning procedure was developed. A laser scanner with a single point accuracy of 10 mm has been used to produce a 3D model based on three overlapping scans with an average resolution of about 1 cm. The same cathodes were also measured manually for comparison. The method developed gives detailed information regarding the wear at different positions within the cell, and it may become a valuable tool for investigating the influence of different parameters on the cathode wear.

the metal. The cells were considered to be run under normal electrolysis conditions during the life time, which were 2088 and 2184 days, respectively. Pictures of the cathode surfaces of both cells are shown in Figure 1. The cathode surfaces were thoroughly cleaned before measuring started.

Cell D 105 (2088 days)

Introduction

Cell D107 (2184 days)

Figure 1. Cathode surfaces of cells D105 and D107.

The very non-uniform wear observed in aluminium cells with graphitized cathodes is today limiting the cell life and it represents a great challenge for the aluminium industry.

Laser Scanning Method The laser scanning method offers the possibility to capture a surface of an object — in this case a cathode — in great detail. It is a non-contact active measuring system which acts by sending and receiving laser light. The scanner shoots a laser beam and collects the reflected light. The result is a point cloud consisting of a resolution dependent number of points with 3D coordinates and often some additional attributes as well. One distinguishes between mainly two kinds of laser scanners which work with different techniques: triangulation and time-of-flight laser scanners.

The so-called w-wear profile is reported by several authors [1,2]. The mechanism for the wear observed is still not fully understood. It seems clear, however, that current densities and metal/bath movements play important roles together with the formation and dissolution of aluminium carbide. Higher wear is observed in areas with high current densities. The ramming paste between the cathode blocks normally shows less wear. The documentation of the observed wear is usually made by photography and manual point measurements. This gives a crude overview of the wear pattern with little preservation of details. Therefore, there is a need for a new method which is more detailed and provides documentation that can easily be studied in retrospect. In this work the laser scanner method is reported and compared with a manual telescope leveling method. Both methods have been used during post mortem investigation of two cells at the Hydro Sunndal aluminium smelter.

A time-of-flight laser scanner, such as the one used (Riegl LMSZ420Î [3]), measures the time t a laser pulse needs to travel from the sender unit to the target surface and back to the receiver unit, and thus the distance c/can be calculated. d = ct/2

c: speed of light

Together with the known horizontal and vertical angles analogue a total station (tachymetry) - 3D coordinates can be computed. A laser scanner beam will be reflected from the first object it hits; hidden objects cannot be measured (shadow effect). Therefore, it is often necessary to work with more than one scan position.

Cells, Measurements, and Procedures Cells Two cathode linings (D105 and D107) from Hydro Sunndal Al plant (SU4) were investigated with respect to cathode wear profile. The cell design for both cells was similar and the surfaces consisted of 19 graphitic cathode blocks with ramming paste in between. Both cells were shut down due to high levels of iron in

The terrestrial time-of-flight laser scanner used in the present experiments has a measurement rate up to 1Γ000 pts/sec; measurement range 2-l'000m; field of view 80 x 360 degrees,

1061

and it consists of a fixed lower part (in this case mounted on a custom-made frame) and a rotating upper part analogous to a tachymetrical total station. The generated laser beam is diverted by a rotating respectively oscillating mirror. After measuring a vertical line with a given resolution/angle, the scanner turns horizontally at a given angle and continues scanning the next vertical line, as illustrated in Figure 2. A calibrated digital photo camera Nikon D200 (f=20mm, 10.2 megapixels) was mounted on the scanner, making it possible to incorporate RGB values for the points obtained by the laser scanner. Figure 4. Laser scanner positioning above the cathode.

((Û i r»Π <^P (>tt"-- „di c ( < 1 A\ ·:μ /I

º

1 À

Figure 2. Principle of the laser scanner. Laser Scanning Configuration and Procedure A custom-made steel frame constituted the platform for the laser scanner (Figure 3). With the help of a crane or a digger it was mounted on the steel shell of the cathode, and the laser scanner was fixed at the top. The construction could easily be moved along the pot shell while the laser scanner was attached.

■ CcfarwSeari>o*0ß1 £3 Cobi.ScariPes«» ft C<*»„Scari>oi002

Figure 5. Three overlapping point clouds - the cathode, the steelframeand the surrounding area are visible. Data Processing To be able to get accurate results from the data processing of the geometrical laser scanning data together with the photos taken by the calibrated camera, calculation of the camera mounting calibration (camera position related to the scanner) was essential. Therefore, these calibration values are calculated for each of the three scan positions by using the scanned targets. Each point cloud (one per scan position) is then coloured by the photos taken at the same scan position. The working step 'registration' links together the different scan positions via the targets they have in common. This registered total point cloud had 3D coordinates (XYZ) in a casual local coordinate system which was then transformed to a cathode-adjusted one. For this step the three single point clouds had to be re-registered to one total point cloud.

Figure 3. Steelframefor placement of the laser scanner. To obtain data for the entire cathode surface with approximately the same resolution and to avoid shadow effects, scans from three positions with overlapping areas were performed (as illustrated in Figures 4 and 5). 20-30 circular tape reflectors (target points) were distributed on and around the cathode to help aligning the three overlapping scans. This configuration also enhanced the point accuracy (the single point accuracy of Riegl LMS-Z420Ì is 10 mm). At each scan position a scan with a resolution of 0.05 degrees (approx. 5 mm at a distance of 6 m) was executed, and photos were automatically taken for the corresponding surface as well.

1062

The transformation results were as follows: Standard deviation [mm] - Cell D105: Point cloud 1 to new coordinate system: 3.2 (16 targets used) 2 " 2.6 (19 ") 3" 3.1 (19") Standard deviation [mm] - Cell D107: Point cloud 1 to new coordinate system: 1.8 (27 targets used)

2 " 2.0 (29 ") 3" 1.8 (25 ") Useless or even disturbing points were deleted and the data were merged to one single 3D point cloud (5.3 million points). By using the scanner related software RiScan Pro and other programs, various types of visualisation could be obtained (using points and/or the triangulated surface). Manual Method with Leveling Telescope The current method for measuring the cathode wear profiles at Hydro Sunndal makes use of a leveling telescope and a measuring stick with a bubble level attached. The leveling telescope is placed on a stable surface outside the cathode shell from where one can read the measuring stick, which is placed at different locations on the cathode surface. On each corner of the steel shell there are reference points, from which the original depth of the newly installed cathode blocks is known. The difference in depth before and after operation can then be measured at several points on each cathode block. The typical number of measuring points is 7 points per cathode block, covering the deepest points on the sides, the highest point in the middle and the local high and low points in between. The cathodes investigated in this work consisted of 19 cathode blocks, giving 133 measuring points. A team of three persons, one operating the leveling telescope, one holding the measuring stick, and one taking notes, needs about one hour to measure one cathode. Results

Figure 7. 3D model colored according to the depth of the cathode. The 3D point coordinates from the laser scan were post-processed using Perl (a public domain data processing language) to resample the data to XYZ values at 1 cm resolution on each axis. The resampled data was converted to a grayscale image, where each pixel XY coordinate (column and row) corresponds to the respective XY position in cm, and the Z value (gray level) corresponds to the respective Z depth in cm. This picture was opened in the public domain image analysis program ImageJ for further visualization and plotting. These programs were chosen for convenience; the data could also have been plotted in other data plotting or image analysis software. The manual method gives a set of XYZ values, where the X and Y values represent the measuring positions on the cathode surface and the Z values are the measured cathode depth. The amount of data is relatively small (133 points in this case), and interpolation is needed to create a good visualization of the wear profile. Just like the data from the laser scans, the data from the manual scans can be plotted like 3D models and contour plots. In this work linear interpolation was used to fill in the gaps between the measuring points. The data in the Z direction is quite reliable; the X and Y coordinates are usually not. For convenience the X values are determined by the cathode block number and the seven Y values on each block are assumed to be the same for every cathode block. In reality these points are not in the same spot on every cathode block. The measuring stick is typically placed at the deepest or the highest point (i.e. the most interesting point) in the relevant area to be measured. It is also worth noting that the less worn ramming paste between the cathode blocks was not measured with the manual method, according to the standard procedure.

Topographic Plots with Laser Scanner and Manual Method Besides the forever reusable 3D point coordinates including additional values (i.e. intensity, RGB), saved in a common ASCII format, various meaningful products were created. For example 3D visualizations of the cathode where the point cloud is coloured by RGB were made (Figure 6), as well as a 3D model coloured according to the height or respectively the depth of the cathode surface (Figure 7). Also better known visualizations as i.e. height curves or horizontally and vertically surface cuts were calculated.

The two methods are compared in Figure 8 and Figure 9. The pictures are sized and coloured using the same scale, from blue (highest) to red (deepest). It can be observed that while the average values are similar, the laser scans are much more detailed. The perimeter of each cathode block can be seen as areas of less wear in the laser scans. Obviously, this information is lacking in the recordings from the manual method. Figure 6. RGB-coloured scan data of cathode (Cell D105).

1063

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1065

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A plot of wear profiles on each single cathode block numbered from 1-19 for the two cells by the manual method and laser scanning is shown in Figures 10-13. The initial cathode level before operation is used as reference and cathode heave has not been taken into account. Positive values mean that the cathode level after operation is higher than the initial level, which is a result of cathode heave. By assuming that the contribution of cathode heave is the same for all the measuring points, the cathode wear plots show good correlation with the real wear pattern of the cell. Line plots of the data set from the manual method in the longitudinal direction of each cathode block reveal a general trend for both cells; a WW-profile, as shown in Figures 10 and 11. The same trend can be seen from the plots obtained from the laser scanning data shown in Figures 12 and 13. Cathode blocks no. 1, 2 and 19 stand out; 1 and 2 being less worn and 19 more worn in the mid section of the block. Cathode block no. 19 is located at the metal tapping end of the cell, while block no. 1 is at the suction end. Figures 15 and 16 compare line plots from the two methods for two single cathode blocks.

Figure 15. Wear profiles of two selected cathode blocks in cell D105, comparing laser scan and manual measurement methods.

A plot of the wear profile in the longitudinal direction of the entire cathode is shown in Figure 14. It is evident that the ramming paste between the graphitic cathode blocks is less worn. Summary The laser scanning method gives several advantages: Much more detailed measurements than the manual method. The detailed wear pattern and shape of the cathode surface can easily be visualized in several ways. The high resolution makes it possible to zoom in and study local areas inside the cell (for instance tap pot holes or other local areas of interest). Wear profile plots can easily be made with high accuracy at different positions within the cathode surface inside the cell. Local areas of high wear (weak points of the cathode) are easily found. If the extent of cathode heave is known, the average cathode wear rate and the total carbon consumption can be easily and accurately estimated.

Figure 16. Wear profiles of two selected cathode blocks in cell D107, comparing laser scan and manual measurement methods.

Thanks to Alf Inge Ulvund and Tore Engen (Alf Engen AS) for design and construction of the steel frame for placement of the laser scanner instrument and all help and good advice during the scanning procedures. We appreciate also the assistance from the following personnel at the Hydro Sunndal plant: Jana Hajasova, Svein Kare Sund, and Anders B0rste.

The more detailed measurements give more information and may provide more knowledge regarding cathode wear phenomena. Especially when considering modelling data of metal flow, cathodic current density, alumina distribution and also type of cathodes and linings the wear can be explained in a better manner than before. We regard that this method will be a valuable tool in deriving the mechanism for cathode wear in aluminium cells.

References Acknowledgement

[1]

M.S0rlie, H.A.0ye, "Cathodes in Aluminium Electrolysis" 3rd edition, Aluminium Verlag 2010. [2] P.Reny, S.Wilkening, "Graphite cathode wear study at Alouette", Light Metals 2000, pp 399-405. [3] http://www.riegl.com (1.9.2010).

Thanks to Hydro and Research Council of Norway for financing the work through the research program Process Innovations for High Current Density (PI-HCD). Permission to publish the work is gratefully acknowledged.

1066

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

COKE SELECTION CRITERIA FOR ABRASION RESISTANT GRAPHITIZED CATHODES Raymond Perruchoud1, Werner Fischer1, Markus Meier1 and Ulrich Mannweiler2 1

R&D Carbon Ltd., P.O. Box 362, 3960 Sierre, Switzerland 2

Mannweiler Consulting, 8006 Zürich, Switzerland

Abstract The high abrasion rate of graphitized cathodes of high amperage cells is a major limitation of the pot lining life time. This is related to the use of low Sulfur anode grade petroleum coke resulting in a too soft cathode after graphitization. A review of the effects of the coke characteristics on the cathode abrasion as well as other pot lining relevant properties has been performed in a pilot scale. The selection of more isotropie calcined cokes produced in a delayed coking plant from appropriate soft pitches allows a decrease in the abrasion rate by a factor of four. No detrimental effects on the graphitization behaviour or the cathode relevant properties can be observed. Dedicated coal tar pitch feedstocks used in optimized delayed coking in combination with shaft kiln calcination have the potential, to solve the "W shaped" wear of graphitized cathodes that is responsible for the short pot life time. Introduction In an excellent overview paper published in 1989 W.R. Hale [1] stated that by measuring the erosion rate of the cathode in the pot and knowing the thickness of the collector bars, plants can determine their ultimate pot life. However, he observed that few cathodes failed due to erosion on a regular basis.

These interesting approaches, however, were not economically successful, as a significant anisotropy of the extruded blocks required expensive needle like coke materials. A variable resistivity through adapted graphitization conditions means that a significant part of the block has an electrical behaviour similar to inferior, cheaper graphitic grade blocks. Therefore an improvement of the cathode material is urgently needed to improve the resistance to the wear. A considerable amount of work on cathode characterization has been performed especially on the wear behaviour over the last two decades. Recently Skybakmoen et al. [6] concluded that the exact mechanism behind the "W" shaped wear pattern can still not be explained by the available laboratory test methods applied to different sophisticated laboratory cells. On top of this they confirmed that the laboratory cell set-ups, described in their excellent overview paper, were not able to distinguish between cathodes having different levels of apparent density. However, clear indications in the industrial practice exist that denser graphitized cathode blocks show a slower wear rate. The shortest and best summary of the mechanisms at work is still given by 0ye and Welch [2]. The chemical wear through AI4C3 formation as well as the physical wear, either by detachment of carbon or A14C3, are operative mechanisms. The chemical wear is non-discriminating with regard to materials but depends on current density and electrolyte composition while the physical wear can be explained by the variable behaviour of different cathode materials.

One decade later cathode wear became the life-determining factor for high amperage pots boosted to high current densities [2]. The maximum erosion area systematically observed at the graphitized block ends, the so-called "W shape" erosion, was responsible for that significant reduction of the pot life time [3]. For high intensity / high current density pots the bench mark figure of 3000 days for semi-graphitic (30% Graphite) blocks was suddenly reduced to less than 2000 days when graphitized blocks were used.

Therefore we took, as a wear ranking criteria of cathodes, the physical abrasion measured on cathode disk specimens rotated with a given pressure on a sand paper surface. The test results overestimate the difference between graphitized and amorphous cathode erosion rate [7] but this physical abrasion sand paper test is a precise and very discriminating tool for judging the merit of any graphitized cathode type.

As the current density is mainly driving the erosion process, different solutions have been studied to obtain a more homogeneous current distribution at the cathode block surface. Dreyfus and Joncourt [4] promoted the manufacture of blocks showing a high anisotropy factor of the block resistivity and later Dreyfus et al. [5] introduced the concept of variable resistivity cathodes for fighting the local W shaped graphite erosion.

It is known that the hardness of graphitized cathode depends on the following parameters:

1067

Type and quality of coke Pitching level Mixing and forming intensities Impregnation Graphitization degree In the studies reported below a fixed pilot plant manufacture of unimpregnated graphitized cathodes were used for investigating the impact of the type of coke, at their optimum pitch content, on the cathode relevant properties. Selection of Cokes Classical raw materials for graphitized cathodes have been tested in order to establish a base line of the graphitized cathode types currently available in the market. This included low Sulfur, low Vanadium anode grade petroleum coke, delayed pitch coke out of low QI tar, low Sulfur petroleum shot coke. For petroleum coke the following candidates were considered: needle cokefromdecant oil, medium Sulfur, Vanadium anode grade coke, shaft kiln low Sulfur, Vanadium anode grade coke, to cover first, the entire scale of the macrostructure (anisotropie needle coke), second the possible positive effect of the less anisotropie medium S/V coke and third of the harder shaft kiln coke related to the nature of the calcining process.

cokes. The other cokes were calcined at a typical degree for anode grade or for recarburizer materials. Experimental Graphitized Cathode Pilot Preparation The calcined cokes were sieved in seven fractions from +8mm to -0.25mm. The oversize +8mm was crushed to -8mm and resieved. The finest fraction -0.25mm was blended with the intermediate material 1-0.25mm for the preparation of fines (3500 Blaine) using an air jet collision mill equipped with a classifier. The following dry aggregate recipe was used 8-4mm 4-2mm 2-lmm

14% 14% 14%

l-0.25mm 0.5-0.25mm Fines 3500 Blaine

14% 14% 30%

An addition of 0.5% Fe 2 0 3 fines was chosen for moderating any puffing tendency during the graphitization process. A coal tar pitch with a medium QI content of 9% and a typical softening point (SP) Mettler of 112°C was used as a binder. The dry aggregates were preheated at 200°C prior to mixing. An intensive impeller mixer (10 1, Eirich) was used. The paste temperature was adjusted to 182°C (70°C higher than the pitch SP). For each coke the pitch content was varied in a 4% range which was selected according to the porosity characteristics measured on the coke. For each pitching two batches of paste were produced and two pilot electrodes of 146 mm diameter and -200 mm height were pressed with 200 bar at 140°C. The paste was cooled by using water injection into the mixer at the end of the mixing process.

From the pitch coke family a second type of material, produced from a tar rich in high molecular weight component (high QI content tar) was selected as it is known and proven [8] that thanks to a more isotropie coke texture harder cathodes can be obtained.

Green pilot electrodes of 5 to 6 kg were baked in an electrical furnace at a rate of 10°C/h during devolatilization between 200 and 600°C. The soaking time was 20 hours at 1100°C. Graphitizing was performed on three cores of 50mm diameter and 130mm length in a lOOkW length wise pilot furnace where samples are pressed pneumatically at 10 bar pressure (Figure 1).

In the same line, a medium temperature tar pitch coke (a byproduct of coal gasification) was also selected. This material is unsuitable for use in prebaked anodes as brittleness, related to high coefficient of thermal expansion and inelasticity was experienced. These two cokes can be named as follows: delayed pitch coke out of high QI tar, medium temperature coal tar delayed pitch coke. A low porosity isotropie coke produced out of oil shale residue by delayed coking was evaluated as well. An acetylene coke resulting from the agglomeration of carbon blacks during the preparation of acetylene through the partial combustion of methane was also tested. These two more exotic cokes, that are however produced on a significant production scale, can be named as follows: delayed oil shale shaft kiln calcined coke, calcined acetylene coke.

Figure 1. Length wise graphitization furnace for 050mm cores

The 10 selected cokes were industrially calcined materials, most of them in rotary kiln except for the shaft kiln anode grade coke and the oil shale coke. The calcining degree was of course high for the needle coke, a normal practice for this type of material, and for the two low and high QI coal tar delayed pitch

The electrical current was adjusted in the range of 600 to 3000°C for maintaining a constant heat-up rate of 500°C/h. The length of the sample column was continuously registered in order to quantify the puffing pattern but also any post baking temperature shrinkage and / or graphitization shrinkage occurring at temperatures above 2000°C.

1068

Testing The calcined cokes were analyzed by using the classical routine quality control ISO standards. Additionally the porosity was measured by Hg pressure porosimetry and the Nitrogen content by gas chromatography after combustion. The hardness of the coke, an important property of the calcined material revealing its macrostructure, was characterized by using the Hardgrove mill with increased number of revolutions (200 instead of 60 in the HGI ASTM D409 standard) and calculating the number of grams per minute of milling passing 75 Mesh [9]. This is reported as a pulverizing factor that is close to 1 for anode grade coke but can reach four times lower values for hard coke. On the baked pilot anodes the sticking propensity of the packing material and the shrinkage of the pieces after baking were measured in order to detect optimum pitching conditions. Non-destructive testing were performed including the specific electrical resistance and the dynamic elasticity modulus that are also excellent indicators for optimum binder content for the graphitization stage. On the dilatometrie curves during the graphitization, the rate and the total extent of puffing, of post baking or post puffing shrinkage and graphitization shrinkage (T>2000°C) were determined.

The volumetric change on the cold samples (before and after graphitization) as well as the graphitization weight loss were measured. On the graphitized cores beside the physical/thermal/electrical properties, the following tests relevant for the cathode behavior in the pots were performed: abrasion sand paper test (RDC-191), Rapoport swelling test (ISO 15379-1), sodium vapour test (RDC-193). Results Table I gives an overview of the data obtained for optimally pitched dry aggregates for which the corresponding graphitized cathode abrasion resistance was reached with the lowest possible level. The results of the coke candidates are ranked from left to right of the Table I in descending abrasion rates of the graphitized cathodes (see also Figure 2). It is interesting to note that the first five cokes with high abrasion rates are petroleum cokes, from needle to shot coke, followed by pitch cokes from coal process. Then the shale oil shaft kiln coke and especially the coke byproduct from the acetylene preparation process reached an unprecedented low abrasion rate which is six times lower than for graphite electrodes made with typical cathode raw materials.

Table I. Coke, baked and graphitized cathode results Source Petroleum Petroleum Petroleum Petroleum Petroleum Feedstock Decant Oil Low S Res. M. S Resid Low S Res. High Asph. Calcination Rotary Rotary Shaft Rotary Rotary Grade Anode Anode Anode Shot Needle

COKE PROPERTIES Sulphur % Nitrogen % Vanadium ppm Pulverizing factor Density in xylene kg/dm3 Crystallite size Le  Tapped bulk density 2-1 mm kg/dm3 Hg porosity % PITCH CONTENT BAKED PILOT CATHODES Green apparent density kg/dm3 Baked apparent density kg/dm3 Baking loss % Shrinkage % Sticking of packing material % Sp. Electrical resistance μΏπι Compressive strength MPa |Dynamic elasticity modulus MPa GRAPHITIZATION BEHAVIOUR Puffing rate lO" 6 ^ 1 Post Baking Shrinkage rate lO"^" 1 Graph. Shrinkage rate 1Ö^K' * Cold volumetric change % Graphitization loss % GRAPHITIZED CORES Graph, apparent density kg/dm3 Sp. Electrical resistance μΩπι Thermal conductivity W/mK CTE lO" 6 ^ 1 Xylene density kg/dm3 Dynamic elasticity modulus MPa Abrasion % Compressive strength (CS) MPa CS after Na vapour test MPa |Rapoport swelling %

Coal LowQI Rotary Cathode

Coal High QI Rotary Cathode

Coal Oil Shale Med. T. Tar Resid Rotary Shaft Recarb. Recarb.

Acetylene 1

-

Rotary Recarb.

0.37 0.28 1 1.38 2.133 35.1 0.901 16.2 16%

1.02 1.04 35 1.29 2.073 28.0 0.862 20.5 16%

1.96 0.96 194 1.14 2.063 28.1 0.794 21.5 16%

0.37 1.66 7 0.75 2.069 26.6 0.926 16.6 14%

1.09 1.42 482 0.56 2.053 26.2 1.087 12.1 12%

0.35 1.18 13 0.53 2.071 35.7 1.010 14.1 14%

0.26 1.40 10 0.35 2.050 38.8 1.042 9.7 12%

0.16 1.50 1 0.35 1.988 27.3 1.064 8.8 12%

0.40 0.30 1 0.25 2.029 23.3 0.980 5.5 12%

0.02 0.30 8 0.40 1.909 19.4 0.971 11.9 10%

1.716 1.664 5.2 2.2 0.06 62 27.0 5.8

1.648 1.591 5.3 1.9 0.10 59 38.3 7.8

1.600 1.557 5.1 2.5 0.27 59 45.6 8.3

1.651 1.612 4.3 2.2 0.10 62 37.4 6.6

1.628 1.617 3.9 3.3 0.05 46 68.9 10.1

1.655 1.625 4.3 2.0 0.19 48 58.7 9.3

1.671 1.657 3.8 3.0 0.22 43 70.0 10.3

1.661 1.599 4.0 0.3 0.33 43 71.3 9.8

1.645 1.629 3.6 2.6 0.20 53 57.9 8.5

1.552 1.552 2.9 2.8 0.33 59 55.9 9.6

1.1 0.7 18.3 -1.56 2.58

12.0 0.8 17.1 -1.43 4.15

55.0 0.0 13.2 -1.53 5.20

17.0 0.0 14.6 -1.97 4.22

0.0 4.1 12.2 -5.58 4.63

5.4 2.1 6.1 -2.76 3.52

6.5 0.9 2.8 -1.80 3.15

0.8 10.7 18.1 -7.27 3.55

0.0 32.9 0.0 -3.97 2.71

0.0 40.8 0.0 -8.54 2.21

1.608 14.0 86 3.27 2.245 2.5 60 12.0 17.0 0.25

1.541 13.7 87 3.48 2.235 3.5 54 15.9 17.1 0.24

1.495 16.7 84 3.70 2.226 3.6 50 20.4 18.8 0.24

1.562 16.3 78 3.54 2.236 2.9 40 16.9 23.5 0.25

1.635 10.7 99 5.62 2.213 6.6 28 37.4 24.4 0.42

1.620 13.3 93 4.64 2.198 5.7 28 32.2 31.7 0.34

1.636 14.7 98 5.33 2.171 6.9 23 42.0 26.2 0.45

1.663 13.3 98 6.32 2.171 7.3 15 50.0 29.9 0.44

1.642 14.7 83 5.79 2.205 5.2 14 30.9 30.6 0.36

1.643 34.7 37 5.92 2.081 6.9 9 40.2 14.4 0.52

1069

This shows how important the selection of the coke material is but also takes into account the behavior during the graphitization step and also the other relevant cathode properties.

skrinkage up to 2100°C followed by pre-graphitization shrinkage up to 2700°C. But on the contrary to the needle coke material there is no end-graphitization shrinkage up to 3000°C, a sign that the shot coke is harder to be graphitized. As a consequence the level of the xylene density of the end product is significantly lower than for the graphitized needle coke artifacts (2.213 vs. 2.245 kg/dm3). The suppression of the puffing for isotropie material along with the continuous shrinkage tendency during the graphitization (despite the high thermal expansion level) results, for the shot coke artifacts, in a volumetric change of the cold specimen after graphitization of -5.6% i.e. three times the values of the other tested petroleum cokes. This graphitization behavior is also favorable for a high resistance to the abrasion of the graphitized cathode but also for a low resistivity level. (%») Change of artifacts length

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The positive effect of the slow shaft calcining process, related to the coking of evolved tars from the green coke on the coke pores, is also quite significant. Compared to the other low S anode grade coke calcined in rotary kilns, which has the same macrostructure characteristics (same CTE and Xylene density of the graphitized cathode), the abrasion value improves from 54% down to 40%.

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2800 <°C) Temperature

Figure 3. Graphitization dilatometrie curves of pet cokes Pitch Cokes from Coal Process The level of abrasion of the low QI coal for pitch coke cathode is close to the one observed for the shot petroleum coke material. This is in line with the level of the pulverizing factor of the cokes though their different behaviour during graphitization. Figure 4 shows that there is a smaller post baking shrinkage from 1200°C to 1600°C followed by an expansion up to 2100°C where a moderate pregraphitization shrinkage process starts up to 3000°C.

The best indicator among the coke properties for predicting the abrasion level is the pulverizing factor. It integrates the microstrength pattern resultingfromthe macrostructure of the coke or from the peculiarities of the calcination process, and this even though there is an intermediate graphitization step that might change the picture. Concerning this aspect, Figure 3 shows the dilatometrie curves registered during the graphitization of the needle coke, low S anode grade and shot petroleum coke baked artifacts.

The high QI coal tar pitch coke with a poorer graphitizability reflected by the lower xylene density of the graphitized artifact reaches only 2.17 kg/dm3. The abrasion value is correspondingly lower (23 vs. 28%); a fact in line with the lower pulverizing factor of this pitch coke as well as its higher CTE (5.3 vs. 4.6·10"6Κ"1 for the graphitized artifacts).

The puffing tendency of the low S/V anode grade coke is totally suppressed for the shot coke material though the S+N contents are quite comparable. There is massive post baking (>1100°C)

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These results are in line with those reported in [8], the presence of high molecular weight components in the pitch feedstock hindering the formation of mesophase during coking. A harder fine mosaic pitch coke results from this selection of high QI tar prior to the soft pitch preparation process. (%o) Change of artifact» length

>

+10

Despite the high CTE value, the graphitized artifacts show a relative high real density level (2.205 kg/dm3 ) comparable to the shot coke or to the low QI pitch coke materials. This is the sign of a reasonable level of graphitability confirmed by the moderate levels of the mechanical characteristics. However with the intrinsic hardness of this oil shale coke the abrasion rate of the graphitized artifacts (14%) is as low as for the best pitch coke tested here. (% ) Change of artifact« length It

r "

y

Fitch coke f romiti w Q Ï :oal tar

**—S \gm\ baking shrinkage '"IOC

1 1

O i l Shale coke 800

1200

HO

2400 ,

2800

F lifting., Prt 2;r»ohit

***

(°Q

H

2800

im tinnì

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0

+*

4

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I

ima <·€) hfl

fQ

Acetvlene coke

-5

1

Post making shrink:ige

800

1200

1600

+s

4 2000

2800 fC) Temperature

Acetylene coke

Pitcl i coke f r o m i nediur n temijeratu re eoa tar

1600

2400

Figure 5. Graphitization curves of shale oil and acetylene cokes

egraph iti/atio n 1

2000

2400

2800 f C ) Temperature

Figure 4. Graphitization dilatometrie curves of pitch cokes The medium temperature tar pitch coke results in interesting low cathode abrasion levels (-15%). This pitch coke shows the same low level of pulverizing factor as the high QI coal tar pitch coke but a more severe post baking shrinkage, no puffing tendency followed by a pronounced pregraphitization shrinkage from 2100°C to 2700°C. The CTE reaches the highest level (6.3·10"6Κ_1) of the graphitized artifacts along with the highest levels of mechanical characteristics. This is related also to the massive volumetric change of -7% observed on the cold graphitized samples after the graphitization step.

The acetylene coke is a unique material having a low porosity that is mainly to be found in between the small carbon black like agglomerated particles. Its remarkably low level of real density (1.909 kg/dm3) demonstrates its resistance to the heat treatment, which is confirmed also after the graphitization (2.081 kg/dm3). The pitch demand is extremely low (10%) and the level of apparent density unchanged after baking followed by an impressive increase after graphitization results from the massive post baking shrinkage (Figure 5) up to 2500°C. The abrasion value reaches here an unprecedented one digit number of 8% which is six times better than for low S anodegrade coke and still three to four times lower than for shot petroleum coke cathodes. The drawbacks of this material are of course its poor level of electrical and thermal conductivities which is two to three times lower than typical. This is related to the nonplanar structure of the aromatic compounds [10].

Compared to more classical coal tar pitch cokes this material is more isotropie as its content in heteroatom (N,0) is higher due to the lower severity of the thermal treatment producing the tars (700°C in coal gasification process vs. 1100°C in the coke oven producing coal tar by-products).

Other cathode relevant properties The cathode relevant properties for selected cokes are summarized in Table II and compared to the typical value of industrial cathodes. Even though there is a scale up factor between the pilot and the industrial scale it appears that the first three candidates gave results falling mostly in the typical range of graphitized industrial cathodes. This is not the case for the high QI coal tar pitch coke, medium temperature tar pitch coke, oil shale shaft kiln coke,

Oil shale coke The oil shale coke shows the lowest pulverizing factor but S/N levels close to be needle coke material. As shown in Figure 5 there is no puffing tendency but instead a strong post-baking shrinkage up to 1600°C followed by a plateau area up to 3000°C where the residual shrinkage compensates the thermal expansion. On the cold specimen the volumetric change is moderate (~ -4%).

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Table II. Cathode relevant properties: selected pilot plant and industrial cores data Source Petroleum Petroleum Low S High Asph. Resid Rotary Calcination Rotary Feedstock

Abrasion Comp. Strength (CS) CS after Na vapour Rapoport swelling Graph, apparent density Sp. Electrical resistance Thermal conductivity |CTE

% MPa MPa % kg/dm μΩτη W/mK lO^K"1

54 15.9 17.1 0.24 1.541 13.7 87 3.48

28 37.4 24.4 0.42 1.635 10.7 99 5.62

Coal

Oil Shale Acetylene

Coal

Coal

LowQI

HighQI

Rotary

Rotary

Med. T. Tar Rotary

23 42.0 26.2 0.45 1.636 14.7 98 5.33

15 50.0 29.9 0.44 1.663 13.3 98 6.32

28 32.2 31.7 0.34 1.620 13.3 93 4.64

which gives better abrasion values together with compressive strength and CTE values on the high side or above the typical industrial range. The drawback of higher Rapoport swelling, observed here generally for the more isotropie coke cathodes, is not an issue as the swelling is still comparable to the graphitic cathodes that are known to be unproblematic for this aspect, The compressive strengths after the Na vapour test is also on the high side of the typical range, an indication that the cathode materials resist well to the sodium attack. The higher CTE and Rapoport swelling might imply some adaptation on the cell lining geometry/design but the reduced abrasion, reaching the level of graphitic cathode, is a must for the pot life time. Abrasion and Pot Lifetime

-

Shaft

Rotary

14 30.9 30.6 0.36 1.642 14.7 83 5.79

9 40.2 14.4 0.52 1.643 34.7 37 5.92

Vertical cores taken in industrial cathodes Graphitized

Graph. Impregn.

32-64 15-34 15-30 0.1-0.4 1.60-1.68 10-14 80-110 2.2-5.0

30-40 25-40 25-35 0.1-0.3 1.66-1.70 8-10 100-120 2.4-3.2

100 % Graph. 30% Graph. 16-24 20-34 15-28 0.4-0.5 1.62-1.70 18-24 25-40 2.6-3.4

1-4 22-38 10-25 0.5-0.8 1.52-1.60 30-48 7-13 2.8-3.8 |

There is a very promising potential especially in the delayed coking of soft pitches produced out of selected tars or out of soft pitches where the heavy molecular weight components are concentrated for the production of hard coke resulting in abrasion resistant graphite cathodes. This last route is quite attractive as it is complementary with the production of needle coke out of QI free soft pitch for the steel graphite electrodes. In the same direction the growth of the bituminous sand processing for oil recovery can be an interesting option for preparing, by delayed coking of appropriate tars free of sand, a high potential coke for cathodes. The option of blending of relative soft cokes with more exotic cokes, like the acetylene coke or others, can also be attractive for improving the wear resistance of the graphitized cathodes.

Impregnation, made industrially only on relative soft graphitized cathodes produced out of classical low S anode grade coke, has much lower potential of abrasion improvement than the one reached through an appropriate coke selection (Table II). Figure 6 gives the trend line for the average pot lifetime as a function of the abrasion of cathodes, as observed in several smelters [11].

Max. expected pot life (years) 11

Resid

References 1.

W.R. Hale, "improving the Useful Life of Aluminum Industry Cathodes", JOM, November 1989, pp. 20-25.

2.

H.A. 0ye and B.J. Welch, "Cathode Performance: The Influence of Design, Operations, and Operating Conditions", JOM, February 1998, pp. 18-23.

3.

D. Lombard et al., "Aluminum Pechiney Experience with Graphitized Cathode Blocks", Light Metals 1998, pp. 653-658.

4.

J.M. Drey fuss and L. Joncourt, "Erosion Mechanisms in Smelters Equipped with Graphite Blocks", Light Metals 1999, pp. 199-206.

5.

J.M. Dreyfus et al., "Variable Resistivity Cathode Against Erosion", Light Metals 2004, pp. 603-608.

6.

E. Skybakmoen et al., "Laboratory Test Methods for Determining the Cathode Wear Mechanism in Aluminum Cells", Light Metals 2007, pp. 815-820.

7.

H.A. 0ye, "Carbon Cathode Materials Approval and Quality Control Procedures", Journal of Metals February 1995, pp. 14-19.

8.

S. Toda and T. Wakasa, "Improvement of Abrasion Resistance of Graphitized Cathode Block for Aluminum Reduction Cells", Light Metals 2003, pp. 647-653.

9.

R. Perruchoud and W. Fischer "Granular Carbon mechanical Properties" Light Metals 1992, pp. 695-700.

10.

I.C. Lewis et al., "Conversion of petroleum feed stocks to coke", ACSDPC 1988, pp 404-412.

11.

Private communications, 1995-2009.

Days

4 2 1 Abrasion (%) Figure 6. Potlife time vs. cathode abrasion Oil shale and acetylene coke cathode showing abrasion values better than 100% graphite based cathodes (non-graphitized) and close to the level of the 30% graphite amorphous cathodes might improve significantly the potlife time by 1 to 2 years respectively. Conclusion The selection of relatively isotropie cokes allows a substantial improvement of the abrasion of the graphitized pilot artifacts.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

DETERMINATION OF THE EFFECT OF PITCH-IMPREGNATION ON CATHODE EROSION RATE Pretesti Patel1, Yoshinori Sato2, Pascal Lavoie1 The Light Metals Research Centre, University of Auckland, Private Bag 92019, New Zealand 2 SEC Carbon Limited, 3-26 Osadano-cho, Fukuchiyama, Kyoto, Japan 620-0853

lr

Keywords: Graphitized cathodes, pitch impregnation, electrochemical wear Abstract A number of smelters have adopted or have trialed high density, pitch impregnated cathode blocks as one measure to counter the trend in decreasing cell life due to line current increases. To date the true benefits of pitch impregnated cathode blocks are not fully understood and therefore a joint collaboration between SEC CARBON Limited and the Light Metals Research Center has endeavored to understand the effect of pitch impregnation on cathode block performance. The initial results of this project showed that pitch impregnated cathode blocks had no benefit in regards to electrochemical wear resistance, and it was proposed that this was due to the pitch impregnation increasing the reactivity of the cathode material [1]. This paper reports on recent work conducted to firstly characterize the difference between the pitch impregnation phase and other phases present in the bulk cathode matrix and secondly to understand the relative reactivity of these phases under electrolysis conditions.

In industry the correlation between the erosion of graphitized cathode blocks and current is not only evident by lower overall pot life but also by the W-shape wear patterns observed in decommissioned pots as shown in Figure 1. This wear pattern is due to local increases in current density as the current takes the shortest path through the low resistivity blocks to the collector bars. In the areas of high localized current density the rate of aluminium carbide formation is higher which when coupled with higher fresh bath recirculation in the same areas will lead to increased localized erosion rates shown.

Introduction The use of graphitized cathode blocks has become wide spread in the aluminium smelting industry as these low electrical resistivity blocks allow the for higher amperages at lower cathode voltage drops. This can result in significant increases in production while minimizing the impact on energy efficiency. Although the benefits of increased line current (production) have been realized it has come at the detriment of decreasing pot life due to increased erosion rate.

Figure 1: W-Shape wear patterns in modern aluminium cells It should be noted that aluminium carbide formation is not solely limited to the surface and it has been found by Rafμei et al. and Patel et al. that electrochemical aluminium carbide formation can occur within the pores of the cathode structure which can lead to internal weakening and particle detachment which can result in pitting and ultimately cathode failure [3, 4, 8].

The correlation between increases in amperage and increased erosion rate has been studied extensively and it is generally accepted that the electrochemical formation and subsequent dissolution of aluminium carbide is one of the dominant erosion mechanisms in graphitized cathode materials. The electrochemical formation of aluminium carbide has been found to be accelerated with increasing cathode current density the mechanism shown in equation 1 [2-6].

Electrochemical aluminium carbide formation and dissolution is not the sole mechanism contributing to cathode erosion. Physical erosion is also a major contributor and is extremely important in graphitized cathode blocks which are generally softer than other cathode types. Physical erosion occurs due to the movement of sludge, bath and metal over the cathode surface. It is the alumina in the sludge which is the most detrimental to the cathode as it is hard and naturally abrasive [9].

(cathode)

4 A1F3 (
In recent years, efforts in graphitized cathode block development have mainly concentrated on increasing abrasion resistance and reducing internal electrochemical wear mechanisms by increasing cathode density, mainly through efforts to reduce open porosity. This has mainly been done through formulation optimization and/or pitch impregnation.

(2)

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Pitch impregnation has been a major avenue of development for a number of cathode manufacturers around the world. It involves forcing pitch into the pores of the cathode material, with subsequent baking and graphitization to reduce the open porosity. The use of this densification technique does result in improved physical and mechanical properties such as increased density, increased flexural strength and reduced electrical resistivity, however it was reported by Sato et al. that it can adversely affect the wear resistance of the material [1]. Sato et al.'s findings showed that when non-pitch impregnated and pitch impregnated cathode samples were tested in laboratory scale electrochemical erosion experiments the pitch impregnated samples generated higher erosion rates (as shown in Figure 2) [1]. This indicated that pitch impregnated samples have higher susceptibility to electrochemical aluminium carbide formation and dissolution.

Characterization To understand the possible reasons for greater reactivity in the impregnation pitch phase a greater understanding on its structure was required, especially in terms of how different it is from phases found in a non- pitch impregnated cathode material. It was decided that comparisons will only be made between the impregnation pitch material and a standard binder pitch material as these are the phases that will likely have the highest reactivity in a cathode block in terms of aluminium carbide formation (via equation 1). The impregnation pitch and binder pitch selected were standard pitch types produced by a Japanese maker. Table I shows that the properties of the two pitch types are significantly different with the impregnation pitch having lower viscosity, softening point and toluene and quinoline insoluble fractions. For impregnation pitches, viscosity, softening point and very low quinoline insolubles content are highly important as this will determine the ease of pitch penetration and target pore size range reduction for the impregnation process. Table I: Properties pitches.

for

2

[ Paralleli

Softening Point ( C) Fixed Carbon (%) Toluene Insoluble ( % ) Quinoline Insoluble ( % ) Viscosity (mPa.s)

Parallel2 BAverage

Figure 2: Laboratory erosion rates with rotating cathode. 48 hour electrolysis at a cathode current density of 1 A/cm2 with rotation speed of 80 rpm. A and B represent different formulation types

Impregnation Pitch 77 51.5 12 <0.1 20 at 200 QC

and

binder

Standard Binder Pitch 97 57.5 33.5 11.5 800 at 160 SC

Table II: Binder and impregnation pitch samples Binder Pitch Calcinated Binder Pitch Graphitized Impregnation Pitch Calcinated Impregnation Pitch Graphitized

The two main areas investigated and reported in this paper were:

•

impregnation

From these pitch types four samples with varying heat treatments were analyzed using x-ray diffraction (XRD) and environmental scanning electron microscopy (ESEM). The four samples investigated are outlined in Table II.

Although with the pitch impregnated samples a definite increase in electrochemical erosion was found, the mechanism leading to the increased erosion rate was still unclear. It was suggested however that the impregnation pitch used could have a higher reactivity than the bulk cathode matrix and therefore cause increased erosion. With this in mind the work reported here was designed as a continuation of the work reported by Sato et al. [1] to try and understand the mechanisms leading to the increased erosion rate after graphitization of cathode materials with pitch impregnation.

•

the

Characterization of the impregnation pitch material to determine any differences between this and regular binder pitch which could help explain an increase in reactivity.

Sample Forming Method Normal binder pitch, heat treated to 5 1000 C and then crushed Normal binder pitch, heat treated to 2800QC and then crushed Impregnation pitch, heat treated to 1000QC and then crushed Impregnation pitch, heat treated to 28005C and then crushed

Characterization - XRD Results XRD was used to determine the degree of graphitization of the samples shown in Table II. As the degree of graphitization increases the carbon peak measured by XRD will become narrower because the sample is more crystalline.

Determine the relative electrochemical reactivity of the pitch impregnation phase and the standard binder pitch phase under electrochemical testing conditions.

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sample was impregnated there will be a definite difference in the phases present in the bulk. This difference in phase properties could lead to the adverse affects noticed in the pitch impregnated cathode samples previously tested. Samples Development for Environmental Scanning Electorn Microscopy (ESEM) and Electrochemical Wear Analysis

Figure 3: XRD pattern for calcined binder and impregnation pitches

Custom cathode samples were required so comparisons could be directly made between the graphitized impregnation pitch and binder pitch. To achieve this, two samples were developed which contained 60 wt% of either the powdered impregnation pitch phase or binder pitch phase (Sample U and V respectively). Each sample type contained the same aggregate filler coke material type and grain size fractions (40 wt%) thus making the primary difference between each sample the powdered pitch phase and therefore any differences in measured electrochemical wear rate could be attributed to this. Along with samples U and V, a third sample W, was produced which contained powdered coke material for its 60 wt% fraction. This sample represented a material which approaches commercial cathode formulations and was used for comparison. The production steps for all test samples are outlined in Figure 5. Sample U

Sample V

Impregnated Pitch B

Sample W

Binder Pitch

Heat treated to 1000°C

Calcination of Impregnated/Binder Pitch

Figure 4: XRD pattern for graphitized binder and impregnation pitches

Crushing/Screening Weighing and Mixing

From the spectra (Figures 3 and 4) the full width half maximum (FWHM) can be used to determine the peak width, infering the crystallinity (Lc). From the results shown in Figure 3 and Table III the calcined pitches have relatively low and similar crystallinities. This shows mat heat treatment to 1000SC has minimal effect on the graphitization of both samples.

Kneading Forming

Baking

The graphitized sample results (heat treated to 2800SC) in Figure 4 and Table III show that the degree of graphitization increases dramatically for both samples. However, it was found that the graphitized impregnation pitch had the highest Lc value, indicating the greatest degree of graphitization and crystallinity.

Graphitizing

Figure 5: Production steps of test samples for determining electrochemical reactivity of impregnation pitch Table IV shows some of the resulting properties for the test samples U, V and W.

Table III: XRD results for binder and impregnation pitches heat treated to 1000 and 2800QC Material B(rad) Lc(ΐ) 1 Binder Pitch Graphitized Impregnation Pitch Graphitized Binder Pitch Calcinated Impregnation Pitch Calcinated

0.00785 0.00606 0.0737 0.0724

Granulometry Coke grain (40%) + Powder (IP/BP/Coke : 60%) + Binder Pitch

Table IV:

Propert ies of specialized samples Sample U Sample V Bulk density (g/cm3) 1.667 1.567 Electrical resistivity (μΩιη) 13.3 12.9 Young's Modulus (GPa) 4.45 6.47

254.3 360.7 20.5 21.4

The high crystallinity found in the graphitized impregnation pitch is particularly interesting when directly compared with a standard non-impregnated graphitized cathode block sample. The Lc for the standard cathode block was found to be 306 A, which is lower than the graphitized impregnation phase, thus showing that if the

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Sample W 1.599 12.1 6.49

Characterization - ESEM Results One property difference thought to have a significant effect on the final performance of the cathode block was the difference in thermal expansion between the graphitized impregnation phase and the graphitized binder phase. Phases with higher degrees of graphitization such as the graphitized impregnation pitch will have higher thermal expansion coefficients and therefore when heated could expand and widen finer pore networks, potentially increasing the exposed surface area and thus the reactivity of the material. To test this hypothesis, ESEM analysis with in-situ sample heating was conducted to determine if any microscopic structural changes could be observed when the samples U (impregnation pitch rich) and V (binder pitch rich) were heated from room temperature to 800 SC. Samples were heated in the presence of reducing gas atmosphere to eliminate oxidation of the sample at high temperature. Figure 6 and 7 show room temperature ESEM images of samples V and U respectively. The images show a difference in texture between the two samples with Sample U having a more ordered structure as evident by distinct and preferentially orientated graphite platelets. This is a direct reflection of the higher graphitization state of the impregnation pitch material as confirmed by the XRD results.

Figure 7: Sample U, graphitized impregnation pitch rich sample room temperature

Although structural differences were seen between the different materials at room temperature, when the samples were heated slowly to 800 9C, very little to no quantifiable structural change could be observed in either sample. This can be seen when comparing the U sample in Figure 7 (room temperature) and Figure 8 (800 QC) where images appear almost identical, with little difference in pore size and no significant evidence of opening up of graphene layers. Although these results disprove the hypothesis that was being tested, they do allow us to conclude that surface area changes due to thermal expansion are not likely the cause of the increased wear rate observed in the impregnated cathode samples.

Figure 8: Sample U, graphitized impregnation pitch rich sample 800eC Determination of Electrochemical Reactivity of Impregnation Pitch Through controlled laboratory electrolysis experiments the electrochemical wear rate (in cm/year) of cathode materials was measured and compared to determine the relative electrochemical reactivity of samples U, V and W. The laboratory electrolysis configuration (Figure 9) used in this set of experiments was an inverted cell configuration where the cathode sample was centrally suspended in a graphite crucible which acts as the anode for the experiment. This configuration was used as it promotes electrochemical wear because the sample is not protected by an aluminium pad and therefore is always exposed to fresh bath. In all experiments, a sintered alumina cone was used to direct aluminum away from under the cathode sample to the anode where aluminium re-oxidation would take place. The

Figure 6: Sample V, graphitized binder pitch rich sample - Room Temperature

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cone also provided an extra source of alumina so the cell did not require feeding at any point during the experiment. During the entire experiment the cathode was rotated at 80rpm to promote mass transfer of aluminium carbide to the bath. :-v."

^:: 'KM ·::.-;Ú'"

Variable speed motor

Cathode ccnnectof rod

Figure 11: U sample - 60 wt% impregnation pitch powder

Sintered alumina cone

Figure 9: Experimental set up for laboratory electrolysis wear testing with rotating cathode Experimental parameters such as bath composition, current density and experiment time can be found in Table V. Table V: Experimental conditions for laboratory electrolysis of cathode materials. Bath Composition Weight Percent (%) 1 78.7 Cryolite 9 Alumina 7.8 Aluminum fluoride Calcium fluoride 4.5 Aluminium added at start Cathode current density Electrolysis time Cathode immersion depth

26g 1 A/cm2 48 hours 38mm

I '

Figure 12: V Sample - 60 wt% binder pitch powder 1

|

1 1 Ì

• - . . . . . · . -"W:ΙkΚÈ

For each sample type, two samples were tested to ensure reproducibility of results. Figure 10 gives a summary of the erosion rates found after 48 hour laboratory electrolysis.

Ιûπιiπ Figure 13: W Sample - 60 wt% coke powder Results show that sample U produced a significantly higher wear rate (nearly double) under identical conditions than samples V and W. This confirms the theory that the graphitized impregnation pitch phase has a higher electrochemical reactivity than the graphitized binder pitch phase and also the graphitized coke phase of sample W. This in agreement with the results reported in Light

Sample Type |

Paralleli

Parallel2 «Average |

Figure 10: Erosion rates for samples U, V and W

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Metals 2010 where it was stated that pitch impregnation had an adverse effect on the results. Although these wear results confirm the proposed hypothesis that the impregnation phase has a higher reactivity than the binder pitch phase or the coke phase, it however also goes against the general accepted theory that highly ordered graphite structures will have lower reactivity than the less ordered structures. The reason for the higher reactivity in the highly ordered graphitized impregnation phase is still unclear and more experimentation will be carried out to determine the acting mechanism. It is clear however, from the ESEM results that the driving mechanism relates to the chemical reactivity of the material rather than structural aspects such as increased porosity and surface area due to differing physical properties of the pitch impregnation material. More work on the actual mechanism causing the increased reactivity will be carried out and reported at the Light Metals Conference.

4.

Rafiei.P, Hiltmann.F, Hyland.M and Welch.B, Electrolytic Degradation Within Cathode Materials. TMS (The Minerals, Metals and Materials Society), Light Metals 2001, 2001. p. 747 - 752

5.

S0rlie.M and 0ye.H.A, Cathodes in Aluminium Electrolysis. 2nd Edition ed. 1994, Dusseldorf, Germany: Aluminium-Verlag.

6.

Wilkening.S and Reny.P, Erosion Rate Testing of Graphite Cathode Materials. TMS (The Minerals, Metals and Materials Society), Light Metals 2004, 2004: p. 597-602.

7.

0degβrd.R, On the Formation and Dissolution of Aluminium Carbide in Aluminium Cells. Aluminum, 1998.

8.

Rafiei.P, Hiltmann.F, Hyland.M and Welch.B, SubSurface Carbide Formation Contributing to Pitting and Accelerated Cathode Wear, in Al Smelting Conference. 2001. Queenstown: Univeristy of Auckland.

9.

Liao.X and 0ye.H.A, Method for the Determination of Abrasion Resistance of Carbon Cathode Materials at Room Temperature. Carbon, 1996. 34(5): p. 649 - 661.

Conclusions •

Graphitized impregnation pitch was found to have a higher degree of graphitization than graphitized binder pitch and also standard graphitized cathode block material. This indicates that within impregnated cathode blocks a definite phase differential exists.

•

Cathode samples which contained 60 wt% powdered impregnation pitch wore significantly more than the powdered binder pitch sample. This implies that the impregnation pitch has a higher electrochemical reactivity than the other samples tested.

•

The higher electrochemical reactivity of the impregnation pitch phase can not be attributed to increased surface area due to structural changes of this phase during heating. It is therefore concluded the increased reactivity is due to the chemical reactivity of carbon phase itself. References

1.

Sato.Y, Patel.P, and Lavoie. P, Erosion Measurements of High Density Cathode Block Samples Through Laboratory Electrolysis with Rotation. TMS (The Minerals, Metals and Materials Society), Light Metals 2010, 2010. p. 817 - 822

2.

Keller.R, Burgman.J.W, and Sides.PJ, Electrochemical Reactions in the Hall-Heroult Cathode. TMS (The Minerals, Metals and Materials Society), Light Metals 1988, 1988: p. 629-631.

3.

Patel.P, Hiltmann.F, and Hyland.M, Influence of Internal Cathode Structure on Behavior during electrolysis, Part II: Porosity and Wear Mechanisms in Graphitized Cathode Materials, TMS (The Minerals, Metals and Materials Society), Light Metals 2005, 2005. p. 757 - 762

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SIMPLIFYING PROTECTION SYSTEM TO PROLONG CELL LIFE Maryam Mohamed Al Jallaf1, Margaret Hyland2, Barry Welch3, Ali Al Zarouni1 1 Dubai Aluminium Company Limited, PO Box 3627, Dubai, U.A.E department of Chemical & Materials Eng. Univ. of Auckland N.Z. 3 Welbank Consulting Ltd, N.Z., and S. of Ch.Sci. & Eng UNSW 2052 Aust. Keywords: Cathode, TiB2, protection, granule erosion[2]. Hence the effort to increase cathode or cell life is relevant. The end result of increasing production of aluminium and lower cell life is more spent potlining waste and a figure of 13.2 kg/t AI is suggested [3].The low solubility of TiB2 in aluminium makes it a suitable material to protect the cathode [4]. See also Figure 2 .

Abstract Cathode materials mainly deteriorate through physical erosion and chemical / electrochemical corrosion. The wear patterns are characterised by a "W-shape" and localised potholes. A variety of materials such as alpha alumina and refractory hard materials have been used for protecting conventional cathodes. TiB2 based refractory hard materials are well suited for this. They satisfy the electrical conductivity requirements, have a very low solubility in molten aluminium and are wetted by aluminium providing potential benefit in modifying design and operation.

Electrolysis Solid Waste Not Recycled Output -SPL Landfilled

-

rn

gjg

\

r>tt.3:sM<3ev14.3|

I

t

g io

/

So M

This study was to explore the potential of increasing life for graphitised cathodes using TiB2 grains without compromising metal purity, cell performance or lining design. Statistical analysis of measurements confirms that cathode life could be increased by at least 2 years. The study revealed that cathode life could be increased by at least 2 years. Increases in B and Ti impurities seen in pot metal were within casthouse tolerance limits.

_ > r+*~+**

,.-^-^~~~~

Aluminium production, cumulative ( milion tonnes)

Figure 2: SPL generation statistics [3] Deterioration mechanisms for cathode are well known [5,6,7]. The cathode wear in modern electrolytic cells is typically of "W" shape pattern, with troughs near the sidewall [8] due to higher electrochemical wear in zones of increased current density ( Figure 3).

Introduction Smelters are pushing for higher productivity and thus increased current densities and this is associated with a decrease in the life of cells as shown in Figure 1 [1].

Amperage vs. cell life (data f r o m ECGA)

Figure 3: "W" shaped erosion [3] 130

180

300

One of the primary routes for cell failure is through the cathode bottom [5,9]. Use of TiB2 to protect the cathode and thus improve cathode life is by no means a new idea and coatings of TiB2 have been tried in the past [10,11,12]. However, in this paper, we are reporting on a new approach: the addition of granular TiB2 to operating electrolysis cells. This approach was chosen as it was simple and expected to be effective provided a layer of sufficient thickness could be maintained. The impact of the

Amperage, kA

Figure 1 : Amperage vs. cell life - data from [1] At the same time due to demands for higher thermal and electrically conducting cathodes these are being made with higher graphite content with less resistance to mechanical

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methods were used to allow comparison with earlier measurements in operating test cells and with cut out cells of similar design which are normally measured using the cathode centreline as reference.

addition of the TiB2 on metal quality will be dealt in part II of this paper [13]. Trial design and implementation

Erosion measurement in operating test and control cell

Two adjoining cells of the same age and design with graphitised cathodes operated at similar operating parameters were selected as test and control cells for the trial. The addition was limited to only one test cell due to high cost of TiB2 and to minimise potential risk to downstream product quality. The quantity of TiB2 to be added was estimated at 290 kg based on solubility product (Ksp) of TiB2. An initial TiB2 layer of about 18 mm on the cathode was expected assuming uniform distribution and no compaction. A minimum thickness of 4 mm would be expected if the material compacted to the true density.

The procedure assumes no cathode heave is taking place. The measurements were performed during routine anode changes and a complete set of measurements took about 26 days. By studying the erosion rates in the maximum erosion zone, the cathode life for the test and control cells could be predicted. Measurements were made using a special tool, based on earlier design [14], that was developed to suit the cell design used in the plant (Figure AY

14 kg of the TiB2 grains were added during each anode change thus taking one full anode cycle of about 26 days to complete the addition. The analysis of TiB2 grains is shown in Table 1 and Table 2. Samples A and B represent two batches and a 50/50 mixture of these was used.

3—CJL Horizontal 3r

Table 1 : Chemical analysis of TiB2 Element, Weight % Boron Titanium Iron Oxygen Carbon Phosphorous

Sample

Sample

A 30.6 67.6 0.11 1.58 0.26 < 1 ppm

B 31.2 67.2 0.19 0.73 0.41

cathode bottom operating cell

(distance from deckplate reference)

Fifteen measurements were performed per anode location with a total of 300 data points for all 20 anodes as shown in Figure 5.

-

Table 2: Physical properties of TiB2

Bulk Density (tapped), g/cm3 Real Density, g/cm3 Particle Size, % + 4.76 mm + 2.36 mm + 1.18 mm - 1.18mm + 45 urn - 45 urn

Sample A

Sample B

1.05

1.05

4.51

4.51

22.1

56.2 (-6.3+4mm)

(all dimensions in mm)

Figure 4: Cathode erosion measurement tool

For illustrating, the locations are shown for anodes 1 and 2 in Figure 6 these being near the tap end and the corresponding cathode numbers were 19 to 16. For anode 1 and anode 20 opposite to each other, there would hence be three columns of data with each column having 10 data points, 5 of them from anode 1 and five from anode 20. For reporting, the columns were numbered from 1 to 30 starting from anode 1 (right). It is seen from Figure 6 that there may be measurements quite far away from the centreline of the cathode block, where normally maximum erosion is observed in industrial cells. Fortunately there were 10 columns of data (100 data points) that were within 50 mm of the centre line of cathode and hence this concern could be addressed.

48.4 69.6 30.4

____ -_..

Four sets of cathode erosion measurements were performed on the operating test and control cells before and during the trial over a period of about two years. In addition, two sets of measurements were done in two other cells of same design and age in order to ensure that sufficient base data was available in the early stages. After about 2 years the test cell was cut out and a partial autopsy done to inspect condition of cell as well any residual TiB2. Detailed erosion measurements were done in the cut out cell Both anode and cathode centre line referencing

Figure 5: Anode shadow on cathode surface

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The surface contours for the first and fourth set of measurements in the operating test cell for the entire set of 300 data points are shown in Figure 7 and Figure 8. Similar data for 300 data points from the operating control cell are shown in Figure 9 and Figure 10. These graphs are with same orientation as in the figure showing measurement positions, namely that tap end of the cell is to the left and upstream locations are at the top. These surface contours reveal clearly that the cathode is eroding over time for both the test cell and control cell but the control cell erosion is much more than for the test cell during the same time interval.

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Figure 10: Erosion contour - control cell- 4th set Erosion rates estimated from the measurements and presented in Figure 11 show that the addition of TiB2 has resulted in significant reduction in the rate. The life for test cell was estimated using a more conservative wear rate of 26 mm/year to be 3.6 years more than the control cell. Based on the expected life of the cell using this rate and the consumption of TiB 2 estimated from solubility product and other factors, it was estimated that adding 450 kg of TiB2 would result in optimum level of protection. If one were to add this at the beginning of cell operation and assume 26 mm/year erosion, an increase of 5 years can be predicted for cathode life.

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Figure 8: Erosion contour - test cell 4 th set Erosion is seen to be higher in the same zones namely near anodes 7 to 10 corresponding to cathodes 2-7. Thus it is seen that that the procedure developed was adequate even though the accuracy of individual measurements was probably ±1 cm. Typically there is maximum erosion at centre line of each cathode block but since there were 300 points measured in the cell, there were more than 30% of the data points that were close to the cathode centrelines and hence the anode referencing method adopted though not ideal was adequate.

Figure 11 : Erosion rates - before & after TiB2 addition Since there were four measurements, the rate of erosion can be calculated for different time intervals and to demonstrate that the rate was always slower for the test cell, the estimates are described in Figure 12.

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TiB2 with metal. The grains seemed to be mobile and not bound to the cathode carbon. TiB 2 particles added originally were coarse aggregates of more than few mm and tended to break up easily. From this it was concluded that nature of TiB 2 has not changed much. A macro view of the cell showing locations of the grains is depicted in Figure 13.

Figure 12: Erosion rates - for different time intervals Part 2 - Cell autopsy A set of measurements was done just prior to cut out and repeated in the cut out cell for validation. The average and maxima were reasonably close to each other. Based on the comparison it was concluded that operating cell measurements can give an idea of the erosion patterns and rates but localised measurements can occasionally be off by about 10 -20 mm/year through special events such as positioning errors or freeze. This has been over come to some extent by multiple measurements during the trial. It can also miss steep points of erosion which may be covered by freeze and in addition areas such as centre channel are not possible to measure in operating cells. The practice of using the centre line of the anode and 400 mm of either side of the same for measurements can lead to these being done at times very close to the seam. However, the erosion data from operating cell measurements was found to be good enough to draw conclusions, namely that TiB 2 addition is improving the cathode life substantially and to provide an estimate of the rate of erosion and cathode life.

Figure 13: View of test cell after cut out The grains are covering all areas of the cell except the dark areas visible on the photograph. Two of the areas where grains are visible are shown by arrows but the grains are present in most parts of the cell. The clear evidence of residual TiB2 in the cell as seen in autopsy was encouraging.

Erosion measured by standard procedure Standard erosion measurements were carried out at 57 points after cut out at the centre line of the cathode for test cell and another cell of similar design. The maximum erosion zone was in the upstream area and the rate of erosion was lower for test cell by 11 mm/year at the maximum erosion point. The cathode life for the test cell was estimated to be more than 1.79 years longer than the control cell but this increase would have been greater if TiB2 was added from start of operations.

Figure 14: TiB 2 as received

The average erosion rate value for the upstream locations was significantly lower by 20 mm for test cell as compared to cell G by Student's t-test of significance at 99% confidence interval.

Figure 16: TiB2 Grains 1 to 3

Residual TiB 2

Figure 15: TiB2 in test cell

Figure 17: Grain 1- zoomed

Theoretical estimate of TiB 2 consumption

Inspection and analysis showed TiB 2 grains to be present in cut out test cell after 2 years of operation. The TiB2 added did not seem to be bound to the cathode carbon or to enhance the wetting of the cathode surface although the particles were wetted by metal. The grains appeared to be mobile and not uniformly distributed on the surface of cathode. From analysis of this agglomerated grains wetted by metal it was concluded that the material is a mixture of

Based on solubility product (Ksp) of 10000 ppm3 units (1.825 x 10 "18 gram moles3/g3 Al), it is estimated that the initial amount needed to saturate the metal in test cell would be about 0.5 kg. The loss of dissolved TiB 2 due to dilution by produced metal and tapping was estimated to be 0.91 kg/day. Based on these estimates the 290 kg of added TiB 2 was expected to last 3180 days. This did not

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take in to account any losses during tapping of bath / metal due entrainment of grains.

3.

The erosion measurement technique employed in operating cells was acceptable but would benefit from further refinements to align with the central portion of the cathode.

4.

A substantial portion of the TiB 2 added to test cell was still present even after 2 years of operation though a mass balance could not be performed.

5.

The TiB 2 grains were not uniformly distributed on the cathode surface and were seen to be clearly wetted by aluminium. Zones of high erosion did not have higher amounts of TiB2 as originally expected prior to the trial. TiB2 grains with a more optimum particle size distribution that would pack better would be desirable in any future studies so as to minimise movement of the grains.

Patterns of erosion By plotting the average and maximum values measured as a function of distance from the reference clearly shows more erosion upstream and downstream close to the sidewall. The wear generally increased from the central channel towards the sidewalls on each side of the cathode block. Upstream erosion is much greater than downstream (Figure 18). Test cell: Comparison of erosion at different locations Distance from reference, mm 400

600

800 1000 1200

I Distance from reference [anode 1 -10 (upstream end of cell)

1200

100

800

600

400

I Distance from reference | anode 11 -20 (downstream end )

A lesson from these observations is that would to be advisable to use TiB2 with a better particle size distribution that will pack well so that the grains will be less mobile. It may be also worthwhile to add the TiB2 as a slurry prewetted with aluminium which could minimise movement of grains. Initial techno economic calculations were done to evaluate the feasibility of this project. These will be confirmed once the population of the test cells is increased for validating the initial results.

Figure 18: Erosion at different locations Cell performance and process parameters

Acknowledgement

Cell performance data was collected for test and control cells and comparison of data showed several features. However there were no adverse effects found. The test cell had more net bath tapped than the control cell and higher A1F3 consumption for test cell as compared to most of the adjacent cells with similar age indicating that TiB2 may be protecting cathode from sodium penetration. Freeze profile data and other process parameters such as metal velocity, anode current distribution, superheat and shell temperature were also compared for control and test cell and no major differences seen.

The authors would like to thank the management of DUB AL for sponsoring and supporting this project. The success of such a large investigation involved many of DUBAL's competent and resourceful engineers. But we want to particularly acknowledge and thank G. Meintjes for coordinating the many facets, N. Al Jabri for ensuring the potline management enabled the achievement of goals. Also the many specialised supporting skills and analysis contributed by teams led by A. Al Jaziri, M.Tawfik, N. Rana, F.Abdulla, S. K. Howaireb, S. Joseph and Dr.K.Venkatasubramaniam.

Conclusions

References

The following can be concluded based on trials 1.

Autopsy observations and operating cell measurements confirmed that TiB 2 addition in test cell has reduced erosion rate compared to the control cell without such addition. The expected increase in cathode life can be at least two to three years and can be as high as 5 years if the quantity of TiB 2 was optimised at 450 kg and added at the start of cell operations.

2.

The reduction in erosion has not altered the pattern of erosion, namely higher rates of erosion at ends of cathode especially in upstream zone; though even at these locations the wear rate is estimated to be less for the test cell as compared to the control cell. The trends and patterns by measurements both in cut out cell operating cell are similar but measurement in the operating cell understates the erosion.

1. From ECGA website (carbonandgraphite.org) aluminium_production.pdf as on 7-3-10. 2. J .M. Dreyfus and S. Lacroix , "6.07 - Improvement of Erosion Resistance of Graphitised Cathodes", Eighth Australasian Aluminium Smelting Technology Conference And Workshops, Yeppoon, Australia 3 - 8 October 2004, 9 pages. 3. "Life Cycle Assessment Of Aluminium: Inventory Data For The Primary Aluminium Industry- Year 2005 Update LCS update 2005 - September 2007", from International Aluminium Institute as on 6-3-2010. http://www.worldaluminium.org/?pg=/Downloads/Publications/Full%20Publ ication&path=269 4. N.J. Finch, "The Mutual Solubilities Of Titanium And Boron In Pure Aluminum", Met. Trans. By 3(10), 1972, 2709-2711. 5. M. S0rlie and H. A. 0ye, "Cathodes in Aluminium Electrolysis", Aluminium-Verlag, Düssedorf ,1994.

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6. P.Brisson, G.Soucy, M.Fafard, H.Darmstadt, and G. Servant, "Revisiting Sodium And Bath Penetration In The Carbon Lining Of Aluminum Electrolysis Cell", Light Metals 2005, 727-732. 7. P. Rafiei, F. Hiltmann, M. Hyland, B.James, and B. Welch, "Electrolytic Degradation Within Cathode Materials", Light Metals 2001, 747-752. 8. A.Schnittker and H.Nawrocki, "Performance Of Graphitized Carbon Cathode Blocks", Light Metals 2003, 641-645. 9. P. R. Tehrani, "Inter-Relating Aluminium Smelting Carbon Cathodes Formulations To Cell Operations And Wear Mechanism", Ph.D thesis, University of Auckland, 2002), 205 pages. 10. H. 0ye, V. de Nora, J.J.Duruz, and G. Johnston, "Properties Of A Colloidal Alumina-Bonded TiB2 Coating On Cathode Carbon Materials", Light Metals 1997, 279286. 11. CE. Ransley, "Producing Or Refining Aluminium, US patent", 3,028,324,1962. 12. J.T. Keniry, "The Economics Of Inert Anodes And Wettable Cathodes For Aluminum Reduction Cells", JOM, May 2001,43-47. 13. Part II of this paper (to be published). 14. P. Reny and S. Wilkening, "Graphite Cathode Wear Study at Alouette", Light Metals 2000, 399-404.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

ALUMINATE SPINELS AS SIDEWALL LININGS FOR ALUMINUM SMELTERS X. Y. Yan1, R. Z. Mukhlis2, M. A. Rhamdhani2, and G. A. Brooks2 talRO PSE, Box 312, Clayton South, Victoria 3169, Australia 2 Swinburne University of Technology, PO Box 218, Hawthorn, Victoria 3122, Australia Keywords: aluminate spinel, sidewall, corrosion behavior

designs and associated materials stable towards surrounding environments are essential to commercialization of the above technologies.

Abstract Ledge-free sidewall is preferred as it potentially reduces the energy requirement of aluminium production by about 30%. However, this approach poses great material challenges because such sidewalls are in direct contact with oxidizing, corrosive and reducing environments at different cell locations. In this study, NiAl204, MgAl204 and Nio.5Mgo.5Al204, were identified and tested in cryolite-AlF3-Al203-CaF2 electrolyte melts at 980 °C under air and C0 2 . Both specimens and baths were characterized using XRD, XRF and SEM-EDS methods. This study revealed that the NiAl204 and MgAl204 spinels had good corrosion resistance toward the melts under C0 2 , with solubility of Ni from NiAl204 being 0.01wt% and Mg from MgAl204 being 0.07wt%. Dissolution of Ni and Mg into the bath was lowered to 0.007wt% Ni and 0.05wt% Mg through the formation of a Ni0.5Mg0.5Al2O4 solid solution.

Carbon blocks are traditionally used as sidewall linings in the HHC [1]. Recently, Si3N4-bonded SiC refractory has been increasingly utilized as the lining materials since it can (i) extend cell life through improved oxidation and cryolite resistances, (ii) increase cell capacity and (iii) facilitate application of large-sized anodes to the cells, thereby leading to an overall increase in cell productivity [1,8,11-12]. However, Si3N4-bonded SiC refractories have higher manufacturing costs than carbon blocks. There are intrinsic problems associated with the HHC when the carbon and Si3N4-bonded SiC sidewall linings are unprotected by a side ledge. The problematic issues for un-protected carbon blocks mainly include reactions with 0 2 or CO/C02 dissolved in baths, bath penetration, chemical erosion due to A14C3 formation, spalling due to Na intercalation, mechanical abrasion due to undissolved A1203 particles in circulation and easier erosion at metal/bath/carbon three phase boundaries [1]. For un-protected Si3N4-bonded SiC, the problems are dissolution of Si3N4 in baths and chemical reduction of SiC by liquid Al [13, 14]. Current practice to minimize these sidewall attacks is to operate the HHC with the side ledge by ensuring that heat loss through the sidewall is carefully balanced, thereby maintaining stable side ledge formation. Such a heat loss due to this practice accounts for approximately 30-40% of the total energy consumption, which is largely responsible for 40-45% energy efficiency of the current HH process [3].

Introduction Aluminum is commercially produced in the Hall-Hιroult cells (HHC). The HHC consists of a steel shell lined with refractory insulating bricks covered with sidewall lining materials, a pool of molten Al on carbon blocks as the cathode at the bottom of the cell, and a carbon block immersed in the electrolyte from the top of the cell as the anode. Molten Al is produced at the cathode while the carbon anode is electrochemically oxidised and, thus, consumed during electrolysis to produce 70-90% C0 2 , with the rest being CO [1]. The overall cell reaction can be expressed as: 2 A1203 (diss) + 3 C (s) = 4 Al (1) + 3 C0 2 (g)

(1)

The downside of the side ledge is that there are occasions where the sidewall may be exposed to a molten bath in the HHC. For example, during cell start-up, the molten bath will be in direct contact with the sidewall lining for a period of time before the formation of a protective side ledge. In other occasions, the protective side ledge may melt away during anode effects or cell instability. In both instances the sidewall is directly exposed to the bath, causing potential degradation and/or sidewall failure. An alternative approach of tackling these difficulties is to develop a sidewall that does not require a side ledge to protect such attacks, i.e., a ledge-free sidewall where the lining is directly exposed to the surroundings during cell operations. With such a ledge-free sidewall, the heat within the HHC needs to be kept rather than removed. This leads to a potentially 30-40% energy savings that would otherwise be lost through the sidewalls with a side ledge. Furthermore, the cell capacity and productivity may be most increased with a ledge-free sidewall cell configuration. Despite these obvious benefits, few efforts have been so far made to develop the sidewall materials that enable the HHC to operate in aggressive ledge-free cell operating environments.

The reversible cell voltage is 1.19 V at a typical electrolysis temperature of 960 °C. Modern HH standard cells today operate at up to 380-400 kA, with the first commercial AP50 line being built (the trial cells require 13.25 kWh/kg Al at 500 kA amperage and 95.7% current efficiency) [2]. In practice, cell sidewalls are covered by a layer of a solidified electrolyte normally termed side ledge or side freeze which protects the side linings from a corrosive cryolite-based electrolyte melt (bath) and a highly reducing molten Al. Either service life of the sidewall materials or the carbon cathode may be the limiting factor in overall cell life, depending on the individual materials employed and operating practice applied. There have been significant advances in enhancing energy efficiency and productivity and reducing greenhouse gas emissions over the past 60 years [3-9]. Noticeably, emerging technologies being developed for the HH process would lead to changes in heat balance within the cell, with the amount of heat dissipated through the sidewalls likely to be lowered [3, 10]. Hence, novel sidewall

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Ultimate Materials Challenge and Opportunity

Thermodynamic Analysis

One of the primary obstacles to the ledge-free sidewall approach is a lack of suitably refractory sidewall materials that are able to withstand all different surrounding environments [1, 3, 15]. Thus, the approach poses substantial materials challenge. On the other hand, the discovery of material that could serve as a ledge-free sidewall, especially if combined with inert anodes and a wettable cathode, would enable major changes in cell design and operating practice. This would improve energy utilization and reduce environmental impacts. Welch and May [15] and Pawlek [16] have summarized general selection criteria for an ideal sidewall material after considering physical, chemical, and economical requirements, as given in Table I.

Equilibrium calculations on the systems MgAl204, NiAl204, and NiFe204 with molten aluminium and electrolyte bath were carried out using FactSage 6.1 thermodynamic package using FTlite, FThall and FToxid solution models and databases [22]. The liquid solution phase in FTlite (light metal database) uses modified quasichemical model taking into account short range ordering tendency of atoms or molecules in liquid solutions. The solid solutions of spinel and monoxide in FToxid (oxide database) takes into account the mixing of various cations on crystallographically different sublattices. The cryolite bath model in the FThall assumes a non-random mixing of elemental cations and anions on their respective sublattices [22]. All calculations were carried out for the temperature range of 900 to 1010 °C using the parameters used in the experimental part (e.g. the amount and composition of spinel, aluminum and electrolyte bath).

Table I. General selection criteria of sidewall refractories for use in Al electrolysis cells [15,16]. High electrical resistivity Physical High thermal conductivity 1 Superior abrasion resistance to sludge 1 High mechanical strength 1 Not reactive to cryolite 1 Chemical Not reactive to molten Al and Na 1 Insoluble in molten cryolite and Al 1 Not oxidised by air 1 Impervious or low porosity 1 Not wetted by bath or metal 1 Low fabrication costs, and 1 Economical Ease of joining To date, no single material has been identified that can satisfy all the chemical requirements in Table I. The efforts have focused on carbon based and non-oxide ceramic based materials, whereas the use of oxide based ceramics as ledge-free sidewall materials has not been fully explored largely because most oxides exhibit unacceptable high solubilities in HH baths [16-18]. There are, however, certain oxides or complex oxides that can provide acceptably high chemical stability towards the baths, typically those currently under investigation for the development of cermet inert anodes in the HHC [6, 19]. The recent study further shows that NiFe204 exhibited a limited solubility in molten cryoliteAlF3-CaF2-Al203 electrolytes at 1000 °C, especially when the melts contained high contents of AI2O3 and atmospheres of the cell had high oxygen potentials, and it could be readily prepared from cheap NiO and Fe203 thus resulting in a potential cost reduction in refractory fabrications compared to those of Si3N4bonded SiC [20]. Hence, opportunities exist for oxide based ceramics as sidewall materials in the HHC.

Figure 1 show the predicted equilibrium solute concentrations of Mg and Ni in molten aluminum when 6 g of each spinel (MgAl204, NiAl204, and NiFe204) is reacted with 224 g aluminum. In the case of MgAl204, the predicted Mg equilibrium concentration in aluminum increased from 0.057 to 0.094 wt% as the temperature was increased. In the case of NiAl204 and NiFe204, the nickel concentration in the aluminum was predicted to be constant, as all nickel dissolved in aluminum at all temperature studied. It should be noted that for the same sample mass, the amount of nickel dissolved from NiFe204 sample was lower than that of from NiAl204 sample. This is because, for the same amount sample mass, NiFe204 brings less Ni into Al melt than NiAl204, if Ni is completely dissolved into Al melt. These results indicate that in molten aluminum; MgAl204 is more stable compared to NiAl204 and NiFe204. Worth to be mentioned that the the typical concentration ranges of impurity elements in primary aluminium metal is 1-80 ppm Ni, 5-60 ppm Mg and 4003000ppmFe[23].

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In the present study, we propose a series of aluminate spinels, including NiAl204, MgAl204 and Ni^Mg(1.^Al204 solid solutions, as candidate materials for use as the sidewall linings in the HHC and present a set of preliminary experimental results to demonstrate their chemical stabilities towards molten cryoliteAlF3-CaF2-Al203 electrolytes at 980 °C Provided that the selected spinels also fulfill the other requirements in Table I except for thermal conductivity, the application of the spinel-type sidewall materials to the HHC may potentially lead to a novel design of the sidewall where no side ledge is required.

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Figure 1. Predicted equilibrium concentration of Mg and Ni in aluminum when 6 g of each spinel (MgAl204, NiAl204 and NiFe204) is reacted with 224 g liquid aluminum. In the case of equilibrium between MgAl204 and electrolyte bath system, MgAl204 was predicted to be unstable in the given temperature range, as all spinel was dissolved in the cryolite. Figure 2(a) shows the predicted equilibrium Mg content and

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concentration in the electrolyte bath. All Mg (-1.03 g) in the spinel was predicted to be dissolved in the bath. The change in Mg concentration in the bath shown in Figure 2(a) was due to the change in the amount of liquid bath due to the dissolutions of other solid phases such as Na3AlF6(s) and Al203(s). This can be seen clearly in Figure 2(b) that shows the equilibrium amount of various phases. At 900 °C, more than half of the cryolite (about 169.6 g) was in the form of solid which decreased as temperature was increased. The cryolite was completely dissolved at 950 °C.

particle sizes less than 10 μπι. The A1F3 and CaF2 used were analytical reagents. The as-purchased synthetic cryolite powder contained at least 96.5% Na3AlF6 and approximately 1% CaF2 and 2% A1203 with the rest being Na5Al3F14, as determined using standard X-ray fluorescence spectroscopy (XRF) and X-ray diffraction (XRD) techniques. To fabricate pure oxide spinels, the oxide precursors were weighed at stoichiometric ratios corresponding to individual oxide spinels. The weighed mixture was ball milled in acetone for 24 h and dried in air. Approximately 14g of the thoroughly mixed oxides were cold-pressed in hard steel dies into a rectangular bar under a uniaxial pressure of 120 MPa. Green bars were then sintered in air at 1500 °C for 3-72 h. To prepare a spinel solid solution, the prepared pure oxide spinels were crushed and ground for 10 min into powders. The two powders were weighed at a desired mass ratio and mixed completely. The well mixed pure spinel powders were pressed into a green rectangular bar under 120 MPa. The green bar was sintered in air at 1500 °C for 72 h. The dimension of each sintered bar was about 45 mm long, 12 mm wide and 6 mm thick. The methods of immersion-type and stirred finger tests were utilized in this work. In an immersion-type test, one test specimen was fully immersed in a molten bath at 980 °C for a given period, during which the bath was sampled at various intervals. The samples taken were analyzed for elemental concentrations. Details of the stirred finger tests have been described previously [20]. In the stirred finger test method, one end of the sintered bar was assembled into a stainless steel holder and an outer surface of the holder was covered with an alumina tube. Space gaps between the holder and the tube were then filled with alumina cement to ensure that the complete holder surface was not directly exposed to molten bath during testing. The bar to be tested was immersed in the molten bath and rotated at a pre-determined rotation speed for a given period of time at 980 °C. The rotation speed was chosen so that the concentrations of dissolved elements from the bar were independent of distances from the sample surface, i.e., the mass transfer was not a rate-determining step in controlling the overall dissolution rate. Under such conditions, the concentrations of the dissolved elements from the bar in the bath are regarded as the solubility of the material in the bath at 980 °C.

Temperature (°C)

Figure 2. (a) Predicted equilibrium amount and concentration of Mg in electrolyte bath; (b) equilibrium amount of phases when 6g of MgAl204 is reacted with 200g cryolite and 22g A1203. The current liquid bath solution models in FThall in FactSage do not take into account nickel. For this reason, in the case of NiAl204 and NiFe204 in electrolyte bath, the Ni concentration in the bath was evaluated using the nickel solubility equation reported by Lorentsen [24]: wi%M 2+ =-0.0755 + p ^ M ]

(2)

After testing, solidified baths were crushed and ground into powders and tested specimens were mounted and sectioned. Both were then examined to determine the extents of reaction. Phases of the specimens and bath samples were identified by XRD. Microstructures and elemental concentrations of the specimens were examined using the scanning electron microscopy (SEM) with an energy dispersive spectrum (EDS). Elemental concentrations of the bath samples were analyzed by XRF.

where x is the amount of alumina in wt% (from 0.9 to 13.1 wt%). Using this equation, the solubility of nickel in the bath, containing 10 wt% alumina, was calculated to be 0.013 wt%. For 6g of NiAl204 and NiFe204 immersed in 22 g alumina and 200 g cryolite, the total amount of nickel in the system was calculated to be 0.0087 and 0.0066 wt%, respectively. These amounts are smaller than the solubility limit. Therefore (from thermodynamic point of view), it may be that all nickel from the NiAl204 and NiFe204 dissolve into the bath. It should be emphasized, however, that the above analyses consider thermodynamic factor only but does not consider critical kinetic factors and thus provide the theoretical limits.

Results and Discussion The aluminate spinels tested were NiAl204, MgAl204 and Ni0.5Mgo.5Al204, with Ni and Mg ferrite spinels also tested for comparison. The experimental conditions used for preparation and bath tests of the spinel specimens are given in Table II.

Experimental The NiO powder used had a mean particle size of 10 μπι, with 7677% Ni. The MgO powder was 98-100% pure on ignited material. The Fe203 powder was at least 99% pure, with mean particle sizes less than 5 μπι. The A1203 powder was 99.7% pure and had

Characterization of prepared test specimens The X-ray diffraction experiments have been performed on a Bruker AXS D8 diffractometer using the Cu Κα radiation

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(λ = 0.15406 nm). The scattering intensities were measured over an angular range of 5° < 2È < 80° with a step size Α(2È) of 0.02° and a scan speed of 0.87min. The result of XRD performed on the sintered oxide precursors powder confirmed that only spinel phase exist in the powder as shown in Figure 3. No other phases were evident from the diffraction pattern. The lattice parameter of the NiAl204, MgAl204 and Nio.5Mgo.5Al204 in Figure 3 was calculated to be 8.048, 8.083 and 8.0836 A, respectively. 3 = Sintered M g O + A l ^ powder b = Sintered NiO+Af 0 3 powder c = Sintered UiN1Oi+M$M2Ol

«—

p

powd«

samples as the relative at% ratio of metallic elements in the samples are very close to the desired stoichiometry ratio of elements of the target spinels. Chemical interactions The interactions between aluminate or ferrite spinels and molten cryolite-AlF3-CaF2-Al203 electrolytes may involve: (i) chemical reactions between component oxides from the spinels and the electrolytes and (ii) dissolution reactions. At A1203 concentrations greater than 5 wt%, the dissolution reactions are mainly expressed as follows [20]: 3 MA1204 (s) + 2 (A1F3) = 3 (MF2) + 4 (A1203)

* = MgAl î 0 1 o = NtALO,

(2)

3 MFe204 (s) + 8 (A1F3) = 3 (MF2) + 6 (FeF3) + 4 (A1203) (3) where M = Ni or Mg. Reactions 2 and 3 imply that Ni and Mg from the spinels dissolve into the electrolyte melts to form respective metal fluorides. Fe from the ferrite spinels also dissolves into the melts. Reactions 2 and 3 can be shifted towards the left by increasing activities of A1203 dissolved in the melts at given temperatures. Dissolution behaviors of the prepared spinels were investigated experimentally under the conditions in Table II.

JUL

Table II shows the measured concentrations of elemental Al, Ni, Mg, Fe and Ca dissolved in the Al203-saturated cryolite-AlF3CaF2-Al203 melts after bath tests of NiAl204, MgAl204, Ni0.5Mgo.5Al204, and NiFe204 under C0 2 at 980 °C. It was found that Ni concentrations in the melt in contact with the NiAl204 specimen stabilized at 0.01 wt% Ni after 30-hours immersion test. This value was much lower than measured for the NiFe204 spinel (0.045 wt% Ni). However, it was slightly higher than the concentration of Ni (0.0027-0.004 wt% Ni) in an Al203-saturated cryolite-A1203 melt in equilibrium with solid NiAl204 and A1203 at 980 °C measured by Jentoftsen et al [25]. This observed higher Ni content can largely ascribed to the presence of CaF2 in the present bath. The XRF results indicated that the NiAl204 was chemically more stable than NiFe204 in terms of dissolution of Ni from the spinels in the melts at 980 °C.

2 È (degree)

Figure 3. XRD pattern of sintered oxide precursor powders

50 ìçé

50ìçé A

O

• = Mgi A = AI ! o=0 i * = Aui

A-Ni ] * =Mgj * = AI ! 0=0 | * = Au

!

Mi ' 1 ' 1 '

L· 8

I'M'f ■ | ■ I ■1 ■*!

10 12 14 0

' 1 ■ i 8 10 12 14 0 Energy (keV)

Figure 4. SEM micrographs and the associated EDS result of as prepared (a) NiAl204, (b) MgAl204, and (c) Ni0.5Mg0.5Al2O4 samples The SEM micrographs in Figure 4 show that the bar samples contain significant porosity. Large and interconnected pores may accommodate bath to penetrate deep into the sample. On the other hand, very dense sample is also not favorable as it tends to easily broken during a thermal cycle. It should be noted that microstructure optimization was not performed in present work. Currently, the spinel microstructure optimization is being carried out by PYROmetallurgical Group at University of Wollongong. According to the energy dispersive spectroscopy, the content of Al and Ni in NiAl204 sample is 26.02 and 12.80 at%, respectively. MgAl204 contains 11.86 at% Mg and 23.40 at% Al. Ni0.5Mgo.5Al204 sample contains 5.95 at% Mg, 6.23 at% Ni and 25.39 at% Al. Associated with the XRD results; the EDS results also appear to indicate that spinel structure has been formed in the

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It was also found that the concentration of Mg from the MgAl204 was 0.07 wt% Mg after 151.12-hours immersion test in the melt in contact with the MgAl204, which was seven times higher than that of Ni from the NiAl204 under the same test conditions but was at least five times lower than that of Mg from the MgFe204 tested at the bath ratio of 1.5 under air. It was interesting to note that the concentrations of Ni and Mg from the Ni0.5Mgo.5Al204 solid solution in the bath were lowered to 0.007 wt% Ni and 0.05 wt% Mg due to reduced activities of NiO and MgO in the spinel solid solution. It was further noted that the Ni content of the bath in contact with the Ni0.5Mgo.5Al204 was more than six times lower than in contact with the pure NiFe204 whilst the content of Mg in the bath in contact with the spinel solid solution had the same order of magnitude as that of Ni from the NiFe204. Given the higher thermodynamic stability of MgO than NiO, the dissolved Mg may be less likely to be co-deposited with Al than the dissolved Ni. Furthermore, the Ni0.5Mg0.5Al2O4 solid solution is also more difficult to reduce by liquid Al compared to NiFe204 due to the reduced NiO activity in the spinel solid solution. Hence, (Ni,Mg)Al204 solid solutions could be potentially promising materials for the ledge-free HHC sidewalls when considering chemical resistances towards the baths and liquid Al.

Table II. The experimental conditions used for studying chemical and dissolution reactions in the spinels/bath system at 980 °C Corrosion Testing of Specimens Run Specimen Fabrication of No. Specimens Bath Rotation Sintering Sintering Time Atmosphere Electrolyte temp. (°C) ratio speed (rpm) time (h) (wt%) (h) 0 1.38 0 1 As prepared melt of n.a. n.a. 82%cryolite/10%Al2O3/ 2 3%AlF3/5%CaF2 90%cryolite/10%Al2O3 90%cryolite/10%Al2O3 1.50 24.00 2 1500 3 25 Air *MgFe 2 0 4 82%cryolite/10%Al2C>3/ 1.38 24.00 24 1450 3 3 *NiFe 2 0 4 [20] 2 3%AlF3/5%CaF2 82%cryolite/10%Al2C>3/ 1.38 24.00 4 1500 6 0 C02 NiAl 2 0 4 3%AlF3/5%CaF2 30.00 96.00 1500 72 1.38 6.08 0 5 82%cryolite/10%Al2O3/ MgAl 2 0 4 2 3%AlF3/5%CaF2 24.25 31.25 48.38 55.38 127.05 151.12 72 1.38 5.00 6 Nio.5Mgo.5Al204 1500 0 82%cryolite/10%Al2O3/ 2 3%AlF3/5%CaF2 23.08 30.08 47.22 54.08 Notes: (a) The well mixed oxides were uniaxially cold-pressed in hard steel dies under 120 MPa into green rectangular bars. (b) * stands for the specimens that were dynamically tested in the stirred finger test apparatus. All others were statically tested in the cups containing the baths.

co

co

co

co

Run No. 1 2 3 4

5

6

Notes:

Table III. XRF analysis of some solidified baths after the corrosion tests at 980 °C Sample Bath Ratio/ Elemental Concentration (wt %) Time Atmosphere Al Ni Fe (h) MR <0.004 As prepared melt of 0 1.5/air 18.6 <0.01 0.056 90%cryolite/10%Al2O3 <0.004 *MgFe 2 0 4 24 1.5/air 0.38 18.1 1.57 *NiFe2O4[20] 24 1.38/C02 18.1 0.045 0.05 0.16 24 1.38/C02 0.02 NiAl 2 0 4 17.8 0.015 0.02 17.4 30 0.02 0.02 0.01 0.02 96 16.7 0.01 0.02 16.4 <0.004 MgAl 2 0 4 <0.01 6.08 1.38/C02 0.01 16.2 <0.004 <0.01 24.25 0.009 16.2 <0.004 31.25 0.01 0.014 <0.004 16.0 0.01 48.38 0.008 <0.004 55.38 16.0 0.01 0.01 15.6 <0.004 127.05 0.05 0.016 151.12 15.4 <0.004 0.07 0.009 5.00 1.38/C02 0.027 0.02 0.02 (Nio.5, Mg0.5)Al2O4 16.9 0.004 0.04 0.04 23.08 16.6 30.08 16.5 0.007 0.05 0.01 47.22 0.007 16.5 0.05 0.01 16.4 <0.004 54.08 0.05 0.01 * stands for the specimens that were dynamically tested in the stirred finger test apparatus. All others were statically cups containing the baths.

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Ca 0.52 0.48 3.05 2.63 2.66 2.79 2.72 2.78 2.77 2.80 2.82 2.87 2.88 2.68 2.69 2.67 2.73 2.74 tested in the

Conclusions In this study, aluminate spinels, i.e. NiAl204, MgAl204 and Ni0.5Mgo.5Al204, have been tested in cryolite-AlF3-Al203-CaF2 electrolyte melts at 980 °C under air and C0 2 . The results indicated that both NiAl204 and MgAl204 had good corrosion resistances towards the melts under C0 2 , with solubility of Ni from NiAl204 being 0.01wt% and Mg from MgAl204 being 0.07wt%. It was also shown in this study that the dissolutions of Ni and Mg could be reduced by employing the (Ni,Mg)Al204 solid solution (with Ni:Mg molar ratio of 1:1). It was suggested that this was due to reduced activities of NiO and MgO in the spinel solution. Further investigations need to be carried out for the detailed mechanisms of the dissolution of these spinels in electrolyte baths. Acknowledgements The authors thank CSIRO Light Metal Flagship and Swinburne University of Technology for the provision of funding. References 1.

K. Grjotheim and BJ. Welch, Aluminum Smelter Technology, Aluminium-Verlag GmbH, Dusseldorf, 1980. 2. B. Benkahla, O. Martin, and T. Tomasino, "AP50 performances and new development", Light Metals 2009, G. Bearne, ed., TMS, Warrendale, PA, (2009), pp. 365-370. 3. W.T. Choate and J.A.S. Green, "U.S. energy requirements for aluminum production: Historical perspective, theoretical limits and new opportunities", prepared under Contract to BCS, Incorporated, Columbia, MD., U.S. Dept. Energy, Energy Efficiency and Renewable Energy, Washington, D.C., 2003. 4. A.T. Tabereaux, "Reviewing advances in cathode refractory technology", JOM, 11 (1992), pp. 20-26. 5. G. Brooks, M. Cooksey, G. Well wood, and G. Goodes, "Challenges in light metals production", Mineral Processing & Extractive Metallurgy, TIMM C, 116 (2007), pp. 25-33. 6. I. Galasiu, R. Galasiu, and J. Thonstad, Inert anodes for aluminium electrolysis, 1st ed., Aluminium-Verlag GmbH, Dusseldorf, 2009. 7. R.P. Pawlek, "Wettable cathodes: An update", Light Metals 2010, J.A. Johnson, ed., TMS, Warrendale, PA, (2010), pp. 377-382. 8. A.T. Tabereaux, "Silicon carbide bricks in aluminum reduction cell cathodes", Light Metals Proc. And Appi., Canadian Inst. of Mining, (1993), pp. 149-163, 9. J. Keniry, "The economics of inert anodes and wettable cathodes for aluminum reduction cells", JOM, 53 (2001), pp. 43-47. 10. R. Mukhlis, M.A. Rhamdhani, and G. Brooks, "Sidewall materials for Hall-Heroult process", Light Metals 2010, J.A. Johnson, ed., TMS, Warrendale, PA, (2010), pp. 883-888. 11. R.P. Pawlek, "SiC in electrolysis pots: An update", Light Metals 2006, T.J. Galloway, ed., TMS, Warrendale, PA, (2006), pp. 655-658. 12. J. Schoennabl, E. Jorge, O. Marguin, S.M. Kubiak, and P. Temme, "Optimization of SÌ3N4. bonded SiC refractories for aluminum reduction cells", Light Metals 2001, J. L. Anjier, ed., TMS, Warrendale, PA, (2001), pp. 251-255.

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13. E. Hagen, M.A. Einarsrud, and T. Grande, "Chemical stability of ceramic sidelinings in Hall-Heroult cells", Light Metals 2001, J.L. Anjier, ed., TMS, Warrendale, PA, (2001), pp. 257-263. 14. M.V. Swain and E.R. Segnit, "Reaction between cryolite and silicon carbide refractories", J. Aust. Ceram. Soc, 20 (1984), pp. 9-12. 15. B.J. Welch and A.E. May, "Materials problem in HallHeroult cells", 8th Int. Leichmetalltag., Leoben-Vienna, (1987). pp. 120-125. 16. R.P. Pawlek, "Methods to test refractories against bath attack in aluminum electrolysis pots", Aluminium, 70 (1994), pp. 555-559. 17. B.L. Gao, Z.W. Wang, and Z.X. Qiu, "Corrosion tests and electrical resistivity measurement of SiC-Si3N4 refractory materials", Light Metals 2004, A.T. Tabereaux, ed., TMS, Warrendale, PA, (2004), pp. 419-424. 18. J.G. Zhao, Z.P. Zhang, W.W. Wang, and G.H. Liu, "Test method for resistance of SiC material to cryolite", Light Metals 2006, T.J. Galloway, ed., TMS, Warrendale, PA, (2006), pp. 663-666. 19. D.R. Sadoway, "Inert anodes for the Hall-Heroult: the ultimate materials challenge", JOM, 53 (2001), pp. 34-35. 20. X.Y. Yan, M.I. Pownceby, and G. Brooks, "Corrosion behavior of nickel ferrite-based ceramics for aluminum electrolysis cells, Light Metals 2007, M. Sorlie, ed., TMS, Warrendale, PA, (2007), pp. 909-913. 21. T.E. Jentoftsen, O.-A. Lorentsen, E.W. Dewing, G.M. Haarberg, and J. Thonstad, "Solubility of iron and Nickel oxides in cryolite-alumina melts", Light Metals 2001, J.L. Anjier, ed., TMS, Warrendale, PA, (2001), pp. 455-461. 22. C.W. Bale et al., FactSage thermochemical software and databases, CALPHAD, 2002,26: p.189-228. 23. J. Granfield and J. A. Taylor, "The impact of rising Ni and V impurity levels in smelter grade aluminium and potential control strategies", Materials Science Forum 630 (2010), pp. 129-136. 24. O-A. Lorentsen, "Behaviour of Nickel, Iron and Copper by application of inert anodes in aluminium production", Doctoral Thesis, Department of Materials Technology and Electrochemistry, Norwegian University of Science and Technology, Trondheim, 2000. 25. T.E. Jentoftsen, O.-A. Lorentsen, E.W. Dewing, G.M. Haarberg, and J. Thonstad, "Solubility of some transition metal oxides in cryolite-alumina melts: Part I. Solubility of FeO, FeAl204, NiO, and NiAl204", Metall. Mater. Trans. B, 33 (2002), pp. 901-908.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

A NEW RAMMING PASTE WITH IMPROVED POTLINING WORKING CONDITIONS Bιnιdicte Allard1, Rιgis Paulus1, Gιrard Billat2 barbone Savoie, 30 rue Louis Jouvet, Vιnissieux, 69200, France 2 Carbone Savoie, La Lιchθre, Aigueblanche, 73260, France Keywords: Ramming paste, Aluminium electrolysis pots, Hygiene, Environment, Characteristics Abstract The ramming paste used in the aluminium electrolysis pots has long been a concern regarding health, safety and environment. While the situation has improved over the last 20 years, pastes on the market today still contain either carcinogenic products as PAH components or phenol, or other hazardous components. Health regulations in many regions require substitution of these hazardous components as soon as an acceptable alternative becomes available. A new paste, NeO2, has been developed which contains no hazardous component, according to the present regulations. Standard physico-chemical properties of the paste have been studied, with some specific characterizations which will be detailed. This paper will provide a summary of paste developments and provide the latest results achieved with the NeO2 paste. Introduction Several tons of ramming paste are used per pot, and since the beginning of the aluminium electrolysis process, people have been concerned by the emissions released by the ramming pastes mainly during the densification operation. These emissions come from the binder of the ramming pastes, which is based on coal tar pitch. As for anode and cathode businesses which also use coal tar pitch binder, the concern is that coal tar pitch releases polycyclic aromatic hydrocarbons (PAH), a number of which are considered as carcinogenic. Ramming paste is the most hazardous product to which most of the people are exposed during the pot building, as the other products have been either heat-treated, or they are not classified as carcinogenic (cements and mortars for example). In this paper the historical evolution in the ramming pastes formulation in order to decrease the level of carcinogenic products will be described. The criteria to qualify the potential carcinogenic effects will be also presented. A new ramming paste will then be presented which does not contain any hazardous component according to current regulations. Its main characteristics, standard physico-chemical ones, but also specific ones will be described. Finally the expected impact on the working conditions during the application in pots will be given.

but also they were exposed to a strong release of emissions during the densification. Therefore they were obliged to be wellprotected by personal protective equipment. In the 1990's strong improvements were made on the working conditions: the typical coal tar pitch binder has been modified to decrease its viscosity and allow the densification at much colder temperatures, typically 40°C. That was the development of what was called tepid ramming pastes. The impact on the PAH evolved during ramming operations was quite impressive [1]. Cold ramming pastes, being densified at room temperature (typically 20°C), went further in the easiness of densification conditions. Among the PAH, which are defined as organic compounds with more than 2 aromatic cycles, about 50 compounds can be present in coal tar pitch, which are considered as more or less carcinogenic. Benzo[a]pyren (BaP), which is classified by the European Union as carcinogenic Category 2 (probably carcinogenic for human) is generally considered as a tracer. The US Environmental Protection Agency (EPA) developed criteria, called the equivalent BaP, the most commonly used being the one proposed in 1994, which takes into account the content of the most hazardous components and their relative weight in the hazard, with BaP being the reference. Table 1 gives the PAH used to calculate the equivalent BaP. Since that time, medical information on the effects of PAH has improved, and a new factor has been developed by INERIS in 2003 [2]: the Equivalent Toxicology Factor ETF, whose calculation is also detailed in Table I. Table I: Criteria with the relative toxicity factors of the most hazardous PAH according to EPA and INERIS. Naphtalene Fluorene Phenantrene Anthracene Fluoranthene Pyrene Benzo(a)anthracene* Chrysene * Benzo(b)fluoranthene* Benzo(k)fluoranthene * Benzo(a)pyrene* Benzo(g,h,i)perylene Dibenzo(a,h)anthracene* Indeno-pyrene

Review of the developments improving working conditions during ramming In the 1970' s, everybody was producing and using hot ramming pastes, based on the same coal tar pitch binder that was used in the anode or cathode manufacture. As a consequence, the working conditions during pot ramming were difficult: the ramming paste, but also the cathode blocks and the sidewalls had to be preheated at 120-140°C, so the paste could be easily densified. People not only worked in a hot atmosphere (and walked on a hot surface),

EquivBaP (EPA 1994)

0.034 0.033 0.26 0.1 0.01 1 1 1.4 0.1

ETF (INERIS) 0.001 0.001 0.01 0.001 0.001 0.1 0.01 0.1 0.1 1 0.01 1 0.1

The most hazardous PAH are generally the heaviest ones, which are released as particulates. The lighter ones are mostly released as gas, like naphthalene. Among the 13 PAH considered by INERIS to quantify the toxicity risk, 6 of them (indicated by *) are particulates classified as carcinogenic by the European Union.

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The improvements in the ramming paste composition can be clearly illustrated by the decrease of the equivalent BaP or ETF. On the binder alone, when considering hot paste and tepid paste, the equivalent BaP went from roughly 29300 ppm to 21600 ppm, decreasing by 26% [3]. After the move from hot to tepid and cold pastes, there have been some developments in the mid 1990's to substitute resin binders for coal tar binders. The best resin candidate to replace coal tar pitch is phenolic resin [4]. But phenol is also a hazardous component which is now classified as carcinogenic. It does not allow the removal of all the carcinogenic compounds from the paste composition. Moreover the final characteristics of the ramming paste show that the shrinkage of the paste during its baking up to 1000°C, could be very important and may induce a risk of infiltration in the pot [5]. In 2002, a new family of ramming paste was introduced on the market, called ecofriendly (EF) [3]. The binder, still issued from coal tar pitch, was specially treated to decrease significantly the most dangerous PAH, and especially BaP. Table II shows the decrease in ETF when moving from tepid and cold pastes to the EF family. Table II: Typical ETF for the different types of ramming paste. Type of Tepid Cold ColdEF Tepid EF paste ETF 2500 2300 300 360 (ppm) The difference of ETF between tepid and cold ramming pastes is not considered as significant, because these average values come from a small number of values. The EF family shows a decrease in the ETF by a factor 6 or 7. This improvement has also been demonstrated by emission measurements during ramming operation and during laboratory baking. In parallel, the final characteristics of the EF ramming pastes were similar to those of standard pastes. Since then, the EC Directive 2004/107/CE of European Parliament was published in Official Journal dated 26.1.2005, and asked for substitution of carcinogenic products when possible. There is an obligation of justification towards local authorities when carcinogenic products are used. As the EF family still contains PAH components with a non-null ETF, there is still a need to medically follow-up the use of these products. They still require individual protective equipment and a specific labelling showing the hazards for the users. The situation is the same for pastes with resin binders.

Table III: PAH profile of the new paste compared to tepid standard and EF ramming pastes. Values for Ne02 are the detection limits. In ppm Tepid Tepid EF NeO2 paste paste paste Naphtalene 400 7000 <58.7 Fluorene 2300 2500 <5.9 5100 Phenantrene 6900 <11.7 600 500 <2.9 Anthracene 2200 <29.4 Fluoranthene 4400 2500 1200 <11.7 Pyrene Benzo(a)anthracene* 800 100 <11.8 <29.4 Chrysene * 800 100 1100 100 Benzo(b)fluoranthene* <11.8 Benzo(k)fluoranthene * 500 43 <5.9 <5.9 Benzo(a)pyrene* 1200 100 100 B enzo(g,h,i)perylene 800 <11.8 200 Dibenzo(a,h)anthracene* 1700 <23.5 100 <117.4 Indeno-pyrene 900 Investigation on the components released during baking up to 1000°C is on the way. It has to be recalled that petroleum tar binders, which do not release PAH at ambient temperature, emit a high amount of PAH, especially BaP, during baking [3], almost 30 times more than a hot ramming paste! In NeO 2 paste, among the list of 14 compounds of PAH, only naphthalene is detected at elevated temperature. Therefore the use of this new paste will modify significantly the working conditions in the pot, as we will see later. Standard physico-chemical characteristics The standard physico-chemical characteristics of the new paste have been determined, in order to compare them to those of standard and ecofriendly ramming pastes, all of them being purely anthracitic grades. The samples have been densified with a hand-rammer at 20°C for the new paste and some cold pastes. The tepid pastes have been densified at 40°C. Table IV summarizes the green density on samples 90 mm in diameter and 150 mm long, and the main properties after baking at 1000°C: volumic expansion and crushing strength. Table IV: Typical characteristics of the new paste compared to standard and ecofriendly ones, tepid or cold ones. 20°C 20°C 40°C 40°C NeO2 std EF std EF paste paste paste paste Green 1.62 1.59 1.60 1.59 1.59 density Baked 1.44 1.44 1.39 1.45 1.46 density Volumic 1.4 expansion 2.6 1.5 1.8 1.1

Characteristics of the new paste developed Chemical Hazards The new ramming paste developed, called NeO2, does not contain any carcinogenic component classified by the European regulations. It is not based on a resin binder or on a coal tar or petroleum tar binder. The PAH profile of the ramming paste (Table III) shows that any PAH present was below the detection limit of the method used, compared to some recent batches of standard and EF pastes. In the case of the new paste, no carcinogenic product is present, and in fact no hazardous component is used. The Safety Data Sheet of each raw material does not mention any hazardous ingredient.

(%)

Crushing strength (MPa)

19

17

18

23

24

The volumic expansion, that represents the difference of volume between the green stage and the densified stage, is a key parameter for the performance of pot. If this volumic expansion is

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negative, it means that the paste will shrink during the pot preheating, and the risk of metal infiltration is very high. As can be seen in Table IV, the new paste presents a typical green density close to that of the other pastes, a good volumic expansion and a crushing strength also of the same order of magnitude. Specific characteristics Rammability index and temperature window from the Fischer Sand rammer (FSR) The rammability index is determined with a Fischer rammer equipment which has been adapted in order to perform automatically 250 strokes and to measure the geometrical density of the ramming paste for each stroke (see Figure 1).

ramming temperature for NeO2 and a 20°C standard cold ramming paste. N2 is comprised between 65 and 130 for a range of temperature of [11-35°C] for NeO2, and for a range [2-11°C] for a cold ramming paste. This is due to the fact that the rammability index is generally higher than the one of standard paste, especially below 30°C, whereas the standard paste exhibits very low values of N2, even at low temperatures. It seems that the temperature window is thus larger for the new paste. But another study will be made on larger samples, to check the characteristics for different ramming temperatures, and determine the practical temperature window.

20

30

Temperature (*€) Neo2 - - -N2= 130

20 "C std paste -NeO2 (model)

- - - N2 = 65 20°C paste (model)

Figure 2: Evolution of the rammability index N2 versus temperature for NeO2 and a 20°C paste. 100

150

200

Expansion during baking curve Besides the volumic expansion, it is important to follow the linear expansion curve of a ramming paste versus temperature, to clearly detect the dimensional evolution that will be met in pot during the preheating. The test is performed on the sample densified by the Fischer Sand Rammer. The linear evolution of the sample versus temperature is recorded with a displacement gauge. Figure 3 shows the curve for the new paste compared to typical ramming paste.

Strokes number

Figure 1 : Examples of density curves obtained for the new paste and for the standard pastes. The paste is densified at 20°C in a mould to get a sample of 50 mm in diameter and 50 mm high. From the densification curve, the maximum green density can be determined (asymptotic value), and also the rammability index N2. This index is calculated by assuming a Weibull's law, as detailed in the ISO standard 17544:2004. When the value of N2 is very high, it corresponds to a dry paste, which requires a lot of energy to be densified. If the value of N2 is very low, it corresponds to a very wet paste, which is easily densified. Table V shows that the new paste presents a higher rammability index than most of the standard pastes presented here, so may requires slightly higher energy (longer ramming time) to get densified. The maximum green density is also higher. Table V: Typical rammability index N2 and maximum green density D max, for different ramming pastes 40°Cstd 20°C std 20°C EF NeO2 paste paste paste N2 72 51 44 48 (strokes) Dmax

1.641

1.622

1.620

T(«C) I-+-NE02 -~40
20qCEFpaste[

Figure 3: Linear expansion during baking curve for NeO2, 40°C pastes (EF and standard), 20°C pastes (EF and standard).

1.605 1

The ISO standard 17544 considers the range of ramming temperature, where the rammability index N2 is comprised between 65 and 130, and defined it as the temperature window of the paste, even if it is not really in accordance with the practical temperature window determined from the characteristics of handrammed pastes. Figure 2 shows the evolution of N2 versus

The curve of the new paste exhibits a peak above 100°C, as the other paste, but the shrinkage after the peak is smaller, and above 500°C (the typical cokιfaction temperature of the coal tar binder), the shrinkage which is already very low for the considered pastes, is even a small expansion for NeO2 paste. This could be result in a decrease in the risk of metal infiltration in the pot.

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Thermal conductivity and oxidation resistance These are some specific characteristics that need to be determined to predict the thermal conductivity of the pot side and the frozen bath ledge position, or to confirm that the paste could stand heating without being oxidised, for example in case of pot preheating with gas burners. These two characteristics are the same for all anthracitic grades, whatever ramming temperature is used. They could be modified for semi-graphitic grades: typically the thermal conductivity is increased by a factor 2, and the oxidation resistance is also slightly increased. Compared to the average values for anthracitic pastes, NeO2 presents the same level of thermal conductivity, as shown on Table VI, and a slightly better oxidation resistance.

Figure 5: Evolution of the volumic expansion versus the months of storage, for the new paste.

Table VI: Thermal conductivity (in W/m.K) and weight loss per oxidation (in %) NeO2 paste Anthracitic pastes Thermal conductivity

5.6

6.0

Weight loss oxidation

6.0

9.5

per

Temperature window The practical temperature window inside which the ramming paste could be densified and still exhibit good properties is determined on large samples (90 mm in diameter and 150 mm long). Different ramming temperatures have been studied from 5 to 40°C, each 5°C. Two samples were densified and baked at 1000°C for each case.

Ageing

Typical ramming pastes present a shelf life of several months. Above this shelf life, due to the volatile departure from the binder, there is a risk of an evolution of the paste characteristics: mainly a drying of the paste (the rammability index increases) and a decrease of the volumic expansion. The ageing of the new paste has been followed on 25kg bags stored at ambient temperature. Each month 2 bags were analyzed and the average values of the rammability index and the volumic expansion are given in Figures 4 and 5. After 12 months, the paste still presents a volumic expansion above 1%.

3,5 -

8

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.*

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4

6

8

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Figure 4: Evolution of the rammability index versus the months of storage, for the new paste.

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Application in pots Impact on the working conditions

Figure 6: Evolution of properties versus the ramming temperature: a) green density, b) volumic expansion, c) crushing strength. Even if the nominal ramming temperature of the paste is 20°C, it appears that it could be densified between 10 and 40°C, and still presents good volumic expansion and crushing strength (Figure 6). Above 30°C, the volumic expansion is higher, in accordance with the lower crushing strength, but the average levels are still correct. At 5°C, the volumic expansion is negative, which could be detrimental for the performance in the pot. Evolution of crushing strength versus baking temperature The ramming paste has been baked between 300 and 600°C, and the crushing strength has been measured at room temperature after cooling. This test is not as representative as a crushing strength measured at the baking temperature, but it gives an idea of the evolution of the mechanical properties of the paste at different steps of baking. The results are illustrated in Figure 7, in comparison with various pastes. For NeO2, the crushing strength is nearly constant since the baking temperature of 300°C, whereas for the other pastes, it increases from 500°C, which corresponds to the binder cokιfaction. In the case of NeO2, the binder cokιfaction may occur at a lower temperature than 300°C. Some further tests will be performed to check this hypothesis.

The use of the new paste could improve significantly the working conditions. The main point is the absence of carcinogenic products, and more generally of hazardous products in the composition of the ramming paste according to current regulations. It could be used by itself without respiratory masks. It does not smell, which means that the protective clothes will not be impregnated by a strong smell, as with coal tar pitch or resin binders. In addition to the suppression of the risk of breathing hazardous compounds, the risk of pitch contact with skin, which requires efficient protection [6] is also removed. Due to its composition, the ramming paste could be destroyed as a common waste. It means that there is no more need to send it to special incineration station. The working temperature window is rather wide, which will be more convenient for the blocks / paste and tool preparation before ramming. Compared to typical 20°C ramming paste, the shelf life of the paste is larger: 11 months at least, instead of 8. The physico-chemical characteristics of the paste are in line with those of the other pastes, thus with a good volumic expansion, and a promising linear expansion curve during baking. The crushing strength is intermediate, and not too high to avoid applying stresses on the cathode blocks. The rammability index is higher than those of the main pastes that have been presented for comparison, and closer to some other pastes on the market. It means it may need a slightly longer ramming time than the reference pastes, but similar to the conditions used for other commercialized pastes. Experience in pots Several industrial production batches of Neo2 paste have been made, and two pots are now running in two different smelters, in two different pot technologies. The ramming operation was similar to the usual one, and the total ramming duration was found to be about the same, not significantly increased. Measurements of emission have been done during ramming. The results are not yet available, but will be compiled in the next future. Conclusion In accordance with the EC Directive 2004/107/CE of European Parliament that asks for substitutes to carcinogenic compounds whenever it is possible, the new ramming paste NeO2 appears to be a good alternative to the usual ramming pastes, as it does not contain any carcinogenic product, neither any hazardous product, according to current regulations. The working conditions during ramming should be easier, with the removal of the masks. The fact that the paste does not smell, can be densified in a rather large temperature window, can be stored during at least 11 months, and can be treated as a common waste are some of the advantages which differentiates the paste from the standard ones. The physico-chemical characteristics are maintained and sometimes improved such as the expansion during baking curve, with a slight expansion during 500 and 950°C. The evolution of crushing strength with baking temperature tends to be a plateau, different from the increase observed with the other pastes. Future work will include measurement of compressive strength at the baking

Baking temperature (°C)

Figure 7: Evolution of the crushing strength measured at room temperature, versus the baking temperature, for the different standard and EF ramming pastes

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temperature to be able to better predict the paste mechanical behaviour in the pots. Results from pots are expected regarding the emissions measurements during ramming and the follow-up of the pot performances. References 1. D. Dumas, S. Meseguer, R. Paulus, "Relevant Properties of Ramming Pastes for Aluminium Smelters" (Int. Conf. Aluminium of Siberian, 1998, Krasnoyarsk). 2. B. Doornaert, A. Pichard, "Hydrocarbures Polycycliques Aromatiques (HAPs) " (rapport final INERIS, 18 dιcembre 2003). 3. S. Lacroix et al., "A new Ramming Paste for the Aluminum Electrolysis Cell Compatible with Technical and Environmental Constraints", Light Metals 2002,413-418. 4. K.R. Kvam et al, "Resin Binders in Ramming Paste", Light Metals 1996, 589-596. 5. F.B. Andersen, M.A. Stam, D. Eisma, "A Laboratory Evaluation of Ramming Paste for Aluminium Electrolysis Cells", Light Metals 2005, 739-744. 6. J GM Van Rooij et al., "Effect of the Reduction of Skin Contamination on the Internal Dose of Creosote Workers Exposed to Polycyclic Aromatic Hydrocarbons", Scand J Work Environ Health, 19 (1993), n°3, 200-207.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

TOWARDS A BETTER UNDERSTANDING OF THE CARBURATION PHENOMENON Martin Lebeuf4, Marc-Andrι Coulombe1, Bιnιdicte Allard2, Gervais Soucy1. dιpartement de gιnie chimique (REGAL), Universitι de Sherbrooke, 2500 boul. universitι, Sherbrooke, Qc, J1K2R1, Canada. 2 Carbone Savoie, 30 rue Louis Jouvet BP 16, 69631 Vιnissieux Cedex, France Keywords: Aluminum Electrolysis, Aluminum Carbide Formation, Cathode Wear, XPS, XPS Imaging Abstract Cathode wear in aluminum electrolysis cells is an undesired phenomenon which decreases potlife. Although it has been the subject of many studies, it is not yet satisfactorily understood. Furthermore, one major factor of this wear is the formation and dissolution of aluminum carbide, for which the mechanisms also still remain to be thoroughly explained. In view of this, laboratory scale electrolysis experiments were performed under different operating conditions, namely the duration and the presence or absence of an aluminum metallic layer on the cathode surface. The aluminum carbide formation was then studied using tomography, optical microscopy, Scanning Electron MicroscopyEnergy Dispersive Spectroscopy (SEM-EDS) and X-ray Photoelectron Spectroscopy (XPS). In particular, the XPS analysis permits further investigation of the chemical species present in bath-penetrated veins of the carbon cathode. Introduction The formation and dissolution of aluminum carbide (A14C3) and its subsequent oxidation at the anode are known to be important mechanisms in regards to the erosion of the cathode blocks, according to the following simplified general mechanism. Initially, A14C3 forms at the cathode-metal interface, i.e. on the cathode surface and in pores near the surface, according to the general reaction (1) [1]. 4A1(1) + 3C (S) = A14C3

improved properties, especially electrical, exhibit higher erosion rates [7]. In previous work [8], laboratory scale electrolysis experiments were performed to examine bath-filled pores near the cathode surface with X-Ray Photoelectron Spectroscopy (XPS). Inside and/or on the periphery of those pores, the analysis of the high resolution C Is spectrum revealed the presence of C-Al and C-OAl bonds. Those findings encouraged the authors to further investigate those supposed C-Al and C-O-Al bonds. Although AI4C3 formation in the carbon cathode pores can be identified with relatively good confidence by optical microscopy (due to its characteristic yellow color) and by X-ray mapping (concordant carbon and aluminum signals suggest the presence of aluminum carbide) [9,10], the use of XPS is justified by the new information it can provide, which could help better understand the AI4C3 formation/dissolution mechanisms. Indeed, as shown in the previous work, it can confirm with very good confidence the presence of different chemical bonds such as C-Al. It can also allow mapping of the chemical elements as well as the chemical species. In an effort to identify new tools to study the AI4C3 formation, some results provided by tomography will also be presented. Experimental Laboratory bench scale electrolysis experiments were performed within a cylindrical graphitic cathode crucible with dimensions, general operating conditions and sample preparation outlined in previous articles [8,11]. In order to study carbide formation, three experiments were performed using the parameters presented in Table 1. In one experiment, a layer of aluminum was initially placed on the cathode surface, but the post-experiment examination revealed that the cathode surface was not completely covered by the metal.

(1)

The carbide can then dissolve in the liquid metal in which it is slightly soluble (probably less than 0.01 wt% at reduction cell temperatures [2]). The much higher solubility of the carbide in molten cryolite (2.1 wt% at CR=1.8 and 1020°C) [3] suggests that the possible presence of a bath film at the cathode-metal interface could enhance its dissolution [4]. In any case, the carbide eventually reaches the electrolytic bath, where it is believed that it dissolves according to reaction (2) [3]. AI4C3 (s) + 5 AIF3 (diss) + 9 NaF (1) = 3 Na3Al3CF8 (diss)

(2)

Run

#

Once the carbide is dissolved in the electrolyte, it can reach the anodic region by convection and diffusion and then be oxidized. The net effect is thus a loss of carbon at the cathode surface.

1 2 3

Among the electrolysis cell parameters known to influence the carbon cathode erosion rate, increased current density is the most important [4,5,6]. Some other known factors are higher aluminum metal pad velocity, excess alumina and higher A1F3 content [5]. Furthermore, graphitized materials, increasingly used due to their

Table 1 : Experimental Conditions Cathode current Electrolysis Initial density (A/cm2) time (h) presence of Al 0.8 5 No 7.3 0.8 No 7.3 0.8 Yes

Inert gas N2 N2 N2

For optical microscopy, SEM-EDS and XPS analysis of the cathode subsurface, small vertical pieces of the cathode were cut in order to obtain a surface perpendicular to the bath-cathode interface (Figure 1).

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Figure 1: Perpendicular cross-section of the cathode: the grey area corresponds to the cathode surface and the striped area is the surface analyzed.

Figure 2: View (from above) of the cathode crucible, with cathode surface in white, cathode walls in black and position of the sawnoff pieces analyzed by tomography in grey.

Tomography scans were performed with a Skyscan 1172 operating at 70 kV, 142 μΑ and using an Al-Cu filter. The camera was set at a resolution of 1000 x 500 pixels and an exposure time of 632 ms. XPS analysis of the cathode material was performed using an AXIS ULTRADLD spectrometer (Kratos Analytical Ltd., Manchester, GB) equipped with a monochromatic Al Κα source operated at 225 W. The elemental composition of the analyzed surface areas was obtained from survey spectra collected at a pass energy of 160 eV. High-resolution C Is and O Is spectra were collected at 20 eV, but for the parallel imaging experiment, the equipment was set at 40 eV for intensity purposes. The pressure in the analytical chamber was lower than 10"6Pa and the apparatus was calibrated against the following lines: Au 4f, Ag 3d and Cu 2p. Since the sample was charging, an electron flood gun was used during the XPS experiments. Atomic concentrations of each component were calculated using CasaXPS (Casa Software Ltd.) by determining the relevant integral peak areas; a Shirley background was used. To compare the high-resolution C Is, O Is and Al 2p peak positions, the spectra were calibrated in reference to the fitted graphitic C-C component at 284.5 eV. The detailed spectra were fitted with several peaks using a Lorentzian asymmetric lineshape with tail damping for the graphitic C-C and a mixed Gaussian-Lorentzian function for other components.

Figure 3: Tomography scan (right) of a crucible piecefromtest #3, showing the bath penetration profile. The carbon cathode block size is 60 mm long by 35 mm large by 20 mm deep. The cathode surface is visible in the upper right part of the image. The bath penetration profile can be seen in the right part of Figure 3, the darker area being bath-impregnated. The density and composition of bath versus graphitic carbon causes this zone to be less transparent to the x-rays. It is then possible to measure the distance between the cathode surface and the bath front. Comparisons between the tests were made by measuring the bath front depth at two different positions for each sample: beneath the cryolite/cathode wall border (left arrow, Figure 3) and the middle of the cathode surface (right arrow, Figure 3). Penetration depth values typically ranged from 5 to 25 mm. The measurements from the three tests resulted in average penetration rates of 1.7 mm/h and 2.6 mm/h for the wall and middle regions, respectively.

Results and Discussion Tomography Tomography scans were generated to monitor the penetration depth and penetration depth profile in a specified location within the cathode blocks (Figure 2). It is known that cryolite might react with the cathode block carbon and intercalated sodium according to reaction (3). 4 Na3AlF6 (1) + 12Na {inQ + 3C (s) = A1 4 C 3 (s) + 24NaF (1)

SEM-EDS A bath-impregnated vein right beneath the cathode surface, from a sample taken from test #2 (as described in Figure 1), was analyzed by optical microscopy (Figure 4), SEM-EDS (Figures 4-5) and XPS (Figures 7-11). The optical microscope image shows the yellow carbide layer on the cathode surface and on the vein edges.

(3)

In view of this reaction, it was relevant to measure the bath penetration front. Visual observations of bath penetration are, though, somehow imprecise. For instance, because of preferential paths caused by cracks and porosities, bath infiltrations much deeper than the true bath penetration front can be seen. A more accurate method was desired, and the tomography scans provided a rapid and reliable way to confirm the true depth of the bath penetration front. Furthermore, it allowed a 3-D analysis, instead of the visual surface observations. Each tomography scan required only five minutes to acquire and process.

The X-ray elemental mapping acquired (Figure 5) reveals a strong Al presence in the carbide layer. The C, F and Na elements are present as expected, with some expected Na in the graphitic matrix. As for Ca, its non-homogeneous distribution in the vein can probably be attributed to the crystallization of the different Ca-species. The strongest presence of O appears to be linked with the Al element, i.e. in the carbide layer. Such a result was reported previously in the literature. Xue and 0ye [12] suggested that the carbides could have reacted with alumina to generate oxycarbides,

1098

according to reactions (4) and (5). Patel et al. [10] suggested that it could be alumina formed by the hydrolysis of aluminum carbide, according to reactions (6) and (7). Zoukel et al. [Il] suggested that handling of the sample in air could have allowed reaction (6) to take place, or that, simply, alumina could have deposited on the surface during cooling. AI4C3 (s) + A1203 (s) = 3 Al2OC (s)

(4)

AI4C3 (s) + 4A1203 (s) = 3AI4O4C (s)

(5)

AI4C3 + 60 2 = 2A1203 + 3C02

(6)

A14C3 + 6H20 = 2A1203 + 3CH4

(7)

Figure 4: Optical microscopy (left) and SEM back-scattered electron (right) images of a bath-impregnated vein beneath the cathode surface from test #4. Figure 5: Bath-infiltrated veinfromtest #2 : EDS mapping of a) C, b) Al, c) F, d)Na, e) Caand f) O.

About oxycarbides, Grjotheim et al. [13], in 1978, identified Al2OC in the reaction product of alumina, cryolite and aluminum placed in a graphite crucible and kept for 12 hours at 1050°C and for another 24 hours at 975°C (no electrolysis was performed). Since then, in the literature related to the Hall-Heroult process, there has not been significant reported indications of the presence of oxycarbides in laboratory or industrial alumina reduction cells (at cooled temperatures, where chemical analysis are possible). Chryssoulakis and Righas [14] did, in an electrochemical study performed at 1300K in acidic cryolitic baths with alumina additions, suggest that the observed reoxidation peaks could be due to the oxycarbides compounds Al2OC and AI4O4C. These two oxycarbide compounds were first described and studied by Foster et al. in 1956 [15]. It seems relevant to indicate that since then, even though it has been described and studied in different works [16,17,18], the very existence of thermodynamically stable Al2OC has been put in serious doubt [19]. In any case, either stable or metastable, this Al2OC compound should not, as pointed out by Lihrmann [20], be observed in significant amounts in an aluminum reduction cell, since it was shown to exist only at temperatures between 1715°C and 2000°C. 0degard [3], in his study of the solubility of A14C3 in cryolitic melts at reduction cell temperatures, also did not see any indication of the occurrence of this compound.

XPS C Is high resolution spectrum interpretation The analysis and deconvolution of a high-resolution XPS spectra can be quite complex and the present study required some further investigation. In the literature, many authors have identified C-Al and C-O-Al bonds in the high resolution C Is spectrum [21,22,23,24,25], but the majority of these studies were performed on homogeneous, pristine surfaces, mostly made by the deposition of a thin film of aluminum on a material by evaporation, sputtering, or chemical reaction. Furthermore, those thin films were often etched to various degrees for surface bondings analysis and so the very last layers of deposited metal were analysed. This fact is mentioned since one can expect the C-Al and C-O-Al species or complexes formed by those methods to be different from the crystalline A14C3 expected in the carbon cathode. Nevertheless, in both cases the BE (binding energy) values for specific chemical bonds should be similar, and those values are discussed below. In genera], in the above-mentioned literature, the XPS charge effect is compensated by calibrating the energy scale in reference to the C-C bond of the C Is at a BE value between 284.2 eV and 285.0 eV. This C-C bond can originate from different types of molecular arrangements (aliphatic versus aromatic, for example), which might explain this relatively large range of BE values. In the present study, as mentioned earlier, a value of 284.5 eV for the graphitic C-C bond of the C Is was used. This is not true for the spectra presented in Figure 6 though, where no charge effect corrections were performed. Graphite was not always present in sufficient amounts to allow a reliable correction.

Therefore, the presence of oxygen in the carbide layer seen in Figure 6 can probably be attributed to A1203 and/or A1404C. Since XPS results obtained in a previous work [8] suggested the presence of oxycarbide species in bath-impregnated pores of the cathode, reaction (5) should be of some importance. No thermodynamic analysis has been conducted for the moment to further investigate the matter.

1099

In the C Is high resolution spectrum, the various XPS studies mentioned in the previous paragraphs attributed a C-O-Al complex at BE values between -1.2 eV and -1.6 eV relative to the C-C bond. It is noteworthy that a true linear chemical bond C-OAl, where the carbon atom is directly linked to the O atom, should rather appear at a higher binding energy than the C-C bond, since the O is far more electronegative than either C or Al [25], In other words, depending on its exact chemical structure, a C-O-Al species could be found on both energy sides of the C-C bond.

which can possibly induce localised charge effects (bath zones versus graphite zones). It can also be difficult to know the exact, precise location of an XPS scan when very small areas are scanned.

As for the C-Al bond, the above-mentioned literature suggests it should appear at BE values from -2.1 eV to -2.6 eV relative to the C-C bond. In order to verify those C-Al BE values in the presence of a graphitic C-C bond peak (versus other types of C-C bonds), powders of A14C3 and graphite were mixed at different concentrations and analysed by XPS (the A14C3 powder, >99% pure, was obtained from Alfa Aesar and the graphitic powder was obtained by crushing pieces of the graphitic cathode materials). The results are shown in Figure 6. The dotted lines are added to help compare the major peaksfromthe different spectra. It is worth mentioning that there is in principle no true chemical C-C bonds in pure A14C3, since in its complex crystal structure the closest distance between two carbon atoms is 316 pm [26]. In spectrum (a) of Figure 6, the characteristic graphitic C Is appears as expected. In spectrum (b), a mixture of 50wt% graphite and 50 wt% A14C3 reveals that the graphitic C Is peak is still dominant, but has shifted about -0.5 eV, probably due to the apparition of the C-Al peak (still mostly hidden by the graphitic C-C peak). In spectrum (c), the top of the two peaks are separated and the graphitic peak is back to its 284.5 eV BE value. In spectrum (d), with only slightly higher A14C3 concentration, the graphitic peak has almost disappeared beneath the C-Al peak. From spectra (c-d), the calculated BE difference between the C-Al and the graphitic C-C is -2.2 ±0.1 eV, in good agreement with the above-mentioned values. The (e) spectrum shows almost no sign of graphite although it represents 20 wt% of the sample, and the last spectrum (f) shows the 100 wt% A14C3. The two small peaks at low BE values that can been seen in (e) and (f) and partly in (b), (c) and (d), at about -1.9 ± 0.1 eV and -3.2 ±0.1 eV relative to the C-Al bond, remained unidentified. Itatani et al. [27] studied A14C3 powders produced by different methods, and the XPS spectra they obtained were quite similar to the one obtained in the present work (f). They suggested the small peaks at low BE values to be caused by alkylaluminum radicals still present in the powder after the reactions steps used to produce A14C3. The current study is concerned with those peaks since XPS spot-checks performed on bath-impregnated pores of a cathode in a previous work showed peaks appearing at lower BE values than expected for C-Al and C-O-Al.

Figure 6: C Is High resolution XPS spectra of powders with different A14C3 (A) and graphite (G) mass concentrations. Scans were performed at a pass energy of 20 eV, a dwell time of 250 ms and 5 sweeps. No charge effect corrections were made. To further investigate the carbide and the so called oxycarbide chemical species in the bath-impregnated cathode, XPS imaging was used to isolate those peaks from the C Is spectrum. The technique consisted of collecting an image of 128 x 128 pixels and transforming this image to the according number of spectra. Those spectra were divided into regions that corresponded to C-F (286.0 - 288.0 eV), graphitic C-C (284.0 - 285.5 eV), C-O-Al (282.0 - 284.0 eV) and C-Al (281.0 - 282.5 eV) bonds. These regions were converted to an image. Because of the presence of a cryolite vein going through the carbon matrix, the regions gave different images according to their intensity. Dividing the intensity scale in relevant ranges called false color (Figures 7-8), the pixel spectra were sorted to generate a representative spectrum of each bond.

The presented spectra and approximate BE values still consist in a preliminary evaluation. A complete deconvolution of the high resolution Cls peak, as well as the Al 2p and O l s peaks (which are more complex, since their respective peaks are less separated), will be necessary to properly characterise the species and draw conclusions. Moreover, the XPS analysis of a bath-impregnated cathode is further complicated by its non-homogeneous nature,

1100

χΐθ'-

of two different C-O-Al bonds represented by one line scan since they cannot be isolated from each other in any part of the image. Figure 9 shows the results associated with the region model development. The graphitic C-C image corroborates Figures 4-5, while C-O-Al, C-Al and C-F bond imaging is new information. The new bond is assumed to be C-F because of its BE of 287.7 eV, but this still needs to be confirmed with element imaging. The presence of that bond was established using the same strategy of averaging spectra of another false color region as demonstrated in Figures 7-8 for C-Al and C-O-Al, respectively. vbO , 0]

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It is difficult, from XPS imaging (Figure 9), to interpret the different bonds distribution since they are solely based on an intensity scale. Linescans (Figures 10-11), however, enable comparison with counts that are proportional to concentration within the C Is spectrum. From those results, it is made clear that in the vein, the supposed C-O-Al, C-Al and C-F bonds exist while there is a much lower level of graphitic C-C bond than in the cathode matrix. The linescans also reveal differences in the distribution of the various chemical bonds within the vein (it seems relevant to indicate that for such XPS imaging, the resolution is about 3 μπι, or 0.5 arbitrary units on the line scan abscissa). However, a band larger than 281.0 to 288.0 eV should have been used to fit a proper background and thus obtain less noisy results. More obvious gradients were observed before [8], but the vein observed in the present work is considerably thinner.

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1101

9. O. Herstad et al. "Precipitation of Alumina and Aluminium Carbide During Electrolysis of Cryolite-Alumina Melts," Light Three experiments were conducted in order to investigate the Metals 1983, TMS, 347-356. formation of aluminum carbide during the reduction of alumina in 10. P. Patel et al. "Influence of Cathode Structure on Behavior a cryolite-containing graphite crucible laboratory electrolysis cell. During Electrolysis Part II: Porosity and Wear Mechanisms in Graphitized Cathode Material," Light Metals 2005, 757-762. Tomography analysis of carbon cathode samples from lab-scale 11. A. Zoukel et al. "Study of Aluminum Carbide Formation in experiments allowed quick and reliable determination of the bath Hall-Heroult Electrolytic Cells," Light Metals 2009, TMS, 1123penetration front depth. 1128. 12. J. Xue amd and H. A. 0ye. "Al4C3 Formation at the Interface Investigation of the XPS high resolution C Is spectrum helped of Al - Graphite and Al - Carbon/TiB2 Composite," Light Metals confirm the BE value of C-Al and revealed the presence of 1994, TMS, 211-217. unknown species in a commercial grade A14C3 powder. XPS 13. K. Grjotheim et al. "Aluminium Carbide and Oxy-carbide imaging allowed to observe various chemical bonds and Formation in Alumina-Containing Cryolite Melts," Light Metals difference in their distribution in a bath-impregnated cryolite vein. 1978,101-117. 14. Y. Chryssoulakis and G. Righas. "Ιtude ιlectrochimique des rιactions d'oxydorιduction des carbures ou oxycarbures Future Work d'aluminium dans les bains cryolithiques acides additionnιs The deconvolution of the high resolution C Is, Al 2p and O Is d'alumine à 1300 K," Bulletin de la Sociιtι chimique de France, spectra for pure A14C3 and spot-checks of bath-impregnated 1990, no. 5, 635-640. cathode samples still needs to be fully performed. This will allow 15. L.M. Foster et al. "Reactions Between Aluminum Oxide and the formal identification and quantification of the different Carbon - The A1203-A14C3 Phase ", Journal of the American compounds, which may include an oxycarbide. The refinement of Ceramic Society, 1956, vol. 39, no. 1, 1-11. the XPS imaging technique will reveal more precisely the 16. M. Heyrman and C. Chatillon. "Thermodynamics of the Al-Cdistribution of the different chemical bonds. Combined with the O System", Journal of the Electrochemical Society, 2006, vol. analysis of bath-impregnated cathode pores from different 153, no. 7,E119-E130. electrolysis conditions, this information will be used to propose 17. R. Yu et al. "Elastic constants and tensile properties of Al2OC reaction mechanisms for the A14C3 formation. by density functional calculations," Physical Review B, 2007, vol. 75, no. 10, 104114-1-104114-5. 18. J.-M. Lihrmann. "Thermodynamics of the A1203-A14C3 system Acknowledgements I. Thermochemical functions of Al oxide, carbide and oxycarbides Financial support from the Fonds Quιbιcois de la Recherche en between 298 and 2100 K", Journal of the European Ceramic Nature et Technologie (FQRNT) is gratefully acknowledged. Society, 2008, vol. 28, 633-642. 19. K. Motzfeldt. "Comment on Thermodynamics of the Al-C-0 We are grateful to Sonia Biais and Irθne Kelsey-Lιvesque from System [J. Electrochem. Soc, 153, El 19 (2006)]", Journal of the the Centre of Characterization Materials (CCM) of the Universitι Electrochemical Society, 2007, vol. 154, no. 3, S1-S2. de Sherbrooke for the tomography, SEM-EDS and XPS analysis. 20. J.-M. Lihrmann. "High-Temperature Behavior of the Aluminum Oxycarbide Al2OC in th eSystem A1203-A14C3 and A special appreciation is given to Neal Fairley for his support with Additions of Aluminum Nitride", Journal of the American with mapping curve synthesis using CasaXPS software. Ceramic Society, 1989, vol. 72, no. 9, 1704-1709. 21. B. M. DeKoven and J. M. White. "XPS Studies of MetalPure aluminum was provided by Neuman Aluminium Canada. Polymer Interfaces - Thin Films of Al on Polyacrylic Acid and Polyethylene," Applied Surface Science, 1986, vol. 27, 199-213. 22. S. Akhter et al. "XPS Study of Polymer/Organometallic References Interaction : Trimethyl Aluminum on Polyvinyl Alcohol 1. Morten Sortie and Harald A. 0ye, Cathodes in Aluminium Polymer," Applied Surface Science, 1989, vol. 37, 201-216. Electrolysis, Düsseldorf, Aluminium-Verlag, 2nd ed.(1994), 408 p. 23. B. Maruyama et al. "Catalytic Carbide Formation at 2. R. C. Dorward, "Aluminium Carbide Contamination of Molten Aluminium-Carbon Interfaces," Journal of Materials Science Aluminium," Aluminium, 49 (1973), 686-689. Letters, 1990, vol. 9, 864-866. 3. R. 0degard, "On the Solubility of Aluminum Carbide in 24. C. Hinnen et al. "An in situ XPS Study of Sputter-Deposited Cryolitic Melts," Metallurgical and Materials Transactions B, 19 Aluminium Thin Films on Graphite," Applied Surface Science, (1988), 441-447. 1994, vol. 78,219-231. 4. K. Vasshaug et al. "Formation and Dissolution of Aluminium 25. L. Sandrin and E. Sacher, "X-Ray Photoelectron Spectroscopy Carbide in Cathode Blocks," Light Metals 2009, TMS,1111-1116. Studies of the Evaporated Aluminum/Corona-Treated 5. S. Wilkening and P. Reny. "Erosion Rate Testing of Graphite Polyethylene Terephthalate Interface," Applied Surface Science, Cathode Materials," Light Metals 2004, TMS, 597-602. 1998, vol. 135, 339-349. 6. H. A. 0ye et al. "Cathode Bottom Wear During Aluminum 26. Norman Greenwood and A. Earnshaw. Chemistry of the Electrolysis," The Electrochemical Society, 2002, 847-856. Elements, Second Edition, Butterworth Heinemann, 1997,297. 7. D. Lombard et al. "Aluminium Pechyney Experience With 27. K. Itatani et al. "Some Properties of Aluminum Carbide Graphitized Cathode Blocks," Light Metals 1998, TMS, 653-658. Powder Prepared by the Pyrolysis of Alkylaluminum," Journal of 8. M.-A. Coulombe et al. "Carburation Phenomenons at the the American Ceramic Society, 1995, vol. 78, no. 3, 801-804. Cathode Block/Metal Interface", Light Metals 1010, 811-816. Conclusions

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

CHARACTERIZATION OF SODIUM AND FLUORIDE PENETRATION INTO CARBON CATHODES BY IMAGE ANALYSIS AND SEM-EDS TECHNIQUES Yuanling Gao, Jilai Xue, Jun Zhu, Kexin Jiao, Gangqiang Jiang School of Metallurgical and Ecological Engineering, University of Science and Technology Beijing Xueyuan Road 30, 100083 Beijing, China Keywords: Cathode material, Pore distribution, Image analysis, Sodium and fluoride penetration various industrial cathode products of semi-graphitic (HC35), full graphitic (HClOO), and graphitized carbons (SMH). All cathode samples were made in cylindrical form and their apparent density, open porosity and compressive strength were measured according to the relevant industrial standards.

Abstract The porous structure of carbon cathodes is materials dependent, which can be related to sodium and fluoride penetration during aluminum electrolysis. This work is aimed at developing a better digital method for characterization of the pores and the penetration resistance of the cathode products. The profiles of penetrated Na, F, AI versus penetration depth in the cathode samples after aluminum electrolysis were obtained using SEM-EDS analysis. The internal porosity, pore size distribution, pore shape and orientation and pore connectivity were also characterized by image analysis. It is found that the graphitization, total and partial porosity, pore orientation and connectivity can contribute to the sodium and fluoride penetration into the carbon cathodes. The penetration behaviors varied with change in cathode materials (from semi-graphitic, full graphitic to graphitized carbons).

Table I. Physical Properties of Cathode Samples Open Apparent Compressive Sample Density Strength Porosity (g/cm3) (MPa) (%) HC35

1.57

14.04

27.84

HClOO

1.61

18.12

33.06

SMH

1.54

22.27

19.17

For further optical microscopy inspection, the cylindrical cathode samples were sliced along their longitudinal axis, and the cutting surface on the cross-section of each segment was ground and polished according to an ASTM standard [10] modified with no polishing agent used.

Introduction Numerous publications have dealt with improving cathode properties and performance in aluminum reduction cells. One of important reasons to use more graphitic and graphitized cathode materials is to have better resistance to sodium and fluoride attack. This can lead to a longer cell service life while a higher cost for cathode materials.

Image Analysis Method Figure 1 is a flowchart illustrating the image acquiring and processing procedures used for the sliced cathode samples. The microscopy photographs of the samples were obtained by a standard polarization microscope (Leica, DMRX). The image analysis was performed using Image J software (a public domain, Java-based image processing program developed at the National Institute of Health). Image J was designed with an open architecture that provides extensibility via Java plugins and recordable macros.

The carbon or graphite cathodes have a porous material structure, which is developed in both the cathode forming and baking stages [1]. Porosity and density are important properties for cathode quality [2]. Sodium and fluoride penetration in aluminum electrolysis have received considerable attention for various carbon, graphite and composite materials [3-4], where higher graphitization is usually recognized for better resistance to sodium attack. In recent years, image analysis methods have been used to study carbon anode porosity [5-9], but there is so far no open literature reporting its application in studying the cathode pores. The aim of this work is to find a computer aided, statistical approach characterizing both pore parameters and sodium/ fluoride penetration properties. In this paper, the samples tested were taken from various industrial cathode products and image analysis and SEM-EDS methods were applied together before and after aluminum electrolysis. The pore parameters obtained were interpreted in association with the results from sodium and fluoride penetration studies. Experimental Preparation of Samples Table I shows some selected physical properties of the cathode samples used in this investigation. These were directly taken from

The digital image processing method was able to eliminate some miscellaneous, irrelevant details and enhance the total image quality. The skeletonization obtained by an image thinning operation was used to reflect the pore connectivity within the carbon cathode materials. The 2D data at each segment obtained from the image analysis were applied to statistically express some essential characteristics of the cathode materials. The total porosity of each cathode sample was determined from the average porosity data of five arbitrary regions (768x582) on one segment. The value of pore size distribution (PSD) is the sum of the areas of pores in a certain interval of pore diameter, where the diameter is a mean length of lines crossing the pore centroid per 5 degree in angle. Pore shape was described by the aspect ratio of the longest diameter to the shortest diameter crossing the pore centroid. Pore orientation and connectivity were also studied for comparison of the three cathode materials.

1103

calculated from the perimeter of the area of the pore according to Saltykov equation [12]: Cathode block and sample

S = 4 Lp/πΑρ

(1)

where Lp is the mean perimeter of pores on the cross-section surface and Ap is the mean area of pores. HC35 cathode sample exhibits a higher total porosity but lower number of pores than HClOO sample, while SMH has the highest total porosity (25.48%) among the three cathode samples. For the partial porosity with Dmean below 600 urn (%), the graphitized cathode sample is also the highest one (15.27%), followed by HClOO and HC35. The partial porosity mainly is the reflection on the inner micropores for the cathode materials, which could be related to some surface properties, such as capillary force, wetting angle, etc.

Optical microscopy

5 Arbitrary regions

Table II. Statistical Results of Image Analysis on the Pores in Cathode Samples Processed images Binary and thinning Final images and analvsis results

Samples

HC35

HClOO

SMH

Total Porosity (%)

21.05

15.35

25.48

Partial Porosity with Dmean below 600 urn (%)

7.31

13.08

15.27

Number of Pores

230

488

438

186.61

135.97

182.31

0.017

0.029

0.023

Average Pore Diameter (Mm) Specific Surface Area of Pore (um2/um3)

Figure 1. Image acquiring and processing procedures for analyzing cathode samples Sodium/Fluoride Penetration and SEM-EDS Examination

Figure 2 further reveals the details of the fraction difference of pores in terms of the pore size distribution (PSD) among the three cathode materials. The major fractions of the pores in HC35 and SMH cathode samples are the meso pores between 300 urn and 900 Mm in mean diameter, while a large part offractionsin HClOO samples are the micropores bellow 300 um. HC35 cathode sample has more fractions in the macropores (>900 urn) than HClOO and SMH samples.

Sodium and fluoride penetration tests during aluminum electrolysis were carried out in a laboratory cell, as described previously [3, 11], at a cathode current density of 0.5 A-cm"2 and temperature of 960 C. The cylindrical cathode sample (dia 15x60 mm) was immersed 2 mm in the cryolitic melt (cryolite ratio = 4) containing 5 wt.% A1203 and 5 wt.% CaF2. Both the sodium generated during aluminum electrolysis and the fluorides (NaF, Na3AlF6, etc) from the bath were penetrated into the cathode body along its axis direction.

2.5x10'

After the electrolysis experiment for 3 hours, the cathode sample was cut open along the longitudinal axis, and the penetration depth and the amount of penetrated sodium and fluoride were determined by SEM-EDS analysis with scanning time of 90 seconds crossing the area for each 1 mm along the axis direction. The penetrated metals and fluorine in the cathode samples were then expressed in term of the relative percentage of F, Na, Al, Ca and C against the penetration depth. Results and Discussion Pore Distribution

200

400

600 D

Table II presents the resulting data from characterization of the pores, where the 2D data could reflect the relevant 3D parameters of the pores. The specific surface area of pore can also be

800

mean(M

m

1000

1200

1400

)

Figure 2. Pore size distribution in cathode samples

1104

In Figure 3, the curves of pore shape distribution are plotted against the aspect ratio for the pores in the cathode samples. HClOO cathode sample has higher content of pores with aspect ratio of 1 to 3 than SMH and HC35 samples. In general, HC35 contains more meso pores than HClOO, while SMH has the majority of pores in the meso pores range. HClOO is mainly composed of the micropores with smaller total porosity than HC35. Both HC35 and SMH have evenly distributed pores with a higher total porosity. For the pore shape distribution, the majority of measured pores have an aspect ratio of 1 - 3 for the all tested cathode samples.

250 hr

-»-HC35 ---•-HC100 A SMH

•: ·

200 h\-

e

150

It was individually observed that part of the pores in the HC35 cathode was aligned in certain directions, roughly showing a higher connectivity. HClOO sample exhibited numerous micropores randomly distributed but a certain orientation massively connected through the cross-section of the sample. However, the pores in SMH cathode samples were evenly distributed and the connection among these pores was also randomly orientated.

3·

[

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i

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i

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Aspect Ratio Figure 3. Pore shape distribution in cathode samples

k3_

Pores Orientation and Connectivity

Figure 5. Skeletonization of pore images for cathode samples

Figure 4 shows the pore orientation of 180 degrees in the inner areas of the cathode samples, in which the reference direction (blue line) is perpendicular to the axis on the cross-section of the sample. In the skeleton of binary pore images, as shown in Figure 5, connected or intersected lines on the 2D segment area are used to represent the connected pores passing that area. These have been used to describe the connectivity of the pores within the cathode samples.

0.30

5s < u Oi)

aG 00

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Sodium and Fluoride Penetration Figure 6 shows the extent of sodium and fluoride penetration within the cathode samples after aluminum electrolysis. Figure 7 shows the micrographs of the cross-sections of HC35, HClOO and SMH, where the porous structures within these carbon cathodes can also facilitate this penetration.

HC35 HClOO SMH

\ \

\\

V

N

v«

M *Λ

0.00 I

-50

0

·

I

50

-J

100

1

150

i

1

0

200

10

15

20

25

30

35

40

45

50

55

Depth (mm)

Angle (°)

Figure 4. Orientation of pores in cathode samples (The upper-left sketch illustrates the orientation of a pore)

5

Figure 6. Penetrated sodium and fluoride vs. the penetration depth in cathode samples after aluminum electrolysis (960 C, 3 hr)

1105

The sodium generation during aluminum electrolysis can be expressed by the reactions below: Al(l) + 3NaF(l) = 3Na(in C) + A1F3(1)

(2)

Na(inAl)^Na(inC)

(3)

Na(in C) + C(Cathode) -> CxNa

(4)

the graphitized cathode sample has higher porosity than the full graphitic one. However, fluoride penetration is more dependent on the porous structures of the cathode samples. The percentage of penetrated fluorine increases in the order of SMH > HC100 > HC35, which is similar to their order in the open porosity (see Table I) and the partial porosity of micropores (see Table II). The micropores may play an important role in providing inner channels and capillary force for the bath penetration during aluminum electrolysis. Within the inner area of the carbon cathodes, the penetrated fluorine containing compounds might take more complex chemical forms, for instance, in NaF, A1F3, Na3AlF6 and CaF2. In addition, the strongfluctuationof penetrated sodium against the penetration depth within the HC35 sample could be due to the uneven pore connectivity and pore orientation (see Figure 4), as well as a fluctuating pathway caused by mixture of coke and anthracite grains in the sample inner area. It is primarily important in this work that some pore parameters, such as pore orientation and connectivity, pore size and shape distribution, etc, have been identified for their possible relevance to the sodium and fluoride penetration and cathode performance. The method is still under development, and more carbon cathode samples will be collected for systematical investigation. Conclusions 1.

Carbon and graphite cathode materials can be characterized using an image analysis method for the porosity, total number of pores, pore size and shape distribution, pore orientation and connectivity. The results can be combined with SEM-EDS analysis to evaluate sodium and fluoride penetration into the carbons.

2.

Graphitized cathode material has high total porosity but even pore size distribution, while it shows a strong resistance to sodium penetration in aluminum electrolysis.

3.

Full graphitic cathode material has higher total porosity and higher partial porosity in micropores than the semi-graphitic cathode materials, showing better resistance to the sodium penetration but higherfluoridepenetration.

4.

Semi-graphitic cathode material has higher open porosity while lower partial porosity in micropores than the full graphitic and graphitized cathode materials, exhibiting poor resistance to sodium penetration but lowfluoridepenetration.

Figure 7. SEM micrographs of the cross-sections of cathode samples: a) HC35, b) HC100, and c) SMH The sodium can continuously migrate into the whole areas within the carbon cathodes. It can be seen that the percentage of penetrated sodium, which could include both metallic Na and NaF or Na3AlF6, increases in the order of HC35 > HC100 > SMH, which is similar to their order in the degree of graphitization. This means that the fully graphitized cathode material, SMH, has the best resistance to the sodium attack, which is in agreement with the sodium expansion results in previous Rapoport tests [13,14]. Here the sodium penetration is found less sensitive to the porosity variation of the various cathode materials investigated, and in fact

Acknowledgement Financial support from National Natural Science Foundation of China (NSFC), and Funding for Doctor Degree Education from Ministry of Education of China are acknowledged References 1. Morten Sorlie and Harald A. 0ye, Cathodes in Aluminum Electrolysis (Dusseldorf, FRG: Aluminium-Verlag GmbH, 1994), 282-361..

1106

2. P. Clery, "Green Paste Density as an Indicator of Mixing Efficiency," Light metals 1998, ed. B. J. Welch (Warendale, PA: The Minerals, Metals & Materials Society, 1998), 625-626. 3. J. Xue, H. A. 0ye, "Investigating Carbon/ TiB2 Materials for Aluminum Reduction Cathodes, " JOM, 1992, 44(11): 28-34

9. S. Rorvik, A. P. Ratvik and T. Foosnaes, "Characterization of Green Anode Materials by Image Analysis," Light Metals 2006, ed. J. G. Travis (Warendale, PA: The Minerals, Metals & Materials Society, 2006), 553-558. 10. ASTM D2797-85 Standard Practice for Preparing Coal Samples for Microscopical Analysis by Reflected Light. ASTM International, 1999.

4. P.Y. Brisson, G. Soucy, M. Fafard. "Revisiting Sodium and Bath Penetration in the Carbon Lining of Aluminum Electrolysis cell," Light Metals 2005, ed. H. Kvande (Warendale, PA: The Minerals, Metals & Materials Society, 2005), 727-732.

11. J. Xue, Q. Liu, and J. Zhu, "Sodium Penetration into CarbonBased Cathodes during Aluminum Electrolysis," Light Metals 2006, ed. J. G. Travis (Warendale, PA: The Minerals, Metals & Materials Society, 2006), 651-654.

5. J. L. Eilertsen et al., "An Automatic Image Analysis of Coke Texture," Carbon, 34 (3) (1996), 375-385. 6. M. Tkac, T. Foosnaes and H. A. 0ye, "Effect of Vacuum Vibroforming on Porosity Development during Anode Baking," Light metals 2007, ed. M. Sorlie (Warendale, PA: The Minerals, Metals & Materials Society, 2007), 885-890. 7. S. Rorvik and H. A. 0ye, "A Method for Characterization of Anode Pore Structure by Image Analysis," Light Metals 1996, ed. W. Hale (Warendale, PA: The Minerals, Metals & Materials Society, 1996), 561-568. 8. A. N. Adams. "The Use of Image Analysis for the Optimization of Pre-Baked Anode Formulation," Light metals 2002, ed. W. A. Schneider (Warendale, PA: The Minerals, Metals & Materials Society, 2002), 547-552.

12. S. A. Saltykov, Stereometric Metallography (Moscow, State Publishing House for Metals and Sciences, 1958), 3nd ed., 1970 13. J. Xue, et al, "Characterization of Sodium Expansion in Industrial Graphitic and Graphitized Cathodes", Light Metals 2010, ed. J. A. Johnson (Warendale, PA: The Minerals, Metals & Materials Society, 2006), 651-654. 14. J. G. Hop, "Sodium Expansion and Creep of Cathode Carbon," (Dr.ing Thesis, Norwegian University of Science and Technology, 2003), 179.

1107

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION Inert Anodes and Wettable Cathodes SESSION CHAIR

Vιronique Laurent Rio Tinto Alcan Voreppe Cedex, France Jilai Xue University of Science and Technology Beijing, China

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PRESSURELESS SINTERING OF TIB2-BASED COMPOSITES USING TI AND FE ADDITIVES FOR DEVELOPMENT OF WETTABLE CATHODES Hamed Heidari1; Houshang Alamdari1; Dominique Dubι1; Robert Schulz2 1

Department of Mining, Metallurgical and Materials Engineering, Universitι Laval, Quιbec (QC), Canada GIV 0A6 2 Hydro-Quebec Research Institute, 1800 Boul. Lionel Boulet, Varennes, QC, Canada J3X 1S1 Keywords: Pressureless liquid phase sintering, Wettable cathode, Titanium diboride (TiB2) composite.

Abstract Titanium diboride is the most promising candidate material for development of wettable cathodes for aluminum smelting. It is considered as an alternative for carbon cathodes in order to reduce the anode cathode distance resulting in higher energy efficiency in electrolysis cells. In this work, TiB2-based ceramic specimens were consolidated using metallic additives followed by pressureless sintering. Different proportions of iron and titanium (< 10 wt%) were used as low melting point sintering additives. Sintering was conducted at 1400-1650°C under controlled atmosphere. The effects of composition, sintering temperature, milling time and pre-alloying of the additives on densification, microstructure, and mechanical properties were investigated. It was found that pre-alloying and milling time have significant influence on densification, microstructure uniformity and bending strength. Uniform crack-free microstructure with even distribution of pores as well as maximum relative density of 91% and bending strength of 300 MPa were obtained using pre-alloyed additives, milling time of 30 min and sintering for 1 h at 1650°C. Introduction Liquid aluminum reacts with almost all materials with only a few having good stability [1]. Most of the stable materials are very expensive metals or non conductive ceramics, which are major obstacles for their application as cathode material. Nevertheless there are some electrically conductive ceramics such as graphite and TiB2. Graphite has been used for more than a century in aluminum electrolysis cells [2]. However, liquid aluminum does not wet graphite. Thus a relatively thick liquid metal pad is kept on top of the graphite cathode to avoid the diffusion of electrolytes through the cathode blocks and to insure good electrical current distribution within the cell. The magnetic fields present in smelting cells apply significant Lorentz forces on the metal pad resulting in wave creation. In order to avoid the short circuits between the metal pad and anode, the anode-cathode distance is kept large (typically 4.5 cm) unnecessarily increasing the bath resistance [3, 4]. TiB2 has very high melting point (about 3000°C), low density, excellent strength, high hardness, and very good thermal and electrical conductivities [5-7]. It is chemically stable in and well wetted by liquid aluminum [3,4, 8]. It has been therefore the most promising in the attempts to find an alternative material for carbon cathodes since the search began in the 1950's [9]. Despite its extraordinary properties, strong covalent bonding and low diffusion coefficient make sintering of TiB2-based ceramics quite difficult [7]. Fully dense TiB2 is probably not necessary for cathode application. It has been shown that using dense TiB2 cathodes results in early failure by cracking [10]. Liquid aluminum reacts with impurities at grain boundaries and, after a period of time,

results in the formation of new phases, internal stress build up and crack formation [11]. A number of techniques have been used to consolidate TiB2 [1215]. Pressureless sintering is a low cost technique to produce large and near net-shape components. However to consolidate pure TiB2 with this technique, very high sintering temperatures are required resulting in exaggerated grain growth and reduced mechanical properties. It has been reported that at temperatures above 1700°C, the presence of titanium oxide at the surface of particles increases both pore and grain size by increasing the surface diffusivity [16, 17]. In industrial conditions, it is difficult to remove this oxide layer. Thus, it is preferred to sinter TiB2based ceramics below 1700°C. At these relatively low temperatures the use of sintering additives are almost unavoidable to provide a liquid phase promoting consolidation of TiB2-based composite. Under these conditions, the sintering of TiB2-based composite requires an appreciable amount of liquid phase, wettability of TiB2 by the liquid phase, and small solubility of the solid phase in the liquid [18]. In this study, titanium and iron were used as additives to reduce the sintering temperature. Titanium was chosen as the principal metallic additive. Upon infiltration of aluminum into the porous cathode, it could react with the titanium additive to produce TiAl3 with a melting point higher than that of aluminum. TiAl3 also shows a very good wettability with respect to molten aluminum. Although T1AI3 is soluble in liquid Aluminum, it has been shown to be stable when formed at the surface of TiB2 [19, 20]. Iron, in turn, was chosen to somewhat reduce the melting point of additives. Ti and Fe have a eutectic at 1078°C and 71.1 mol% Ti [21]. These metals wet the surface of TiB2 [22] while TiB2 has a small solubility in liquid Ti-Fe [23]. The TiB2-Fe system is characterized also by a eutectic at 1340°C and 6.3 mol% TiB2 [24]. The reaction between TiB2 and Fe may accelerate densification and lead to the formation of Fe2B. This phase could cause deterioration in the mechanical properties as well as the resistance to liquid aluminum. The presence of titanium in the liquid mixture could prevent the formation of Fe2B [25]. Fe can also react with liquid aluminum to form FeAl3 having a melting at about 1160"C [26]. In this work, the effects of different compositions, processing conditions, and sintering temperature on density, microstructure and mechanical properties were investigated. A total amount of 10 wt% Ti and Fe was selected as metallic additives to provide about 10 vol% of liquid phase during sintering. Materials and methods Commercial TiB2, Ti and Fe powders (Atlantic Equipment Engineers Inc., Bergenfield, NJ) were used as starting materials. The particle size of TiB2 powder was between 2 and 10 μπι with a mean size of 6 μπι. Its purity was >99.7 % and the impurities were

1111

C, N, O and Al. For Ti powder the particle size was <20 μπι and for Fe powder, the particle size was between 1 and 9 μπι. The TiB2 powder was mixed with selected proportions of Ti and Fe powders and then milled in a high energy ball mill (SPEX 8000, Spex Industries, Inc., Edison, NJ) using hardened steel vial and balls with a ball-to-powder weight ratio of 4:1. Ti and Fe powders were added to TiB2 either separately or after being prealloyed. Pre-alloying of Ti and Fe powders was performed, first by mixing titanium and iron in a 70:30 Ti:Fe weight ratio and then by pressing the powder mixture at 400 MPa using an uniaxial steel die. The compacted specimens were then sintered at 1150°C for 1 hour. The resulting pellets were subsequently crushed and milled using high energy ball milling for 1 h to obtain the pre-alloyed additive in powder form. XRD analysis showed the presence of aTi and the FeTi intermetallic compound in this powder. The TiB2-Ti-Fe powder mixtures were pressed at 150 MPa in a uniaxial die to form pellets of 16 mm in diameter and bars of 38 x 13 mm. The specimens were then heated in a tube furnace (INP15-20, Norax Canada inc., QC) at a rate of 6°C/min from room temperature to the specified sintering temperature. Sintering of pellets was performed under an Ar/5%H2 protective atmosphere for 1 hour. The various experimental conditions of this study are given in Table I.

method (XRD; Siemens D5000) with a Cu Ka radiation at a scanning rate of 1° min"1. Wettability of sintered specimens was studied by placing a 90 mg aluminum piece at the surface of sintered specimen surface and heating under high vacuum (1.2E5 mbar) to 960°C. The wetting angle was monitored using an instant imaging system. From these images, contact angles can be measured as a function of time. Results and discussion Effect of additive composition The relative density was measured for specimens with two different compositions after sintering at 1400, 1600 and 1650°C. As shown in Figure 1, at all sintering temperatures, specimens containing 7wt% Ti and 3% Fe (T7F3M10) had higher density than that of specimens containing 8% Ti and 2% Fe (T8F2M10). According to the Fe-Ti phase diagram [28] the additive with a mass ratio of 7Ti:3Fe is closer to eutectic. A lower melting point leads to earlier formation of liquid phase during sintering, therefore promoting better densification.

Table I- Experimental conditions used for consolidation of specimens Additives No.

Code*

Sintering

PreMilling wt%** Temperature (°C) alloying time (min) Ti Fe

1 T8F2M10 8 2 1400-1600-1650 10 2 T7F3M10 7 3 1400-1600-1650 10 3 T8F2M30 8 2 1650 30 4 T7F3PM10 7 3 1600-1650 yes 10 5 T7F3PM30 7 3 1650 yes 30 6 T7F3PM60 7 3 1650 yes 60 7 T7F3PM120 7 3 1650 yes 120 8 T7F3PM240 7 3 1650 yes 240 *T: Titanium, F: Iron, P : Pre-alloying, M : milling time **Unless noted, all compositions in this article are in weight percent

The green density of specimens was evaluated by measuring their weight and geometrical dimensions. The bulk density of sintered specimens was determined using the Archimedes method with isopropanol as the immersing medium. Theoretical density was estimated using the rule-of-mixtures calculations that assumed the nominal compositions of the powder specimens as specified. Relative densities were calculated by dividing the measured bulk density by the calculated theoretical density. Based on replicated measurements on identical specimens, the uncertainty on relative density was estimated to be less than ±1%. Hence, no error bar was included in the figures. The bending strength of the sintered specimens was measured using the three-point bending test with a 26 mm span at a loading rate of 0.5 mm/min according to the ASTM Cl 161 standard [27]. The dimensions of the test specimens used for bending strength measurements were 38 mm x 13 mm x 4 mm. For microstructural investigations, the specimens were cut with a diamond saw and polished to 0.1 μιη surface finish using successively finer diamond abrasives. The microstructure of specimens was investigated using a scanning electron microscope and the chemical microanalysis was performed by energy dispersive X-ray spectroscopy (EDX; PGT Avalon, Princeton, NJ). The crystalline structure was determined by X-ray diffraction

1112

1350

1400

1450

1500

1550

1600

1650

1700

Sintering Temperature (*C)

Figure 1- Comparison of the relative density as a function of sintering temperature and composition of sintering additives (T8F2M10: TiB2+8%Ti+2%Fe; T7F3M10: TiB2+7%Ti+3%Fe). Effect of sintering temperature As shown in Figure 1, for both compositions (T7F3M10 and T8F2M10), there is not much densification after sintering at 1400°C. However, sintering at 1600 and 1650°C resulted in an appreciable densification to approximately 70%. Between 1600 and 1650°C, there is no further densification. An SEM backscattered (BS) micrograph of the T8F2M10 specimen, sintered at 1400°C, is presented in Figure 2. It shows that the additives formed segregated phases. A temperature of 1400°C, well below the melting point of Fe and Ti, is not high enough to provide the liquid phase required to promote densification. Moreover, these additives were not locally mixed in the proper ratio and were not in intimate contact with each other. When Fe and Ti particles are in intimate contact, solid state diffusion occurs at the interface resulting in the formation of a thin liquid layer in between. Increasing the sintering temperature to 1600°C resulted in the formation of significant amount of liquid and consequently in higher densification of specimens.

specimens milled for 30 min showed a maximum density of 91% after sintering. Further milling resulted in a slight decrease in density. Figure 4 shows that the three-point bending strength of sintered specimens follows a similar trend: a maximum bending strength of 300 MPa was achieved for specimens milled 30 min.

-••After sintering Hk-Green Samples

Figure 2- Backscattered SEM micrograph of T8F2M10 (TiB2+8%Ti+2%Fe) specimen, sintered at 1400°C for 1 h (The arrows show segregated phases containing the additives).

90 120 150 Milling time (min)

Effect of pre-alloving additives Besides the uniform distribution of additives, it is important to have Ti and Fe particles in contact with each other to promote the formation of the liquid phase. Hence, the addition of pre-alloyed additives, instead of adding Ti and Fe separately, was considered to achieve this goal. The pre-alloyed additives were prepared by mixing, pressing, and sintering Ti and Fe powders with a mass ratio of 7:3. Table II compares the relative densities of specimens with prealloyed additives (T7F3PM10) with those obtained by adding the additives separately (T7F3M10) after sintering at two different temperatures. Under the same processing and sintering conditions, pre-alloying of the additives resulted in better densification. The difference is significant at 1650°C.

Figure 3- Influence of milling time on relative density of green and sintered specimens (TiB2+7%Ti+3%Fe) using pre-alloyed additive and sintered at 1650°C for lh.

£g

Sinter. 1600°C

Sinter. 1650°C

T7F3M10

59

72

72

T7F3PM10

62

74

80

'

240 «

V

ω

180 H ISO

/ / f

\ \

V

Y

____

_

>*--g-1

-4

140 -

«

120

150

Milling Time (min)

Figure 4- Influence of milling time on bending strength of specimens (TiB2+7%Ti+3%Fe) prepared using pre-alloyed additive and sintered at 1650°C for lh.

Relativek density (%) Green Density

260

S

Table II- Relative density of specimens with separate (T7F3M10) and pre-alloyed (T7F3PM10) additives

Specimen

/ /

2 X

In order to understand the influence of milling time on sintering, the particle size distribution was determined and XRD analyses were performed on milled powders while the microstructure of sintered specimens was investigated by SEM. The effect of milling time on the particle size distribution of powders is shown in Figure 5. After 10 min of milling, the powder mixture is mainly composed of particles with a mean size of 6 micrometers similar to that of the starting TiB2 powder. However, a wider particle size distribution was observed. This distribution widening as well as the appearance of a shoulder at around 2 micrometers is most likely due to the TiB2 particlefracturingand refining during milling. In addition, EDX analysis of the very large particles showed that the peak at 200 micrometers is related to the Ti and Fe pre-alloyed particles. Milling of mixed powders for 10 min reduced the particle size of additive powders, but some large additive particles remained. Milling for 30 min, however resulted in a quite different particle size distribution. The distribution is much wider and shifted toward the small diameters. The quantity of particles smaller than 0.7 micrometer increased, and the peak related to the metallic additive particles disappeared. This suggests that milling for 30 minutes results in a good refining and dispersion of metallic additives within the powder mixture

Effect of milling time Milling was performed in order to achieve a uniform distribution of additives. However, milling time should be as short as possible to reduce costs and to prevent oxidation of powders. To investigate the effect of milling time on the densification process, powder mixtures were milled for different times prior to sintering. Preliminary results showed that by increasing the milling time from 10 to 30 min, density after sintering increased. These preliminary experiments suggest that the milling time has an important influence on densification. The effect of milling time was further investigated in a systematic way for specimens containing 90 wt% TiB2 , 7 wt% Ti, and 3 wt% Fe using five different milling times (10, 30, 60, 120 and 240 minutes) followed by compaction and sintering at 1650°C. (Specimens 4-8, Table I). The densities of specimens (green and after sintering) were plotted as a function of milling time in Figure 3. No significant influence of milling time on green density was observed. However

1113

and provides partial refining of TiB2 particles (particles smaller than 0.7 micrometer). By increasing the milling time to 60 min, a second peak appeared in the particle size distribution at around 100 micrometers. Since the powder samples are deagglomerated using an ultrasonic bath prior to analysis, the presence of this second peak suggests the formation of strong agglomerates in the powder. By further increasing the milling time, the quantity of these strong agglomerates increases but no significant increase is observed in the quantity of small particles. EDX analysis of large agglomerates showed that they were rich in Ti and Fe. They are usually formed by plastic deformation and cold welding of smaller particles and are very difficult to deagglomerate even in an ultrasonic bath. The typical shape of the large agglomerates observed after 240 min of milling is shown in Figure 6. These large particles are basically composed of TiB2 particles welded together by metallic additives. ^«

240 min

M c O

u

ÜULJ 25

~*~T7f3PM10

~»-T7f3PM60 \

55

65

75

~H-T7F3I»M240

SEM micrographs of polished sections of specimens containing pre-alloyed additives and sintered at 1650°C are shown in Figure 8 to Figure 10. In Figure 8, the microstructure of the specimen milled for 10 min revealed a highly porous structure. The high level of porosity explains the low density (84%) and reduced bending strength of this specimen (197 MPa). By increasing the milling time to 30 min (Figure 9), a more uniform and denser microstructure was achieved after sintering. A significant increase of density (91%, Figure 3) and bending strength (300 MPa, Figure 4) was observed for this specimen compared with the previous one. As shown in Figure 5, after 30 min of milling, there was a refinement of the particle size and a broader distribution was observed leading to higher densification. During liquid phase sintering, densification is mostly caused by rearrangement of particles upon formation of the liquid phase [29]. Further densification is achieved by the solution-precipitation process: small particles promote this stage due to their higher surface energy and therefore higher solubility in the liquid phase [29]. A broader particle size distribution increases the overall contact area between particles and eases the particle rearrangement in the early stage of sintering. Moreover, the presence of small particles can help the solution-precipitation stage. As a result, higher densification and bending strength could be achieved after 30 min of milling. By further increasing milling time, densification and bending strength decrease. As revealed by the microstructure of T7F3PM120 specimens ball milled for 120 min (Figure 10), cracks were formed after sintering. These cracks explain the dramatic decrease of bending strength (Figure 4). Crack formation was attributed to the presence of hard agglomerates in the powder mixture. As shown in Figure 6, these large agglomerates are formed by the cold welding of TiB2 particles with additive particles. In the early stage of milling, the TiB2 particles are partially refined and stick to the additives. Upon further milling, the TiB2 particles become embedded in additives and form large and dense agglomerates. Upon sintering, these large agglomerates shrink initiating cracks around them in the compact. This phenomenon has also been reported previously [18, 29]. These

fK M 7\

ff\^\/ L

If ^ ^ v j \ Å>»>Å<^*ΤΛ

.

7%m**&^

WI>ΜNiΙ»t»»»»»i

Diameter (μητι)

Figure 5- Effect of milling time on the particle size distribution for powder mixtures containing the 70%Ti and 30%Fe pre-alloyed additive.

*

45

û L* x

20 Figure 7- XRD analysis of powders containing pre-alloyed additives after 10 and 240 min milling. (Cu Κα).

-«•»r/f-iPMiO

J^Hw

35

10 min

"

<>

Figure 6- SEM micrograph of a large agglomerate formed after 240 min milling in T7F3PM240 powder (TiB2+7%Ti+3%Fe). The XRD patterns of the TiB2 powders milled for 10 and 240 min are shown in Figure 7. The peaks were slightly broadened after 240 min of milling while the intensities decreased owing to the overall decrease of the crystal size of TiB2. At the resolution of these x-ray scans, the minor phases corresponding to the prealloyed additives were not detected.

1114

cracks limit the densification of the sintered specimens and reduce their strength.

temperature reached 960°C (t=0) shows that the liquid Al drop was almost spherical. The contact angle between liquid aluminum and the specimen at this moment was 169°. After a while, the contact angle started to decrease. After 60 min the contact angle was 96° and only 14° after 175 min. After 185 min, liquid aluminum was completely spread over the surface which indicates that wetting occurs quite rapidly on this specimen and it has good wettability for liquid aluminum.

Figure 8- Backscattered SEM micrograph of T7F3PM10 specimen (TiB2+7%Ti+3%Fe) milled for 10 min and sintered 1 h at 1650°C.

Figure 11- Behavior of liquid Al drop over T7F3PM30 specimen (TiB2+7%Ti+3%Fe) during the wettability test at different time. (The time from beginning of test are reported in minutes) The stability and reactivity of T7F3PM30 specimen in liquid aluminum have also been investigated. The specimen keeps its integrity after being exposed to liquid aluminum during 24h. Detailed results of these experiments will be published in an upcoming report.

Figure 9- Backscattered SEM micrograph of T7F3PM30 specimen (TiB2+7%Ti+3%Fe) milled for 30 min and sintered lh at 1650°C.

Figure 10- Backscattered SEM micrograph of T7F3PM120 specimen (TiB2+7%Ti+3%Fe) milled for 120 min and sintered 1 h at 1650°C.

Conclusions TiB2-based composites with 10% of Ti and Fe additives were consolidated using pressureless sintering. Specimens with 7%Ti+3%Fe additives showed better densification due to the formation of liquid phase during sintering. Best results were obtained for a sintering temperature of 1650°C. Pre-alloying of Fe and Ti before addition to TiB2 powder significantly improved the densification. The milling time has also a marked influence on densification and on the properties of the sintered TiB2 specimens. A maximum relative bulk density of 91% and maximum bending strength of 300 MPa were achieved with specimens milled for 30 min and sintered at 1650°C for lh. The micrographs of specimens milled for 30 min reveal a uniform crack free microstructure with an even distribution of pores while those milled for longer times show the presence of numerous cracks in the specimens. The sintered specimens showed some resistance in liquid aluminum although more tests are needed. The resistance of specimens against aluminum infiltration and erosion are under investigation.

Wettabilitv and stability in liquid aluminum Since specimens milled 30 min with pre-alloyed additives (T7F3PM30) show the best density, strength and uniform microstructure, their wettability by liquid aluminum was investigated. Figure 11 shows images of the aluminum drop on the surface of T7F3PM30 specimen at 960°C during the wettability test. From these images, the contact angles were measured as a function of time. The first image taken as the

Acknowledgement The authors wish to acknowledge the kind collaboration of the technicians of the Dept. Mining, Metallurgical and Materials Engineering of Laval University and of Sylvio Savoie from Hydro-Quebec. The financial support of this project was provided by Hydro Quebec and the Natural Sciences and Engineering Research Council of Canada (NSERC). The research project was also partially financed by the "Fonds Quιbιcois de la Recherche

1115

sur la Nature et les Technologies (FQRNT)" via the Aluminum Research Centre - REGAL. References [I] J.E. Hatch, Aluminum : properties and physical metallurgy, American Society for Metals, Metals Park, Ohio, 1984. [2] F. Habashi, Extractive metallurgy of aluminum, in: G.E. Totten, D.S. Mackenzie (Eds.) Handbook of aluminum, CRC Press, 2003, pp. 1-46. [3] M. S0rlie, H.A. 0ye, Inert and wettable cathodes, in: Cathodes in aluminium electrolysis, Aluminium-Verlag, Düsseldorf, FRG, 1994, pp. 66-70. [4] D.W. Richerson, Aluminum industry, in: D.W. Freitag, D.W. Richerson (Eds.) Opportunities for advanced ceramics to meet the needs of the industries of the future, US Office of Energy, 1998 [5] R. Telle, L.S. Sigi, K. Takagi, Boride-based hard materials, in: R. Riedel (Ed.) Handbook of ceramic hard materials, Wiley-VCH, Weinheim; Germany, 2000, pp. 802-945. [6] R.A. Cutler, Engineering properties of borides, in: S.J. Schneider (Ed.) Engineering Materials Handbook, Ceramic and Glasses, ASM International, Metals Park, OH, USA, 1991, pp. 787-803. [7] R.G. Munro, Material properties of titanium diboride, J. Res. Nati. Inst. Stan., 105 (2000) 709-720. [8] H. Zhang, V. Nora, J.A. Sekhar, Materials used in HallHιroult cell for aluminium production, TMS, 1994. [9] K. Billehaug, H.A. Oye, Inert cathodes for aluminum electrolysis in Hall - Heroult cells. Pt. I, Aluminum, 56 (1980) 642-648. [10] C.J. McMinn, Review of RHM cathode development, in, Pubi by Minerals, Metals & Materials Soc (TMS), San Diego, CA, USA, 1991, pp. 419-425. [II] M.S. Jensen, M. Pezzotta, Z.L. Zhang, M.A. Einarsrud, T. Grande, Degradation of TiB 2 ceramics in liquid aluminum, Journal of the European Ceramic Society, 28 (2008) 3155-3164. [12] M.S. Jensen, M.-A. Einarsrud, T. Grande, Preferential grain orientation in hot pressed TiB 2 , J. Am. Ceram. Soc, 90 (2007) 1339-1341. [13] T. Graziani, A. Bellosi, D.D. Fabbriche, Effects of some iron-group metals on densification and characteristics of TiB 2 , Int. J. Refract. Met. Hard Mater., 11 (1992) 105-112. [14] S. Torizuka, K. Sato, H. Nishio, T. Kishi, Effect of SiC on interfacial reaction and sintering mechanism of TiB 2 , J. Am. Ceram. Soc, 78 (1995) 1606-1610. [15] A. Petukhov, I. Khobta, A. Ragulya, A. Derevyanko, A. Raichenko, L. Isaeva, A. Koval'chenko, Reactive electricdischarge sintering of TiN-TiB2, Powder Metall. Met. Ceram., 46 (2007) 525-532. [16] S. Baik, P.F. Becher, Effect of oxygen contamination on densification of TiB 2 , J. Am. Ceram. Soc, 70 (1987) 527-530. [17] M.G. Barandika, J.J. Echeberria, J.M. Sanchez, F. Castro, Consolidation, microstructure, and mechanical properties of a TiB2-Ni3Al composite, Mater. Res. Bull., 34 (1999) 53-61. [18] W.D. Kingery, H.K. Bowen, D.R. Uhlmann, Grain growth, sintering, and vitrification, in: Introduction to ceramics, Wiley, New York, 1976, pp. 448-515. [19] P.S. Mohanty, J.E. Gruzleski, Mechanism of grain refinement in aluminum, Acta Metallurgica et Materialia, 43 (1995) 20012012.

[20] T.E. Quested, Understanding mechanisms of grain refinement of aluminium alloys by inoculation, Materials Science & Technology, 20 (2004) 1357. [21] H. Okamoto, Fe-Ti (iron-titanium), Journal of Phase Equilibria, 17 (1996) 369-369. [22] V.N. Eremenko, Y.V. Niadich, I.A. Lavrinenko, Capillary phenomena and wettability in liquid phase sintering process, in: Liquid phase sintering, Plenum Publishing Corporation, 1970, pp. 16-21. [23] L. Ottavi, C. Saint-Jours, N. Valignant, C.H. Allibert, Phase equilibria and solidification of Fe-Ti-B alloys in the region close to Fe-TiB2, Zeitschrift für Metallkunde, 83 (1992). [24] V. Raghvan, B-Fe-Ti (Boron-Iron-Titanium), Journal of Phase Equilibria, 24 (2003) 455-456. [25] T.H. Jungling, R. Oberacker, F. Thummler, L.S. Sigi, K.A. Schmetz, Pressureless sintering of TiB2-Fe-materials, Powder Metall. Int., 23 (1991) 296-300. [26] M. Palm, G. Inden, N. Thomas, The Fe-Al-Ti system, Journal of Phase Equilibria, 16 (1995) 209-222. [27] ASTM_Standard, C1161-02c Standard test method for flexural strength of advanced ceramics at ambient temperature, in, ASTM International, West Conshohocken, PA, 2002. [28] H. Baker, S.D. Henry, H. Okamoto, M.P.O.H.A.P.D.a.C. Asm International, ASM handbook. Vol. 3 : Alloy phase diagrams, ASM International, Materials Park, OH, 1992. [29] R.M. German, Liquid-phase sintering, in: Sintering theory and practice, John Wiley & Sons, Ltd., 1996, pp. 225-313.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

FURAN RESIN AND PITCH BLENDS AS BINDERS FOR TiB2-C CATHODES Jinlong Hou, Xiaojun LÜ, Hongliang Zhang, Yanqing Lai, Jie Li School of Metallurgical Science and Engineering, Central South University, Changsha 410083, China Keywords: aluminum electrolysis, furan resin, pitch, TiB2-C cathodes, physical properties, electrolysis expansion The commonly used pitch, however, expands during carbonization to produce porous and cracked carbons with random lamellar structure [7]. These pores and cracks not only degrade the mechanical and electrical properties of the TiB2-C cathode materials, but also become the pathway for infiltration and erosion of electrolyte and molten aluminum during the electrolysis process. This makes the cathode materials more susceptible to deformation and expansion, with a shorter life. To overcome these shortcomings, a wide range of research has been completed. A variety of methods such as thermal oxidation, flash distillation, thermal polymerization or use of additives have been tried to modify pitch in order to enhance its performance as binder [8-12]. Only the process of using additives to modify the pitch was simple and easy to implement. Condensation resin, as an additive to pitch, was expected to substantially improve the effect of the modification of pitch binder for aluminum electrode. It contains reactive groups which can form chemical bonds linking with the pitch to produce a three-dimensional network structure on carbonization [13]. Therefore, further research on modifying pitch by resin for TiB2-C cathode is necessary. In this study, blends of furan resin and pitch were used for the binder of TiB2-C cathodes. Characteristic properties of the binder and TiB2-C cathodes were studied to assess the advantages of the furan/pitch binder.

Abstract Blends of furan resin and pitch were applied as the binder of TiB2-C cathodes. Characteristic properties of both the binder and furan/pitch based TiB2-C cathodes were studied. The actual coking value of furan/pitch binder was higher than the theoretical value calculated from the coking values of the two components. There was a synergistic effect between pitch and furan resin. This leads to TiB2-C cathode materials with lower open porosity and electrical resistivity and higher bulk density and compressive strength than those of pure pitch based TiB2-C cathode materials. In addition, the low-temperature electrolysis expansion of TiB2-C samples was tested. Furan/pitch based TiB2-C samples showed lower electrolysis expansion than pitch based TiB2-C samples (1.68% and 1.87%, respectively). Molecular structure (FTIR) and semicoke morphology (SEM) analysis results indicated that after adding an appropriate amount of furan resin, the pitch binder had improved cohesion and flexibility. Introduction The poor wettability of carbon cathodes used in aluminum electrolysis requires the retention of a 19~30cm layer of molten aluminum and a high anode-cathode distance (ACD) in the electrolytic cell. The resulting smooth and steady surface of molten aluminum avoids short circuits but leads to a high energy consumption. However, it has been proved that applying wettable cathode materials of titanium diboride or composites based on TiB2 can narrow the ACD effectively. The wettable cathode allows a lower ACD with a decreased thickness (3-5mm) of molten aluminum. The secondary reaction of Al and the interference of magnetic field are reduced, lowering the energy consumption substantially [1-4]. TiB2-C composite material is considered one of the most promising inert wettable cathode materials [5] because it can reduce the usage of TiB2 while retaining the good wetting with the molten aluminum. In addition, it can be prepared by baking at a temperature much lower than the sintering temperature of TiB2. Pitch was the binder of first choice for TiB2-C cathode because of its high carbon yield, good fluidity, easy graphitization, and good compatibility with TiB2 aggregate [6].

Experimental Pretreatment of pitch The furan resin and pitch blends were mixed in the required ratio by forming a slurry in acetone. The solvent was then evaporated before moulding. Preparation of cathode samples The raw materials including TiB2 power with an average particle size of 10 μιη, petroleum coke (106-15Ομπι) and binder paste were mixed by kneader for 30 minutes, and then formed into cylindrical samples (O20mm) under a pressure of 160 MPa by a hydraulic press. After moulding, the samples

1117

were baked in a temperature controlled furnace according to the heating curve shown in Figure 1. 1000 P

800 -

|

600

/ i

K 400 S ^

Binder Coking values/%

\

i

\\

Furan resin

59.50

47.66

- actual value theoretical value

62 r k

I 60

1

/

200

Pitch

\

y

S

Table I. Coke values of pitch and furan resin

58 1

0

1

20

40

60

80

100

120

56

Time/h

8

Figure 1. Heating curve for TiB2-C cathode baking.

12

16

Furan resin content/%

Research methods

Figure 2. Dependence of actual coking value and theoretic coking value on furan resin content of furan/pitch binder.

The coking values of binders were measured according to GB 8727-88. The molecular structure of binders after curing at 150°C for 2h was studied using FTIR. SEM was also used to evaluate the topography of baked TiB2-C composite cathodes and semi-cokes obtained from binders carbonized at 500°C for 60 min. The bulk density, open porosity and compressive strength of baked TiB2-C composite cathodes were determined according to ISO 12985-2: 2000. The electrical resistivity of baked samples was measured using the SZT-90 four point probe measuring system. A modified laboratory Rapoport device [14] was used to test the electrolysis expansion of samples, the electrolysis was performed in [ Κ ^ Α Π ν ^ Α η ^ Ι - Α à ν Α Ι ^ electrolyte melts at 923TC , the current density was 0.8A/cm2 and the whole electrolysis process was taken in high-purity argon atmosphere.

On mixing the furan resin with pitch, a series of complex condensation and cross-linking reactions occurs as the temperature is increased. As the carbonization temperature of resin is lower than pitch, the furan resin is carbonized first to produce a microporous carbon with a certain extent of three-dimensional network structure produced. With a further increase in temperature, the pitch become more fluid and can infiltrate into the micropores of resin carbon. This combination suppresses the expansion and volatilization of the lighter components of the pitch. Therefore, the actual coking value of the modified pitch increased. FTIR analysis: To ensure the reactions between furan resin and pitch, and the cross-linking reaction of furan resin is completed, the furan/pitch mixed binder has been held at 150°C for 2 hours. The same process of holding at 150°C for 2 hours was used during the TiB2-C cathode baking. The samples were the subjected to FTIR (Fourier Transform Infrared Spectroscopy) analysis. Figure 3 shows the FTIR spectra of pitch, furan/pitch (12wt% resin content) and furan resin binders after curing at 150°C for 2hours. The spectra show that several reactions take place between the furan resin and pitch that causes the C=0 adsorption band of furan resin (a) at 1715cm"1 to almost disappear. A new strong secondary alcohol C-0 adsorption band (b) at 1010cm"1 appeared in the spectra of the furan/pitch binder, and the aromatic ring skeleton C=C adsorption band (c) at 1593cm"1 of modified pitch increased in intensity. The conversion of C=0 to C-0 and the increase in intensity of the aromatic ring skeleton C=C adsorption band of modified pitch show that this compositing treatment of furan resin and pitch promotes the condensation of the

Results and discussions Properties of binders Coking value: The coking values of furan resin and pitch are shown in Table I. The theoretical coking values have been calculated using the individual component coking values and the furan/pitch ratio in the blend. The actual coking value and theoretical coking value of the furan/pitch binders tested are shown in Figure 2. The actual coking values of the furan/pitch binders were higher than theoretical values. This is related to a certain synergistic effect between pitch and furan resin during carbonization.

1118

aromatic rings. The cross-linking between the rings enhances the degree of aromatization of the binder. This may also explain why the actual coking value of furan/pitch binder was higher than theoretical value.

| (ai)

(32)

3000

2500

2000

Wavebumbers rw^n

3000

1500

1000

( cài)

(2)í+¥Α\\^ΛÀ

2500

2000

Wavebumbers

500

1500

1000

^

500

( αÞα)

Figure 3. FTIR spectra of binders after curing at 150 °C for2h: (1) Pitch; (2) furan/pitch (12wt% resin content); (3) furan

SEM analysis of semi-cokes: Figure 4 shows the SEM image of semi-cokes obtained from binders carbonized at 500°C of furan resin for 60 min. From Figures ai and a2 many large pores can be seen in the semi-coke of pitch. Some of the pores are connected in series, and furthermore, many cracks and random lamellar structures are seen in Figure a2. In contrast, the surface of the furan/pitch (12wt% resin content) semi-coke shown in bi and b2 are smoother, with the pores mostly small or closed and surrounded by concaves. This shows that after modified by furan resin, the cohesion and flexibility of pitch binder were significant enhanced.

Figure 4. SEM image of resultant semi-cokes obtained from binders carbonized at 500°C for 60 min: (a) Pitch; (b) furan/pitch (12wt% resin content) (al, bl - 50 times magnification; bl, b2 - 200 times magnification). Physical properties of TiB2-C cathodes From the results above we know that there is a synergistic effect between pitch and furan resin during carbonization.

1119

These improvements in the coking value of the binder will have a large effect on the physical properties of TiB2-C cathodes. As shown in Table II, the furan/pitch binder provided TiB2-C cathodes of improved physical properties with lower open porosity and electrical resistivity and higher bulk density and compressive strength than those of pure pitch based TiB2-C cathodes. However, the highest resin content produced brittle carbon which increased the electrical resistivity and reduced the compressive strength of the

cathodes. The optimal physical properties of the TiB2-C cathodes are produced at a furan resin content of 12wt%. Figure 5 shows an SEM image (1000 times magnification) of the pitch and furan/pitch (12wt% resin content) based TiB2-C cathodes. Many large pores are observed in the cathode using pitch alone as binder. In contrast, furan/pitch blending suppressed the expansion of pitch, resulting in lower porosity in the cathode.

Table II. Properties of baked TiBr-C cathodes in the presence of 14 wt% binder of pitch and furan/pitch. BD (g/cm3)

OP (%)

PO

2.25

PO*

2.50

4%FA+96%P

ER (μΩ-m)

CS (MPa)

36.99

51.41

22.78

29.61

48.68

29.07

2.54

29.41

44.31

30.84

8%FA+92%P

2.53

29.30

43.58

36.52

12%FA+88%P

2.54

28.82

38.75

41.28

16%FA+84%P

2.55

28.44

39.76

37.59

Binder

P0: Mixed without acetone. P0*: Mixed in acetone. BD: bulk density, OP: open porosity, ER: electrical resistivity, CS: compressive strength.

(a)

(b)

Figure 5. SEM image of TiB2-C composite cathodes: (a) pitch (P0*) based; (b) furan/pitch (12wt% resin content) based. Electrolysis expansion of TiB^-C cathodes Figure 6 shows the electrolysis expansion curves of pitch based and furan/pitch (12wt% resin content) based TiB2-C cathodes tested in [K^AUVNasAlFτJ-AlFs-AkOs melts. It can be seen that both the electrolysis expansion curves are a

9-

parabolic shape. At the beginning of electrolysis, the expansion of the cathode occurred rapidly. As the electrolysis continues, the rate of cathode expansion reduces and tends to a constant after a period of time. There was a significantly

UJ

2.0 1.8 1.6 / 1.4 1.2 ! 1.0 0.8 0.6 0.4 - / 0.2

a

different behavior between the pitch and furan-based TiB2-C cathodes in the electrolysis expansion experiments. After electrolyzing for 1.5h, the expansion rates of pitch and furan/pitch based composite cathodes were 1.87% and 1.68%,

tφ>

1

/

" '

f

—P 88%PH2%FA

1

1

1

1

1

1

1

1

0.2 0.4 0.6 0.8 1.0 1.2 1.4 1.6

Time/h

Figure 6. Electrolysis expansion of pitch, furan/pitch (12wt%

respectively.

resin content) based TiB2-C composite cathodes.

1120

Pitch is a kind of soft carbon material. During cathode baking, it expands and produces porous and cracked carbons with a random lamellar structure. This makes the TiB2-C composite cathode more easily penetrated by electrolyte. The contact area of electrolyte and cathode increases, resulting in more sodium and potassium metal being produced on the surface of cathode. The K and Na are more likely to penetrate into the cathode through the lattice and grain boundaries of the carbon materials and form alkali metal-graphite intercalation compounds [CXM (K, Na)] [15]. As a result, the layered structure of the carbon may deform or rupture, leading to a large expansion of the cathode. In contrast, furan resin is a hard carbon material. It undergoes less expansion during carbonization than pitch. The micropores and three-dimensional network structure produced during cathode baking reduced the formation of the intercalation compounds [CXM (K, Na)]. Therefore, the furan/pitch based TiB2-C composite cathode showed a smaller expansion rate after electrolysis.

56(11):713-718. [5] M. Dionne, G Lespιrance, A. Mirchi. "Microscopic characterization of a TiB2-carbon material composite: raw materials and composite characterization," Metallurgical and Materials Transactions A. 32A:2649-2656. [6] John T. Baron, Stacey A. McKinney, Robert H. Wombles. "Coal tar pitch - past, present, and future," Light Metals. 2009:935-939. [7] M. Sorlie, H. A. 0ye. Cathodes in aluminum electrolysis, 2nd edition. Aluminum-Verlag. 1994. 7. [8] L.M. Manocha et al. "Carbon/carbon composites with heat-treated pitches I. Effect of treatment in air on the physical characteristics of coal tar pitches and the carbon matrix derived therefrom," Carbon, 2001, 39(5): 663-671. [9] Li Tongqi et al. "Structural characteristics of mesophase spheres prepared from coal tar pitch modified by phenolic resin," Chinese J.Chem.Eng, 2006, 14(5):660-664. [10] VE. Yudin, et al. "Carbon/carbon composites based on a polyimide matrix with coal tar pitch," Carbon, 40(2002): 1427-1433. [11] S. M. Oh, Y. D. Park. "Comparative studies of the modification of coal-tar pitch," Fuel, 1999, 78(15): 1859-1865. [12] LÜ Xiaojun et al. "Effects of pitches modification on properties of TiB2-C composite cathodes," Light metals. 2009:1145-1149. [13] STEVENS D A. "Mechanisms for Sodium Insertion in Carbon Materials" (Ph.D. thesis, Dalhousie University 2000). [14] Li Jie et al. "Effect of TiB2 content on resistance to sodium penetration of TiB2/C cathode composites for aluminum electrolysis," Journal of Central South University of Technology, 2004, 15(1): 400-404. [15] Fang Zhao et al. "Electrolysis expansion performance of TiB2-C composite cathode in [KsAlFτ/NasAlFeJ-AlFs-A^Os melts," Light metals. 2010:901-906.

Conclusions (1) There was a synergistic effect between pitch and furan resin during carbonization, with the actual coking value of furan/pitch binder being higher than the theoretical value. (2) After modification by furan resin, the cohesion and flexibility of pitch binder were significant enhanced. (3) The synergistic effect between pitch and furan resin improved both the physical properties and electrolysis expansion resistance of TiB2-C cathodes, with the optimal content of furan resin in binder being 12wt%. Acknowledgement Project

Supported

by

Major

State

Basic

Research

Development Program of China (2005CB623703), National High-Tech Research and Development Program of China (2008AA030502) and National Science & Technology Pillar Program (2009BAE85B02). References [1] J Thonstad et al. Aluminum Electrolysis, 3rd edition. Dusseldorf: Aluminum-Verlag. 2001. 328. [2] S. Briem et al. "Development of energy demand and energy-related C0 2 emissions in melt electrolysis for primary aluminum production," Aluminum, 2000, 76(6): 502-506. [3] K Billehaug, H. A. 0ye. "Inert cathodes for aluminum electrolysis in Hall-Heroult cells (I ) , " Aluminum, 1980, 56(10): 642-648. [4] K Billehaug, H. A. 0ye. "Inert cathodes for aluminum electrolysis in Hall-Heroult cells (II)," Aluminum, 1980,

1121

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INFLUENCE OF COBALT ADDITIONS ON ELECTROCHEMICAL BEHAVIOUR OF NI-FE-BASED ANODES FOR ALUMINIUM ELECTRO WINNING Vivien Singleton1, Barry J. Welch2, Maria Skyllas-Kazacos1* Centre for Electrochemical and Minerals Processing (CEMP), School of Chemical Engineering, University of New South Wales, Sydney NSW 2052 AUSTRALIA 2 Welbank Consulting Ltd., PO Box 207, Whitianga 3542 NEW ZEALAND Key words: inert anodes, aluminium electrolysis, Ni-Fe superalloys, corrosion resistance Abstract The anodic behaviour of surface-oxidised Ni-Fe-Co alloys was investigated over short-term periods of aluminium electrolysis. Additions of 10wt% Co were found to significantly improve the anodic wear resistance, due to suppression of FexO formation. Anodes having cobalt contents >30wt% exhibited poor performance due to rapid outwards diffusion of cobalt to the reaction interface. In general, the protective ability of the preformed oxide scales was greatly affected by the level of porosity and surface adhesion. In electrolytes containing <4wt% AI2O3, catastrophic failure of the anodes was observed due to concentration polarisation at the reaction interface. Under these conditions, the metal was rapidly destroyed by a combination of dissolution and fluoridation events.

oxidation of the metal. Steps must therefore be taken to reduce the activity of iron in the alloy, and thereby the driving force for FexO formation. It was recently suggested that this may be achieved by the addition of ternary or quaternary alloying elements [8]. In this study, the influence of cobalt additions on the anodic behaviour of Ni-Fe-based alloys is investigated. Experimental Anode preparation Three compositions of Ni-Fe-Co alloy were prepared, outlined in Error! Reference source not found.. In all cases, the Fe:Ni mass ratio was fixed at 1:1.85. The Co concentration was set at either 10, 30 or 50 wt%. For simplicity, the alloys shall henceforth be referred to as NiFeCo10, NiFeCo30 and NiFeCo50 respectively.

Introduction

Table 1. Ni-Fe-Co alloy compositions

Metals are arguably the most attractive candidates for inert anodes for the Hall-Hιroult process due to their good electrical conductivity and ease of fabrication. Of the metals considered, Ni-based superalloys have shown particular promise, largely owing to:

Alloy NiFeCoio NiFeCo3o NiFeCo 50

(a) The high oxidation resistance of Ni-rich alloys; and (b) The low solubility of Ni-containing oxides in cryolite-alumina melts [1-3].

Fe:Ni (wt) 1:1.85 1:1.85 1:1.85

Nominal Composition (wt%) Ni Fe Co 58.5 45.5 32.5

31.5 24.5 17.5

10 30 50

Alloys were prepared by arc melting using nickel, iron and cobalt lumps. Nickel (99.9wt%) and cobalt (99.99wt%) were obtained from Alfa Aesar; iron (99.9wt%) was obtained from Sigma Aldrich. The appropriate amounts of each metal were melted under reducing atmosphere (95% Ar, 5% H2), producing buttonshaped ingots of approximately 30g. The buttons were cast into rods using a copper crucible with a cylindrical mould. The rods were annealed in argon for 24h at 1100°C. The rods were machined to produce cylindrical anodes of diameter 8mm and length 10mm. The anode surfaces were abraded using 80 grit SiC paper to produce a rough surface for pre-oxidation. A female 3mm screw thread was machined into the top of each anode to facilitate connection to the cell power supply.

Unfortunately, nickel and nickel-rich alloys are poorly resistant to fluoridation in the Hall cell environment. Anodes having nickel contents of >75wt% have been shown to undergo surface fluoridation after only a few minutes operation [4-6]. NiF2 films have high electrical resistance and invariably cause passivation and failure of the anode. To reduce the risk of NiF2 formation, it is necessary to lower the nickel content of the anode by alloying with other metals. Iron has been a popular choice due to the low solubility of the nickel ferrite spinel (NixFe3.x04) [1]. Under oxidising conditions, Ni-Fe alloys develop layered scales consisting of one or more of the following phases [7]:

Pre-oxidation

1. Iron oxides, FexO: 2. Nickel oxides, NixO 3. Nickel ferrite, NixFe3.x04

Prior to electrolysis, the anodes were oxidised in air at 800°C for 24h to produce a protective semi-conducting scale. The oxidation products were identified using X-ray diffraction (XRD)

Of these, FexO offers the least protection to the underlying metal from the corrosive Hall cell environment. The problem is twofold: not only are iron oxides highly soluble in the bath, but they also serve as relatively poor diffusion barriers to prevent further

Electrolysis The cell configuration and electrolysis procedure have been described in detail elsewhere. The laboratory-scale cell consisted of a vertical tube furnace containing a vitreous carbon crucible

1123

cathode. Both anode and cathode were connected to the power supply using threaded stainless steel rods, sheathed using sintered alumina tubes. All electrolysis experiments were conducted at 960°C under an atmosphere of dry nitrogen. An interelectrode distance of 4cm was employed in all cases.

Η 10o

(1)

Electrolysis under high alumina concentration.

NiFeCoio, NiFeCo30 and NiFeCo50 anodes were subjected to galvanostatic electrolysis for 2h at 0.8A/cm2 in a bath consisting of 79.6wt% cryolite, 8.75wt% A1F3, 4.65wt% CaF2 and 7.0wt% A1203. To ensure stable metal pad formation, approximately 5g aluminium metal was added to the base of the crucible prior to electrolysis. (2)

Materials characterisation Following electrolysis, the anodes were removed from the bath and allowed to cool slowly to room temperature. Once cool, the anodes were immediately mounted in epoxy resin to prevent absorption of moisture. The mounted specimens were gently abraded using fine-grit SiC paper to expose the anode/bath interface. XRD analysis was carried out using a Philips X'Pert PRO Materials Research Diffractometer with resolution =0.02°2Θ, divergence slit =1/4°, anti-scatter slit =ιΛ0 and 0.04 radian soller slits in the incident and receiving positions. A mono-capillary was fitted to the instrument for incident beam conditioning and focusing.

1

1

1i

1

1

1

i

'

1

1

1

9-

"j J

£ c o

6-

]

ÎT3 i>-

4

6 4-

η

J

'1

1

0

1

i

·

1

«

i—'—i—1

10 20 30 40 Cobalt content in alloy (wl%)

50

Figure 1. Mass gain versus cobalt content for Ni-Fe-Co alloys oxidised at 800°C for 24h.

Extended electrolysis without alumina replenishment

To investigate the anode behaviour under conditions of low alumina concentration, NiFeCo3o was subjected to extended electrolysis at 0.8A/cm2. The same initial bath composition was employed as in experiment (1). The bath was not replenished with alumina for the duration of the experiment.

|

c 8"cφ σ> -, to 7 -

|)

Two types of electrolysis experiments were performed:

1

Unfortunately, there is very little information available in the literature about the oxidation behaviour of Ni-Fe-Co alloys. Previous studies have largely focussed on Fe-rich systems such as Kovar [10] or maraging steels [11-13]. Oxidation products were identified using XRD analysis. X-ray diffraction spectra are shown in Figure 2. Scales on alloys having >30wt% Co were cobalt-rich, consisting principally of Co304, NixCo3.x04 and CoxFe3.x04. In the case of NiFeCoio, the major oxidation product was the nickel ferrite spinel (NixFe3.x04), with small amounts of Co304. Cobalt (II) oxide was observed only on NiFeCo50. 2,3

3M

2,3 2,3

The carbon crucible cells containing frozen bath were potted in epoxy resin and sliced in half to reveal a vertical cross-section. Energy dispersive x-ray spectroscopy (EDX) analysis was performed on selected areas of the crucible cross-section.

ÜLJ

-ST 2kJ

NiFeCo u

2,3

13

o o

2,3

Results and discussion

4

2 c

Pre-oxidation

lk-l

Anodes were pre-oxidised at 800°C for 24 h in air, producing adherent scales approximately 10-30μπι thick. To assess the extent of oxidation, the anodes were weighed before and after oxidation to determine the net mass gain. For comparison, a binary Ni-Fe alloy having the equivalent Fe:Ni mass ratio (1:1.85) was oxidised under the same conditions. The results are shown in Figure 1.

I 21 —i—

20

It is clear that oxidation resistance decreases strongly with additions of >30wt% Co. Interestingly, addition of 10wt% Co results in an increase in oxidation resistance compared to the binary alloy. It is possible that the addition of small amounts of cobalt to the alloy significantly reduces the iron activity, lowering the driving force for Fe diffusion through the scale according to Fick's laws [9].

23

4

— i —

40

jj —i—

60 80 2Theta (degrees)

100

Figure 2. XRD spectra of oxidation products on Ni-Fe-Co alloys following 24h oxidation at 800°C in air. l-NixFe3.x04, 2-Co304, 3-NixCo3.x04, 4-CoxFe3.x04, 5-CoO.

1124

It has been well documented that hematite (Fe 2 0 3 ) is formed on Ni-Fe alloys containing as little as 15wt% Fe under the chosen conditions [7, 14-15]. In the present study, no perceptible amounts of Fe 2 0 3 were observed following pre-oxidation. It is likely that Fe 2 0 3 formation is suppressed due to preferential oxidation of cobalt, resulting in scales containing cobalt oxides and cobaltite spinels. Of course, it remains to be seen whether these cobalt-rich phases offer any advantages as protective barriers during electrolysis. Promisingly, NiFeCoio appears to offer the dual advantages of high oxidation resistance and a strong preference for nickel ferrite formation.

0

20

40

60 80 Time (min)

100

120

20

40

60 80 Time (min)

100

120

Electrolysis under high alumina concentration Galvanostatic electrolysis was performed using pre-oxidised Ni-Fe-Co anodes for 2h at 0.8A/cm2 (absolute current = 1.41 A). The initial alumina concentration in the electrolyte was a nominal 7.0wt% (the saturation point was estimated as 7.9wt% using models developed by Grjotheim and Welch [16]). Assuming a 90% current efficiency, it was predicted that the alumina concentration would decrease by no more than lwt% throughout the experiment.

Where

Å^^ Ecathode Ranode Rcathode Rbath Rext i η

+

^ b a t h + i^ext + ¹

3.5

g

2.5

LU

2.0 1R

I

I

I

d

Ecell = E a n o de ~" ECathode + i^anode + iRorthode

<

I

,

> 4·°

Potential versus time plots for each anode are shown in Figure 3. Potentials are reported with reference to the aluminium deposition potential, with the graphite/aluminium cathode serving as a quasireference electrode. It should be noted that the reported cell voltages include ohmic contributions. Since the electrical conductivity of the bath does not change dramatically with A1 2 0 3 concentration [17-19] and the cathode reaction does not change, it may be reasonably assumed that fluctuations in the cell potential reflect changes in the anode processes or condition. The electrochemical potential was estimated by approximating the voltage contributions due to non-faradaic processes. The total cell voltage is given by Equation 1:

I

0

20

I

40 60 80 Time (min)

I

100

120

Figure 3. Potential vs. time plots for: A-NiFeCoi0, B-NiFeCo30 and C-NiFeCo50 anodes. Temperature: 960°C, current density: 0.8A/cm2, interelectrode distance: 4cm.

(Ό

Table 2. List of possible reactions on Ni-Fe-Co anodes in cryolite-alumina melts E at 1000°C (V) mO(*f Reaction Theoretical* Measured Kei

is the reversible anodic potential is the reversible cathodic potential is the anodic resistance is the cathodic resistance is the bath and bubble layer resistance is the resistance due to external components such as leads and connections is the cell current is the overvoltage

Additional voltage requirements due to anode and cathode and resistances are expected to be low: the individual materials resistivities are in the order of Ι-ΙΟμΩ.αη at the operating temperature. Voltage requirements due to contact resistances were estimated to be in the order of 0.1V. The bath resistivity was estimated at 0.5Ω.αη using models developed by Hives' et al. [1719], contributing an additional 0.15V. The overpotential at 0.8A/cm2 was estimated between 0.3-0.5V, based on data published by Thonstad et al [20]. The remaining potential difference, representing the electrochemical potential, changes considerably depending on the anode material and condition. In most cases, potentials above the oxygen evolution potential (2.2V vs AIF3/AI) are observed. As such, we can be reasonably confident that oxygen evolution is one of the primary anodic reactions.

3Ni +2A1F3 -+ 3NiF 2 + 2A1 3Ni + A1 2 0 3 - * 3NiO + 2A1

1.69 1.55

1.40 1.48f

[21]

3Fe +2A1F3 -> 3FeF 2 + 2A1 3Fe + A1 2 0 3 — 3FeO + 2A1 2Fe + A1 2 0 3 — Fe 2 0 3 + 2A1 9Fe + 4A1 2 0 3 —3Fe 3 0 4 + 8A1 3Co +2A1F3 — 3CoF 2 + 2A1 3Co + A1 2 0 3 -> 3CoO + 2A1 9Co + 4A1 2 0 3 ->3Co 3 0 4 + 8A1

1.27 1.26 1.35 1.27

0.85

[21]

1.52 1.45 1.65

1.25

2A1 2 0 3 - * 3 0 2 + 4A1

2.19

2.20

1.16*

-

[4]

[4] [21]

[22]

2A1F3 - * 3F 2 + 2A1 4.07 $ Calculated using HSC software, licensed to Outokumpu Research. t Measured at 970°C Table 2 provides a non-exhaustive list of possible cell reactions and their theoretical and measured potentials at 1000°C. While oxygen evolution is likely to be the primary anodic reaction at most stages of electrolysis, it is probable that a number of other reactions occur - most notably, oxidation and fluoridation of the anode metal. It is clear that the nature of the primary anodic

1125

reaction is significantly influenced by the relative activities of AI2O3 and AIF3 in the melt. Under high A1203 activity, oxygen-rich products - such as metal oxides and oxygen gas - are likely to dominate. Under low A1203 activity, fluoridation reactions dominate, usually leading to accelerated anode wear and in most cases, passivation. This indicates why it is necessary to maintain a high alumina concentration during cell operation with inert anodes. As well as reducing the dissolution rate of the protective oxide scale, a high alumina activity suppresses fluoridation reactions, which in turn, extends the lifetime of the anode. Of course, without the assistance of additional analytical techniques, it is difficult to speculate about the types of reactions occurring at the anode surface. Figure 3 suggests that NiFeCoio and NiFeCo30 maintained steady oxygen evolution for the duration of the experiment. In each case, stable anodic voltages were observed - approximately 3.25V and 3.75V respectively. It is unclear why a significantly higher voltage was observed for NiFeCo30. It is possible that the thicker pre-formed scale on NiFeCo30 causes a higher anodic resistance. Alternatively, the formation of an electrically resistive species may have occurred at the anode surface. For NiFeCo10, intermittent rapid drops in potential were observed during the first hour of electrolysis. The potential drops can be seen as small spikes on the voltage plot in Figure 3. In each case, the potential drops rapidly by 0.1-0.2V, then returns slowly to the steady-state voltage over 20-30sec. It is likely that these events represent spallation of parts of the scale from anode surface, possibly as a result of gas evolution. The potential returns to the steady-state value as the oxide is re-formed in situ. Interestingly, the spalling events cease in the final hour of electrolysis, possibly indicating complete destruction of the original scale. For improved anode performance, it may be necessary to significantly reduce thefrequencyof spalling by improving the scale adhesion. NiFeCo50 performed poorly during electrolysis, exhibiting highly unstable behaviour. For the first 45min of the experiment, the anode maintained a steady potential of approximately 3.35V. Following this, the voltage increased sharply to ~4V, then declined slowly to -2.1V with significant noise. Oxygen evolution ceases in the last 20min of electrolysis as the potential drops below 2.2V. It is likely that the primary anodic reaction during this period is metal dissolution.

1

The relative thicknesses of the remaining scales offer information about the solubilities of the oxidation products in the cryolitealumina bath (or, more accurately, the relative rates of oxide formation and dissolution). Despite having significantly lower oxidation resistance, the cobalt-rich anodes NiFeCo30 and NiFeCo50 have relatively thin reaction scales (50-200μιη). In comparison, NiFeCo10 has a thick scale (400-500μπι): over 25 times greater than its starting thickness, as assessed by optical microscopy. This suggests that oxide phases formed on NiFeCo10 - principally NixFe3.x04 and NiO - are significantly less soluble than cobalt-containing oxides. XRD analysis was performed on the scale/bath interfaces following electrolysis. The spectra are shown in Figure 5. It is clear that the surface compositions have changed considerably over the course of electrolysis. In particular, there is a notable absence of cobalt-rich phases (Co304, CoO, CoxFe3.x04, NixCo3.x04) in preference for nickel-rich phases (NiO, NixFe3.x04). This reflects the significantly lower solubility of Nibased oxides in the electrolyte. Iron (II) fluoride, FeF2, was found in significant quantities on the surface of the NiFeCo30 anode. Most metal fluorides are poorly conductive, and can contribute to anode passivation if deposited in significant concentrations on the anode surface. The average anode potential during electrolysis was not unusually high (see Figure 3), suggesting that FeF2 has minimal effect on the scale conductivity when present in these concentrations.

m%\

||_A_ 1

Following electrolysis, the anodes were mounted in epoxy resin and sliced vertically to reveal a cross-section of the electrode centre. Optical micrographs of the cross-sections are shown in Figure 4. While no significant changes in the anode dimensions were observed, it was clear that the protective scales had become severely damaged during the experiment. In all cases, the scales had delaminated from the anode surface, causing massive bath penetration and metal wear. The bath appears to have penetrated 500-800μηι into the alloy, leaving the metal etched at the grain boundaries. Importantly, it appears to be the action of the bath, and not thermal shock due to re-heating of the scales, that causes delamination of the oxide layer. Inspection of the portion of the anodes which had not been in contact with the bath showed that the oxide scales in these regions were still well adhered to the metal.

* X

*p|

jfl BJf B%g SOOum

K Μ S

■ 111 M SOOum

WM

m I

^^β SOOum

Alloy >

Figure 4. Optical micrographs of Ni-Fe-Co anode cross-sections following 2h electrolysis. A-NiFeCo10, B-NiFeCo30, C-NiFeCo50.

1126

1

were observed, including NiO and Ni x Fe 3 . x 0 4 . From these results, it is clear that cobalt oxides have relatively high solubility in the bath.

15kH 2,4

NiFeCcu <

UWJWJLMHXIMMJVIA^

^ I ^ ' > H ^ W ^ » M *V W *AJ

NiFeCo

9

9

NiFeCo i n — i — « — i — ' — i — " — i —

20

30

40

50

— i — i — i —

60

70

— i — ' — i — · — i — ·

80

90

100 110

2"ftieta (degrees)

Figure 5. XRD spectra of scale/bath interface on Ni-Fe-Co anodes following 2h electrolysis at 0.8A/cm2. 1-Cryolite, 2-Chiolite, 3-NixFe3_x04, 4-Co 3 0 4 , 5-CoAl 2 0 4 , 6-CoO, 7-NiO, 8-FeF2,9-FeNi3.

Following electrolysis, the graphite crucible cathode was potted in epoxy resin and sliced in half to reveal a vertical cross-section of the cell. A photograph of the halved cell is shown in Figure 7. To gain information about the distribution of elements within the cell, EDX point analysis was performed on selected areas of the crosssection. There are several regions of interest: first, the "U" shaped area of dark material marked 3 on the photograph. This area appears to represent regions of reduced metal which have been deposited on the sides of the crucible, and are trickling down toward the large mass of metal at the base of the cell. Interestingly, the EDX spectrum indicates that the metal deposit does not contain appreciable amounts of aluminium, but is instead composed of iron from the anode and small amounts of chromium from the stainless steel connecting rod. The lump of reduced metal at the base of the cell is approximately twice the size of the aluminium button added to the crucible prior to electrolysis. The round - almost circular - shape of the deposit indicates the very high surface tension between the liquid aluminium pad and the graphite crucible. The deposit appears to consist of two distinct regions: an outer, more porous region and an inner dense region. EDX analysis revealed that the outer region is predominantly aluminium, while the inner region is represented by re-deposited metal from the dissolved anode. At the cell operating temperature of 960°C, the intermetallic compounds of nickel, iron and cobalt are well below their liquidus temperatures. Re-deposition of the dissolved anode metals would result in a solid product, having much higher density than liquid aluminium. As a consequence, we see most of the metal contaminants being segregated from the aluminium product.

Extended electrolysis without alumina replenishment Extended electrolysis using pre-oxidised NiFeCo30 was terminated after 4h due to catastrophic failure of the anode. Figure 6 shows the potential versus time plot for the cell. The anode maintained a steady potential of approximately 4.5V for the first 3 hours of electrolysis. It can be reasonably assumed that oxygen evolution occurs during this time. In the final hour of electrolysis, the potential increased steadily to >10V. This indicates a dramatic rise in anode resistivity, almost certainly caused by fluoridation of the metal. Based on an assumed current efficiency of 90%, the alumina depletion rate was predicted as approximately l.lg/hr, or 1.2wt%/hr. Assuming the prediction is accurate, the bulk alumina concentration at the onset of anode failure (approx. 3h electrolysis time) was around 3.4wt%. However, it is likely that the oxide concentration at the electrode surface is much lower, giving rise to concentration polarisation and the onset of an anode effect. It has been shown that the solubility of iron and nickel oxides is 3-10 times higher in baths containing 3wt% A1 2 0 3 than 7wt% A1203 [3, 23]. Under these conditions, complete destruction of the protective scale is anticipated. As a consequence, the anode metal becomes exposed to the bath and is subsequently dissolved.

Interestingly, the lump of re-deposited anode metal appears to contain only a small amount of cobalt. Clearly, the Ka peak intensities of the EDX spectrum (Point 1 in Figure 7) are not proportional to the ratio of Ni, Fι and Co in the original alloy. Further EDX analysis revealed that cobalt was undetectable in other regions of the crucible. Furthermore, the XRD analysis of the damaged anode revealed virtually no cobalt remained in the alloy. A crude mass balance for the system would suggest that a 12-

10H

>

Following electrolysis, the anode was cooled and removed from the cell for visual inspection. As expected, the dimensions of the anode had been significantly reduced, indicating wholesale dissolution of the metal in the bath. In comparison, the NiFeCo30 anode dimensions were virtually unchanged after 2h electrolysis. Hence, it is clear that the majority of anode wear occurs in the 24h period, when the alumina concentration drops below 5wt%.

§

8

-

>

/ * " ■ * -

LU

XRD analysis was performed on the remaining part of the destroyed anode. It was found that the anode consisted almost entirely of metal oxide species and solidified bath material. This helps to explain the high anodic resistance observed in the final hour of electrolysis. Interestingly, despite the alloy's high initial cobalt concentration, the remaining part of the anode contained no detectable amounts of cobalt oxide. Instead, nickel-rich oxides

1127

0

1

2 Time (hours)

Figure 6. Potential versus time plot for oxidised NiFeCo30 anode, subjected to extended electrolysis at 0.8A/cm2 without alumina replenishment.

—i—i—i—i—|—i—I—i—I—i—I—

—T

1

1

1

1

1

1

1

r—

Al

.

ttUL&y

Ni.

Energy (KeV)

3 4 5 Energy (KeV)

6

7

Figure 7. Left- photograph of crucible cross section following 4h electrolysis without alumina replenishment. Right- EDX spectra of selected points on the cell cross section. [6]

significant amount of cobalt has been lost, probably in the form of a volatile product. Conclusions

[7] [8]

The anodic behaviour of pre-oxidised Ni-Fe-Co alloys was investigated during short-term aluminium electrolysis. A general correlation exists between the oxidation resistance of the anode metal and the short-term anode performance. In particular, anode compositions with a preference for NixFe3.x04 formation must be considered promising, owing to the very low solubility of the spinel in the electrowinning bath. It was demonstrated that the preference for NixFe3.x04 formation can be strongly increased by the addition of 10wt% Co to the Ni-Fe system. Alloys containing >30wt% Co performed poorly due to high corrosion rate and relatively high solubility of cobalt-containing oxides.

[9] [10] [11] [12] [13] [14]

Visual inspection of the anodes following electrolysis showed that all alloys were inadequately protected by the pre-formed oxide scale. Bath penetration through pores and defects in the scale was shown to result in irreversible anode damage. Therefore, there is a strong need to improve scale adhesion and compactness, possibly by appropriate physical pre-treatment of the anode surface or incorporation of small quantities of rare earth elements to the alloy. The need to maintain alumina concentration at close to saturation was highlighted during extended electrolysis.

[15] [16] [17] [18]

References

[19]

[1] D.H. DeYoung, Light Metals 1986, R.E. Miller Editor, TMS, Warrendale PA (1986) 299. [2] S. Pietrzyk and R. Oblakovski, Light Metals 1999, CE. Eckert Editor, TMS, Warrendale PA (1999) 407. [3] T.E. Jentoftsen, O.-A. Lorentsen, E.W. Dewing, G.M. Haarberg, and J. Thonstad, Metall. Mater. Trans. B, 33 (2002)901. [4] P.G. Russell, J. Appi. Electrochem., 16 (1986) 147. [5] E.V. Antipov, A.G. Borzenko, V.M. Denisov, A.Y. Filatov, V.V. Ivanov, S.M. Kazakov, P.M. Mazin, V.M. Mazin, V.l. Shtanov, D.A. Simakov, G.A. Tsirlina, S.Y. Vassiliev, and Y.A. Velikodny, Light Metals 2006, T.J. Galloway Editor, TMS, Warrendale PA (2006) 403.

[20] [21] [22] [23]

1128

D.A. Simakov, E.V. Antipov, M.I. Borzenko, S.Y. Vassiliev, Y.A. Velikodny, V.M. Denisov, V.V. Ivanov, S.M. Kazakov, Z.V. Kuzminova, A.Y. Filatov, G.A. Tsirlina, and V.l. Shtanov, Light Metals 2007, M. S0rlie Editor, TMS, Warrendale PA (2007) 489. R.T. Foley, J. Electrochem. Soc, 109 (1962) 1202. B.J. Welch, Light Metals 2009, TMS, Warrendale PA (2009)971. A. Fick, Phil. Mag., 10 (1855) 30. D.W. Luo and Z.S. Shen, Acta Metall. Sinica, 21 (2008) 409. I.E. Klein, A.V. Yaniv, and J. Sharon, Oxidation of Metals, 16(1981)99. J. Wu, L. Zhang, J. Zhou, and Y. Xu, /. Mater. Sci. Tech., 16 (2000)509. B.-S. Kim, B.-G. Kim, H.-W. Lee, and W.-S. Chubg, Metals and Materials International, 8 (2002) 367. G.L. Wulf, T.J. Carter, and G.R. Wallwork, Corr. Sci., 9 (1969) 689. I.A. Menzies and J. Lubkiewicz, /. Electrochem. Soc, 117 (1970)1539. K. Grjotheim and B.J. Welch, Aluminium Smelter Technology: A Pure and Applied Approach. 1980, Düsseldorf: Aluminium-Verlag. J. Hives, J. Thonstad, Â. Sterten, and P. Fellner, Light Metals 1994, U. Mannweiler Editor, TMS, Warrendale PA (1994)187. J. Hive§, J. Thonstad, Â. Sterten, and P. Fellner, Metall. Mater. Trans. B, 27B (1996) 255. J. Thonstad, P. Fellner, G.M. Haarberg, J. Hive§, H. Kvande, and Β. Sterten, Aluminium Electrolysis: Fundamentals of the Hall-Hιroult Process. 3rd ed. 2001, Düsseldorf: Aluminium-Verlag. J. Thonstad, A. Kisza, and J. Hives, Light Metals 2006, T.J. Galloway Editor, TMS, Warrendale PA (2006) 373. L. Cassayre, P. Chamelot, L. Arurault, and P. Taxil, J. Appi. Electrochem., 35 (2005) 999. J. Thonstad, Electrochim. Acta, 13 (1968) 449. T. Utigard, Light Metals 1993, S.K. Das Editor, TMS, Warrendale PA (1993) 319.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECTS OF THE ADDITIVE Zr0 2 ON PROPERTIES OF NICKEL FERRITE CERMET INERT ANODE Xiao Zhang, Guangchun Yao, Yihan Liu, Jia Ma, Zhigang Zhang School of Materials and Metallurgy, Northeastern University, No. 11 Lane 3 Wenhua Road, Shenyang, Liaoning, 110004, China Tel.:+86-024-83686462, E-mail: [email protected] Key words: Cu-Ni-NiO-NiFe204 composite ceramic, Zr0 2 , bend strength, relative density, corrosion resistance of the produced metal Al can not be guaranteed and more Abstract

importantly, the use of inert anodes is not economically feasible. More recently, Aluminum Company of America (Alcoa)

The objective of this paper is to study a new attempt on

conducted, with support by US Department of Energy, having

preparing Cu-Ni-NiO-NiFe204 ceramic inert anodes by adding

arrived

Zr0 2 (0~1.50wt %), carrying out high temperature solid-state

17Cu-42.91NiO-40.09Fe2O3 to provide acceptable electrical

process. NiFe204 spinel, the matrix material, was prepared

conductivity combined with good corrosion resistance [5].

firstly with extra 18wt% NiO and Fe 2 0 3 as the raw materials.

Steven C. Marschman applied for a U.S. Patent 4,871,437 in

The product was crushed to fines, and then mixed uniformly

1989, having arrived at a new cermet material consisting of a

with Cu-Ni and Zr0 2 powders to prepare Cu-Ni-NiO-NiFe204

nickel ferrite-nickel oxide ceramic containing copper and nickel

ceramic inert anodes by cold-pressing sintering method. The

as metals phase to provide a stable electrode to have

impact of Zr0 2 addition on the relative density, the bend

significantly improved electrical conductivity [6]. A U.S. Patent

strength and the corrosion resistance of Cu-Ni-NiO-NiFe204

5,794,112 applied by Siba P. Ray in 1998, describes the

at

a

new

cermet

material

consisting

of

ceramic inert anodes was investigated. The results show that,

sintering atmosphere preparation of cermet. The gaseous

with the addition of 0.5wt% Zr02, the relative density slightly

atmosphere has an oxygen content that is controlled at about

increases and the corrosion resistance decreases a little while the

5-3000 ppm, preferably about 5-700 ppm and more preferably

bending strength improves remarkably from 55.50MPa of the

about 10-350 ppm in order to obtain a desired composition in

sample without Zr0 2 additive to 105.26Mpa.

the resulting composite [7]. The former studies mainly focused on the conductivity and corrosion resistance of cermet but little on the additive and mechanical properties. In this article, we use

Introduction

Zr0 2 as additive to improve the bending strength of inert anodes and to study various properties of inert anodes by the

In conventional aluminum electrolysis, using carbon anodes has

experiment.

many disadvantages such as the consumption of carbon anode and the emission of carbon dioxide and fluorocarbon. So the

Experimental

appearance of inert anodes draws people's attention [1]. The electrolyzing inert anodes will release oxygen [2, 3], being

Preparation of the matrix material

friendlily to environment, which makes the use of inert anodes commercially attractive. Numerous researches have been done

The matrix material NiFe204 spinel was prepared using NiO

on finding the appropriate material for inert anodes in aluminum

with excessive content of 18wt% and Fe 2 0 3 as the raw materials.

production. Nowadays inert anodes can't still meet some basic

The two oxides were firstly cold compacted to produce billets to

requirements [4]. For example, inert anodes exhibit a high

the pressure of 60MPa. Secondly the billets were sintered with

corrosion rate in the high-temperature cryolithe melts, low conductivity and environment of high oxidizability. The purity

the temperature of 1000°C for 6 h.

1129

The changes of relative density and porosity with the content of Zr0 2 are shown in Table.I. It can be seen from Table.I that the

Preparation of Cu-Ni-NiO-NiFeoOA ceramics

relative density decreases to some extent while the porosity NiFe204 spinel was grinded to fines particles. Subsequently the

gradually goes up in the range of 0.5-1.5wt.% Zr0 2 . The rules

granular spinel was uniformly mixed with Cu-Ni (17wt.%) and

of the changes were not obvious, however, it can be determined

the additive of Zr0 2 with different contents. The contents of

that the additive of Zr0 2 is favorable for increase of anodes'

Zr0 2 were 0.0 wt.%, 0.5 wt.%, 1.0 wt.% and 1.5wt.%

relative density and decline of their porosity.

respectively. Finally the mixed powders were cold compacted using uni-axial compaction to the pressure of 200MPa. The

Effect of ZrQ2 content on ceramics bending strength

Cu-Ni-NiO-NiFe204 ceramic inert anodes were made through the sintering process at the temperature of 1200°C for 6 h. The final products were rectangular billets (60 mmxl2 mmx8 mm).

The effect of Zr0 2 content on bending strength is shown in Fig.l. It is obvious that bending strength is dramatically enhanced by adding Zr0 2 . Bending strength of samples without Zr0 2 is 55.5MPa while the bending strength of sample with 0.5wt.% Zr0 2 is 105.26MPa which is 1.897 times as much as

Performance tests

that of sample without Zr0 2 . And it is the highest bending The density was measured by Archimedes' principle. The crystal

strength of all the samples.

structure of samples doped with Zr0 2 were identified by powder X-ray diffraction (XRD, PW3040/60, Holland) with CuKa radiation, 2Θ in the range of 10°~90°with a step of 0.04°. The fractured surfaces were examined by scanning electron microscopy (SEM, SSX-550, Japan). The samples were immerged in molten cryolite with the temperature of 960 D for 10 h to determine the static state corrosion rate. The electrolyte was made up of reagent grade CaF2 and A1F3, technical grade NaF and A1203. The CR (molar ratio of NaF to A1F3) was kept to be 2.8, and the concentrations

0.5 1.0 Content of ZrO (WT.95

of CaF2 and A1203 were both kept to be 5wt.%. The eroded samples were washed in 30wt% A1C13 solution at 100 D to

Fig.l Effect of Zr0 2 on bending strength of samples.

remove the adhering melt. The corrosion rate was determined by mass loss measurement [8].

l-NiFe 2 0 4 (Ni x Fe 3x 0 4 ) 2-NiO 3-Cu3gNi

Results and analysis

4-Zr0 2 4 2

Effect of ZrQ2 content on ceramics relative density Table.I Effect of Zr0 2 on relative density and porosity of samples. the content of

relative density/%

porosity /%

0.0

99.37

0.63

|

0.5

99.63

0.37

|

1.0

99.50

0.50

|

1.5

99.56

0.44

|

Zr02/% 2È ( ·

)

Fig.2 XRD pattern of the sample with Zr0 2 .

1130

Fig.2 shows the XRD pattern of the sample with Zr0 2 .

crystal system. The furnace temperature gradually goes down after the

It is clear from Fig.2 that Zr0 2 does not enter into the matrix

completion of the sintering progress. Volume expansion of Zr0 2

material, and the crystal transformation is taken by itself.

is made impossible due to the compacted matrix material. Therefore, a part of Zr0 2 can be preserved in the matrix material as tetragonal crystal system. It can be concluded that part of Zr0 2 at room temperature still belongs to tetragonal crystal system. The bending strength is improved mainly due to Zr0 2 having tetragonal crystal system. The crack begins to appear as the sample containing tetragonal crystal Zr0 2 is stressed. It can be sure that conversion of Zr0 2 from tetragonal crystal system to monoclinic crystal system is inevitable due to appearance of the crack. The phase transformation of Zr0 2 results in the expansion of Zr0 2 and the extension of crack is effectively prevented. A greater force is employed to destroy the sample. Nevertheless, the content of Zr0 2 of the sample is finite. The sample would be destroyed due to excessive volume expansion of Zr0 2 particle if the content of Zr0 2 exceeds the upper limit. The optimum content of Zr0 2 is 0.5wt.%.

h Fig.3 SEM photographs of the sample containing Zr0 2 with (a)

0.5wt.% Z r 0 2 /

/

I

0.0wt.% and (b) 0.5wt.%.

/ 1

The microstructures of samples without and with 0.5wt.% Zr0 2 are shown in Fig.3 (a) and (b) respectively. The grain sizes with

Without Zr0 2

/

different samples are analogous to each other. However, the intergranular porosity of the sample without Zr0 2 is much

1

greater than that of the sample with 0.5wt.% Zr02, as market by

I

0.0

the arrows in Fig.3. The Zr0 2 particles are mainly distributed at intercrystalline.

y

. 0.2

0.4

1

^

/ s^"'' I

/

0.6

•

0.8

Deflection(mm)

I

1.0

I

I

1.2

Fig.4 Three-point bending curves of samples.

As is well known, the melting point of Zr0 2 is 2715 D and Zr0 2 has three different crystal types including cubic crystal system,

The three-point bending curves of samples without Zr0 2 and

tetragonal crystal system and monoclinic crystal system.

with 0.5wt.% Zr0 2 are shown in Fig.4. Under the force, the

Monoclinic crystal system of Zr0 2 at ambient pressure is

sample without Zr0 2 begin to crack under the force while a

converted into tetragonal crystal system at 1170°C and

platform on the three-point bending curve of the sample with 0.5wt.% Zr0 2 is visible. Volume expansion of Zr0 2 particle starts to be taken due to the transformation from tetragonal

tetragonal crystal system to cubic crystal system at 2370°C.

crystal system to monoclinic crystal system. The sample with 0.5wt.% Zr0 2 can undertake a much greater force.

Zr0 2 crystal converts to and fro between monoclinic crystal system and tetragonal crystal system as the sintering temperature is 1200°C, which is lower than that of tetragonal

Effect of ZrO? content on corrosion resistance

1131

The samples with different content of Zr0 2 eroded by molten

increase of the corrosion rate having Zr0 2 . Zr0 2 particle has

salt are shown in Fig.5. And the effect of Zr0 2 content on

enough space to change from tetragonal crystal system to

corrosion rate is shown in Fig.6. It is obvious that the corrosion

monoclinic crystal system as the corrosion of the matrix

resistance can be degraded by adding Zr0 2 . The corrosion rate

material result in the extension of crack. Meanwhile the Zr0 2

2 _1

of the sample without Zr0 2 is 0.015g-cm" h . The corrosion

particle starts volume expansion lead to more matrix materials

rate of the sample with 0.5wt.% Zr0 2 is 0.028 gcm" 2 h _1 and

crack. Thus corrosion rate of the sample apparently increases.

2 _1

with 1.5wt.% Zr0 2 is 0.040g-cm~ h . And the higher the Zr0 2 content, the higher the corrosion rate. As the content of Zr0 2 is 1.5wt.%, the corrosion rate is the highest.

Fig.5 Photographs of the samples eroded by molten salt containing Zr0 2 with the content of (a) 0.0wt.%, (b) 0.5wt.%, (c) Fig.7 SEM micrograph of the distribution on the polished

1.0wt.%and(d) 1.5wt.%.

section of NiFe204 without Zr0 2 eroded by molten salt.

Cont ent

of

Zr Q, (

Fig.6 Corrosion rate of the samples with different content of Zr0 2 . The volume expansion of Zr0 2 particle is the reasons for the

1132

respectively. The higher the Zr0 2 content, the higher the corrosion rate is. Acknowledgements The authors gratefully acknowledge financial support from National Natural Science Foundation of China (No. 50834001) and the National High Technology Research and Development Program of China (No. 2009AA03Z502).

Fig.8 SEM micrograph of the distribution on the polished section of NiFe204 with 1.5wt.%Zr02 eroded by molten salt. The SEM micrographs of the distribution on the polished section of NiFe204 without Zr0 2 and with I.5wt.% Zr0 2 are

References

shown in Fig.7 and Fig.8 respectively. As seen from the two figures, the diffused layer is dramatically reduced by adding Zr0 2 . The diffused layer has fine corrosion resistance. The diffused layer is the destroyed by volume expansion of Zr0 2 particle. The corrosion resistance decrease resulting from Zr0 2 .

[2] R. P. PAWLEK, "Inert anodes: An update," (Light metals

1) The relative density of the sample with Zr0 2 decreases to some extent while the porosity is on the contrary. The relative density of the sample with 0.5wt.% Zr0 2 may arrive at 99.67%. 2) The bending strength is dramatically enhanced by adding Zr0 2 . Bending strength of sample without Zr0 2 is 55.5MPa. The bending strength of sample with 0.5wt.% Zr0 2 is up to 105.26MPa and is 1.897 times as much as that of sample without Zr0 2 . 3) The corrosion resistance is not significantly improved by adding Zr0 2 . The corrosion rates of the samples containing Zr0 2 with the content of 0.0wt.%, 0.5wt.% and I.5wt.%is are 0.028gcm_2-h-1

materials as inert anode," (Light metals proceeding of sessions, TMS Annual Meeting, USA, The Minerals, Metals and Materials Society, 1996), 249-257.

Conclusions

0.015g-cm-2h_1,

[1] E. Olsen, J. Thonstad, "Behaviour of Nickel ferrite cermet

and

0.040g-cm"2-h~1

proceeding of sessions, TMS Annual Meeting, USA, The Minerals, Metals and Materials Society, 2002), 449-456. [3] D. R. SADOWY, "Inert anode for the Hall-Hιroult cell: The ultimate materials challenge," JOM, 2001,53(5), 34-35 [4] E. Olsen, J. Thonstad, "Nickel ferrite as inert anodes in aluminium electrolysis (Part I): Material fabrication and preliminary testing," Journal of Applied Electrochemistry, 1999, 29(3), 293-299. [5] J. D. Weyand, D. H. DeYoung, S. P. Ray, G P. Tarcy and F. W. Baker, "Inert Anodes for Aluminium Smelting, Final Report," (Aluminium Company of America, Alcoa Laboratories, Alcoa Center, 1986), DOE No. DOE/CS/40158-20, Department of energy, Idaho operations office, Idaho Falls, ID (1986)

[6] C. S. Marschman, C N. Davis, "Cermet anode with continuously dispersed alloy phase and process for making," US Patent, 4871437, 1989. [7] S P. Ray, R W Woods, Controlled atmosphere for fabrication of cermet electrodes, US Patent, 5794112,1998. [8] J. H. Xi, Y. J. Xie, G C. Yao, Y. H. Liu, "Effect of additive on corrosion resistance of NiFe204 ceramics as inert anodes," Transactions of Nonferrous Metals Society of China, 18 (2) (2008), 356-360.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Effect of sintering atmosphere on phase composition and mechanical property of 5Cu/ (10NiO-NiFe2O4) cermet inert anodes for aluminum electrolysis Zou Zhong, Wei Chenjuan, Tian Zhongliang, Liu Kai, Zhang Hongliang, Lai Yanqing, Li Jie School of Metallurgical Science and Engineering, Central South University, Changsha 410083, China Key word: inert anode, 5Cu/ (1ONiONiFe204) cermet, sintering atmosphere, phase composition, mechanical property inert anode because of its high melting point, excellent

Abstract

corrosion resistance, and stable thermal and chemical properties [1 4]

" . With the support by US Department of Energy, Aluminum

5Cu/(10NiO-NiFe2O4) cermet anodes were prepared by the cold-pressing sintering method in different

Company of America (Alcoa) conducted a considerable work

atmospheres.

about cermet inert anode material. However, the test results of

Furthermore, their phase composition, microstructure and

6kA pilot cell scale showed that both the corrosion resistance

mechanical properties were also investigated. The results reveal

and mechanical properties of NiFe 2 0 4 cermet could not meet the

that 5Cu/ (10NiO-NiFe2O4) ceramic material can be obtained by-

requirements [>].

sintering in the vacuum (oxygen content is 0.02ppm) or atmospheres of Ar + air with oxygen content of lOppm, 200ppm, 2000ppm, and lOOOOppm respectively, and phase composition

As well known, the mechanical properties and corrosion resistance of ceramic which are closely related to its phase

of the ceramics varies with oxygen content of the sintering

composition and microstructure are affected by the sintering

atmosphere, greatly. Content of NiFe 2 0 4 in the ceramic

atmosphere of the preparation process.. Therefore, to improve

materials increases with the increase of oxygen content in the

the properties of NiFe 2 0 4 cermet, the materials were prepared in

sintering atmosphere, But content of phase Cu decreases with

different sintering atmospheres. The results showed that ceramic

the increase of oxygen content. When oxygen content in the

phase of NiFe 2 0 4 would be probably decomposed if the material

atmosphere is lOppm, the grain size of 5Cu/ (10NiO-NiFe2O4)

was prepared in vacuum or a reductive atmosphere. While the

cermet is 5.43 μπι, meanwhile the bending strength reaches

utilization of inert atmosphere not only improves the

80.05 MPa at the room temperature.

densification of material but also obtains the expected material. [6-10]

Introduction Nevertheless, investigations of mechanical property of NiFe 2 0 4 The traditional industrial production of aluminum has been

cermet inert anodes for aluminum reduction in sintering

challenged by its high energy consumption and huge amount of

atmosphere were few. In this paper, 5Cu/(10NiO-NiFe2O4)

greenhouse gas emission. The application of inert anode

cermet was prepared in the vacuum (vacuum degree is 1.0 X

replacing carbon anode in Hall-Hιroult electrolysis cells has

10"2Pa,and oxygen content is 0.02ppm) and atmospheres of Ar +

been considered as an effective approach to solve these

air with oxygen content of lOppm, 200ppm, 2000ppm, and

problems, as well as it can increase the production efficiency.

lOOOOppm (to make the results easy to discussion, the above

In recent years, cermet has become one of the most promising

respectively), and then, the phase composition, microstructure

inert anode materials, because it not only has good electrical

and mechanical property were studied. This paper aims to

conductivity but also can resist to high temperature corrosion of

provide technical guidance for optimization of preparation

atmospheres were expressed by G0, Gi, G2, G3 and G4

molten. NiFe 2 0 4 is often utilized as ceramic matrix for cermet

1135

process of NiFe 2 0 4 based ceramic materials. with 9% in G0. Moreover, in G3, content of phase NiFe 2 0 4 is

Experimental

71%, the highest of all, but content of Cu is only 6.2 %, which is lower.

Fabrication of anodes

ONiFe 2 0 4 ACu

c

A mixture of Fe 2 0 3 and NiO with the molar ratio of 1.35 was

°Δ

*

1

Λ

h

. » I.

A .

. »-

1

.A

K

.

i

i

A

.

JU . . -

*

ë

A

o

O

Q

D -I

1

prepared, and then calcined in a muffle furnace at 1200°Cfor 6h in a static air atmosphere to form 10NiO-NiFe2O4 ceramic powder. The synthesized powder and Cu powder were ground in a medium containing dispersant and adhesive. Finally, the mixed ceramic-metal powder was dried and was statically pressed into some bars (6mm X 5mm X 42mm)

D

Q

under 2 X

.

108Pa. Then they were sintered at 1350°C for 4h in different

I

atmospheres.

,

10

,

1

20

1

A .

,A

Λ

/U

,

1

30

L~ —

_*_Λ__Λ

k 1

1

40

1

50

1

,

60

_

. 1

1

1

70

G0 1

80

1

1

90

2È

Fig.l The X-ray diffraction patterns of qualitative analysis of

Measurement methods

5Cu/(10NiO-NiFe2O4)

Phase composition of ceramic material was analyzed by

(G0, Gi, G2, G3 and G4 in the figure repress samples obtained

Rigaku3014 X-ray diffraction system, and microstructure was

in different atmosphere)

analyzed by JSM-6360LV scanning electron microscope. In

Table 1 Relative content of phases by XRD quantitative

addition, the phase composition was measured for three times

analysis in 5Cu/ ( 10NiO-NiFe2O4) cermet

to insure the reliability of the results.

Phase

Three-point-bending strength of the sintered specimen was evaluated with a CSS-44100 electrical universal testing machine by using a span of 30mm and a cross-head speed of 0.5mm/min. Each bending strength data was achieved by the average value with testing for five bars.

Relative content of phases (%)

species

GQ

G]

Q2

Q3

Q4

NiFe 2 0 4

64.9

66.3

68.5

71

69.1

NiO

26

26.1

24.3

22.7

25.3

Cu

9

7.6

7.2

6.2

5.6

(G0, Gi, G2, G3 and G4 repress samples obtained in different

Result and discussion

atmosphere. Data of phase content was the average value of three repeated experiments under the same condition.)

Phase and microstructure

The reason of the above results can be expressed by the

The XRD phase and its quantitative analysis results of

following reaction (1),

5Cu/(10NiO-NiFe2O4) cermet sintered in different atmospheres (G0, Gi, G2, G3 and G4) are shown in Fig. land table 1,

NiFe 2 0 4 NiFe204_x+ — 0 2

2

respectively. From the fig. l,it is found that all the samples

(1)

obtained in five different sintering atmospheres contain phases

It needs to point out that the stoichiometric compound

of Cu, NiFe 2 0 4 and NiO. Table 1 reveals that phase contents of

NiFe 2 0 4 and the nonstoichiometric compound NiFe204.x

NiFe 2 0 4 and Cu vary with the changing of 0 2 content in the

cannot be distinguished in the XRD analysis. Fe in NiFe204.x

.

has two forms of Fe2+ and Fe3+. When oxygen content in the

Moreover, content of phase NiFe 2 0 4 in the cermet sintered in

atmosphere increases, Fe2+ is oxidized to Fe3+, then Fe3+

atmosphere (G3 or G4) is higher than that in the cermet obtained

together with NiO transforms to NiFe204, as a result, phase

in other atmospheres. It is also noted that content of phase

content of NiFe 2 0 4 increases.

sintering atmosphere, which is also found in literatures

[1M2]

NiFe 2 0 4 is lowest with 64.9%, while content of Cu is highest

Furthermore, table 1 shows that phase content of NiO in G0 or

1136

G! is higher than that in G2, G3, and G4. The variation of phase

Moreover, phase of Cu almost disappears in the outer layer of

content of NiO among G2, G3 and G4 are very small. The

the cermet, and layered phenomenon takes place in the cermet

reason is that NiFe 2 0 4 phase will decompose partly in G0 and

when 0 2 content in the atmosphere is lOOOOppm.

Gi, which can be expressed by the following reaction (2). NiFe 2 0 4 <->NiO+Fe 2 0 3

(2)

Content of phase NiO is related with not only quantity of raw material, but also oxygen content in atmosphere. When oxygen content in atmosphere is higher, such as G2, G3 and G4 NiFe 2 0 4 phase will decompose no longer. Thus the variation of phase content of NiO from G2 to G4 is very small. In addition, phase content of Cu decreases with the increase of oxygen content in atmosphere. That is probably because that Cu phase is oxidized more easily in atmosphere with high oxygen content. Cu reacts with 0 2 , and forms Cu 2 0 or CuO, which can be expressed by reactions (3) and (4) as follow. Cu + I Oz «-> CuO 2

(3)

2Cu+ l o 2 <-> Cu 2 0 2

(4)

.Table 2 Oxygen partial pressure of different reaction by thermodynamic calculation at 1350°C Code name of

Oxygen partial

Oxygen content in

reaction

pressure /Pa

atmosphere /ppm

(3)

1428

2856

(4)

49

100

(Oxygen partial pressure was obtained by thermodynamic calculation, and oxygen content in atmosphere was transformed from oxygen partial pressure.) Oxygen partial pressure of reaction (3) and (4) at 1350°C are shown in table 2. It reveals that Cu will be partly oxidized to Cu 2 0 when oxygen content in atmosphere is between lOOppm and 2856ppm. When oxygen content in atmosphere is higher than 2856ppm, Cu phase will be partly oxidized, and form both Cu 2 0 and CuO. The microstructures

of

samples obtained

in

different

atmospheres are illustrated in Fig.2. The dark-gray region, light-gray region and white region represent NiFe204, NiO and Cu respectively. It shows that metallic phase of Cu and NiO in 5Cu/ (10NiO-NiFe2O4) cermet present polygon graphic, and distribute in phase NiFe204, independently. This phenomenon was also reported in literatures [13-15]. Moreover, metallic phase Cu in 5Cu/ (10NiO-NiFe2O4) cermet is well distributed when oxygen content in atmosphere is less than 2000ppm.

follow. Generally speaking, there are many factors affected on the bending strength, such as material grain size. Relative density and phase content. And the relation of all the factors can be expressed in the mixing rule [16] ,

af=ar+(am-ar)Vm

(5)

Vr+Vm=l

(6)

a f - Material bending strength, MPa

Fig.2 SEM images of 5Cu/(10NiO-NiFe 2 O 4 ) cermet in

a m - Bending strength of matrix phase, MPa

different sintering atmosphere ((a), (b), (c), (d) and (e) represent samples obtained in G 0 , G1? G2, G 3 and G4, respectively)

σ γ - Bending strength of strengthening phase, MPa Vm- Volume fraction of matrix phase, % Vy - Volume fraction of strengthening phase, %

Bending strength As is shown in table 3, the shrinkage of ceramic material

The bending strength decreases obviously with increase of 0 2

obtained in G 0 is the largest (12.8%), but the shrinkage

content in the atmosphere, though shrinkage and average grain

variation of material obtained in G b G 2 , G 3 and G 4 are very

of ceramic materials obtained in different sintering atmosphere

small. The result also reveals that grain sizes obtained in G 0

vary inconspicuous. It is the result of comprehensive influence

(5.76μιη), and G] (5.43μπι) are larger than that obtained in

of metallic phase

( a m ) , hard phase

( σ γ ) and grain size etc.

other atmospheres. Meanwhile gain sizes of cermet sintered in Conclusion

G 2 , G 3 and G 4 almost have no change. Therefore, oxygen content in atmosphere hardly influenced shrinkage and grain size of ceramic material. That is because that liquid phase sintering is used to prepare 5Cu/ (10NiO-NiFe 2 O 4 ) ceramic materials in the study.

Ceramic materials containing the promising phase can be obtained in sintering atmospheres with oxygen content of lOppm, 200ppm, 2000ppm and lOOOOppm, respectively. Phase content of NiFe 2 0 4 and Cu are affected greatly by the sintering atmosphere. In a word, Ceramic materials tend to have higher

Table 3 Shrinkage, grain size and bending strength of

phase content of NiFe 2 0 4 sintered under atmosphere with

5Cu/(10NiO-NiFe 2 O 4 ) cermet Performance

higher oxygen content, and higher phase content of Cu sintered

Sintering atmosphere

under atmosphere with lower oxygen content. When oxygen

test

G0

G!

G2

G3

G4

Shrinkage/%

12.8

11.65

11.9

11.35

11.35

5.76

5.43

4.74

4.78

4.77

when sintering atmosphere is the vacuum.

80.05

73.96

65.02

61.87

With the increase of oxygen content in sintering atmosphere,

Average grain diameter /urn Bending

72.8

strength /MPa

4

content in atmosphere is 2000ppm or lOOOOppm, content of phase NiFe 2 0 4 is higher. Phase content of Cu is the highest

the bending strength decreases significantly, but both the

(G 0 , Gi, G2, G 3 and G 4 repress samples obtained in different

shrinkage and grain size of 5Cu/ (10NiO-NiFe 2 O 4 ) almost have

atmosphere. Each test in table 3 was reproduced for five times.

no changes. Results of the above experiments show that when

Data in table 3 was the average value of these five repeated

oxygen

content

in

the

atmosphere

is

lOppm,

5Cu/(10NiO-NiFe 2 O 4 ) ceramic material presents the good

experiments. )

mechanical property, and the bending strength reaches the The bending strength decreases with the increase of oxygen content

in

atmosphere.

And

bending

strength

highest (80.05MPa).

of

5Cu/(10NiO-NiFe 2 O 4 ) obtained in G 0 is highest with a value of

Acknowledgement: The authors are grateful for the

80.05MPa. It is could be interpreted by the mixing rule as

financial support of the National High-Tech Research

and Development Program of China (2008AA030503). Reference

Central South University, (2005) (in Chinese) [10]A. Ghosh et al, "The effect of ZnO addition on the

[1] Y. X. Liu, "Advance on the research and development of

densification and properties of magnesium aluminum spinel".

inert anode and wettable cathode in the aluminum electrolysis".

Ceramics International, 2000, 26(6):605-608.

Light Metals, 5(2001), 26-29 (in Chinese)

[11] B. Li, R. Zhang, S. Wang. "Effect of sintering atmosphere

[2]D. R. Sadoway, "Inert anodes for the Hall-Hιroult cell: The

on the the dialectric properties of Er-Mg and Er-Mn doped

ultimate materials challenge". JOM, 53(5) (2001), 34-35

BaTiCVJournal of Inorganic Materials,22 (5),(2007), 821-827

[3]R.P. Pawlek, "Inert anodes: an update". Wolfgang Schneider.

[12] Yu Mao. "Effect of doping and sintering atmosphere on the

Light Metals, Warreudale PA, USA: TMS, 2002: 449-456

properties

[4]H. Xiao et al, "Studies on the corrosion and the behavior of

Acoustooptic, 2006, (28)1, 79-81

of

BiNb0 4

materials".

Piezoelectric

and

inert anodes in aluminum electrolysis". Metallurgical and

[13] S. Q. Zhou. "Microstructure and properties of A1203/A1

Materials Transactions B: Process Metallurgy and Materials

ceramic matrix composites". Special Metal Casting and

Processing Science, 27 (2) (1996), 185-193

Non-Ferrous alloys. (4) (1998), 41-43

[5]T. R. Alcorn, A. T. Tabereaux, N. E. Richards, "Operational

[14] W Li. "Study of microstructure and properties of Mo2FeB2

results of pilot cell test with cermet "inert" anodes". In: DAS S

cermet coating materials". Ceramics, (8) (2008), 18-20

K, eds. Light Metals, Warrendale, PA: TMS, 1993: 433-443

[15] M. Castillo-Rodriguez et al, "Effect of atmosphere and

[6]J. D. Weyand, "Manufacturing process used for the

sintering time on the microstructure and mechanical properties

production of inert anodes". Light Metals, TMS, Warrendale,

at high temperatures of -SiC sintered with liquid phase

Pa, 2 (1986), 321-339

Y203-A1203". Journal of the European Ceramic Society, 26

[7] L. Zhang et al, "Effect of atmosphere on the densification in

(2006), 2397-2405

sintering nickel ferrite ceramic for aluminum electrolysis".

[16] B. D. Flinn et al, "Fracture Resistance Characteristics of a

Science and Engineering of Powder Metallurgy, (9) (2004),

Metal-toughened Ceramic". Journal of American Ceramic

65-71

Society, 76 (1993), 369-375

[8] G Zhang et al, "Effect of sintering atmosphere on the relative density and conductive properties of Ni-Fe spinel ceramics". Functional Materials, 36 (11) (2005), 1709-1711 [9] X. G Sun et al, "Study of densification and mechanical properties of Ni-NiFe204-NiO cermet inert anode". Changsha:

1139

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011S ELECTRODE TECHNLOGY for ALUMINUM PRODUCTION Poster Session - Electrode SESSION CHAIR

Alan Tomsett Rio Tinto Alcan Brisbane, Australia

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

INFLUENCE OF ULTRAFINE POWDER ON THE PROPERTIES OF CARBON ANODE USED IN ALUMINUM ELECTROLYSIS Xiao Jin, Deng Songyun, Li Jie, Lai Yanqing, Liu Yexiang School of Metallurgical Science and Engineering, Central South University, Changsha, Hunan, 410083, RR. China

Keywords: Carbon anode, Ultrafine Powder, Blaine Abstract

ball mill product content and powder purity, combined with

The ultrafine powder (<400 mesh) content of ball mill

the analysis of Blaine Number and BET specific surface area,

product has a great influence on the carbon anode quality.

the effect of ultrafine powder content of ball mill product on

Experiments based on the aggregate recipe and raw material

the properties of carbon anode was studied.

of an aluminum carbon anode plant in China have been conducted. The ultrafine powder was separated from the ball

Experimental

mill product by screening. On the premise of unchanged

Method to measure coke powder

powder purity and content, anode samples with different

The BET specific surface area and Blaine Number were used

ultrafine powder content were prepared and the properties

to characterize the properties of the powder materials with

were analyzed. Ultrafine powder materials were characterized

different ultrafine content. Powder purity and ultrafine

by Blaine Number and BET specific area. Results showed

powder content of ball mill product were controlled by

that good electrical conductivity, thermal conductivity and

screening.

air/C02 reactivity behavior could be obtained with low ultrafine powder content (3100-4000 Blaine, 2.510-3.734

The Blaine Number used in the anode carbon industry is

m2/g specific area). Anode properties such as baked density

based on a testing method for specific surface-Blaine which

and air permeability were closely related with the ultrafine

is widely used in cement industry. Fischer [1] outlines how

powder content and the Blaine value of powder material. The

the Blaine air permeability apparatus can be adapted to

results can be used to optimize the properties of carbon

measure and compare coke dust granulometries. The

anode.

calculated value of the specific surface is referred to as the Blaine Number, a dimension index. It is designed for

Introduction

comparison of similar materials and is not intended to be an Carbon anodes used in aluminum electrolysis are a mixture of

absolute measure of specific surface area [2].

a various grain sizes of calcined coke, recycled spent anode butts and coal tar pitch, which are kneaded together, formed

The Blaine number and specific surface area of graphite,

to the desired shape and then baked. The mixture is a

anthracite coal and calcined coke powder of different particle

distributed structure system. In this system, carbon particles

size were measured, with the RDC 155 Dust Fineness by

and powder are bonded by intermolecular forces and

Blaine, and BET specific surface instruments. Results are

mechanical adhesion. Coke powder, especially the ultrafine

shown in Figure 1.

powder (<400 mesh; 37 μπι) of anode formulation has a large It can fill the open

The Blaine Number of non-porous and ball-like graphite and

porosity in the distributed system structure, which increases

influence on anode performance.

anthracite coal powder has a linear relation with the BET

density. It also has a high specific surface area which directly

specific surface area. The calcined coke powder does not

influences the pitch content of the anode. Thus, ultrafine

have a linear relation because of its porous and

powder control is very important to the stability of anode

irregularly-shaped structure. When the specific surface area is

quality.

more than 5.5m2/g, the Blaine Number increases rapidly.

In this paper, the formulation of a carbon anode plant in

So, while the BET specific surface area has some referential

China was used as the basis for the study. Based on fixed

value, it can not be used alone to characterize the

1143

granulometries of ultrafine coke powder. A combination with

A detailed plan was developed to increase the ultrafine

analysis of the B laine Number is required to characterize the

powder content of the ball mill product material in

granulometries of the powder more in detail.

increments of 7 wt%, from 0 wt% to 70 wt%. The anode properties of the samples produced were then investigated. An assessment of the pitch content to be applied to all recipes was initially determined. It was then fixed to decrease the

28000

Anthracite Coal

effect of pitch content on the results. An increase in ultrafine

24000

powder content will cause the optimum pitch content to

20000

increase, because of the increase in the specific surface area.



m

Accordingly, the recipe with an ultrafine powder content of 70 wt% was chosen to evaluate the pitch content to be used

120001-

for all experiments. The pitch content was increased in 0.5 wt% steps from 16 wt% to 19 wt%. Ten anode samples

BET specific surface (m 2 /g)

prepared from each recipe. Table 1 Dry formulation of carbon anodes

4

6

8

10

Proportion (wt%)

Coarse particle

3-6

25

Particle

1-3

24

Small particle

1-0.3

21

Coke Powder

<0.15

30

12

BET specific surface ( m /g)

Results and Discussion

12000

Determination of Pitch Content Calcined Coke

10000ί

Ö C

Particle Size(mm)

The properties of the baked anodes with 70 wt% ultrafine powder content and the different pitch contents are shown in

8000 f-

Figure 2. The shadow graph shows the distribution of the properties of all anode samples tested.

JS 6000 h Cu

Baked density, electrical resistivity, compressive strength and 3.0

3.5

4.0

4.5

5.0

5.5

6.0

6.5

7.0

BET specific surface ( m2/g) Figure 1 Blaine number vs. BET specific surface area Design of Experiment

air permeability deteriorated with increasing pitch content. According to R&D Carbon work on optimum pitch content [3], the pitch content of 17 wt% can be considered as excess. A pitch content of 17 wt% was therefore used as the fixed pitch content in all later experiments.

The dry aggregate formulation of the carbon anode is shown

Variation trend of baked anode properties with different

in Table 1. The sizing of the coke powder was controlled to

ultrafine powder contents

70% passing 74μπι by screening. Anode samples were prepared with the size of Φ50* 100mm. Kneading, paste

The Blaine Number and BET specific surface area of the calcined coke powder material with different ultrafine

cooling, forming and the highest baking temperature were set

contents are shown in Table Π. The trend of the Blaine value

as 180°C, 160°C, 155°C and 1200°C respectively.

with specific surface area is similar to Figure 1. There was no linear relationship between these two parameters.

1144

Blaine, 4.17 m2/g specific surface area) because of the improved particle packing. But as shown above, BAD was also influenced by the pitch content. Table II Blaine and BET Specific Surface Area Test Results of Powder Materials Powder

Ultrafine

Blaine

BET specific

sample

powder (%)

Number

surface(m2/g)

0

2427

2.127

2

7

3151

2.51

3

14

3523

2.904

4

21

3860

3.321

5

28

4011

3.734

6

35

4321

4.174

7

42

4720

4.548

8

49

5120

4.945

56

5886

5.393

63

7508

5.750

70

10484

6.113

1

1io9 1n 1

The baked anodes with Blaine values of 3100-4000 showed good electrical resistivity. At higher Blaine Numbers,

electrical

resistivity

increased

rapidly.

Theoretically, with ultrafine powder content increasing, anode electrical resistivity should improve, since the optimum pitch content of the recipe would have increased, approaching to the actual value of 17 wt%. The deterioration of electrical resistivity could not be due to the lack of pitch. It is suggested that large numbers of boundary layers were produced by ultrafine powder and pitch within the anode and caused the increase in electrical resistivity. t.O i

\ 16

»

\ 17

»

1 18

·

1 19

Air permeability decreased with the increase of ultrafine

Pitch (%)

powder content with a linear relationship. It indicated that ultrafine powder can efficiently fill the pores in

Figure 2 Baked anode properties vs. Pitch content

carbon anode and improve the anode structure.

The test results of anode properties are shown in Figure 3. The red points represent the properties of anodes adopted for

When the ultrafine powder content was more than 28

the original recipe of the plant. Comparison of baked anode

wt% (Blaine 4000), thermal conductivity decreased

properties with different ultrafine powder content yields the

shapely with the rise of ultrafine powder content.

following observations: Baked Apparent Density (BAD) showed a peak when the ultrafine powder content was about 35 wt% (4300

1145

10

20

30

40

50

Ultrafine Powder(%)

60

10

x:

3.6

>

3.3

3 3.0 T5 C 2.7

o O 2.4 Τä 2.1

E

0

10

20

30

40

50

60

Ultrafine Powder (%)

10

20

30

40

50

30

40

50

60

1

t„„ 1

4.2 3.9

0

20

Ultrafine Powder (%)

60

Ultrafine Powder (%)

70

- a m

ir

_ :

1.8 i

1 M,

,,*„. 1

nnl

»

1

1

t

>

10

20

30

40

50

60

70

0

10

20

30

40

50

60

70

Ultrafine Powder (%)

Ultrafine Powder(%)

70

Figure 3 Properties of Baked Anode with Different Ultrafine Powder Content

1146

>,"„!

0

recommended to be combined with specific surface area to characterize the granulometries of coke fines

The baked anode without ultrafine powder had a high

more precisely.

coefficient of thermal expansion (CTE) - 5.03*10~6/K. As ultrafine powder was added to the recipe, the CTE

2.

With the same aggregate formulation, powder purity and pitch content, optimizing the ultrafine powder

decreased significantly.

content had a positive impact on the anode baked The baked anodes with a lower ultrafine powder content

density, electrical resistivity, air permeability, thermal

had the lowest reactivity. However, different from its

conductivity, CTE, and air/C02 reactivity. Based on an

reactivity to air, a recipe with no ultrafine powder had a

overall consideration of various factors, it is suggested

negative impact on the anode C0 2 reactivity. The baked

that the Blaine Number of powder material be

anode with 7% ultrafine powder content (Blaine 3100)

controlled between 3100 and 4000.

had the lowest reactivity to C0 2 . Vichtus and Cannova [4] indicated that anode air and C0 2 reactivity should

3.

Increasing the ultrafine powder content of the ball mill

decrease with the decrease of air permeability. However,

product will decrease air permeability. However, it will

in this work, the trend is different because ultrafine

increase the reactivity of the baked anode to 0 2 and

powder has two effects on anode reactivity. On the one

C0 2 .

hand, ultrafine powder can fill the porosity in anode to prevent 0 2 and C0 2 infiltration . On the other hand,

4.

The influence of ultrafine powder on carbon anode properties is complex and multiband is influenced by

powder, especially ultrafine powder, is easier to react

many factors. Ultrafine powder content should be

with 0 2 and C0 2 , which can generate additional porosity

controlled according to the requirements of the plant.

when the anode is in use in the cell.

Increasing or decreasing the ultrafine powder content without careful experimentation would probably result in deterioration in the properties of carbon anode.

Through optimising the ultrafine powder content of ball mill product, there is still large scope for the carbon anode plant to improve its anode performance.

References [1] Werner K. Fischer, "The Interdependence of Pitch

Conclusion

Content, Dust Fineness, Mixing Temperature and Kind of On the basis of unchanged powder purity and content, the

Raw Materials in Anode Formulation" TMS 1980, 80-73

ultrafine powder content of the ball mill product has a

[2]

significant impact on carbon anode properties.

S.Nascimento, et al., "Finer Fines in Anode Formulation"

During

Francisco

E.O.Figueiredo,

Ciro

R.Kato,

Aluisio

anode production, the stability of anode properties can be

Light Metals 2005, p 665-668

improved if the concept of Blaine Number is used for the

[3] Werner K. Fischer, and Raymond C. Perruchoud, Anodes

powder purity control.

for the Aluminum Industry R&D Carbon Ltd., P.O. Box 157, CH-3960 Sierre, Switzeland-1995

In this paper, the main results were obtained as follow:

[4] Bernie Vitchus, Frank Cannova. "Practical Air Reactivity Impacts on Anode Performance" Light Metals 2002, p.

1.

The Blaine value of non-porous graphite and anthracite

553-559

coal powder had a linear relation with their BET specific surface area, while the calcined coke powder showed different behaviour because of its porosity and irregularly-shape

structure.

Blaine

Number

is

1147

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

PREPARATION NiFe204 MATRIX INERT ANODE USED IN ALUMINUM ELECTROLYSIS BY ADDING NANOPOWDER Zhigang Zhang, Guangchun Yao, Yihan Liu, Xiao Zhang School of Materials and Metallurgy, Northeastern University, No. 11 Lane 3 Wenhua Road, Shenyang, Liaoning, 110004, China E-mail: [email protected] Keywords: NiFe204, Inert anode, Nanopowder, Particle gradation After many years' researches, there are mainly three types of inert anode for aluminum electrolysis, that is, a metal or alloy

Abstract

anode, a metal oxide ceramic anode, and a cermet anode [3]. Two-step sintering process was adopted to prepare

Alloy anode is good at ductility and conductivity, but bad at the

NiFe204 matrix inert anode in this research. In the process of

thermal stability and corrosion resistance [4], while ceramic

synthesizing NiFe204 spinel, Fe203 and NiO powders as raw

anode is bad at brittleness and conductivity, but good at the

materials added additives were synthesized at 1000D. Through

refractoriness and corrosion resistance. Among them, nickel

crushing and screening, adding NiFe204 nanopowder, particle

ferrite-matrix inert anode has become the key study object for

gradation and compression molding, the nickel ferrite matrix

its high strength, good stability at elevated temperature, good

ceramic inert anode was sintered secondarily at 1250D for 6h.

corrosion resistance and so on [5]. Considering the high

The effect of addition level of nanopowder on the density and

brittleness and poor thermal shock resistance of NiFe204~matrix

porosity, bending

and impact toughness was

ceramic inert anode, it has not been used in electrolytic

investigated emphatically. The results showed the addition of

strength

production of aluminum for it doesn't meet the requirements of

NiFe204 nanopowder had significantly increased the bending

aluminum electrolysis process. There are many ways to toughen

strength and impact toughness of NiFe204 matrix ceramic inert

ceramic materials such as particle reinforcing, fiber reinforcing

anode. Inert anodes had the best comprehensive properties while

and so on. Nickel ferrite spinel nanopowder was added to

adding 60wt% nanopowder. The values of density and porosity

improve the mechanical properties of inert anode in this

3

were 4.62g/cm and 5.2% respectively, the value of bending

research.

strength was 53.39MPa and the value of impact toughness was

In this paper, NiFe204-matrix inert anodes were prepared

2

3.19J/cm .

by cold-pressing and sintering progress adding nickel ferrite spinel nanopowder. The effect of addition level of nanopowder on the density and porosity, bending strength and impact

Introduction

toughness was investigated emphatically. In recent years, numerous works have been actively focused on the development of inert (non-consumable) anodes

Experimental Section

for replacement of consumable carbon anodes in Hall-Hιroult electrolysis cells for the production of aluminum. Comparing

Materials

with the consumable carbon anode used in aluminum production, the use of inert anode can bring economic and environmental

All of the chemical reagents were of analytical grade and

enormous benefits due to the elimination of harmful emissions

were used without further purification. Double distilled,

into atmosphere (such as carbon dioxide, sulfur compounds,

deionized water was used as a solvent. Manipulations and

fluorine, etc.) and carbon factory, decreasing the intensity of

reactions were carried out in air without the protection of

labour for anode changing consumedly, and being of acceptable

nitrogen or inert gas. Nickel ferrite spinel nanopowder was

cost [1,2].

synthesized via solid-state reactions at low temperature based on

1149

my own previous work [6]. The precursor prepared by rubbing

Fe^O^owder Disstill water Additive NiO powder

FeS(V7H 2 0, NiS0 4 »6H 2 0, NaOH and dispersant sufficiently at

*

room temperature was calcined at 800°C for 1.5h. The average

Ball milling

particle size of the nanopowder used in this experiment is about 75nm.

i

Preparation Progress Two-step

sintering process

was utilized

.

to prepare

NiFe 2 0 4 -matrix inert anode in this research. In the first step, a proper amount of Fe 2 0 3 and NiO powders as raw materials and

.

(KQM-X4, China) for twenty-four hours using distilled water as

—i

Moulding

A

T

Presintering

.

^

J '

jz

I Crushing and screening | I NiFe^Oa nanopowder

dispersant. Following by drying, the dried mixture with 4% PVA

Ball milling

I

binder was molded by cold pressing under 60MPa pressure and

l·^ ' PVA

-—iJT

calcined at 1000°C in air for six hours to form the NiFe 2 0 4

I

Moulding

spinel matrix material. Through crushing and screening, the matrix material was separated to different size of particles, that

l·^—'

7

Sintering

is, 0.50mm~0.35mm, 0.2lmm~0.15mm, and 0.11mm~0.08mm. Particle gradation design according to the theory of the most compact is showed in Table I. The mixtures adding different

Inert anode

levels of nanopowder were put into glass beaker, with vigorous stirring and ultrasonication at room temperature.

PVA

L

NiFe?Q4 spinel matrix material

additives of M n 0 2 and V 2 0 5 were mixed in planetary muller

anhydrous alcohol as dispersant, were mixed enough under

Drying 1

Fig.l Flow chart of preparing NiFe 2 0 4 -matrix inert anode Characterization

Adding with 4% PVA binder, the dried mixtures were made into

The bulk density and porosity of samples were tested by

60mmxl2mmxl0mm blocks by cold pressing with 160MPa and

liquid displacement method in double distilled water medium

sintered subsequently at 1250°C in the air atmosphere for six

using Archimedes' principle. Mechanical properties such as bending strength and impact toughness were measured by using

hours to produce NiFe 2 0 4 -matrix ceramics. The flow chart of

INSTRON4206-006 electron mechanical experimental machine

preparation process is shown in Fig.l.

(USA). The fracture morphology of samples were obtained on SSX-550 scanning electron microscopy (SEM) equipped with

Table I Design table of particle gradation (mass fraction)

Numb er

Main particle

Filling particle

0.50mm~0.35 0.21mm~0.15 0.11mm~0.08

an energy dispersive X-ray spectroscopy.

Nanopowd

Results and Discussion

er

mm

mm

mm

1

52.5%

10.5%

7%

30%

2

45%

9%

6%

40%

3

37.5%

7.5%

5%

50%

4

30%

6%

4%

60%

anode material. Given low density, there exist lots of pores,

4.5%

3%

70%

which will lead to poor properties of the material such as poor

5

22.5%

Effect of nanopowder content on density and porosity

I think there is no error for this sentence using "Effect of. The bulk density is of great importance to desirable inert

electric conductivity, poor mechanical properties and corrosion resistance and so on. With the poor electric conductivity, it will

1150

result in great anode ohmic voltage drop, high cell voltage and energy consumption as well as be detrimental to the corrosion resistance which decreases the used life of inert anode and the quality of as-product aluminum. The changes of density and porosity with the content of nanopowder are shown in Fig.2. From Fig.2, it can be found that in the range of 30-60% nanopowder, the higher the nanopowder content, the higher the density and the lower porosity. However, as the content of nanopowder ranges from 60% to 70%, the

40

50

60

Mass fraction of nanopowder/%

changing trends of density and porosity are contrary to the

Fig. 3 Effect of nanopowder on bending strength and impact

former. The composite with 60% nanopowder content achieves 3

toughness

a maximum value of density of 4.62 g/cm and a minimum value of porosity of 5.2%. In the particle gradation, with the increase of nanopowder content, large particles reduced result in

In general, the bending strength of ceramic is related to

fewer pores which werefilledby nanopowder fully. So the bulks

several factors such as ceramic grain size [7] and porosity [8].

were more compact after cold-pressing. What's more, more

W. Duchworth [9] presented the relation between ceramic

nanopowder in a proper range provided greater sintering force

fracture strength and porosity as shown in formula (1):

due to higher Gibbs free energy for large specific surface area.

ó = ó0 exp(-θP)

That promoted the gains to develop completely and made the

(i)

bulk more compact as the nanopowder content increasing.

where P is porosity; ó is strength for sample with porosity of P;

Whereas, greater volume shrinkage made the materials cracked

ó0 is strength for sample without pores; and b is a constant.

easily while the content of nanopowder over 60%.

From formula (1), the significant inverse relation is found between porosity and bending strength, which is the lower the

-

\

-

porosity is, the higher bending strength is. Ceramic fracture strength versus grain size presented by

V\ -

Hall-Petch [10] is shown in formula (2):

—o— Density —*— Porosity

óÑ = ó0+êÜ*

(2)

where aF is intensity of ceramic; ó0 is yield intensity; K and m

/

are constants; and d is grain size. 1

1

Mass fraction of nanopowder/%

Fig.2 Effect of nanopowder on the density and porosity

The fracture micro-appearances of with various contents of nanopowder were characterized by SEM, as shown in Fig.4. From Fig.4, as the content of nanopowder increases, the grain size decreases. The average grain size decreases from 6μπι to 1.5μπι with the content of nanopowder increasing from 30 to

Effect of nanopowder content on mechanical properties

70%. The toughening of this material mainly depends on grain The effect of nanopowder content on bending strength and

size. The crack always extends through the way with low energy.

impact toughness is showed in Fig.3, from which it is obvious

It is difficult for crack to extend through the complete grain but

that the changing trend of bending strength and impact

easy to extend across the interface between grains due to the

toughness is consistent with the density's and contrary to the

relatively low fracture energy. In certain extent of research, the

porosity's. As well as the density, the material with the content

smaller the grain size is, the more the grain boundary is. Hence,

of 60% nanopowder has attained to the maximum values of

with the decreasing of grain size, the crack needs more energy

bending strength and impact toughness of 53.39MPa and

to extend across the increasing of grain boundary and it

3.19J/cm2, respectively.

increases the resistance of fracture growth. When the content of

1151

nanopowder is over 60%, though the grain size is decreased,

make the impact toughness decreased slightly.

some deficiencies are produced from sintering progress, which

D

T^f

"*» A

· "-f?'Jr,€ff'm

?

M

_ ί

:

1

r

:

e Fig.4 SEM photographs of NiFe204-matrix inert anode with various contents of nanopowder: a-30%; b-40%; c-50%; d-60%; e-70%

Conclusion

over 60%. 2) The changing trend of bending strength and

1) As the content of nanopowder is increased

impact toughness is consistent with the density's and

from 30 to 60%, there is a sharp change in density and

contrary to the porosity's. The bending strength

porosity of NiFe204-matrix inert anode. A maximum

increases from 32.46MPa to 53.39MPa and impact

3

density of 4.62 g/cm and a minimum value of

toughness increases from 1.82 Jem 2 to3.19 Jem 2

porosity of 5.2% are presented when the content of

with the content of nanopowder increasing from 30 to

nanopowder is 60%. The density decreases and the

60%. Both values of them decrease slightly when the

porosity increases when the content of nanopowder is

content of nanopowder is over 60%.

1152

Acknowledgements

[5] E. Olsen and J. Thonstad, "Nickel Ferrite as Inert Anodes in Aluminium Electrolysis: Part I Material

The authors gratefully acknowledge financial

Fabrication and Preliminary Testing," Journal of

support from National Natural Science Foundation of

Applied Electrochemistry, 29(1999), 293-299.

China (No. 50834001) and the National High

[6] ZHANG Zhi-gang et al., "Synthesis of NiFe204

Technology Research and Development Program of

Spinel Nanopowder via Low-Temperature Solid-State

China (No. 2009AA03Z502).

Reactions," Journal of Northeastern University (Natural Science), 31(6)(2010), 868-872(in Chinese). [7]. GUAN Zhen-duo, ZHANG Zhong-tai, and JIAO

References

Jin-sheng,

Physical

capability

of

inorganic

[1] R.P. Pawlek, "Inert Anodes: An Update," Light

materials(Bei)ing: Tsinghua University Press, 2002),

Metals, 2002, 449-456.

103 (in Chinese).

[2] J. Keniry, "The economics of inert anodes and

[8]. S. Schicker et al., "Microstructure and mechanical

wettable cathodes for aluminum reduction cells,"

properties of Al-assisted Sintered Fe/Al203 Cermets,"

JOM, 53(5)(2001), 43-37.

Journal

of

the

European

Ceramic Society,

[3] D.R. Sadowy, "Inert anode for the Hall-Hιroult

19(13/14)(1999), 2455-2463.

cell: The ultimate materials challenge," JOM,

[9] W. Duckworth, "Discussion of Ryshkewitch

53(5)(2001), 34-35.

Paper," Journal of the American Ceramic Society,

[4] G. Mark and H. Margaret, "Laboratory-scale

36(3)( 1953), 65-68.

Performance of a Binary Cu-Al Alloy as an Anode for

[10] N. J. Petch, "The Cleavage Strength of

Aluminium Electrowinning,"

Polycrystals," Journal of Iron and Steel Institute,

Corrosion Science,

174(1)(1953), 25-30.

48(9)(2006), 2457- 2469.

1153

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

COLD WATER MODEL SIMULATION OF ALUMINUM LIQUID FLUCTUATIONS INDUCED BY ANODIC GAS IN NEW TYPE OF CATHODE STRUCTURE ALUMINUM ELECTROLYTIC CELL Yan Liu, Ting'an Zhang, Zhihe Dou, Hongxing Wang, Guozhi Lv, Qiuyue Zhao, Naixiang Feng, Jicheng He School of Materials and Metallurgy of Northeastern University, Key Laboratory of Ecological Utilization of Multi-metal Intergrown Ores of Education Ministry, Shenyang, 110004, China Keywords: new structure cathode, cold water model, anode gas Figure 1 shows the structure of a new type of cathode reported

Abstract Compared to the conventional cathode structure electrolytic

by Feng et.al where the carbon block surface is not flat but with

cell, the new type of cathode structure electrolytic cell can

ridges. Characteristics of the electrolysis technology and

effectively inhibit the fluctuation of liquid aluminum, which

operations for the new cathode structure aluminum reduction

can reduce the polar distance, decrease cell voltage, and save

cell are as follows. The effective area for dissolved loss of

energy. A cold water model was designed according to scale of

cathode aluminum will be reduced due to stability improvement

1:3 with 160KA industrial electrolytic cell to examine

of cathode aluminum surface, so it can increase the current

aluminum liquid fluctuation induced by anodic gas in the new

efficiency[1].

type of cathode structure electrolytic cell. High-speed photographs of the liquid interface waves show the effect of

This research includes three variations of new structure

several cell parameters such as anode-cathode distance,

cathodes whose size is below.

electrolyte level, and gas flow rate. The study shows that in the new type of cathode structure electrolytic cell, the largest

Lengthxwidthxheight ( mm ) :

interface wave height will be reduced significantly.

570x70x60 570x80x60 570x90x60

Introduction In order to develop and verify theoretical models, physical

High-speed photography was used to shoot the liquid surface

models play an important role. In many industrial processes the

waves for the two different cathode structures, varying anode

conditions do not allow direct measurements of many

cathode distance, electrolyte level, aluminum liquid level, and

parameters, and information which is crucial in order to verify a

gas flow rate. The bubble images were analyzed by using

mathematical

special image analysis software (Image-Pro-Plus) and the

model

cannot

easily

be

obtained.

The

Hall-Heroult process for aluminum production is a typical

maximum aluminum liquid fluctuation heights were obtained.

example, due to the lack of flexibility and due to the highly corrosive high temperature system. Water model experiments of

Experimental Principle and Equipment

are commonly used to simulate aluminum electrolysis in industry.

Water Model Principle Geometrical

Similarity

Geometric

similarity

considers

prototype and model similar in the main dimensions. Using water model made of glass instead of industrial aluminum cells, we use a model with water and vegetable oil to represent metal and bath, respectively, to study gas induced bath flow in the outer channel and the occurrence of waves at the metal-bath Fig. 1 Schematic Diagram of Cathode for new cathode structure

interface. The model and the prototype have the same proportion, but the water model is smaller than the prototype.

1155

Fig. 2 shows the location of the three points 1 to 3, where

6

interfacial deformations were measured. Also shown in the

7

figure is the location of the origin and the unperturbed interface (horizontal black line). The size of physical model was designed according to scale of 1:3 with 160KA industrial electrolytic cell Table 1 The size of physical model and the prototype

Anode size (mm)

Cell

Model

1520x585x535

506x195x250

1 2

3

1 Anode gas 2 Model of new cathode cell 3 Oil and water 4 Anode supporting device 5 Anode gas channels 6 Flow meter

Cell area (mm)

8945x4220

2981x1406

Cell depth (mm)

600

560

Distance anode to

450

150

7Gas cylinder Fig. 3 Sketch of Water model

endwall(mm) Distance anode to

Dynamical Similarity The physical properties of the fluids in a real cell are unfortunately difficult to match[2] .The molten metal phase in an industrial cell is characterized by a low

400

133

sidewall (mm)

viscosity and a high surface tension, properties which are not easily obtainable with common fluids. The fluids used in the current model are water to simulate the metal pad, light

Center channel (mm)

280

94

Distance between

45

15

anodes (mm)

vegetable oil for the bath, and air for the gas formed under the anode. The oil-water combination is chosen to match the density and viscosity ratios as closely as possible, using safe and inexpensive room temperature fluids. A summary of the relevant parameters, compared with the industrial cell parameters (cell data from Chenosis and LaCamera[3] and Haarberg et al.[4]) is given in Table 2.

Table 2 Parameters of physical model and the prototype

Fig. 2 Locations of the measurement points

1156

Parameter

Model

Cell

pb , kg/m3

0.9

2.1

pm , kg/m3

1.0

2.3

vb , mm2/s

46.5

1.3

vm , mm2/s

1.0

0.4

obm , N/m

0.03

0.55

Gbg , N / m

0.02

0.14

wave propagation velocity. Assuming that the characteristic Here, pb is the density of bath and pm is the density of molten metal. vb is the kinematics viscosity of bath and vm is the

velocity is given by Eq. [3] and that the wave is both capillary and gravity driven, Frw is approximated as

kinematics viscosity of bath, σ b m, and σ b g is the bath-metal FrM,oo ,

and bath-gas surface tension, respectively.

[5]

*

As seen from Table 2, the parameters are entirely different

1ð Pm +Pb

from those encountered in a real cell. However, it is the

In Figure 4 the ratio Frw

combination of the mentioned parameters that determine whether dynamic similarity can be claimed. The bubble Froude

Pm +Pb model

/Frwce11 is shown for the

expected wavelengths.171

number FrB is critical for an accurate representation of the

As seen from Figure 4, the Froude number in the model is

bubbles. Fortunately, the Froude number can be defined from

over predicted for short wavelengths (because of the low

geometrical considerations as follows:

surface tension) and under predicted for long wavelengths

K2 Fr = B gL 1

r

(because of the low density ratio). Over the expected range of [i]

wavelengths, Frwmodel is ±25 pet of the values expected in a real cell. Though not matched exactly, the parameters are believed to be close enough so that dynamic similarity can be claimed, at

Where vg is the gas velocity, g is the acceleration of gravity,

least for a narrow band of wavelengths.

and L is the typical anode size. Hence, by choosing a 1:1

2.5

representation of the anode and the realistic gas rates, similarity

2.0

the other nondimensional groups are of secondary order. Here, the influences of viscosity and surface tension are grouped into

1.5 I

a Morton number, defined as follows:

Mo =

VW

-model //cF Jcell| 1 r w w 1

pr

can be claimed for FrB. As noted by Zhang et al[5] the effects of

\

1.0

[2]

°Ì0

Inserting values from Table 2 yields a factor of 500 in

0.1

—·

0.2

1

0.3

i

1

0.4

difference between the model and the cell Morton numbers.

ë

Though significant, it is not expected that this large difference

Fig. 4 The ratio Frwmodel/Frwcdl for various wavelengths ë.

is critical because the FrB will dominate the overall picture. According to Solheim et al., [6] the typical velocity in the interpolar is given as follows:

v i n t oov/ 0 7 M.- 3

Results and Discussion Comparison Aluminum Liquid Fluctuation of Conventional

[3]

Cathode Structure Electrolytic Cell with the New Type of Cathode Structure Electrolytic Cell

Hence, one should expect that the bath velocities in the

Experimental condition: anode-cathode distance 3cm,

water model are up to 25 pet too low, compared with that in the

electrolyte level 16cm, aluminum liquid level 17cm, gas flow

cell. The interface deformation is assumed to be governed by its

rate 1.0m3/h

corresponding Froude number, defined as follows:

Fr.

=V-

[4]

Where V is a characteristic velocity and c is a characteristic

1157

Effect

of Operating Parameters on Aluminum Liquid

Fluctuation in New Type of Cathode Structure Aluminum Electrolytic Cell Effect of Gas Flow Rate on Aluminum Liquid Fluctuation Experiment study on the effect of different gas flow rate on aluminum liquid fluctuation, respectively is 0.6m3/ru 0.9m3/lu 1.2m3/ru 1.5m3/h Experimental condition: anode-cathode distance 3cm, electrolyte level 14cm, aluminum liquid level 17cm Fig. 5 Aluminum liquid fluctuation of new cathode structure electrolytic cell (After image analysis)

0.6

I 0.8

i

I 1.0

i

I 1.2

i

1 1.4

1

1 1.6

3

Rate of flow (m /h)

Fig.7 Effect of gas flow rate on the maximum aluminum liquid fluctuation height at different points In figures 7-11, C means conventional cathode structure; N means new cathode structure, and the numbers 1-3 refer to locations where wavefluctuationswere measured.

Fig. 6 Aluminum liquid fluctuation of conventional cathode

Analysis: As the gas flow rate increases, the liquid surface

structure electrolytic cell (After image analysis)

waves become more intense. The maximum aluminum liquid fluctuation height will increase in both types of cell, but the

In the pictures (Figs 5 and 6), Dl is the liquid surface

maximum fluctuation is greater for the conventional cell at all

before experimental (standard line), and D2 is the maximum

three locations.

aluminum liquid fluctuation height at location 3. From figs 5 and 6, we can see D2 is 4.8 mm in the new

Effect of Anode-cathode Distance (ACD) on Aluminum Liquid

cathode model compared to 5.9mm in the conventional cell

Fluctuation Experiment study on the effect of different ACD

model.

on aluminum liquidfluctuationrespectively is 3cm, 4cm, 5cm.

Obviously , D2 is smaller when we use new cathode

Experimental condition: electrolyte level 16cm, aluminum liquid level 17cm, gas flow rate 1.0m3/h.

structure meaning that fluctuation is weakened and it may be possible to decrease cell voltage and save energy

1

Analysis: with the electrolyte level increase, aluminum liquid fluctuation is eased up and the maximum fluctuation height is decreased in both types of cell. From the chart we also get Cl is above NI, C2 is above N2 and C3 is above N3. The new type of cathode structure is better than the conventional cathode structure.

4 I

I 3.0

i

I 3.5

i

I 4.0

i

I 4.5

i

Conclusion

I— 5.0

We can find out the best operation parameters by the

Anode-cathode distance (cm)

Fig. 8 Effect of ACD on fluctuation height

comparison of the maximum aluminum liquid fluctuation height in conventional and the new type of cathode structure electrolytic cell. With the analysis and the summary of the

Analysis: The change of slope is gentle as the ACD is reduced from

results, we can get the conclusions below:

5 to 4 cm, but a larger slope appears between 3 and 4 cm.

(1) Under different experimental conditions, compared

When ACD is 3cm, the fluctuation is very severe and the

with the conventional cell, the maximumfluctuationheight will

maximum fluctuation height is very large. With the ACD

be reduced significantly in the new type of cell. This indicates

increase, aluminum liquid fluctuation is reduced.

the new type of cell should have a more stable metal pad.

As in the gas flow test the largest interface wave heights are

(2) In summary, in both conventional and new type of

reduced significantly for the same locations in the new type of

cathode structure electrolytic cell, the maximum aluminum

cell compared to the waves in the conventional cell.

liquid fluctuation height will be increased significantly along with the increasing of gas flow rate. With the decreasing of

Effect of Electrolyte Level on Aluminum Liquid Fluctuation

electrolyte level, the aluminum liquid fluctuation will be

Experiment study on the effect of different electrolyte level on

violent.

aluminum liquid fluctuation, respectively is 15cm% 16cnu

with the ACD of 3cm in both types of cell. When the

(3) The aluminum liquid fluctuation is extremely violent anode-cathode distance increases to 4cm, thefluctuationheight 17cnn 18cm

is reduced significantly, and the effect of the interface wave height will be reduced. In new type of cell this effect is smaller,

Experimental condition : anode-cathode distance 3cm ,

so the anode-cathode distance can be reduced. In conclusion, in the new type of cathode structure

aluminum liquid level 17cm , gas flow rate 1.0m3/ho

electrolytic cell, the largest interface wave height will be reduced effectively. The choice of anode-cathode distance is important. The decrease of anode-cathode distance will improve current efficiency. Considering the aluminum liquid fluctuation, the anode-cathode distance should not be too small, so in the new type of cathode structure electrolytic cell, the anode-cathode distance can reduce to 4cm. Within reasonable parameters the height of the electrolyte should be increased.

15.0

15.5

16.0

16.5

17.0

17.5

Electrolyte levels (cm)

Fig. 9 Effect of Electrolyte levels on the maximum aluminum liquidfluctuationheight at different points

Reference

18.0

1. Feng Naixiang, Tian Yingfu, Peng Jianping et.al. Light Metals, TMS, (2010), 406. 2. K. E. Einarsrud, "The Minerals, "Metals&Materials Society

1159

and ASM International", Metallurgical and Materials Transactions £,41 (2010), 560-573. 3. D. C. Chenosis and A. F. LaCamera, Light Metals, TMS, (1990), 211-20. 4. T. Haarberg, E. Olsen, A. Solheim, M. Dhainaut, P. Tetlie, and S. T. Johansen. Light Metals, TMS, (2001), 475-479. 5. W. D. Zhang, J. J. Chen, and M. P. Taylor. CHEMCEA, (1990), 1-8. 6. A. Solheim, S. T. Johansen, S. Rolseth, and J. Thonstad. Light Metals, TMS, (1989), 245-252. 7. S. K. Banerjee and J. W. Evans, Light Metals, TMS, (1987), 247-255. Acknowledgement This research was supported by the National Natural Science Foundation of China (No. 50934005) and a grant from the National High Technology Research and Development Program of China (No. 2009AA063701). National Natural Science Foundation of China (No. 50974035) National Natural Science Foundation of China (No. 51074047); the doctoral fund of EDU gov(20050145029)

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

EFFECTS OF PHYSICAL PROPERTIES OF ANODE RAW MATERIALS ON THE PASTE COMPACTION BEHAVIOR Kamran Azari1, Hany Ammar1, Houshang Alamdari1'2, Donald Picard1'2, Mario Fafard2, Donald Ziegler3 1 Department of Mining, Metallurgical and Materials Engineering, 1065 avenue de la Mιdecine Laval University, Quebec, QC, G1V 0A6, Canada 2 NSERC/Alcoa Industrial Research Chair MACE3 and Aluminum Research Center - REGAL Laval University, Quebec, QC, G1V 0A6, Canada 3 Alcoa Canada Primary Metals, Aluminerie de Deschambault, 1 Boulevard des Sources, Deschambault-Grondines, QC, GOA ISO, Canada Keywords: Anode paste, Formulation, Compression behavior, Particle size, Particle shape Abstract The current study investigates the effects of coke particle characteristics and paste formulation on the flowability and the compression behavior of anode pastes. Shape factor and texture of different fractions of cokes were characterized using an image analysis system where the characteristics of each coke were correlated to its vibrated bulk density (VBD). A compression test was designed to study the effects of particle characteristics and paste recipe on the compactability of pastes. The test was applied on four anode pastes, prepared from different coke types, particle size distributions and pitch contents. It was observed that the compression test is significantly sensitive to any changes in raw materials characteristics and formulations. Consequently, the compression test may be used as a tool for evaluating anode quality in relation with material variations.

In the current research study, a compression test was used to study the paste compression behavior. Sensitivity of this test to variations in materials and formulations was evaluated. When using this specific test, it is possible to investigate the effects of coke type, particle size distribution and formulation on the formability of paste. In addition, shape factor and texture of coke particles were measured using an image analysis system to investigate the influence of coke particles characteristics on their flowability and compaction behavior. Experimental procedures Two commercially available calcined petroleum cokes and a coal tar pitch were used as raw materials for anode manufacturing. The cokes used were Conoco coke (A) and ZCGG coke (B) with a real density of 2.075 g/cm3 and 2.063 g/cm3, respectively, and with a chemical composition, as shown in Table I. The softening point of pitch was 109°C and its quinoline insoluble content was 15.5%. The cokes were crushed using jaw and roll crushers and sieved to seven size fractions. Four formulations were prepared with different particle size distributions, as listed in Table II, to reveal the effect of size distribution. Particle size distributions consisted of a reference distribution, which is used in the industry, and two other distributions; one with 5% more large aggregates and the other with 5% more fine particles than reference. In addition, one formulation of coke B with reference granulometry was used to reveal the effect of particle properties in different coke types. The coke types and coke/pitch ratios are listed in Table III. Pitch contents were chosen based on the fine fraction in the formulation. Large amount of fine particles increases the specific surface area of coke particles and consequently for wetting these fine particles, a higher content of pitch is required.

Introduction Consistent high quality anodes are a basic requirement of anode plants. Anode properties are influenced by several factors including raw materials properties, anode formulation and anode making parameters. Variations in raw materials properties are considered as one of the most significant challenges in anode manufacturing industry which affect anode quality and consistency. The influences of coke properties, particle size and distribution in the anode formulation have been reported by several authors [1-4]. Such an enormous number of variables involved in anode manufacturing process makes difficult the control and optimization of the final anode quality. Thus intermediate quality indices are required to control each step of the process and to take the corrective actions within the subsequent steps in order to keep the anode quality consistent.

Table I. Chemical composition of Conoco (A) and ZCGG (B) cokes

Green and baked anode properties are conventionally used as indicators of anode quality. Green anode density and permeability are measured as quality indices. However, these indices are still influenced by a number of variables including raw materials properties and mixing and compaction parameters. By defining specific paste quality as an intermediate process quality index, it would be possible to limit the number of variables to be controlled in the subsequent manufacturing steps. These indices would be used to correlate materials properties, paste formulation and mixing factors with anode properties. These paste indices could also be used as indicators for paste compaction behavior for the purpose of modelling the anode forming, fabricated either by hydraulic pressing or by vibro-compaction.

Coke Conoco ZCGG

Na ppm <50 <50

Si ppm 10 20

%S 1.1 3.2

Ca ppm 130 <10

V ppm 120 260

Fe ppm 70 60

Ni ppm 90 150

Vibrated bulk density (VBD) was measured for four fractions of cokes A and B, and the blended coke fractions, to study the effects of coke particle characteristics on their packing behavior. ASTM D-4292 standard test method [5] with vibration time of 2 minutes was used for VBD tests.

1161

Table II. Particle size distribution of pastes 5% More 5% More Particle size Reference (mesh) size fine large 14.4% 12.8% +6 13.6% 16.4% -6+14 14.8% 15.6% 15.4% 16.9% -14+30 16.3% 9.8% 11% -30+50 10.4% 7.5% 8.6% -50+100 8.1% -100+200 9.7% 10.7% 10.5% 22% 30% -200 25.5%

Table III. Coke type and pitch content of pastes Reference 5% More 5% More size large fine Coke AandB A A Pitch, wt% 13.8 15.8 13

*Coke A: Conoco and coke B: ZCGG

Single coke fractions were impregnated by a polishing resin under mechanical vacuum and polished for microscopic analysis. In the current study, a Nikon Ephiphot optical microscope equipped with an image analysis system (Clemex, vision) was used to analyse the particles characteristics which include aspect ratio, compactness, roughness, sphericity, roundness and porosity. These parameters describe various aspects of a particle which may be correlated to packing and flow behavior of the particle. Based on the definitions of these parameters, given in Table IV, a more circular particle has lower aspect ratio and higher values for compactness and roundness. Higher values of roughness show smoother surface of the particle.

A Micromeritics helium pycnometer (AccuPyc II 1340) was used to measure the volume of green anode samples where a core sample, with a diameter of 2.5 cm, was drilled in each green sample for this purpose. The percentage of porosity in green samples could be determined using the green apparent density and pycnometry results. Results and discussion The current section presents the results related to the effects of granulometry and particle properties on the flowability of the particles based on the results of vibrated bulk density tests. Compression behavior of anode paste then will be used to study the effects of size distribution and coke type on the compactability of the pastes and the utility of this test will be emphasized. The green apparent density and the vibrated bulk density results will be correlated to the measured porosity of the green samples. The vibrated bulk densities of blended coke fractions used to prepare the samples are shown in Figure 1. Bulk density of coke A has increased with reducing the large fractions (>300 μπι) and increasing the fine portion (<74 μπι) in the formulation. These results had been expected since the smaller particles fill the interparticle voids without increasing the volume [6, 7]. It has also been reported by other authors [8] that higher amounts of fine particles, results in higher VBD of dust fraction and multi-fraction mixtures, up to a maximum value. In addition, the bulk density of particles increases with decreasing the particle size, because large pores are destroyed with size reduction [6].

Table IV. Particle characteristics measured bv imase analvzer Factor

Aspect ratio Compactness Roughness

Roundness

Definition Ratio of longest dimension to shortest dimension Ratio of area over convex perimeter Quantifies the jaggedness of object's edges and is the ratio of convex perimeter to perimeter. Roundness of object's edges

Aspect ratio

1.24 i

Convex perimeter

(TO 0 Roughness

Roundness

Reference Reference 5% more 5% more size-Coke B size-Coke A large-Coke Afine-CokeA

A rigid steel cylindrical mold with the internal diameter of 89 mm was used to study the compression behavior of anode pastes. Coke fractions and pitch were preheated at 185°C for 120 and 30 minutes, respectively, and then were mixed at 185°C for 10 minutes using a domestic Hobart N50 mixer. The paste was then cooled to 125-130°C to allow the fumes to be escaped before pressing. Compression test was carried out at 130°C, using an MTS Servohydraulic press working at a constant displacement rate of 10 rnm/min and a maximum force of 220 kN. The MTS machine provides displacement-force data at a rate of 10 readings per second. The apparent density of pressed samples was measured based on the final volume and mass. Having the final height of the sample and displacement rate of the punch during the test, the apparent density was calculated for each displacement value as a function of the applied force. The evolution of the paste density was then plotted as a function of applied pressure and presented in the form of pressure-density curves.

Figure 1. Vibrated bulk density (VBD) of blended coke fractions With regard to blends of cokes A and B with reference particle size, it may be observed that they exhibit different bulk densities, as shown in Figure 1. This difference is an indication of the influence of physical properties of coke particles on the flowability and rearrangement during vibration. This different behavior may be related to irregularities on the particle surface which result in inter-particle friction and bridging of the particles. These effects will be elaborated in the following paragraph. Coke particles were characterized using an image analysis system to study the correlations which may exist between particle characteristics and their flowability. Table V shows the image analysis results and VBD for each coke fraction. Regarding coke A compared to coke B, for the same size range, it may be observed that vibrated bulk density of the particles increases with

1162

lower values of aspect ratio; with lower level of porosity; and with higher values of compactness, roughness, sphericity and roundness. More spherical particles are expected to display better packing behavior and higher bulk density values than platelike or needlelike particles [6, 9-11]. Although coke B displays lower real density than coke A, it provides higher VBD values for all fractions except for -200 mesh particles. Coke B particles revealed lower aspect ratio and higher compactness and roundness values than did coke A. Since these factors affect particle flowability, coke B has better packing behaviour which most likely compensates the effect of coke density and results in higher VBD values. For -200 mesh particles, on the other hand, coke A with better flowability factors and lower porosity resulted in higher VBD. Table V. Shape factors, texture and VBD of coke particles Coke type anc size (mesh) Particle -6+14 -14+30 1 properties A B A B 1 Aspect ratio 1.69 1.87 1.74 1.86 Compactness 0.607 0.657 0.719 0.761 Roughness 0.802 0.766 0.915 0.957 Sphericity 0.406 0.4 0.678 0.794 Roundness 0.41 0.474 0.527 0.569 %Porosity 28.73 25.98 23.41 19.51 VBD 0.841 0.953 0.927 1.02 Coke type ancI size (mesh) 1 Particle -200 -30+50 properties A B A B 1 Aspect ratio 1.99 2.02 2 1.75 Compactness 0.707 0.766 0.663 0.638 Roughness 0.95 0.963 0.93 0.885 Sphericity 0.707 0.804 0.583 0.507 Roundness 0.5 0.567 0.437 0.427 %Porosity 21.34 20.74 3.3 4.64 VBD 0.919 1 0.969 0.945

The pastes with the same particle size and pitch content, but different sources of coke, showed quite separated curves. This may be related to particle characteristics, as observed in the case

0.9

1.1

1.2

1.3

1.4

1.5

1.6

1.7

3

Apparent density (g/cm ) Figure 2. Compression curves with variations in paste formulations and coke source (K: Coke A, 5% large more than reference; L: Coke A, 5% fine more than reference; M: Coke A, reference granulometry; N: Coke B, reference granulometry). of VBD results. In the present study, the interesting result is that coke A with more elongated and rough particles and lower bulk density provided samples with higher green apparent density, as shown in Tables V and VI. These results agree with other studies where it was reported that although aggregate bulk density is an indication of porosity, packing and size distribution in a powder bed, there is no consistent correlation between aggregate bulk density and green apparent density [7]. On the other hand, other studies reported that irregular shape and rough surface of particles up to 12 mm prevent the movement in a viscous medium and affect the final anode properties [12]. This misleading may be related to the fact that the flow behavior of particles is different in the presence of binder matrix. The inconsistency between particle porosity and bulk density is another reason for this contradiction [11]. In addition, some open pores of coke A may be filled with binder matrix which increases the green apparent density. The results show that VBD cannot be used alone to show the compression behavior. In other words, cokes with higher vibrated bulk density will not necessarily result in denser anodes.

Compression tests were carried out for four formulations to study the influences of size distribution on the compaction behavior of paste. The formulations were prepared using two coke types with different particle size distributions and three coke/pitch ratios, as shown in Tables II and III. Figure 2 shows the average curves obtained from three compression tests for each paste formulation. It may be observed that there is a meaningful difference among these curves and the test is capable of illustrating the effects of materials variations on the paste compression behavior. When using coke A, the final achieved density decreased by increasing the content of large fractions; while the formulation with higher percentage of fine particles revealed higher apparent density compared to the reference granulometry. These observations agree with other studies where it was reported that green apparent density increases when using higher amounts of fine fraction and using suitable amount of pitch [7], Small andfineparticles display better filling ability to the inter-particle voids which lead to improved density. In addition, by increasing the fine content, more binder matrix is available to fill the coke pores. However, fine content beyond a specific upper limit may not have any significant effect on the green density [8] or may reduce it due to more inter-particle friction and particle bridging [9]. For the three pastes made with coke A, the trend of VBD of the coke blends is the same as that of green apparent density. Both VBD and green apparent density are observed to increase with increasing the fine fraction.

Table V][. Density and porosity of green samples VBD Helium GAD Sample of %Porosity 3 pycnometry g/cm blend Reference 1.875 1.498 1.18 20.11 size-Coke B

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Reference size-Coke A

1.88

1.572

1.162

16.38

5% more large-Coke A

1.893

1.551

1.147

18.07

5% more fine-Coke A

1.847

1.592

1.219

13.81

3. C. Jonville et al, "Influence of Coke Source on Anode Performance," Journal of Metals, 47 (8) (1995), 23-24.

Generally, the compression tests used in this study could be used to investigate the influences of materials properties and paste formulation on the compression behavior of paste during forming process. The data and curves obtained from this particular test provide a better knowledge of the paste behavior when it is formed by hydraulic pressing and may be used for modelling the compaction behavior.

4. R.C. Perruchoud, M.W. Meier, W.K. Fischer and W.H.P. Schmidt-Hatting, "Anode Properties, Cover Materials and Cell Operation" (Paper presented at the 130th TMS Annual Meeting, Warrendale, Pennsylvania, 2001), 695-699. 5. ASTM D4292 - 06, "Standard Test Method for Determination of Vibrated Bulk Density of Calcined Petroleum Coke", Annual Book of ASTM Standards, 05.02 (2006), 558-561.

The percentage of porosity in green samples was determined using the green apparent density and pycnometry results. Table VI shows the density of green core samples measured by helium pycnometry compared with the green apparent density. This comparison could reveal the volume fraction of open pores within a porous material. Porosity of green samples can therefore be calculated using equation (1): n / n

.

%Porosityy =

(He Pycnometry)-GAD — He Pycnometry

w Λ nri

x 100

6. D. Belitskus, "Standardization of Calcined Coke Bulk Density Test" (Paper presented at the III t h AIME Annual Meeting, Warrendale, Pennsylvania, 1982), 673-689. 7. Kirstine Luise Hülse, Anode Manufacture: Raw Materials, Formulation and Processing Parameters (Sierre, Switzerland, R&D Carbon, 2000), 122-147.

/1ë

(1)

where He pycnometry and GAD are density measured by pycnometer, and green apparent density, respectively. The results of helium pycnometry confirmed the green apparent density measured for pressed samples. It was observed that, increasing the porosity level in the pressed samples results in reducing their green density, as shown in Table VI. However, sample with 5% more fine particles has higher GAD but lower helium density than sample with 5% more large particles. This may be related to the fact that some pores may be blocked by the binder matrix to create closed pores which increases the volume of the material measured by pycnometer and reduces the helium density.

8. T. Vidvei, T. Eidet and M. Sorlie, "Paste Granulometry and Soderberg Anode Properties" (Paper presented at the 132nd TMS Annual Meeting, San Diego, CA, 2003), 569-574.

Conclusions

11. B. Vitchus, F. Cannova, R. Bowers and S. Ningileri, "Understanding the Calcined Coke VBD- Porosity Paradox" (Paper presented at the 137th TMS Annual Meeting, New Orleans, LA, 2008), 871-873.

9. Randall M. German, Particle packing characteristics (Princeton, NJ, Metal Powder Industries Federation, 1989), 56-59. 10. R. Bowers, S. Ningileri, D.C. Palmlund, B. Vitchus and F. Cannova, "New Analytical Methods to Determine Calcined Coke Porosity, Shape, and Size" (Paper presented at the 137th TMS Annual Meeting, New Orleans, LA, 2008), 875-880.

Paste characterization may be used as an indicator to estimate anode quality. Paste compression test is sensitive to variations of raw materials properties. Compression behavior, as a paste property, may be used to correlate paste formulation and physical properties of coke to its compactability. Shape factor and texture of the particles influence the bulk density of coke and may be used to describe the compactability of the particle bed. Vibrated bulk density, however, is not the only factor which controls the density after compaction.

12. V.A. Sverdlin, G.F. Vedernikov and V.K. Fyodorov, "Optimization of Technological Parameters of Aluminum Production Pot Anode Block Vibration Forming" (Paper presented at the 121st TMS Annual Meeting, 1992), 725-730.

Acknowledgement Authors would like to acknowledge the financial support and collaboration of Alcoa. A part of the research presented in this paper was financed by the Fonds Quιbιcois de la Recherche sur la Nature et les Technologies (FQRNT) by the intermediary of the Aluminium Research Centre - REGAL. References 1. D.L. Belitskus "An Evaluation of Relative Effects of Coke Formulation and Baking Factors on Aluminum Reduction Cell Anode Performance" (Paper presented at the 122nd TMS Annual Meeting, Warrendale, Pennsylvania, 1993), 677-681. 2. W.K. Fischer et al, "Determining Prebaked Anode Properties for Aluminum Production," Journal of Metals, 39 (11) (1987), 4345.

1

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011

FURNACE EFFICIENCY ENERGY AND THROUGHPUT

ORGANIZER

Thomas Niehoff Linde Gas Unterschleissheim, Germany

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Light Metals 2011

FURNACE EFFICIENCY ENERGY AND THROUGHPUT

Session I SESSION CHAIRS

Russell Hewertson Air Products and Chemicals Inc. Allentown, Pennsylvania, United States Thomas Niehoff Linde Gas Unterschleissheim, Germany

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Furnaces Designed For Fuel Efficiency TMS David W. White *TMS (The Minerals, Metals & Materials Society); 184 Thorn Hill Rd.; Warrendale, PA 15086, USA The Schaefer Group, Inc. 1500 Humphrey Ave. Dayton, OH 45410 USA

Furnace efficiency, energy saving furnaces, aluminum melting and holding furnaces

Introduction The purpose of this paper is to enlighten you about the various designs of aluminum melting and holding furnaces and how they relate to energy efficiency. The word efficiency is thrown around with little regard to its true meaning. There are three types of efficiencies in every furnace. There are burner efficiencies, furnace efficiencies and melting and holding efficiencies. They are not the same and cannot be looked at as individual items when making a decision on a new furnace, they are equally important. The first one that should be looked at is furnace efficiency. It is the only one we will talk about today. Does the furnace design make sense from an overall melting and holding standpoint? You can buy four different radiant roof melting furnaces and end up with three different actual BTU's/# of metal melted. Does it take more net BTU's in one gas fired furnace than the other to melt a pound of aluminum? No! But the amount of Gross BTU's required to maintain that net BTU's has to do with furnace efficiency. Adding pre heat hearths and circulation to the furnace will increase BTU's/# melted efficiencies but not combustion efficiencies. In breaking these down one by one we can better understand the importance of each type of efficiency.

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TMS: Furnaces Designed for Fuel Efficiency

Furnace Efficiency Three different furnaces all rated at 7,000#/hr natural gas fired can be designed differently to effect efficiency of the melter. In the first two examples the furnace design is a radiant roof melter with 6 flat flame burners in the roof about 24" off the bath. Total connected capacity 14,000,000 BTU's/hr. These units have a standard 3 component refractory lining with good insulating qualities. They all have standard cold air combustion systems and operate very close to this Sankey Diagram, (Fig 1)

Energy Diagram Cold Air Burner y^

F \

FURNACE

This is a 7,000#/hr high headroom furnace with two sidewall fired burners @ 7,000,000 BTU/hr each. It has a preheat hearth and a circulation pump in the charge well. It will melt for between 1400-1500 BTU's/# of metal melted. Because of the higher sidewalls required to place the burners up out of the metal this furnace requires more BTU's/# to melt because its fixed heat losses are greater. Also since transfer of heat in this case is mainly convection instead of radiant the transfer of BTU's is slower. If this did not have circulation the BTU's/# of metal melted would be about 1700BTU's/#. Radiant heat is the fastest transfer of available heat into aluminum. Stephan-Boltzmann law of radiant heat transfer stated that the closer the heat source is to the object being heated the faster the transfer of BTU's. Also the greater the temperature differential between the heat source and what you are heating the faster the transfer of BTU's. By running a low headroom with a thermal head of 1950 degrees F you can transfer the BTU's into the metal at a faster rate than in a higher headroom furnace. In doing so you do create more dross and that is why the circulation is so important. The flowing metal under the surface strips away the BTU's at a faster rate than the stagnant pool of metal. Kirchhoffs law explains why this works and is really just common sense: When radiant energy hits any particular material it has to do one of three things:

Cold Air Combustion System Fig 1 Slide courtesy of Bloom Engineering

1) Be absorbed (a perfect black body with emissivity of 1 absorbs all the radiant energy)

Furnace #1

2) Be reflected (the opposite of a perfect black body - a material with low emissivity)

This is a 7,000#/hr radiant roof furnace with a charge well and a preheat hearth and no circulation of the metal. You can preheat sows or ingots on the hearth to save about 15% in fuel. You can melt about 45% of the total load in scrap. This melter will hit about 1400BTU's/# of metal melted depending upon the mix of scrap and new material. This furnace will lose about 3% in metal melt loss.

3) Pass through the material (transparency) the energy is being pulled through the reflective metal by the flow.

Furnace #2

Energy Saving Options:

This is a 7,000#/hr radiant roof furnace with a charge well and a preheat hearth and circulation of the metal by mechanical means with a pump in the charge well.

In addition to these relatively inexpensive options to furnaces there are a few other ways to save more energy when buying a new furnace.

This furnace will melt for about 1250 BTU's/# of metal melted due to the increase in efficiency form the circulation process. This circulation actually helps pull the BTU's from the hottest metal on the surface to the middle of the bath and heat up anything being melted under the bath much faster. Metal melt loss will be between 2 and 2.5%

1. Refractory Linings: Buy the most energy efficient lining you can buy. You can control fixed heat loss. The insulating materials available now will reduce casing temperatures to 145 degrees F under the metal level at the casing and about 170 degrees F above the metal line at the casing. This is significantly cooler than furnaces built just ten years ago. This represents a payback on the higher insulating material costs of less than one year in energy saved. Not to mention the cooler building and less make up air required to cool down the melt room.

The energy of the radiation absorbed, reflected and transmitted through the material must equal the energy hitting the material. This is why the larger surface area and bath capacity are so important in radiant roof fired aluminum melting furnaces.

Furnace #3

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TMS: Furnaces Designed for Fuel Efficiency

2. Heat Exchangers: Look at putting heat exchangers on the flue of the large furnaces. This can save you on average about 23% in fuel (fig 2) and represent a payback of less than 2 years. If you have a series of large furnaces then waste heat recovery becomes a very viable option to see a great return on your investment. While this will be much more capital intensive and depending upon current utility rates for recoverable energy purchased, the payback is swift and very beneficial

Energy Diagram Recuperative Burner

Fig 3 Regenerative Combustion 4. Well Covers: Covering up open wells will reduce heat loss substantially. An open well of 1300 degrees F aluminum with a .5 emissivity of the bath surface in the well will lose 8,214 BTUs/square foot/hr. If the average charge well is 8' long by 3 feet wide that is 24 sq. ft. or 197,136 BTUs per hour being lost off that well. That is 197 cubic feet per hour or 4,731 cubic feet a day x 340 days a year is 1,608,540 cubic feet lost.

BASIC REGENERATIVE FURNACE Fig 2 Slide courtesy of Bloom Engineering

5. Preheat Hearths: By ordering a furnace with a preheat hearth designed into the unit you can save enough money to pay for it in less than two years. By placing two 1,000# sows on the hearth and letting them pre-heat to a sweat, the internal temperature of that sow will be 900 degrees F.

3. Regenerative Burners In very large high headroom melters, regenerative burners are still saving metal casters about 40% in fuel usage over conventional burners. The ROI payback this year is extensive because of the low fuel costs but we all know the price of natural gas is only going to go up so an investment in regenerative burners today that has an ROI of 50 months might actually end up being 30 months if the gas prices hit double digits again soon.

This step uses the heat that normally goes up the flue and some radiant absorption from the burners to pre-heat these sows. When placed into the bath as the next two sows are loaded, the stored BTUs in the bath help finish the melting process, taking the sows very quickly from 900 degrees to melting temperature. This will save 15 percent in fuel for that 2000# of aluminum. If you were melting at 1800 BTUs/# and you now pre-heat, you are at 1530 BTUs for every # you melt through the pre-heat hearth. If we use that same 7,000 #/hr unit as in other studies above, that will save you 21,600,000 BTUs per day. So the net savings is 7,344,000 cubic feet of gas per year saved. That alone saves $47,736/year.1 Especially in the case of tilting or barrel type furnaces you need to do what ever you can to improve their efficiencies. These will typically melt for 1600-2,000 BTU's/# and have extreme heat losses. Many are lined with brick which is one of the most inefficient linings available and is 40 year old technology. Flat roofs are more efficient that rounded roofs in these furnaces because of the law of radiant heat. Flat roofs allow the heat to be more evenly transferred into the bath of metal. Electric tilting furnaces (fig 3) are much more efficient that gas melting for less than 900BTU's/#

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TMS: Furnaces Designed for Fuel Efficiency make up some drop in temperature if cold metal is delivered to them or if they are being drawn down to low before refilling them. Remember once you make this furnace have to raise metal temperature you have turned it into a mini melter and if it does not have the power connected to raise the metal temperature quickly then you will shut down the casting process waiting on the holder to catch up. Aluminum holding furnaces whether gas or electric should have highly insulated linings in them. Four component linings are the most efficient; this can either be an all insulating non wetting board lining or a dense castable with non-wetting back up calcium silicate boards, micro porous silica board and block insulated linings. Lids that seal and the capability of repairing that seal easily (anytime it is needed) will also conserve energy. They should have well covers and possibly automated well covers if the cycle times are greater than 1.5 minutes. Electric Immersion Holding Furnaces The Electric Immersion furnaces are being designed today with casing temperature at or below 105 degrees F. They are much more efficient than the radiant electric holders since the heat source is under the bath of metal the heat can radiate off all sides of the protection tube and distribute more heat where it is needed most in the metal. Because the heat is generated under the bath of aluminum there is very little dross created. The lack of high thermal head temperatures virtually eliminated the need to clean these furnaces more than twice a week.

Fig 4 Electric flat roof tilting furnace Furnace #4. The Stack Melter: The most energy efficient gas fired aluminum melting furnace is still the Tower or Shaft melter. From a strictly BTU's/# of metal melted there is nothing in gas that can come close without auxiliary equipment like heat exchangers or regenerative equipment installed on the basic units. Stack melters can melt as low as 950BTU's per pound of metal melted if you keep the stack completely full and the burners are in ratio. But if you let the stack empty then it is no better than a dry hearth furnace and that melts at about 17001800BTU's/#. However, things like up front costs and higher metal melt loss and maintenance can affect the ROI for these furnaces.

itΙΚΚ

Furnace #5 The Radiant Roof Electric Melter:

m

The absolute best and most efficient furnace from a BTU's/# of metal melted standpoint is still the radiant roof electric melting furnace. This unit will melt for less than 790 BTU's/# of metal melted. Electric melters melt at .21KW/# simple math says that .21 x 3,412 (BTU's/KWH) = 750.76 BTU's/#. If you add circulation and Micro porous super insulated lining to this electric melting furnace you will be down to about 687 BTU's/# of metal melted. No one can even come close to that. Now move the elements under the bath with immersion electric melting and you get almost 77% efficiency at 655 BTU's/# melted. With the uncertainty of what will happen here in the U.S. on carbon footprint and cap and trade more and more die casters are looking at in cell electric melting and holding at the die cast machine as an alternative to gasfiredcentral melting. Holding Furnaces: Holding furnace at die cast machines or at foundry casting lines are designed to do one thing...hold the metal at casting temperature. They are still being designed with the lowest possible BTUs or KW connected to do the job. This is such a dis-service to the industry. It does not cost that much more to connect a passing gear to these units to allow them to

||j|e

1

m

« a

^ * ~'

04/03/2007« The bottom line is yours. Do everything you can to enhance it by buying the most energy efficient furnace that fits your needs and has the best ROI. None of these energy saving ideas are free. They all cost money to do. However your return on investment over the next 3 years could make the difference in "being in a profitable business" or going through a "going out of business" sale of all the old gas guzzling furnaces you kept.

1172

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

LATEST TRENDS IN POST CONSUMER AND LIGHT GAUGE SCRAP PROCESSING TO INCLUDE PROBLEMATIC PROCESSING MATERIAL SUCH AS UBC, EDGE TRIMMINGS AND LOOSE SWARF Franz Niedermair1, Guenther Wimroither2 1,2

Hertwich, Weinbergerstrasse 6, Braunau, A-5280, Austria Recycling, UBC, Scrap, Furnace Environmental compliance

Introduction

Minimizing emissions and pollutants like NOx, Dioxin, VOC, without complex post-combustion, lowest possible and in most cases no addition of salts

Developed in the late 1990's by Hertwich, a member of the SMS Group, the Ecomelt furnace has been the best available technology for the processing of post consumer scrap aluminum for many years. Traditionally, charge materials such as painted extrusion profiles, extrusion profiles containing thermal break, painted siding (new and old) and litho scrap, have been charged into the melting unit and turned into molten aluminum whilst minimizing melt losses and attaining the best fiiel efficiencies in the industry.

Minimizing risk of accidents, automation Low operating costs Classification of Problematic Scrap Types

With a greater demand being put on manufacturers and suppliers of aluminum products to produce a 100% recyclable material, the Hertwich furnace is being continually developed to have the ability to process the traditional problematic materials such as loose chips, edge trimmings and used beverage cans (UBC'S).

TypeA
For many years there have been technologies on the market for the processing of these problematic materials, but with more and more pressure being put on processors of these materials with respect to fuel consumptions, overall metal recoveries and environmental impacts a more efficient process needed to be found. The Ecomelt was able to process these material in a very efficient and safe manner; resulting in the best fuel efficiencies and best overall metal recoveries in the industry.

Swarf briquettes, cut wire, foil, profiles, lithographic sheets, ... Impurities are typically oil, grease, ink, paint, lacquer,... Type C, VOC > 5 %

Aluminium Industry Requirements1 For melting of Aluminium scrap the main emphasis is on meeting the ideal ecological and economic conditions in the best possible way. Ever more stringent requirements can be met only with the most up to date melting technologies, to meet the following criteria: •

High product value and quality, avoiding downgrading of scrap

•

High yield/recovery, low metal loss

•

Low energy consumption

•

Post consumer scrap such as dirty profiles, UBC, coated sheets »foil, tubes, wire,... Typically impurities: high amounts of grease, paint, powder coating,... Type D, VOC > 10% Dross, Twitch, Engine blocks, profiles with thermal break,...

Oil, plastic and paint = hydrocarbons to energy

s s

^ν-Λ '

Typically impurities: iron parts, high amounts of plastic, rubber, ...

A high percentage of non metallic contaminants of scrap must be permissible. The recycling equipment must be capable of processing non metallic contaminants, the inherent energy of which is reclaimed. Net energy input is considerably reduced.

Although there are four classifications of scrap listed above, for the purpose of this article we will focus on the two main problematic scrap types; Type B and Type C which contain between 1% and 10% V.O.C.

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Functional Solution to Effective Recycling The solution to effectively and efficiently recycle these materials is a melting furnace with integrated scrap preheating and gasification. The HE-Ecomelt (economy, ecology = environmentally sound) Furnace. This furnace type is designed for remelting of scrap contaminated with oil, paint and plastic.

(Preheating and Combustion) During preheating the scrap charge is heated to a temperature of approx. 500 °C, which at any rate is way below the melting temperature of aluminium. Surface fusing and unnecessary increased oxidation of aluminium in the furnace atmosphere must be minimised.

The total non-metallic, organic substances may reach 10 % of weight, subject to calorific value. Typical main components of the system: Charging system Preheat gasification chamber Melt chamber Furnace main chamber

It is imperative that hot gases flow through the scrap load rather than passing along the outer contours of it. The hot air flow through the scrap load allows simple control of the heat front within the scrap load and with that the gasification rate and quantity of pyrolysis gases. Sudden occurrence of uncontrolled peaks of pyrolysis gases, which are typical for conventional dry charging, can thereby be controlled easier.

(Two-Chamber Furnace) Charging Charging is done automatically with several different mechanisms depending upon application. The volumetric capacity is determined in relation to the specified scrap density, so that no more than three charging cycles per hour are necessary. In case more than one container is to be used, a dedicated buffer storage system can be integrated. Full containers are moved into charging position by a lift system. During charging a lock mechanism prevents escape of flue gases and dust. The scrap load is emptied into the preheat/gasification chamber, upon opening of the containers' bottom flaps.

TEMPERATURE GRADIENT WITHIN THE SCRAP LOAD AT TIME INTERVALS AFTER CHARGING

β I «

Scrap preheating and combustion of VOC particles In the gasification chamber the scrap load is exposed to an intense hot gas flow and organic contaminants are transformed into combustible gases. Part of these pyrolysis gases is precombusted at the burner system of furnace chamber 1 (melting chamber).

1

1

2

i

3

1

4

i

5

1

*

1

7

1 »

1

DISTANCE PROM HOT GAS ENTRY ÈMETER]

Longitudinal section of gasification compartment. The diagram shows the temperature of &e scrap charge at time Intervals after start of heating. Since the heat front moves gradua^ torn entry of hot gases lo the exit» gasification also moves gradually through the charge {see green and yellow hand), this avoids peaks in the generation of pyrolysis gases.

Since no scrap is charged into chamber 2 of the furnace, a constant high furnace temperature of more than 950 °C can be maintained. Possible toxic chlorine compounds, which may occur upon gasification of organic contaminants, are no longer stable at such high temperatures and are disintegrated. All flue gases from the furnace are ducted from chamber 2 through the flue gas duct to two alternating regenerators, which cool down the flue gases to below 200 °C within split seconds. The fast

Thereafter pyrolysis gases are ducted to chamber 2 (main chamber) where they are completely combusted with the adjusted excess oxygen of 2 - 3 %. During peak generation of pyrolysis gases some 50 - 70 % of the energy supply come from combustion of pyrolysis gases and merely 30 - 50 % of the energy needs to be supplied as natural gas, to ensure thorough combustion.

1174

cooling down prevents recombination of dioxins in the critical temperature range 600 °C to 250 °C.

Ecomelt PR - Two Chamber Melting Furnace with preheat ramp

mmmmmmmmmimm*»**

PROS: suitable for contaminated scrap (class I + II, VOC<5%)

(Regenerators) In the regenerators combustion air is preheated to approx. 800 1000 °C. Such intense air preheating in combination with the thermal utilization of pyrolysis gases, leads to extremely low energy consumption values.

low energy consumption (<550 kW/t) with regenerative burners minimum melt loss (submergemelting)

During melting of moderately contaminated scrap the typical gas consumption is merely 400 - 500 kWh/t. After passing through the regenerators flue gases are ducted to a filter plant.

feed back on alloy composition immediately after each charge

The Melting Process The melting process of the clean preheated scrap is as follows: • • • •

CONS: investment cost long preheating time for proper gasification of organics > 30 min for loose scrap > 45 min for bailed scrap melting rates (max. 5-6 t/h in relation to furnace size sophisticated operation required (cleaning, maintenance, automation)

Ecomelt PS - Two Chamber Melting Furnace with preheat shaft

Placing of the preheated and de-coated scrap into the melt of the melting chamber is achieved, subject to furnace design, by: Opening of scrap flaps in the gasification chamber Raising of the melt level to above the dry ramp level (flooding) Pushing the scrap heap into the melt by activating an integrated pusher mechanism

In any case a strong melt flow is induced between melting- and main chamber, by an electromagnetic pump. Preheated metal is melted in the liquid metal of the melting chamber or on the flooded ramp, by the hot metalflowfromthe main chamber.

PROS: suitable for contaminated scrap (class I +11 +III, VOC up to 10 %) !

optimum preheating (> 45 min) for proper gasification ofVOC low energy consumption 1 (<550 kW/t) with regenerative burners

(Melting) Hertwich Ecomeit Solution for Type 2 and 3 Scrap Materials

minimum melt loss (submerge-melting)

To address the concerns when recycling Type 2 and Type 3 scrap HE have developed two specific types of melting furnace over the past decade. Both of these furnace types have been proven in the industry over the past few years.

high melt rates even with low density scrap (up to 250 t/day)

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CONS: investment cost sophisticated operation required (cleaning, maintenance, automation)

The ultimate eco-melting system for light gauge scrap and chips

The swarf dryer and melting furnace is combined into one thermo-mechanical system. Energy for heating of swarf is introduced to the dryer in the forni of hot gases from the furnace chamber. These hot gases are mixed into the circulating gas flow of the dryer. Exhaust gases from the dryer, which also contain hydrocarbons, are ducted back into the melting furnace, where hydrocarbons are burned in the hot atmosphere of approx. 1000 °C and slight oxygen excess. Therefore an expensive, energy intensive after burner system is not required, but burning of hydrocarbons from the swarf drying contributes substantially to the energy requirement of the melting furnace.

Due to the increasing use of aluminium castings and forgings in the automotive industry, the generation of swarf is on the rise. The worldwide volume of swarf generation is estimated to soon exceed two (2) million tons per year. Today's common practices and routines for remelting of swarf are associated with severe disadvantages and also with negative impact on cost effectiveness such as: • Low metal yield • High energy costs • Extensive handling • Emission problems of fumes and solid residue (e.g. salt cake)

Swarf is heated to some 400 °C in seconds, and hydrocarbons and moisture are evaporated. Further oxidation of swarf due to lengthy exposure to high temperatures is prevented.

A new concept developed by Hertwich eliminates these disadvantages and especially ensures cost effectiveness.

Swarf is introduced directly to the melting process in hot condition. The inherent thermal energy is retained as dryer and melting furnace is combined into one process. Swarf is free from moisture and hydrocarbons. For instance when swarf is left to cool down to ambient temperature in a covered building, the swarf again absorbs 0,1 % of its weight in moisture in a short period of time (= 1 kg per ton) (van der Waal's force3). In case this moisture gets into the melt, it supports the increase in hydrogen content as well as oxidation losses of aluminium. (Plant Configuration - Swarf Remelt Plant)

•

After the usual swarf treatment steps such as certifuge, breaker and iron separator, swarf gets to a high speed swarf dryer and is subsequently charged and melted in continuous operation. The entire process is characterized by the following features:

During charging into the furnace side well, hot swarf meets a vertical, downward melt flow generated by an electromagnetic linear-motor, and is quickly submerged into the melt. Swarf melts by direct contact to the liquid metal without any oxidation.

Thefirstplant of this type was commissioned for a prominent German producer of forgings, where swarf is delivered to the swarf dryer directlyfromthe machining centres.

Z3 cmm (Swarf Plant equipped with charge hopper, pre-treatment high speed co-flow dryer)

(Swarf Drying)

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The second, and obvious one was, once shredded into a light gauge scrap component, would the shred allow the through processing into molten aluminium in the submergence melting system that is employed in the melting furnace.

The Industry Dilemma; Efficient and Effective UBC Recycling. Used Beverage Cans (UBC's) have always presented the aluminium industry with a problem with respect to effective recycling. By definition, the UBC is a "hybrid" product; consisting of thin gauge aluminium made from varying alloys, various coatings, and due to the nature of the product varying amounts of trap contaminants. When processing in a conventional furnace the product is virtually "unrecoverable". When processed in a rotary salt furnace the resultant salt cake is a disposal issue and when processed with a decoating system and subsequently charged into a reverbatory furnace, the resultant gas usage from both processes is relatively high in comparison.

The unique airflow and process that Hertwich developed ensured that the fuel efficiency of the furnace would be superior to that of any other technology available. As a hot gas stream is passed through the post-consumer scrap charge, contaminants in the form of vaporized paint and lacquers are driven off. The resultant gas is then passed through a regenerative burner system. This ensures superior fuel efficiencies. The anticipated fuel usage was expected to be half of the fuel usage when using conventional technologies. Additionally, since the preheat chamber has a controlled oxygen atmosphere, it was hoped that the furnace would operate as expected.

In order to get a full understanding of the complexities involved in the successful decoating and melting of UBC's Stein Atkinson Stordy4 did considerable research into the make-up of a typical UBC. The experiments focused on being able to determine the amount of the Volatile Organic Compounds (V.O.C.) evolved from the coatings when subjected to different temperatures in a controlled oxygen atmosphere. Additionally, it was necessary to establish the amount of weight loss relative to the coatings. This information would give a good insight into what temperature was optimal for the successful decoating of the UBC.

A product furnace was built and the process was evaluated for performance. A computer generate schematic of the furnace is shown below.

The experiments consisted of cutting the UBC's into 3' x V£" strips, weighing the strips and heating at the desired temperature for one minute. The shredded UBC was then allowed to cool, and re-weighted to assess the actual weight loss and hence the VOC loading on each sample. Temperature increments of 200 degrees Fahrenheit from 300 upwards were conducted up to 1,000 degrees Fahrenheit.. At each temperature the total weight loss for each shred was measured. The results are shown in the chart below:

Air TemoeraturefF) 300 500 700 900

% Weight Loss of VOC 3.2 59.9 23.8 13.1

It was noted that when the furnace was running with briquetted product (class scrap used to emulate UBC) as the charge it was operating very efficiently and without problems. However, when the system was running with shredded UBC it was noted that the required production requirement (molten metal output) could not be obtained due to the density of the shredded scrap and bridging of the scrap within the preheat/decoating chamber. After many experiments with briquetted product and briquetters the decision was made to briquette the shredded UBC to a density of between 40 and 60 lbs/cubic ft.

It can be seenfromthe above chart that all V.O.C is removed by the time the peak metal temperature is 900 degrees F. It was also noted during these test that when the peak metal temperature was raised to 1,000 degrees F that the shredded sample began to put on weight, or rather oxidise.

Prior to this decision being made several experiments were made to be certain that the briquetted product could be correctly delacquered prior to melting. Conventional wisdom dictated that the ability to decoat a heavily compacted briquette was nonexistent due to the surface of the product not being exposed to the hot air stream.

Based upon this information it was deemed that the Hertwich Eco-melt furnace was ideally suited to not only decoat the UBC but also to melt it in one single process. However, there were some challenges to overcome with the processing of the UBC. The first one was the need to have the product shredded prior to processing to remove the tramp material from the product.

A series of experiments were conducted within the furnace to disprove this theory. It can clearly be seen in the next series of pictures that in fact, high density briquettes can be efficiently decoated in this form.

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The picture below shows the decoated product in the preheat shaft. Clearly the outside of the product has been "cleaned" but the inside cannot be viewed.

Conclusion and Observations In conclusion we can clearly state that over the past decade the advancement in the processing of post consumer scrap has increased dramatically. It is now possible to process many different kinds of post consumer scrap aluminium in one single unit. The Hertwich Ecomelt furnace continues to be developed and redesigned to enable operators the ability to process many types of scrap. This they can do while simultaneous meeting the following criteria:

Once the briquettes were broken open as shown in the pictures below it was evident that the decoating efficiency of the H.E. melting furnace was exceptional. Based upon these test results it is clear that a dedicated briquetting press to process the shredded UBC prior to charging the melting furnace is necessary.

•

Achieve the lowest possible fuel consumptions in the industry

•

Realize the best possible metal recoveries in the industry

•

Meet and exceed all environmental requirements regarding emissions to atmosphere

References Published Papers 1.

F. Niedermair, "Latest Trends in Scrap Remelting" (OEA 2009, Berlin, Germany, March 2009)

Research Documents Theorems 2.

Johannes Diderik van der Waals (November 23, 1837 -March 8,1923)

Unpublished Reports 3

The recoveries from this furnace when processing UBC are in line with industry standards for short periods and it is hoped that these figures will be improved upon in the upcoming months. The fuel usage obtained with this furnace is 600-800 Btu/Lb (400-500 kWh/t). This is considerably less than conventional melting furnaces. Throughputs of up to 400,000 lb/day of molten metal are readily attainable.

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O.H.Perry, "UBC Coating and Resins Report" (Internal Report January 1991)

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Investigation of heat transfer conditions in a reverberatory melting furnace by numerical modeling Andreas Buchholz1, John Rodseth2 ^ydro Aluminium Rolled Products GmbH, R&D, P.O. Box 2468, D-53014 Bonn, Germany 2 Hydro Aluminium as, R&D Materials Technology, Karmey, N-4265 Hβvik, Norway Keywords: melting, reverberatory furnace, numerical modeling Abstract A numerical model based on the commercial software ANSYS FLUENT was used to analyze the heat transfer conditions in an industrial reverberatory melting furnace. The model comprises different physical phenomena as gas flow, chemical reactions, i.e. combustion, conduction, radiation and latent heat release in the metal. The gas circulation was analyzed for different metal loads and burner arrangements. The melting process inside the furnace is inherently time-dependent, and a complete transient analysis is very time-consuming. To study the impact of varied burner positions on the energy utilization, stationary solutions were applied, where the effect of melting heat was approximated by artificial heat sinks inside the metal. The results of the stationary solutions were compared with fully transient results to guarantee the transferability of the cost-efficient steady state calculations. The results stress the dominant effect of radiative heat transfer in the melting process. Introduction Rising energy prices, the economic crisis, national commitments for carbon footprint reductions and the need for a better image of the aluminium industry have put a constant pressure to reduce energy consumptions in casthouses worldwide, but especially in Europe. Since the resources for new investments are limited, the existing casthouses have to achieve energy savings by moderate modifications of existing equipment and improved procedures in the operations. In recent years Hydro Aluminium R&D has developed a furnace model to analyze impacts of furnace operations on melt quality [1][2]. Since this model covers all essential heat transfer mechanisms required to describe the heat flow in a reverberatory furnace, it was decided to investigate the melting process in an existing melting furnace and to analyze various options to improve the utilization of the fuel energy. The furnace has a nominal capacity of 35 t and is run in a batch mode. About 20 t of aluminium are charged as cold metal and later 12 t as potroom metal. The rest is the heel of the previous melting process. The furnace is heated by two cold air burners with a nominal power of 2.5 MW and 1.0 MW. The study should answer, what factors dominate the heat transfer, whether a modified burner position could achieve a better utilization of fuel energy and how the metal charging and burner operation could be optimized to reduce the fuel consumption.

Numerical model The furnace model is based on the commercial software ANSYS FLUENT 12.1. The model covers the following physical phenomena: 1. 2. 3. 4. 5. 6.

gas flow combustion (chemical reaction) heat conduction and advection in the gas heat conduction in the metal and the walls radiation heat of fusion.

The gas flow is modeled as incompressible turbulent flow using the RNG version of the K-epsilon model. The mixing rates calculated in the turbulence model are the basis to model the chemical reactions of the combustion process. This so-called Eddy-Dissipation model assumes a fast chemistry, where the speed of reaction is exclusively determined by the turbulent mixing rate. The oxidation of natural gas is approximated by a simplified 2-step reaction scheme for methane gas decomposition: (1) CH4 + 1 V2 0 2 => CO + 2 H 2 0 (2) CO + V2 0 2 => C0 2 For each chemical component in the reaction scheme an individual transport equation has to be solved. The radiation is calculated with the so-called Discrete Ordinate approach, where radiation transfer equations are solved for a finite number of space angles in each computational cell. Besides calculating the radiation between surfaces this model is also capable to capture absorption effects of radiative gas components. To account for heat of fusion, a latent heat term was implemented into the heat transport equation by user defined functions, since FLUENT'S built-in solidification model is not compatible with the combustion model. The latent heat was introduced as an effective specific heat into the temperature formulation [3] of the heat transport equation. Specific measures guarantee that no latent heat is lost in the iteration process. The model was validated in experiments in a lab scale furnace [4]. One major limitation of the model at the moment is that the metal shape will not change during the melting process. Therefore, redistribution of the liquid metal will not be described fully correctly. To encompass the consequences of this shortcoming, two different metal arrangements were considered. In the first configuration the metal is considered as a melt pool spread at the bottom of the furnace cavity, in the second version the metal is arranged as an ingot assembly including gaps, which allow for the circulation of combustion gases between the ingots and thus increasing the effective metal surface. Furnace geometry The furnace geometry was strongly simplified. Details of the steel construction were neglected. To capture heat capacity effects of

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the refractory material the walls were modeled as thick walls. In the original configuration the 2.5 MW burner is opposite to the front door and the smaller 1.0 MW burner left hand of the door. Figures 1 and 2 show the two basic cases of this study. Figure 1 assumes that 20t of metal are spread at the bottom of the furnace, in Figure 2 the same amount of metal is arranged as a stack of 27 ingots. The majority of the geometry was meshed using tetrahedral cells and later converted into a polyhedral mesh. The final meshes consisted of about 450000 cells and 2 million nodes.

Figure 1. Melt pool configuration A: 201 of metal spread at the bottom of the furnace. The colors represent temperatures in °C.

Figure 2: Ingot arrangement, configuration B: 201 of metal distributed as a stack of 27 ingots. Besides the two different metal distributions several modifications of burner positions were investigated. The main idea was to find a configuration with a longer residence time of the combustion gases. It was assumed that the extended residence time could lead to a higherfractionof energy transferred to the metal. The options to find a reasonable burner arrangement are limited. Besides the original configuration one promising approach seemed to place the burners to the flue side. An overview of the different arrangements and development of the related gas flow patterns is shown in Figure 3.

Figure 3. Overview of investigated furnace configurations and development offlowpatterns. Blue pathlines are generated by the 2.5 MW burner, the orange patterns by the 1.0 MW.

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Configuration A and B represent the original furnace layout, in A the metal is spread at the furnace bottom as shown in Figure 1, while in B an ingot stack as in Figure 2 is considered. In configuration C the burners were shifted to the flue side. In this setup the larger burner was slightly tilted towards the furnace center. Since the further flow pattern showed an overweight of gas circulation close to the front door, this setup was further developed into configuration D: The 2.5 MW burner and the 1.0 MW burner were exchanged at their positions and adjusted to be parallel to the flue axis. This measure provided a more uniform circulation pattern. Finally, this burner arrangement D was applied to the ingot stack, as shown in configuration E. The geometry in configuration F is identical to A. The only difference is a changed composition of the oxidizer to imitate oxy-fuel burners, as will be discussed later.

Table I. Thermo-Physical Data and Process Parameters Property/Parameter Value Unit aluminium density 2350 kg/m3 J/(kgK) aluminium specific heat 1080 W/(mK) aluminium heat conductivity 200 aluminium emissivity 0.3 aluminium latent heat of fusion 390000 J/kg refractory density 2320 kg/m3 refractory specific heat J/(kgK) 1138 refractory heat conductivity W/(mK) 0.5 refractory emissivity 0.6 air flow, 1 MW burner 1200 m3/h natural gas flow, 1 MW burner m3/h 100 m3/h airflow,2.5 MW burner 1970 natural gasflow,2.5 MW burner m3/h 170 computed combined burner power MW 2.3

Steady state calculations The flow patterns shown in Figure 3 are generated from steady state calculations, since it turned out that fully transient calculations of a complete melting cycle are very time consuming and costly. When calculating the steady state of a furnace running with burners at melting power the solution will finally show a completely superheated furnace with an unrealistic temperature distribution. The steady state solution needed in this work should resemble a snapshot somewhere in the middle of a transient melting process. Therefore, it was necessary to place an artificial heat sink into the metal, which imitates absorption of heat of fiision in the real transient process. It was assumed that an appropriate description of a volumetric heat sink term in the metal can be provided by a formula of the type: Q = -a S ink(T-T Ref )

Figure 4 shows a comparison of average metal and flue gas temperatures and of the average fraction liquid of the transient calculations of configurations A and B. Transient calculations: Melt pool A versus ingot stack B 1200 -i

^τ 1.2

(1)

The advantage of this formula is that the metal temperature can be used as an indicator to compare the efficiency of heat transfer conditions in different setups. A higher average metal temperature would point out a more efficient furnace configuration. Some estimates based on the real melting performance and subsequent adjustments lead to the following choice of heat sink parameters, which were used throughout the steady state calculations: ocSink= 355 J/(m3s) and TRef= 500°C. It was assumed that this heat sink together with identical boundary conditions could provide the means to calculate temperatures and heat flux distributions of the different geometric configurations, which can be used to compare the relative performance of the different setups.

o

„....._ 0

5000

... 10000

-r 15000

o 20000

25000

30000

35000

40000

time [s]

Figure 4. Results of transient calculations of configurations A (melt pool) and B (ingot stack): Average flue gas temperatures (orange/red), average metal temperatures (light green/dark green) and averagefractionliquid (light blue/blue). The average temperature of the ingot stack is always ahead of the melt pool configuration. This is a consequence of the increased specific metal surface of the ingot stack, the direct impact of the flame onto some ingots and the circulation of the combustion gases between the ingots. Therefore, the metal starts earlier to melt and the flue gas temperature is lower - at least in the beginning - compared to the melt pool. After about 20000s this situation changes: the melting of the melt pool is accelerating and finally the average fraction liquid catches up with the ingot stack at 33000s. Finally, the fully liquid state is reached slightly earlier (after 37965s) than in the ingot arrangement (39720s). This result appears to be unrealistic, especially since the average temperature of the ingot stack is still higher. It follows from the model limitation that all metal stays in its initial location and cannot flow away after melting. Finally some parts of the stack are very hot, while other regions are shielded from the direct impact of burner the burner flame and remain comparatively cold. In this respect conclusions have to be made with care from the computed data at the end of the melting process. Generally, it can be concluded that the ingot stack exhibits more favorable heat transfer conditions compared to the melt pool configuration.

Transient calculations To assess how far the steady solutions are in agreement with a fully transient analysis, configurations A and B were calculated as transient cases. These simulations were based on previous steady state solutions, but the temperature values of the metal were reinitialized to 20 °C. The artificial heat sink term in the metal was switched off, while the transient latent heat term was activated. Table I gives an overview of typical thermo-physical data and process parameters used in the steady state and transient calculations. To avoid over-prediction of the temperatures in the combustion gas a modified polynomial description of gas specific heats was used as recommended in the FLUENT documentation [5]. The composition of the natural gas was assumed to be 86% methane, 1% carbon dioxide and 13% nitrogen.

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When integrating the fuel flow until the fully liquid state is achieved, a specific melting energy can be calculated for both configurations. The melt pool calculation yields a value of 1200 kWh/t versus 1295 kWh/t for the ingot stack. This compares rather well to the typical value of 1200 kWh/t which is usually reported for cold burner installations. Actually the calculated specific consumption may be even too pessimistic. Since certain areas of the metal are very much superheated, when the last solid fraction is molten, the complete melting could have probably been achieved, even when the burner power had been reduced earlier and superheat utilized for melting of the remaining solid. One question, which should be answered by the transient calculations, was how much the earlier described steady state computations are representative for the complete transient melting process. When comparing average flue gas and average metal temperatures of transient and steady state calculations it turns out that they match at different times (see Figure 5). When the flue gas temperatures are identical at about 26000s, the metal temperature of the steady state solution is about 70 K higher than in the transient solution. A better match could have been probably achieved when choosing lower values of ocSink and TRef in equation (1). Nevertheless, looking at the heat fluxes it can be noticed that the total, radiative and convective heat fluxes through the metal - gas interface are very similar for both the steady state solution and a broad period of time in the transient solution (see Figure 6). It is very remarkable that the radiative component of the heat flux dominates the heat transfer between gas and metal by far.

Comparison of heat transfer conditions in steady state calculations The reasonable agreement of heat fluxes in steady state and transient calculations justifies the usage of the heat fluxes as a criterion to compare the efficiency of heat transfer conditions in the steady state calculations of the different configurations shown in Figure 3. An overview of the total heat flows is given in Figure 7. As mentioned earlier, configuration F uses the same geometry as A, but the composition of the combustion air was changed to imitate oxy-fuel burners. It turns out that the shift of the burners from the side walls to the flue side in configurations C and D has virtually no benefit compared to the original layout A. The increased effective surface area of the ingot stack implies some small gain in total heat flow in configurations B and E. On the other hand, it is remarkable that the average heat flux (heat flow per area) of these configurations is significantly reduced as shown in Figure 8. C o m p a r i s o n of total h e a t f l o w

Furnace t e m p e r a t u r e s : T r a n s i e n t v e r s u s s t e a d y s t a t e c a s e C

D

Configuration

Figure 7. Comparison of total heat flow into the metal in steady state calculations of the configurations shown in Figure 3. C o m p a r i s o n of a v e r a g e total heat flux

80.0

15000

20000

25000

30000

35000

FT700

40000

g 60.0

time [s]

Figure 5. Comparison of steady state solution (horizontal lines) and transient solution of melt pool arrangement A.

|

50.0

J

40.0

Ο £

30.0 20.0 10.0

Heat f l o w into m e l t p o o l : T r a n s i e n t v e r s u s s t e a d y state c a s e

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W 800

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700

|

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500

I 400 o

15

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± a

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^^^"7^S4-

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j

j

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F

Most outstanding is the behavior of the imitated oxy-fuel configuration F, which reveals superior heat transfer properties. Several factors are responsible for mis advantage: 1. Effectively, the nitrogen was removed from the oxidizer. This means that less gas volume needs to be heated. Theflamegets hotter. 2. A hotter flame can emit more heat by radiation. It was noticed earlier that radiation is the dominant effect in the heat transfer mechanisms and will take full advantage of a temperature increase.

300

C 200

CO

D

Figure 8. Average heat flux through the metal surface of configurations A - F shown in Figure 3.

>™~~*°~**^

Ö

'S 100

C

Configuration

. . . .

L

B

_ ^

r ·"·:.:.:: - ~\ 10000 l·'*

convective heat flow

j 1

15000

J 20000

~~

time [s]

*

25000

l — « i >

30000

35000

» «

» «

40000

Figure 6. Comparison of heat flow (total, convective and radiative contributions) through gas - metal interface in steady state (horizontal lines) and transient calculation of melt pool configuration A.

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3.

A higher fraction of combustion gases consists of radiative components C0 2 and H 2 0, which will facilitate the distribution of heat. 4. Due to the reduced amount of oxidizer the combusted gases remain longer inside the furnace and have more time to yield heat. The last item can be clearly seen in an analysis of gas residence times of the considered configurations, see Figure 9. In this context it is also remarkable that the ingot stack cases B and E also show a significantly increased average residence time. It seems that the impact of the metal stack is less to reflect the gas towards the flue than to divert it from a regular pattern. The diverted combustion gases remain longer inside the furnace. Comparison of gas residence times

burner operation and addition of liquid metal, see Figure 11 (green line). The actual state of the metal inside the furnace is described by point A on the conversion curve. The target value for metal transfer is represented by point C. The superheated potroom metal has a enthalpy value B further up on a curve. By weighing with the different metal amounts it can be calculated that the potroom metal together with the furnace metal would reach the target value C, if the furnace metal were at point D. Thus the difference of point A and D on the x-axis gives the remaining amount of energy, which still has to be generated by the burners. Knowing the actual burner power the remaining up-time of the burner can be calculated easily. At the moment production data of flue, roof and metal temperature are analyzed to investigate the feasibility of this kind of control scheme in the real installation. Optimized furnace operation scheme 4000 3500

- mei pool - ingot stack

3000 2500

S^

2000 1500 1000

*^t±

-optimal point for addition of potroom melt - actual enthalpy level inside furnafce

energy still requird

500

Figure 9. Average gas residence times of the configurations A - F.

0

Optimization of charging schemes and burner operation As mentioned in the introduction a part of the metal is added as superheated liquid potroom metal. An interesting question is when the liquid metal should be added to utilize both the better heat transfer conditions of the initial ingot stack and the superheat from the potroom metal. In this scope it is important to measure the amount of energy, which is really transferred to the metal. The average temperature and the average fraction liquid as shown in Figure 1 are not helpful. It is more convenient to look at the average enthalpy density of metal inside the furnace (see Figure 10). In this diagram the former time axis is now converted into an energy axis by multiplication with the burner power. The enthalpy density covers sensible heat and heat of fusion. It is interesting to see that the average enthalpy density curves of the melt pool and of the ingot stack do not differ very much.

Figure 11. Control scheme to optimize liquid metal addition and burner operation. Conclusions A numerical furnace model was successfully applied to analyze the heat transfer conditions in a reverberatory melting furnace. • It turned out that the heat transfer is by far dominated by radiation effects. • Due to the approximately spherical shape of the furnace, modifications of the burner positions showed almost no improvements concerning gas residence times and energy utilization. The potential to achieve energy reductions by this kind of modifications seems to be limited. • A more promising approach to improve the heat transfer inside the furnace could be the application of oxy-fuel burners. Of course, the much higher efficiency has to be weighed against the additional costs for the oxygen supply. • Considering different metal distributions inside the furnace, it can be noticed that the heat transfer into an ingot stack is higher than into a melt pool. • It is recommended to heat the ingots as long as possible and add liquid metal not before the optimal amount of energy was brought into the cold metal. • A timing scheme for the burners based on the heat transfer characteristics inside the furnace could be used to control and minimize the amount of energy, which is transferred into the metal.

Energy utilisation characteristic of furnace _

4000 T-

c

3500 -j-

jfr c =· 2500 0) «? 70 e >. = 2000 Jr — 1500

1 i

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1000 I

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500 1

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Ä | — m e It pool [W^ M—ingots stackfll

I 1

ol0

10

20

30

40

50

60

70

80

90

100

energy generated by burners [GJ]

Figure 10. Average enthalpy density of metal inside the furnace depending on consumed energy. One attractive option to optimize the fuel consumption could be to assume a unified energy transfer curve to control the heat content,

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Acknowledgments The authors would like to thank casthouse manager Pβl T. Endresen and the crew of Hydro Aluminium Karmoy Rolling Mill for enduring stimulation and support in this project.

Literature [1] S. Instone, A. Buchholz, G.-U. Grün: TMS Light Metals 2008, 811-816 [2] G.-U. Grün, A. Buchholz, TMS Light Metals 2009, 735-742 [3] V.R. Voller, C.R. Swaminathan, Numerical Heat Transfer B, 19(2), 1991, 175-189 [4] J. Furu, A. Buchholz, T.H. Bergstrom, K. Marthinsen, TMS Light Metals 2010, 679-684 [5] ANSYS FLUENT 12.0 User's Guide, ANSYS, Inc. 2009, chap. 15.1.3, p 15.24-25

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Oxyfuel Optimization using CFD Modeling Thomas Niehofο4, Sreenivas Viyyuri1 ^he Linde Group, Linde Gas, Carl-von-Linde-Strasse 25, 85716 Unterschleissheim, Germany Keywords: Oxyfuel, CFD, furnace optimization, modeling, burner, emissions, heat transfer,

Abstract

GEOMETRIES OF FURNACE AND BURNER

Before converting production furnaces to different combustion technologies it is essential to understand all related changes and side effects. An experienced team will be able to successfully conclude a conversion like this. However, CFD modeling will enable to make informed decisions in terms of effort and results of furnace retrofitting with new combustion equipment. This paper will give insight of how oxyfuel together with CFD can impact energy balance and productivity of production furnaces.

FLUE GAS

TEMPERATURES, etc..

BURNER FUEL

V

OXIDIZER BATH CONDITIONS

FIRING RATE FIRING CONDITIONS

Introduction SCRAP TYPES AND GEOMETRIES

Aluminum recycling and re-melting is a very competitive industry area. Global markets and globalization of aluminum melting technologies and aluminum trade brushes up the dust in every corner of the business. Aluminum producers with large melting furnaces are constantly under pressure to bring production cost down and hence to use the latest available technology. Any change is associated with risks. Furnaces with larger production capacities face higher risk as compared to small ones. Being able to estimate process changes before capital money is spent allows to even optimize a technology towards temperature and heat profiles as well as productivity and energy usage before the system is build.

PHYSICAL PROPERTIES

Figure 1: Basic elements of aluminum melting process. Model Set Up The melting or heating operation to be modeled typically consists out of a specific furnace geometry (rotary furnace, reverberaotory furnace, tower furnace, shaft furnace and many more). This furnace geometry together with refractory material, flue openings, burner positions and geometries, bath level and other protruding elements define the combustion space. Geometries, thicknesses, and physical properties of the materials then are put together to the model. The mesh size of the model describes how detailed the combustion process will be described in a specific location of the model. The mesh size can vary across the furnace. Typically it needs to be very small (detailed) where rapid changes in either geometry or chemistry is expected. The mesh is refined near the region of the burner to capture the gradients effectively. Homogenizing with slow dynamics areas will have wider mesh sizes, when there is an expectation of reduced activity. A typical aluminum reverberatory furnace model with a capacity of 301 and a footprint of 450 sqft (50 m2) will have 500.000 knots with an average distance of lA foot (0.15 m). Where each knot represents that all equilibrium equations (heat, energy and mass) will be solved. The model works its way through the furnace system by moving from knot to knot. It can take days or weeks to simulate one single steady state operating point in a way that the results make sense and are good enough for verification. Recent advancements have enabled to solve the governing equations in parallel using multiple CPU's which reduces the computing time significantly. Typical combustion systems involve high speed flows. Hence choice of proper turbulence models is very important to predict the solution accurately. In most of the cases, reaction is mixing controlled hence turbulence chemistry interaction needs to be

Basic Elements A melting operation consists of various specific elements that need to be transferred into the computer model. However not all can be transferred and not all can be modeled. Hence the model will only be a model and not real. Aspects like charge material quality, quantity, composition and mixtures are very difficult to describe in a model. Charging material storage and altering process conditions as well as the melting itself with all changing physical properties are hard to specify and model. The combustion space of a furnace can be modeled with justifiable effort. The system of the combustion space then needs to be defined and described well enough to come to reasonable results that can reflect the reality. The model will preferably describe steady state conditions, i.e. altering firing rates and flame shapes cannot be characterized in one single step. The model needs to be connected to reality and will need to be verified with the current operation.

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resolved correctly to accurately predict the flame shape and temperature distribution. Also as high temperatures prevail inside the furnace, radiation heat transfer plays significant role in transferring energy from the combustion space to the metal bath. Hence choice of the radiation model will also play a vital role in the accuracy of the solution.

......

M B

^'

1 -=-

\ Figure 3: Conventional and round oxyfuel flame and combustion space temperature profiles.

Figure 4: Flat Jet oxyfuel burner flame and combustion space temperature profiles.

Figure 2: Example of mesh size and geometry. Combustion System The combustion system is a very complex system of geometries, kinetics and chemical reactions. When modeling the modeler has to decide which condition to model and why. Very often typical and characteristic operating conditions are being modeled. Like high temperature intervals when looking at specific heat transfer and temperature profiles. Flame shapes and sizes for refractory investigations. NOx and CO profiles for gaseous emissions from combustion processes.

Figure 5: Flat Jet oxyfuel burner flame and combustion space temperature profiles. By comparing such different flame specific characteristics in one single furnace geometry there will be differences in heat transferred into metal bath area and towards the walls. When changing burners and/or burner locations and comparing the effects to the metal bath and the walls will lead to and optimization process withoutfiningthe new combustion system.

CFD in Aluminum Melting Traditional cost intensive CFD modeling has been used for major furnace projects, where the modeling cost was only a fraction of the capital investment. The simulation was used to make informed decisions about heat transfer, temperature profiles, emissions, burner and flue locations and efficiencies/cost. Large capacity steel making plants, power plants and large scale glass plants have been using CFD modeling for a long time and have good experience in getting useful results for decision making processes. In aluminum and metals melting and recycling there are only few examples how and where CFD is being used. The Linde Group has put effort in understanding and detailing combustion processes with CFD for melting of metals like aluminum and copper. Examples

Figure 6: Example of temperature profile above metal bath level.

By simulating combustion processes the effects of different flame shapes can be analyzed and compared. Conventional oxyfuel is often regarded as pipe in pipe and round flame geometries (Fig. 3). The flat flame of an oxyfuel burner is shown in Fig. 4.

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Air fuel operated combustion processes for melting aluminum has benefits and disadvantages that are listed below: Benfits:

Disadvantage:

Preheated Air Fuel Burners

115

- Low flame temperature - No oxygen cost - Good convective heat transfer - Slow reactivity - High gas volumes - High energy demand - Low efficiency - Noise emissions - Dust emissions

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Linde has developed a combustion technology that combines the advantages of air fuel combustion by avoiding the disadvantages at the same time. This technology was developed from the need of the aluminum industry to avoid local overheating and hence reduce oxidation of the metal. The Linde response is low temperature oxyfuel combustion technology. The low temperature oxyfuel combustion process combines the benefits of air fuel and oxyfuel combustion process. This means the low flame temperature and high convective heat combined with high energy efficiency from oxyfuel. CFD modeling is used to describe and evaluate the changes that oxyfuel would bring in such a situation. The experience from converting to oxyfuel at a casthouse from Hydro Aluminium in Norway has been described in III. There an oxyfiiel burner has been used to compensate hot pot room metal with ambient temperature solid scrap. This is an example of how oxyfuel lifts existing limits and borders to the next level.

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Figure 7: CFD comparison of temperature contours of pre heated air fuel burners and oxyfuel burners in a reverberatory furnace. In other publications 111, 131 and 141 the energy and emission impacts of oxyfiiel in comparison to air fuel are described. CFD modeling does reflect these changed conditions and helps to get a better and deeper understanding.

These days the analysis of combustion processes especially in the aluminum area focus on highly efficient heat transfer without local over heating. Comparing various oxyfuel cases leads to the following conclusions: oxyfuel and oxyfuel can be different flame shape and volume are important flame temperatures matter furnace gas recirculation has an impact on heat transfer

Summary CFD Modeling has shown to have many benefits and it avoids the often applied "trial and error" approach. It can also be used to maximize oxyfuel benefits and minimize emissions. The Linde Group is pioneering the way into CFD modeling for non ferrous metals melting in combination with oxyfuel combustion technologies.

When evaluating all these different parameters - oxyfuel can be optimized versus oxyfuel. The next exciting question is: How does optimized oxyfuel compare to regenerative pre heated air fuel operation? Here linde has done extensive research to better understand the specifics and details that are important to keep all benefits from air fuel firing and all benefits from oxyfuel firing.

References III

Figure 7 shows a comparison of a regenerative air fuel fired furnace with a low temperature oxyfuel (LTOF)firedcase.

Ill 131 141

1187

H. Gripenberg, et al.: Optimised re-melting by the use of oxyfuel at Hydro Aluminium's primary aluminium cast house, Φvre, Ardai, Norway; TMS 2010, Seattle, WA. T. Niehoff "Oxyfuel - Solutions for Energy and Environmental Conservation", TMS 2008, Feb. 2008, New Orleans, LA. T. Niehoff "Oxyfiiel - Energy efficient melting", TMS 2009, Feb. 2009, San Francisco, CA. T. Niehoff, "Optimised Aluminium Melting", 10. OEA International Aluminium Recycling Congress, Berlin, Mar. 2009.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Operational Efίciency Improvements Resulting From Monitoring and Trim of Industrial Combustion Systems Jim Oakes, Damian Bratcher TMS (The Minerals, Metals & Materials Society); 184 Thorn Hill Rd.; Warrendale, PA 15086, USA Super Systems Inc., 7205 Edington Drive, Cincinnati, OH 45249

!

Combustion is the exothermic chemical reaction (a reaction in which heat is given off) of hydrogen and carbon atoms contained in fuels with oxygen. Excess 0 2 makes heating inefficient, thus requiring more gas for the same results. In addition, excess air also allows for the formation of pollutants such as Nitrous Oxide (NO) and Nitrogen Dioxide (N02).

Optimizing operational efficiency, minimizing production costs and maximizing utilization, is a competitive advantage in good economic conditions. In leaner times, it is a basic necessity. Periodic checking and resetting of air-fuel ratios is one of the simplest ways to get maximum efficiency out of fuel-fired process heating equipment. In heat treatment facilities, the customer would find potential efficiency improvements on generators, radiant tubes, furnaces, ovens, heaters, and boilers.

It is estimated that precise control of air to fuel ratio will yield 5 to 25% or more savings in heat generation. The air gas ratio can be determined by analyzing the flue gas and the mixture for combustion can be altered to produce the most clean and efficient heat for the process.

The two main areas where heat treatment facilities benefit from combustion optimization are fuel savings and throughput improvements. Combustion optimization will be reviewed first. Next, the impact these improvements have on throughput and utilization will be explored.

Periodic checking and resetting of air-fuel ratios is one of the simplest ways to get maximum efficiency out of fuel-fired process heating equipment. In heat treatment facilities, the customer would find potential efficiency improvements on generators, radiant tubes, furnaces, ovens, heaters, and boilers.

Combustion Efficiency

Abstract

Introduction According to the Department of Energy, most high temperature direct-fired furnaces, radiant tubes and boilers operate with about 10 to 20% excess combustion air at high fire to prevent the formation of dangerous CO and "soot" deposits. It is estimated that precise control of air to fuel ratio will yield 5 to 25% or more savings in heat generation. The air gas ratio can be determined by analyzing the flue gas and, with this information, the mixture for combustion can be altered to produce the most clean and efficient heat for the process. Figure 1 displays estimated volume of by product gases based on % of oxygen used when mixing with CH4. Our studies have shown that burners are typically running with excess 0 2 greater than 4% in the flue gas.

Most high temperature direct-fired furnaces, radiant tubes, and boilers are designed to operate with 10 to 20 percent excess combustion air at high fire. This excess air helps prevent the formation of carbon monoxide and soot deposits which can affect heat transfer surfaces and radiant tubes. For the fuels most commonly used in the US, including natural gas, propane, and fuel oils, approximately one cubic foot of air is required to release 100 British Thermal Units (BTUs) in complete combustion. Process heating efficiency is reduced considerably if the air supply is significantly higher or lower than the theoretically required air. In a September 1997 Process Heating magazine, Mr. Richard Bennett provided calculations for an available heat chart which was republished in May 2002 by the Department of Energy. This chart is an excellent basis to determine potential savings in a combustion process. To determine the potential savings, you will need the following information: Exhaust gas temperature as it exits the furnace, tube, etc. % excess air or oxygen in the flue gas (actual) % excess air or oxygen in the flue gas (target)

m M2% by Volume »CQ% by Volume

The available heat chart is shown in Figure 2.

m C02% by Volume *H20% by Volume m 02% by Volum«

Figure 1 - Estimated products of combustion versus excess 0 2 .

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$27,750 per year. At full utilization, the savings reach $98,550 for this furnace. A side benefit to the fuel savings is a documented C0 2 reduction. For each MCF CH4 burned completely, 117 pounds of C0 2 is produced. In this particular case, the customer was able to document a reduction of 175,500 pounds or 87.75 tons of C0 2 . At full utilization on this one furnace with a 1% reduction in excess 0 2 , the reduction would be 630,006 pounds or 315 tons. If the customer has similar success on other fiirnaces and is able to achieve the 0 2 target, his potential C0 2 reduction is 8000+ tons. Batch Furnace Utilization and Fuel Savings The initial R&D on batch furnaces was initiated with John Keough at his Applied Process' Wisconsin and Kentucky facilities. Mr. Keough recognized the value in operating burners at an optimum level to save on fuel. He also recognized the even greater value in creating operational efficiencies by increasing load throughput based on increasing the available heat produced during high fire with the optimal gas/air ratio. Mr. Keough and Applied Process challenged Super Systems with deriving a system that would monitor high fire air/gas ratios and provide operators with alarms and trending to monitor the burner performance. The two test sites demonstrated and proved out the sensor and provided the initial data regarding combustion efficiency and utilization improvement. These results led us to further testing at Queen City Steel Treating.

Figure 2- Available Heat Chart Using the chart, determine the percent available heat under actual and target conditions. The intersection of the measured exhaust gas temperature and % excess air (%02) curves provides these values. The potential fuel savings would be calculated as follows: % Fuel Saving = 100 x (( %AH Target - %AH Actual) / %AH Target) Documented Savings To illustrate the value of combustion optimization, two case studies will be presented.

Queen City Steel Treating in Cincinnati, Ohio worked with Super Systems to document savings relative to varying 0 2 levels in combustion exhaust gases. The tests were conducted on a batch furnace with 4 radiant tubes using the same load density with identical initial conditions. Each burner is rated 250,000 BTU/hr.

Forge Heat Furnace A 6 mmbtu/hr open die forge reheat car bottom furnace was equipped with a high-temperature SuperOX oxygen sensor. Baseline readings of excess 0 2 and fuel consumption were collected over a three month period. Based upon this data, monthly fuel consumption was determined as was the average high-temperature 0 2 readings (6.5% average at 2200° F).

Four tests were conducted with excess 0 2 levels ranging from 2 to 5%. The test results are shown in Table 1. Table 1 - Furnace trial test results

The controls and operation's personnel were concerned about over-trimming the excess 0 2 level. Lowering 0 2 levels can lead to reduced uniformity on the heated ingot. Thus, the 0 2 levels were lowered incrementally to ensure that no impact occurred to product quality.

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At the end of the first incremental change and after process verification, the customer has lowered his excess 0 2 from 6.5% to 5.5%. After numerous runs at this 0 2 level, the customer documented a 20.5% reduction in metered gas consumption. Using the data shown in figure 1, the % available heat at 5.5% 0 2 and 2200F is -25%. Similarly, 6.5% and 2200F, it is -20%. The potential savings is -20% [100 x (25-20)/25]. The actual results are very similar to the expected results.

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For this customer, the 20.5% reduction in fuel cost corresponds to a $15,000 per year savings for this single furnace based upon its current utilization rate. At full utilization, the savings would be $53,874.

The two significant highlights evidenced by the data are the significant improvement in ramp rate (8.82 vs. 6.75 7min) and the reduction in the amount of high-fire time. The improved heating rate shortens the time required for the load to reach heat and shortens cycle time by 15 minutes per load. This come-to-heat time was calculated based upon the 2% and 5% excess 0 2 rate and was consistently more than 15 minutes.

The customer has a goal of reducing the excess % 0 2 several percent. At his target level, he would reduce his fuel costs by an estimated 37%. At the current utilization, the savings would be

1190

The reduction in high-fire time reduces fuel and operating costs along with minimizing CO2 emissions. Table 2 provides a summary of C0 2 and fuel savings for the reduction in high-fire time. The burner's total demand on high fire is 1,000,000 BTU or one (1) dekatherm. The calculations are based upon a dekatherm cost of $5 and a 90% uptime availability. Excess Ö» Jewel Soak Cost per hour CQfclfesperhour Soak cost per day CQj, lbs per day Soak cost per year CO* lbs per year

m

m

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97.60

$3.48 84.73

$3.20 77.96

$2.81 68.44

$96.00 2342.4 $35,536.00

$83.40 2035,0 $23396.90

$75.68 ÎS710 $25,189.38

$67.32 1642.6 $22,114.62

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Table 2 - C0 2 production and fuel cost The cost to maintain temperature is reduced by 30% as are the CO2 emissions. Over the course of one year, the savings will exceed $9,000 and 200,000 pounds of C0 2 by reducing the excess 0 2 from 5% to 2% in the combustion process. Table 3 provides a summary of the improved utilization that is achieved by reducing the excess 0 2 in the radiant tubes for various cycle times. The calculations are based upon a 15-minute savings in come-to-heat time. Cycle times will impact the improvement in utilization and the number of additional loads that can be pushed through the furnace on an annual basis. Cyde time {in hours| 15 minute savings, % of cyde utilization improvement Optimal: Loads per year Max increase loads per year

3

4

5

8

91.67% 109.03% 2920

93,75% 30667% 2190

95.00» 305.26% 1752

96.88% 10323» 1095

265

146

52

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Table 3 - Utilization improvement As the cycle times decrease, the utilization improvements become more significant. For a typical one hour come-to-heat and 3 hour soak (4 hour total cycle), The improvement is 6.67% and 146 additional loads per year. Summary Continuous monitoring and adjustment of excess 0 2 levels in combustion applications provides significant fuel savings, reduced emissions and improved utilization. The savings and improvements will vary from facility to facility and from furnace to furnace depending upon how the combustion system is currently tuned and maintained. As process temperatures increase, the fuel and emissions savings rise exponentially. Several state governments currently offer grants and credits that help further reduce the cost of 0 2 monitoring and reduce the payback time. Even without these grants and credits paybacks are achieved from fuel savings in short periods of time along with gains in utilization. By optimizing combustion efficiency, companies will minimize production costs and maximize utilization and have a competitive advantage over those who overlook this part of their process.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Technology for Electromagnetic Stirring of Aluminum Reverberatory Furnaces Alan Peel CEng, ALTEK LLC, Castle Donnington, Derbyshire, UK James Herbert, CEng ALTEK LLC, Exton, Pensylvania, USA

EM Pumps and Stirrers - Market Growth Abstract

The benefits of circulation in aluminum reverberatory furnaces are well documented and include higher productivity, and reduced fuel consumption and dross generation. One popular method to achieve the benefits of circulation is electromagnetic stirring. That said, this technology has certain drawbacks preventing its universal acceptance: high capital cost, high operating cost (especially power consumption and maintenance of water cooled copper tubes), reluctance to use water in close proximity to molten metal, and in some cases the inability to operate through full thickness refractory hearths. This paper describes a technology which amounts to the reinvention of traditional electromagnetic stirring devices, effectively addressing all of the above negative aspects of traditional systems. In this paper we will describe how these issues are addressed and results documented in recent installations.

Introduction It has long been known of the significant advantages of circulating furnaces to improve melt rate, reduce energy consumption, minimize dross formation and obtain excellent chemical and temperature homogeneity. [1,2,3,4,5,9,10,11]. The application of Electromagnetic stirring (EM Stirring) which was initially introduced into the Aluminium industry in the I960's has grown significantly since the late 1990's. The graph below in Fig 1 shows the huge growth in application of different types of Electromagnetic stirring or pumping technologies to circulate aluminium furnaces reported some years ago [3] but which has continued to grow significantly as the market recognizes the huge benefits to be gained from stirring furnaces.

pre 1996

1999

2001

2004 Years

2007

2009

Fig 1 Showing the growth in application of electromagnetic stirring devices since late 1990's. There are many different types of device available to circulate a furnace today as is shown in the above graph, and all with have more or less the same effect on the operational benefits within the furnace. Over the past 3 years ALTEK have been working on enhancing the MHD air cooled stirrer technology, which they supply under exclusive licence, to meet these specific operational objectives but also focusing on the importance of the capital cost, operating costs, ease of installation and also reliability of the stirring equipment. The first MHD stirrers were developed by the MHD Centre in Krasynarsk, Russia [4] in the early 1990's for application throughout the RUSAL (and former SU AL groups). These companies had the desire to circulate their furnaces to achieve the well understood benefits of metal circulation in their melting and holding furnaces. The original MHD inductor stirrer design was

1193

developed from 1st principles in Krasynarsk as, having reviewed similar technology on the international market they believed it had some inherent weaknesses. After installing 40 of these stirring units on furnaces throughout Russia, a collaboration was formed with ALTEK to further develop the technology, introduce new design and drive control techniques and introduce the new electromagnetic stirring technology, called SIBER FORCE to the whole world. This was initiated in 2007 with a careful step by step introduction into the wider market. One of the main driving forces for aluminium operations to consider the air cooled electromagnetic stirring is safety by removing the water cooling requirement of conventional stirring devices from the basement under a furnace or next to a furnace. EM Stirrer Design The design of the ALTEK SIBER FORCE electromagnetic stirrer features several fundamentally different aspects and we will discuss each of these in turn below. Cooling Medium Due to the design of the inductor coils being made from solid copper bar (not hollow tubing), of certain dimensions, this has the effect of reducing the heating effect of the electrical current applied to the coil. The heating effect from a current passing through an inductor is represented by PR, where R = r/s (r = resistvity of copper and s = the surface area faced by the passing current). A larger surface area therefore presents a lower R in this calculation and a reduced heat generation.

Frequency Converter and Main Control Panel

Fig 2 shows the control architecture of a bottom mounted stirrer installation. There has long been a myth that you need to use huge amounts of electrical energy with this type of technology to obtain the mixing effect within the furnace. This may be the case with alternative designs but with the SIBER FORCE technology this has been overcome significantly reducing the input power requirement The key focus of this stirring system is to deliver the Lorenz Force within the aluminium bath [3] in the proximity of the stirrer, and due to the viscosity of the liquid aluminium, you get a very effective mass flow of aluminium throughout the whole furnace. Some of the smaller models of SIBER FORCE stirrers are being utilized on furnace capacities of >60 tonnes. The size of stirring model chosen is based upon the following factors:1. 2.

This allows for the use of air in a specially applied way within the inductor to be used to remove the lower heat generated by the current applied to the inductor coil. As is discussed later in this paper, when looking at the operating results, Company A monitored the electrical power consumption at between 3 and 5 kWh per tonne. A secondary benefit of this type of coil construction also provides for a very long life of the inductor and reduces any risk of serious damage due to overheating of the coil as would occur with a hollow copper tube type inductor coil. As an example, the oldest inductor installed in 1994 is still in operation today). Control and Inductor Drive System The system comprises of 3 key components:1. 2. 3.

Inductor Control System Cooling Fan

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3. 4.

Size of furnace Type of operation (Melter, holder, liquid, dry hearth) Refractory thickness - to obtain correct penetration of the magnetic flux into the bath. Available space

By using a specially designed and patented technique for driving the inductor coil, a technology utilized in the induction heating industry, allows for a relatively low input current to the control system from the clients facility (the input kVA range is between 80 and 120 depending upon the stirrer model size). Installation The SIBER FORCE stirrer can be installed on many different furnace types and is fitted either on the side of the furnace or underneath the furnace hearth in a small basement [7]. Due to the fact the stirrer is air cooled this removes the necessity for costly water cooling systems and associated pipe work. The air cooling is provided by an air cooling fan which can be located nearby to the inductor and connected by air ducting. The stirrer is separated from the bath through the furnace's refractory thickness and the outer casing in the area of the stirrer is removed and replaced (as normal carbon steel will not allow

1194

transmission of magneticflux),with a stainless steel plate which allows the magnetic flux to penetrate into the aluminium bath. Behind the stainless steel plate the existing refractory at its normal thickness remains or if necessary new refractory is installed in that area. This maintains full integrity of the furnace. There are different stirrer sizes for the different non magnetic gap thicknesses (the thickness of the refractory, insulation, stainless steel plate and air gap) from 350mm (quite typical in a furnace wall) to up to 700mm that may exist on direct dome charge type furnaces. The control system is fed by a 3 phase 50Hz (or 60Hz) supply sized based on the stirrer model chosen. On a side mounted installation the inductor is fitted to the furnace by sitting on rails allowing it to be easily moved in and out and then is clamped to the furnace leaving a small 10mm air gap during normal operation. Fig 3 below shows the installation of a TYPE 400 on the side of a 70 tonne stationary furnace at Company A where the stirrer was rented - hence the temporary location of the cooling fan.

Fig 4 Showing the stirrer and associated scissor lift/trolley mechanism applied to a dome type furnace. This also allows for the introduction of a trolley system that allows the stirrer to be moved in and out of the basement area or even to operate between multiple furnaces as is shown in Fig 5 below.

Fig 5 showing the SIBER FORCE installation for multiple furnaces. Modelling Fig 3 Showing the stirrer installed on the side of a 70 tonne stationary reverb.

Before application and if required it is possible to model the furnace to show effectiveness of the stirring effect on homogeneity. The model in Fig 6 below shows cfd modeling of the impact of a side mounted device fitted to a stationary reverb.

On the bottom mounted installation the stirrer can simply reside on a concrete plinth under the furnace allowing the stirrer to locate into the non magnetic stainless steel window or if the basement is deeper, the use of a scissor lift will allow the stirrer to be raised to its operating position or lowered when not in use. Fig 4 below shows the concept on a direct charge dome furnace now being supplied to a customer.

Fig 6 showing the location and preparation of the cfd grid ready for modeling the stirrer on a side mounted application.

1195

the stirrer running (highly significant effect, risk <1%). The difference (with stirrer - without stirrer) is +2.1 + 1.4 t/h. Mo

Yes

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Fig 7 showing the cfd modeling on a large (>100T) tilting type melting furnace to indicate potential temperature homogeneity benefits.

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m% m m m m m Fig 8: Showing time chart of melting rate, and Statistical average melting rate

Operational Benefits A TYPE 400 side mounted SIBER FORCE stirrer was installed on the side of a 70 tonne stationary reverbatory furnace at Company A in June 2009. A production monitoring study was performed in May 2010 by Company A. To assess the effect of stirring a protocol was put in place designed to reduce the wide variability in the melting process, whether it be intrinsic to production (charge composition,) or to the installation (combustion settings, unidentified degraded operating modes) or of human origin (working practice).

Energy Saving The impact on the quantity of energy needed for melting only was assessed, and the results shown below in Fig 9 i.e. from the end of charging up to the transfer temperature (730°C). Analysis of the data indicates a significant reduction (risk <1%) in energy consumption from 906 to 777 kWh/t (-14%). The difference (with stirrer - without stirrer) is -130 + 90 kWh/t. The specific consumption of the stirrer was measured at between 3 and 5 kWh/t.

An alternating sequence of heats with and without stirring were carried out over 5 days at the rate of 3 to 4 heats per day. A 3xxx alloy series was used throughout the casting run. Production personnel (charging and melting operatives) were asked to observe particular operating procedures which we considered key to the variability of the melting process: - Charge loading: in a single operation with a minimum crown temperature of 1050°C - Bath breaking: breaking at a minimum crown temperature of 1100°C - End of casting regulation: 730°C The furnace operators were also asked to take particular care to open the furnace doors only when strictly necessary, as opening the doors is associated with major energy loss. A summary of the results of the study are shown below. Melt Rate The melting rate represents the tonnage melted on average by the furnace in one hour. It is calculated by dividing the tonnes charged by the melting time between the end of loading of the charge and the moment when the liquid bath reaches a temperature of 730°C. The opening times of the furnace doors are determined. Fig 8 below shows the gross results in the form of a time chart and thefigureon the right is a statistical representation after correction for the effect of the initial temperature of the furnace. On average, the melting rate increased by from 10.7 to 12.8 t/h (+20%) with

urn m m 2Jk 3/ë m Fig 9: Time chart of specific energy consumption and Average specific energy consumption corrected for the effect of crown temperature. Dross Generation The weight of the slag after each skimming of the bath was measured. Over the week of the tests, and taking into account the composition of the charges (small amount of compact materials in particular), it was found little slag variations. Fig 10 below compares the weighed quantities of slag with and without stirrer. It is inconclusive in terms of identifying a stirrer effect. In order to reach a definitive conclusion it would be necessary to compare

1196

complete runs with full melts (70t) over a longer period with drainage of the furnace, and to determine a full material balance.

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In terms of energy consumption in the melting phase, an improvement of 130 (± 90) kWh per tonne was observed (the energy consumption of the stirrer in the order of 3 kWh/t to be deducted from this figure).

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References

It is reported [6] that a large impact on dross generation is the ratio of solid to liquid charge. The more solid charge (less liquid charge added) the higher the dross generation which is as you would expect as most of the dross is generated in the early part of the cycle on a dry hearth furnace operation as a consequence of the exposed scrap to the burners.

1. 2. 3.

Other similar studies with electromagnetic stirrer have demonstrated there can be a reduction of dross formation with a stirrer application [6]. Summary Due to the simplicity of design the capital costs for installation of the equipment can be lower than a water-cooled stirring technology whilst also removing the water from the proximity of the melting furnace. With the solid design of the stirrer and specially designed control and drive system, low operating costs through low energy consumption can be achieved.

Bamji P. J., Circulation of Molten Metal, December 3 1975 Thibault M. A., Tremblay F, Pomerleau J.C., Molten Metal Stirring: The Alcan Jet Stirrer,, Light Metals 1991 Peel, A. (2003) A look at the History and Some Recent Developments in the use of Electromagnetic Devices for Improving Operational Efficiency in the Aluminium Casthouse. 8th Australasian Aluminium Casthouse Technology Conference.

4.

R.M.Khristinich, V.N.Timofeyev, V.V.Stafievskaya, A.V.Velenteyenko, "Molten Metal Elecgtromagnetic Stirring in Metallurgy", International Scientific Colloquim, Modelling for Electromagnetic Processing, Hannover, March 24-26, 2003

5.

Bamji, P.J., Pierson, F.W., "Electromagnetic Circulation of Molten Aluminium: Production tells the story", Light Metals 1983, pp 953-961 A.F.Saavedra, "Electromagnetic Stirrers - Their Influence on Melter Operation and Dross Generation", Light Metals 1993.

As the study at Company A's facility has shown they were able to show that electromagnetic stirring has a significant measurable effect on the melting performance of one of their melting furnaces.

6.

The performance was quantified under test conditions (controlled pre-heating of the furnace, good operating practice).

7.

Eidem. M, Taallback. G, Hanley. P. J., Side Mounted EMS for Aluminium Scrap Melters - A Comparison Between Side and Bottom Mounted EMS, Light Metals 1996 pp861 - 867

8.

Peel, A Alchalabi, R. and Meng, F. (2002) Furnace Operation Optimisation with EMP System, Light Metals 2002. N.Hayashi et al., "Applications of Electromagnetic Stirrer to Aluminium Melting Furnace" Sumitomo Light Metal Met. Tech. Rep., (2), April 1985,91-97.

In terms of intrinsic melting rate, an improvement of 2.1 (±1.4) T/h was observed.

Average without stirrer

Specific melting rate (kWh/t) 10.7

9.

1197

10. Sjoden. O, Lehman A., Developments in Electromagnetic Stirring for Aluminium Furnaces; 71 International Recycling Congress, Munich, 2003 11. Petho. S., Experience of Alcoa Kofem with MHD Induction Stirrer, Light Metals 1996 pp 857 - 860

1198

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

Evaluation of Effects of Stirring in a Melting Furnace for Aluminum Kunio Matsuzaki1,Toru Shimizu1, Yoichi Murakoshi1 and Kenzo Takahashi2 Japanese National Institute of Advanced Industrial Science and Technology AIST Tsukuba East, 1-2-1 Namiki, Tsukuba, Ibaraki 305, JAPAN 2 Zmag Japan, Ltd. 1143-3 Takayanagi, Kashiwa, Chiba, JAPAN

Keywords: Permanent Magnet, Molten Aluminum Stirring Abstract A new type of permanent magnet stirrer was installeted in an existing furnace and the effect of magnetic stirring on molten alloys was examined. The end goal is to improve the quality of aluminum alloys through bath temperature homogenization, and to reduce energy usage and dross generation while providing a reduction in melt loss. Traditionally, molten aluminum has been mechanically or electromagnetically stirred; however the former is inefficient and dangerous while the latter requires excessive energy and complicated components. Using a permanent magnet stirrer, the temperature homogenization of the molten Al alloy was achieved safely and efficiently. This report also discusses thoughts on more effective shapes of furnaces, compared to ordinary furnaces.

The following tests were conducted: •

•

•

Purpose

•

This research report will examine the effects of stirring molten aluminum alloy by use of a permanent magnet based stirring technology. Using a permanent magnet stirrer solves problems associated with stirring mechanically. Some of these problems include maintenance, downtime and safety issues due to replacement and repair in a high temperature environment. There is also an electromagnetic method for stirring molten aluminum; however there are heavy energy consumption requirements as well as a supporting infrastructure and systems necessary to operate the stirrer. These issues have cost considerations as well as maintenance and other related complications. Permanent magnets, not requiring energy to generate a magnetic field, water for cooling nor other components not directly related to stirring, avoids these problems. Since a permanent magnet stirrer can be configured in various ways, there are implications on future furnaces that can be designed in order to fully maximize the benefits of this technology. This topic will also be discussed.

Temperature was measured at 21 points in the molten aluminum bath for evaluation of temperature homogeneity. The 21 points consisted of 7 locations, at 3 depths for each location. The depths were 25mm, 250mm (center), and 475mm from the bottom of the furnace. The temperature was measured and compared at each point both with and without rotating magnetic fields. In addition, the amount of gas consumption to keep the temperature constant was also measured. Samples were taken from various parts of the furnace and their chemical compositioins were then analyzed using ICP (Inductively Coupled Plasma), both with and without rotating magnetic fields. Content of hydrogen on the samples was measured by Gravi-Mass, an outgas measurement device, at NIKKIN FLUX INC. Mechanical properties of the cast samples (consisting of aluminum stirred by magnetic fields) were compared to the original ingot (prepared by conventional, non-stirring, means).

Method Permanent magnets were set beside and beneath an existing furnace. The technology used for the purposes of this test was a Tsunami Series III™ permanent magnet stirrer from Zmag Japan, Ltd. The shape of the furnace was 1,500 by 1,500 mm, and the height of the molten aluminum was 500 mm. Aluminum alloy (AD4C) was melted by a burner attached to the furnace roof; the molten alloy was simultaneously stirred by rotating those magnets. The furnace was regulated to keep the temperature at a certain level, 740°C, in the furnace. When the temperature was less than 740 °C, the burner fired, and when the temperature exceeded 740X3, the burner stopped.

Figure 1

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Result

Furnace Temperature (Stirring)

1) Result of Stirring

(1) Temperature of molten aluminum alloy Figure 1 indicates changes in temperature at each measuring point in the furnace where the temperature was kept at 740 °C for 12 hours without stirring by magnetic fields. When the temperature dropped under 740°C, the burner started, followed by a rise in temperature at each measuring point. A rise in temperature was highest at the upper layer of molten aluminum alloy, and the temperature lowered as the measuring point descended to the furnace bottom. When the temperature rose to 740°C, the burner stopped, followed by a fall in temperature in the molten aluminum. When the temperature dropped under 740°C, the burner started again. This cycle was repeated to keep 740 °C in the furnace. Figure 2 is an enlarged part of this cycle, and indicates changes in temperature on three points: 25mm (from the bottom), 250mm (center), and 475mm (top).

0

3

7

10

13

17

20

Figure 3 While rotating the magnetic fields, noise was observed on the thermometer due to the influence of magnetic fields. In this case, one cycle took 17 minutes, including 5 minutes to heat the furnace. This compares to a 120 minute cycle in the previous test without stirring. The highest temperature was 760 °C around the upper layer of molten aluminum. The temperature was kept about 740°C around the center and bottom layers so no difference was shown between the center and bottom layers. The difference between the upper and bottom layers was at most 20°C, so it could be stated that the temperature in the furnace was very homogeneous compared with the case of no magnetic fields. At other measuring points, the results were similar. There were no noticeable differences according to the direction of the magnetic fields. Furthermore, there was no remarkable difference when the rotating speed was changed to 87rpm. The experiment above shows that homogeneity of temperature is caused by stirring molten aluminum through the rotation of the magnetic fields. (2) Comparison of the amount of gas usage Table 1 shows how much gas the gas burner used for an hour to hold the temperature in molten aluminum. When the magnetic fields didn't rotate, the amount of gas usage was 3.30 m3. However, when the magnetic fields rotated at the speed of 60rpm, the amount of gas usage was reduced to 2.91 rrf, therefore leading to an expectation of energy savings. Energy savings is realized because the temperature is homogeneous due to the stirring of molten aluminum via rotating magnetic fields.

Figure 2 Once the burner started to work, the temperature rapidly rose in the upper layer of molten aluminum, followed by a rise in temperature in the center and bottom layers. When the temperature reached 740°C, the burner stopped and the temperature in the upper layer rapidly started to fall, followed by a gradual fall in temperature in the center and bottom layers. As time passed further, the temperature in the upper layer fell below the temperature of the bottom layer. At that point, the burner started again. One cycle took about 120 minutes, including 30 minutes to heat the furnace. When the temperature reached the maximum measured at 822°C, the upper layer in the center was measured at 764°C and the bottom layer, 740°C. The difference was about 80°C between the upper and bottom layers of molten aluminum. With these results as a baseline, next a permanent magnetic circuit was set 250mm apart from the wall of the furnace, and rotated at a speed of 60rpm to stir molten aluminum. Temperature was measured in the same points. The result is shown in Figure 3.

Rotation Speed

Gas Usage per Hour (m»/h) 0 rpm 3.3 30rpm 2.98 60 rpm 2.91 87.5 rpm 2.93 Table 1 Amount of gas usage per hour to keep temperature in furnace at 740°C (3) Speed of the molten aluminum flow To evaluate the flow speed of molten aluminum stirred by magnetic fields, the speed of a float placed on the surface of the molten aluminum was measured. When molten aluminum was stirred at a rate of 30rpm, 60rpm, and 87rpm,

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the speed of the float was respectively 0.05m/s, 0.2m/s, and 0.2m/s. However, a film of aluminum oxide was on the surface, interfering with the float, so molten aluminum flow may have been higher than indicated. In addition, the float moved fast near the stirrer, but slower further away from the stirrer. This is due to either magnetic fields that influenced the float or due to surface oxide that slowed the float. Whatever the case, a more accurate method of flow measurement needs to be considered, for example, measuring flow in an oxide free environment. (4) Hydrogen content

Sample

Original Ingot 0.616

0 rpm 0.69 1

30 rpm 1.12 3

Hydrogen content (cc/100g) Table 2 Hydrogen content in each test sample

60 rpm 0.93 8

(6) Mechanical properties The Vickers hardness of the cast sample which was stirred by magnetic fields was HV60. The Vickers hardness of the original ingot was HV61. Results are virtually the same. The tensile strength and the elongation to fracture of an original ingot at room temperature were respectively 82MPa and 1.4%. This is in comparison to 95MPa and 2.4% for a cast sample which was stirred by magnetic fields. Therefore, it can be said that the strength and the elongation were improved because of stirring by magnetic fields. A possible reason for this improvement is that the elements in the sample became more homogenous.

87 rpm 1.15 5

Sample Ingot Casting

(%) 1.4 2.4

Table 4 Mechanical properties of sample

^Optical microstructure of cast sample

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(5) Analysis of Chemical Composition Table 3 shows a composition of Si and Mg measured by ICP Optical Emission Spectrometry in samples which were picked from points in the molten aluminum alloy. In the case of stirring a molten aluminum alloy, the difference in composition of Si between the upper and bottom layers of the molten aluminum alloy was less than the difference in the case of not stirring. However, because the ingot used for melting was an alloy, the difference was fairly small. It can be expected that stirring would be most effective in the following cases: Where an alloy element is being added to the molten aluminum. Where the molten aluminum contains a large amount of elements in the alloy. Where the specific gravity of the elements is high. Si (wt%) 4.48 5.22 6.21 6.63 5.02 4.48 6.03 5.90

Elongation to fracture

(7) Microstructure Figure 4 shows the optical microstructure of a cast sample stirred by magnetic fields. It is seen that the sample consists of a matrix of aluminum and preciptates, and the microstructure is not so different from an original ingot.

Table 2 shows the amount of hydrogen in each test sample, measured by Gravi-Mass. As a basis of comparison, the amount of hydrogen in an original ingot before melting was measured. When magnetic fields weren't rotating, the amount of hydrogen was 0.69cc and as the rotation of magnetic fields increased, the amount of hydrogen tended to increase slightly. This increase could be due to the film of aluminum oxide breaking by the stirring action. However, the amount of hydrogen shown in Table 2 is quite low since, for example, AC2B and ADC 12 alloys already include 1.20cc and 2.00cc respectively, according to Gravi-Mass. To reduce the amount of hydrogen, stirring should be done in a way that does not break the surface film of aluminum oxide. In addition, simultaneously degassing while stirring could be expected to lead to good results.

Sample Orpm Upper Bottom 30rpm Upper Bottom 60rpm Upper Bottom 87rpm Upper Bottom

Tensile Strength (MPa) 82 95

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In summary, it was found that the temperature of the molten aluminum alloy was more homogeneous and the gas consumption was less with stirring by magnetic fields than in the case of ordinary stirring. These effects are clearly the result of stirring by magnetic fields. Furthermore, stirring by magnetic fields controls a rise in temperature of the surface of molten aluminum alloys. This effect could save energy and prevent the surface of molten aluminum from oxidizing.

Mg(wt%) 0.25 0.26 0.25 0.27 0.25 0.29 0.31 0.26

2) Possibility of new furnace shapes

The contemporary melting furnace for aluminum is an open hearth furnace which melts aluminum by heating through the burner, placed above or on the side. The depth of molten aluminum is shallow in order to keep homogeneity of temperature between the top and bottom; in this research the depth was 50cm. Thus, when the amount of molten aluminum is increased, the furnace must be enlarged horizontally with a corresponding increase in the number of burners. In this research, it was found that stirring by magnetic fields resulted in an improvement in homogeneity

Table 3 Composition of molten aluminum alloy sample

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of temperature. Given these results, the possibility of more efficient furnace shapes might well be considered. Here, the burner's heat contribution to melting aluminum is, for the purposes of this section, not considered. It is assumed there is sufficient available heat capacity. The blue line in Figure 5 shows a difference of temperature from the upper to the bottom in the furnace without stirring by magnetic fields. The temperature is 820 °C around the surface, 760°C around the center, and 740°C around the bottom. On the other hand, the red line shows the temperature difference by stirring through magnetic fields. Due to the effects of stirring, the temperature around the upper layer of molten aluminum is held at 760^0 and at the bottom it is measured at 740 °C . The difference in temperature is minimized. Figure 6 Figure 6 schematically shows the amount of molten aluminum alloy in the existing furnace and new furnaces with a magnetic stirrer. "L" shows depth and width, and the "H" shows height. Molten aluminum was 50cm high in the existing furnace without stirring by magnetic fields. The volume of the molten aluminum ("V") was as follows V = 50 * L * L In the case of stirring by magnetic fields, the molten aluminum can be increased by 25cm while maintaining the same volume, as shown in the following equation, with "LI" indicating depth and width.

Figure 5 When the furnace is not stirred by magnetic fields, the temperature is 760°C around the center layer (25cm from the surface). Therefore, it can be inferred that if stirring by magnetic fields is adopted for an existing furnace, the depth of molten aluminum can be increased by 25cm. Building furnaces specifically designed for permanent magnet stirrers may be a way to maximize the benefits of stirring. A new shape of furnace, roughly illustrated, follows with approximate figures for the purpose of discussion.

(50 + 25) * LI * LI = 50 * L * L L1 = >T2/3*L = 0.81*L Therefore the size of the furnace can be reduced by 20%. Furthermore, if only the width is shortened in a vertical direction from the magnet ("L2"), the range of magnetic force could be maximized for efficiency. In this case, the width ("L2") can be reduced by two-thirds compared to an ordinary furnace. It follows then that a tall type furnace could be a more efficient configuration compared to a flat type furnace. An important comment is that these results were estimated as though the upper layer was not stirred by magnetic fields; however in reality there was a stirring effect in the upper layer so the width could be reduced further. Adjusting placement of the stirrer and adjusting the size and strength of the stirrer could allow a furnace to be taller still. The reduction in depth and width has some advantages including space savings for more flexible furnace placement. Also, the space reduction provides temperature homogeneity in a horizontal direction when the gas burner is on. This saves energy. Furthermore, when the reduced surface area is heated by gas burner, the amount of molten aluminum which contacts air and oxygen is reduced and the formation of oxides are also reduced. This reduction of oxides leads to lower melt loss. It can be expected that inclusions and hydrogen levels will drop, so mechanical properties of a cast are improved. Stirring by magnetic fields causes homogeneity of temperature, which controls excessive heat near the surface of molten aluminum. Consequently, the materials used in furnaces can be

replaced by refractory for this lower temperature range; this change contributes to lower cost. The results above show that by using a permanent magnetic stirrer, tall type furnaces could be used, the resulting space savings might also provide benefits in energy reduction through improvement in heat efficiency, improvement of production through controlling oxidization, and improvement in the metal quality through reduction of inclusions and hydrogen in molten aluminum. In addition, excessive heat can be minimized allowing for a change in ingredients used in the refractory. This contributes to lower the cost of furnaces. Summary It was found that stirring by magnetic fields led to excellent temperature homogeneity while simultaneously reducing the amount of gas consumption. Furthermore, surface heat could be controlled by stirring; reduced surface heat minimized oxidized dross while saving energy. Maintaining temperature homogeneity by stirring provides an opportunity to use tall type furnaces. Using a tall furnace, the area of molten aluminum surface is reduced, further reducing oxides and further saving energy. In addition, this change can improve production by reduction of melt loss, as well as an increase in material quality through the reduction of inclusions and hydrogen in the molten aluminum.

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

BUSINESS ANALYSIS OF TOTAL REFRACTORY COSTS Cynthia K. Belt1 ^ayetteville, AR 72701, USA Keywords: Energy, Refractory, Furnace Abstract Refractory is a critical component of a furnace. Both planned and unplanned downtime due to refractory wear or failure can detrimentally affect production. Refractory materials, design, and maintenance can improve or degrade energy efficiency and melt rate. In addition, the cost of refractory is typically the largest maintenance expense of a melt or hold furnace. Total refractory costs include refractory materials, installation, energy, and downtime. Many attempts have been made by companies to hold down these costs in various ways. This paper compares several methods and gives relative cost savings based on a benchmark furnace.

furnace. The second is that production and profit were lost. The profit was estimated to be $0.05 per pound aluminum. Of course, the costs and life of a particular furnace can vary greatly. The results of this modeling will not cover all furnaces or situations. However, the general concepts can be used to understand business results of refractory changes. Total Refractory Costs Refractory material costs can be very high. As high as these costs are, other related costs can be equal or higher. Many of these costs are hidden and not necessarily charged to repairs. Figures 14 show the relative percentage of these costs for the baseline furnaces. These costs are based on the total life of the furnace. For instance, materials would include the materials used during the total reline of the furnace and smaller repairs done over time.

Introduction Over the life of the furnace, refractory costs are typically the largest maintenance expense of a melt or hold furnace. The condition of refractory and the downtime due to repairs greatly affects furnace utilization and energy efficiency. Different theories as to the best method to reduce costs abound. Data was collected to model and compare these methods using a standard furnace over the lifetime of the furnace.

1)

Materials - This includes bricks, castable, gunned, and shotcrete material. It includes materials for direct material contact, insulating materials and anchors. Some of the material is lost during application (Shrinkage). For bricks and blocks, minimal shrinkage or material loss is expected. This type of material has a long shelf life with only a small loss due to bricks being cut or breakage. Castable material will have a higher loss. Material is lost within the mixer and pump system or will be thrown away due to a limited refractory life. Bags can be damaged or can get wet. Finally, gunning typically has the same losses as castable but with extra losses from rebound (material that does not adhere during application). 2) Labor - This is the cost of labor and equipment used to repair or reline refractory within the furnace. Material and labor costs are essentially equal for a melt furnace. Because hold furnaces are smaller but the labor is almost the same, labor costs are a higher percentage than materials costs for these furnaces. Labor costs include: a) Tear-out of old refractory materials - For major repairs and relines, equipment must be brought in to break up and haul out the old refractory. b) Installation - Installation costs include labor and materials needed to complete the work. Along with labor, costs for both wood and steel needed for the forms can be high. c) Dry-out - It is best to use external burners that are controlled with multiple thermocouples instead of using the furnace's burners. This process is critical to a good installation! 3) Furnace steelwork repairs - Frequently steelwork must be repaired during a reline. The damage may be due to excessive heat over the life of the furnace or to damage during the tear-out. 4) Downtime - Downtime can be long during a reline (over 30 days). Even smaller repairs require time to cool down then dry-out the refractory. The importance of downtime greatly

The various methods that were compared included: reduction of refractory material costs, reduction of installation costs, reduction of installation time, energy efficient insulation, repair versus total reline, and extension of the life of the refractory. Total Refractory Cost Model For the purposes of this paper, the first standard furnace is an 180,000-pound [82 metric tons] reverberatory melt furnace that melts 5,000,000 pounds of aluminum per month [2,268 metric tons]. The standard life of the furnace used in this model was 4 years. The second standard furnace is an 85,000-pound [39 metric ton] reverberatory hold furnace with a 15-year life and throughput of 7,000,000 pounds of aluminum per month [3,175 metric ton]. It is expected that operations will damage jambs, lintels, and the sill surrounding the door causing holes to form and reducing energy efficiency. Other damage and/or wear will occur over the lifetime of the furnace that will be major, but not enough to completely reline the furnace. External labor for repairs and reline is assumed, however internal labor is used to plan and oversee the work. The cost of downtime can widely vary between furnace, plants, companies, and depending on current sales. In some cases, production can be smoothly moved to other furnaces while for others, unplanned downtime may cause the loss of an important customer. For the purpose of this paper, two outcomes were modeled. First was that production could be moved to another furnace but that extra energy would be required to start that

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varies. When other furnaces can take the production, heating up a spare furnace is a minor cost. When production is lost, it can become equal to the value of the material or the labor. In the case of a hold furnace, downtime becomes even more important. Many hold furnaces are connected to multiple melt furnaces and are directly tied to a casting machine. Downtime here is critical! 5) Material Storage - Space is needed to store new refractory material on-site. Castable and gunning materials must be kept dry. Depending on the season, it may need to be heated or cooled to keep refractory at the right conditions. 6) Disposal - The costs of old material disposal include hopper rental, transportation, and landfill fees. These costs have risen over the years. 7) Internal labor - Internal time is needed to plan, quote, and manage a job of this size.

Shrinkage

Tear-Out Dry-Out

Material

Installation

Materials

8)

Material storage, disposal, and internal labor are a relatively small part of total refractory costs. Energy - While most energy in a furnace is used for production, some is directly related to refractory. Within this paper, this term does not include the energy used to melt and hold metal, flue loss, conveyor loss, or most opening losses. a) Dry-out - This is the energy used during dry-out for burners and blowers. b) End-of-Life - This is the heat loss towards the end of life due to openings occurring in the roof, wear in the walls, or heat lost around the doors. c) Wall Heat Loss - These are the heat losses through the walls, floor, and roof during the life of the furnace. Wall heat loss is an especially high component of Energy over the life of the furnace. This is particularly true for hold furnaces given their long life.

Total Refractory Life Cost - Hold Furnace

Low Downtime Costs High Downtime Costs

Labor

Mat'l Storage

Figure 1. Total Refractory Costs - Melt Furnace Shrinkage

Tear-Out Dry-Out

Material

Installation

Materials

Labor

Total Refractory Life Cost - Melt Furnace

Low Downtime Costs High Downtime Costs

Steelwork

Mat'l Storage

Figure 2. Total Refractory Costs - Hold Furnace

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Energy

installation problems, or due to dry-out issues? When the job is turnkey, this problem is solved since the contractor is liable. On the other hand, if contracts are carefully written and the work is clearly understood, this issue is less likely to occur. This extra liability is difficult to assess and is not included in the model.

Methods to Reduce Total Refractory Costs The various methods that were compared are the following. 1. Reduction of refractory material costs The first method a company typically uses to reduce costs is to negotiate costs of the refractory materials. This works to a limited extent. While there is certainly profit within the sales of these materials, it is not extremely high. Refractories are, in general, a commodity so that a refractory company that demands a higher cost without improved properties will quickly lose business.

4. Energy efficient insulation Better insulation can reduce heat loss. Extra insulation or different types of insulation (such as microporous) can increase the cost of insulation and insulation. However, over the life of the furnace, improvements in insulation are typically easily cost justified.

More importantly, care must be taken so that low cost and low quality materials do not affect overall performance. Indeed, higher material costs may be preferred if it can reduce installation downtime or extend the life of the furnace with materials with greater strength, wear, or release properties.

5. Reduction of installation time Several methods are available to reduce installation time. Material solutions may include big block, pumpable, shotcrete, or quick dry-out materials. Installation methods include longer shifts, weekend, and holiday work. Both the material and the installation method increase costs. For the purposes of the model, it is assumed that installation costs increase by 5 % to shorten downtime by 10 %.

As long as critical properties (Hot Modulus of Rupture, Cold Crush Strength, etc.) are understood and maintained, good negotiations can bring down the costs 5 % and perhaps 10 %. However, since lower costs may reduce the quality of the project, it is assumed within this model that life of the furnace stays the same for a 5 % material reduction and is reduced 2 % for a 10 % reduction. While that reduction is a mere one-month for a 4-year life, it has high implications for the total refractory costs.

While shorter dry-out schedules are available, it is extremely important to follow material supplier specifications. A 'steam event' (explosion) is far more expensive than the extra time for a proper dryout! Even without a catastrophic event such as an explosion, shorter dry-out's in some cases may reduce the properties and life of the material.

2. Reduction of installation costs Similar to material costs, installation costs can be negotiated. However, the low cost installer may not be the best!

6. Repair versus total reline There are two different viewpoints on refractory repairs. To one viewpoint, repairs should be scheduled approximately every year to repair the portions of the refractory showing wear or damage. Over the years, all parts of the furnace will be replaced with some portions replaced morefrequently(such as jambs and lintels).

Again, since lower costs may reduce the quality of the project, it is assumed within this model that life of the furnace stays the same for a 5 % installation cost reduction and is reduced 2 % for a 10 % reduction. 3. Direct contracts versus turnkey Installations can be turnkey or separately contracted. This includes materials, tear-out, and dry-out. We know that the installation company profits by marking up sub-contacts. However, there are trade-offs to be understood.

The second point of view is to perform minor repairs and then totally reline the furnace when the majority of the furnace requires repairs. The first method means that good sections of the furnaces are not replaced. It also means that refractory is periodically damaged due to the cool-downs and heat-ups of this process. Energy is used for extra dry-outs. Multiple tear-outs and dry-outs occur over time so that downtime will be higher overall.

The biggest savings is by purchasing refractory directly. However, additional internal time is needed to estimate the proper quantities and to track usage. Frequently additional material is needed or extra material is left at the end of a job. A contractor is able to stockpile materials to be used on another job but extra material or lack of material will add to costs when purchased directly. In the model it is assumed that normal shrinkage is 5 % while directly purchasing material will increase this to 7 %.

The model assumes that the same amount of material is used but that that extra downtime, dry-out energy, and project management time will be needed. Since repairs are down more often, the energy loss toward the end of the furnace life will be reduced. However, since the refractory is susceptible to cracking from thermal shock, the life of the furnace is reduced by 10 % (4.8 months).

The biggest issue is the question of who is liable if there are problems with the furnace refractory. Did a refractory failure occur due to the material, due to

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7. Extension of the life of the refractory Much work has been done to extend the life of these types of furnaces. A long-term refractory program should include failure analysis to understand the exact needs in certain areas along with in-furnace trials of refractory. Refractory costs may increase by selecting higher strength materials, but costs can be kept down by using these materials in the most critical areas.

There are fewer savings by directly contracting the tear-out and dry-out, but it will save some money. Again, liability for problems must be decided before the job starts. Energy efficient insulation While payback is over a year, installing better insulation will save money over the life of the furnace. This is especially true for hold furnace given their long life. Short-term savings with less insulation will lose the company money with increased heat loss. Invest in insulation!

In this model, it is assumed that overall refractory costs increase 4 % for every 10 % increase in refractory life. As an example, for a furnace using $250,000 materials, an increase of $10,000 would increase the life of a furnace 4.8 months. A second model was run assuming 10 % increase in material costs to receive the extra 10 % life.

Reduction of installation time If downtime causes a loss in production, extra costs can be easily justified. However, there is no purpose in reducing installation time for furnaces where production can be moved to another furnace. Extra costs to reduce installation time will end up costing more for the project. Sometimes management pushes for reduced installation time based on previous experience when plant production or profit per pound was higher. Current conditions must be evaluated.

Modeling Results Figures 3-6 show the potential savings from the specified changes. Please note that the scaling varies to show the data better. Reduction of refractory material and installation costs Hard negotiation is always a good idea on large projects and equipment that is critical to production. Saving 5 % for a $500,000 project on materials and installation is well worth the time.

Repair versus total reline Based on this analysis, it does not make sense to take a furnace cold multiple times to repair sections of a furnace. Each cycle shortens the furnace life by thermal shocking the refractory. While multiple smaller repairs will give short-term savings, it costs more over the life of the furnace. Instead, a scheduled total reline is best for total refractory costs.

Key to this is a bid package that details the specific materials and methods to be used. In this way, everyone is bidding on the same thing. The quality of the installation and materials will be maintained. Poor practices include asking for quotes to "Reline the furnace" without details and bidding using only the contactor you used the last time!

Note that thermal shocking can also occur when cold cleaning a furnace. Cleaning is best done hot. When possible, minor repairs should be done hot or 'warm'.

A warranty is needed of at least one year. The bid package should detail who will be supplying each item (ex. refractory disposal, water, compressed air, etc.). If possible, important installation details should be specified. For instance, this may include keyed sections or whether pumpable materials are acceptable. Of course, a refractory print is important!

The analysis was done assuming a 10 % shorter life (4.8 months). It was also rerun (results are not shown) using an assumption that refractory life remains the same. While the total lifetime cost was better, it still was more expensive to repair sections. This was due to extra costs from multiple tear-outs, dry-outs, and internal labor in managing the jobs.

However, if cutting costs reduces the life of the furnace, there is little benefit. Cheap materials or cutting corners during installation saves the company no money over time. Based on this modeling, it is better to receive only a 5 % discount then a 10 % reduction that reduces the life of the furnace by 2 %, particularly when downtime costs more.

The analysis was not completed for hold furnaces. Most companies will do the total reline and not do patches. This may be due to the long life of the furnace, the difficulty in scheduling downtime on hold furnaces, or that without heavy wear from charging most sections of the hold furnace wear at an equal rate.

Direct contracts versus turnkey Directly purchasing the materials will save money as long as care is taken to calculate the right amount of materials. However, it is important to keep in mind that liability needs to be discussed and documented in purchase orders. It is best to have someone from the refractory company present during as many pours as possible. They can ensure that proper practices are being followed.

Extension of the life of the refractory The best method to save total refractory costs is by extending the life of the furnace. And, the more that downtime is worth to the company, the more important longer furnace life becomes. Two different assumptions were made. In one, each 4 % increase in material costs resulted in 10 % increase in life. Understanding the critical areas and refractory properties can make this possible. For example,

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scenario shows much higher savings then just reducing the material costs by proper bidding. For the same assumptions as above material costs would increase by roughly $125,000 resulting in a furnace that lasts 6 years instead of 4 year. This type of results is very possible.

stainless steel needles can be added to the jamb and lintel section or silicon carbide can be used in ramps and ledges. Adding radii while forming may help. Upgrades in areas such as the roof, upper sidewalls, or sub-hearth rarely are needed. To get a feel for this, on a $250,000 project (material only) we are assuming material costs would increase by roughly $50,000 resulting in a furnace that lasts 6 years instead of 4 year. This takes work, but can be done.

Another point is that many companies will save money by reducing the quality of refractory on hold furnace since wear is less. In fact, using high quality material is very important since downtime tends to be more critical in these applications.

In the second assumption, each 10 % increase in material costs increases the life by 10 %. Even this

I Low Downtime Costs I High Downtime Costs

Melt Furnaces

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h Lί Reduce Material

Il Reduce nstallation

Reduce Material

Direct Contracts - Total

Direct Contracts - Material Only

Direct Contracts - Tear-out Only

Direct Contracts - Dry-out Only

Improved Insulation

Figure 3. Total Refractory Savings - Melt Furnace Hold Furnaces

I Low Downtime Costs I High Downtime Costs

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Direct Contracts Direct Contracts Direct Contracts Direct Contracts -Total : M a î ! ? n l É . 0 „ n l y . - Tear-out Only - Dry-out Onty

Figure 4. Total Refractory Savings - Hold Furnace

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Improved Insulation

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.

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Ji

Extend Life Extend Life Extend Life Extend Life Extend Life Extend Life Extend Life Extend Life Extend Life Extend Life -io%w/ -?n%w/ -3n%w/ -40% -50% -m% -?n% -an% -40% - so% 4% 8% 12% w/16% W20% w/10% w/20% w/30% w/40% w/50%

Figure 5. Total Refractory Savings - Melt Furnace Hold Furnaces

Γ

15.0%

ú Low Downtime Costs I High Downtime Costs

111

σ> | 10.0% (0

1

5.0% 0.0%

IsH r DT -

-5.0%

Reduction

%>

Extend Life - Extend Lif8- E (tend Life - Extend Life - Extend Life - Extend Life - Extend Life - Extend Life - Extend Life - Extend Life 1CPL w / 4 % 9 Ð % w / ft ?n%w/?n% ^n%w/.^n% 4n%w/dn% Rn%w/fin% PA W/...1 O°L dn%w/iR% fin%w/pn% m%w/in%

-10.0%

Figure 6. Total Refractory Savings - Hold Furnace Conclusions Costs related to refractory in a furnace are high. While material and labor costs for a specific repair may seem high, it is important to keep in mind all costs associated with the furnace and the total life of the furnace. Based on the results from this study, the most attention should be spent in extending the life of refractory versus reducing the cost of refractory. Higher cost repairs may be the best business decision to improve energy efficiency, reduce downtime, and reduce overall costs for the life of a furnace.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

IMPROVED FURNACE EFFICIENCY THROUGH THE USE OF REFRACTORY MATERIALS James G. Hemrick1, Angela Rodrigues-Schroer2, Dominick Colavito2, and Jeffrey D. Smith3 ^ a k Ridge National Laboratory, 1 Bethel Valley Road, Bldg. 4515, MS-6069, Oak Ridge, TN 37831-6069, USA 2 MINTEQ International, Inc., 640 North 13th Street, Easton, PA 18042, USA 3 Missouri University of Science and Technology, 1870 Miner Circle, Rolla, MO 65409-0330, USA Keywords: Refractory, Furnace, Corrosion, Wear associated processes through such impediments as necessitated process shut downs, impurities created in the metal or glass being processed, need for the implementation of filtering or purification of the furnace bath, the need for additional melting, and greater heat losses through the furnace walls, floor and roof as the lining deteriorates or forms less insulating phases.

Abstract This paper describes efforts performed at Oak Ridge National Laboratory (ORNL), in collaboration with industrial refractory manufacturers, refractory users, and academic institutions, to improve energy efficiency of U.S. industry through increased furnace efficiency brought about by the employment of novel refractory systems and techniques. Work in furnace applications related to aluminum, gasification, and lime are discussed. The energy savings strategies discussed are achieved through reduction of chemical reactions, elimination of mechanical degradation caused by the service environment, reduction of temperature limitations of materials, and elimination of costly installation and repair needs. Key results of several case studies resulting from a US Department of Energy (DOE) funded research program are discussed with emphasis on applicability of these results to high temperature furnace applications.

Therefore, there is a need to develop innovative refractory compositions to address these issues. The work described in this paper intended to develop improved refractory compositions based on novel compositions, new aggregate materials, alternative bond systems, protective coatings, novel phase formation techniques (in-situ phase formation, altered conversion temperatures, accelerated reactions, etc.), and alternative application techniques, (castables, gunnables, shotcretes, etc). The developed materials were tailored for use in specific industrial environments such as those found in the aluminum, cement, chemical, and forest products industries.

Introduction

The intent of the work described in this paper was to improve energy efficiency in the targeted applications and industries through three approaches. The first was through the identification of materials capable of operating at higher temperatures (goal of increasing operating temperature by 100-200°C over current operating temperatures depending on the process). The second was through identification of materials capable of operating for longer periods of time (goal of twice the life span of current materials or next process determined service increment). The third was through alternative refractory application techniques that could lead to less expensive or faster installation of refractory linings. Such materials and techniques could lead to less process down time, greater energy efficiency through more heat kept in the process, and materials that could be installed/repaired in a more efficient manner. The overall goal of the project was a 5% improvement in energy efficiency (brought about through a 20% improvement in thermal efficiency) resulting in a savings of 3.7 TBtu/yr (7.2 billion ft3 natural gas) by the year 2030 as predicted through an analysis performed using DOE Government Performance and Reporting Act (GPRA) analysis software.

Refractory materials are called upon to function throughout industrial applications as insulation and/or containment vessel linings in high-temperature and corrosive environments. Therefore, they must possess properties suitable for exposure to extreme environments for extended periods of time. Indeed, it would be difficult to identify an industrial process that does not use refractory materials in one aspect or another. In a furnace application, these materials must not only be capable of performing these tasks at elevated temperatures, but may also be called upon to bear mechanical loads and transfer heat. As such, refractories are a vital class of materials for the sustaining of the world industrial economy related to the manufacturing of the products that we rely on in our modern world. Many factors can affect the applicability and performance of refractory materials in a furnace environment. These can include chemical reactions (corrosion) between the service environment and the refractory material which may lead to depletion of the material or formation of other compounds on the refractory surface, mechanical degradation (wear and erosion) of the refractory material by the service environment, penetration of molten material into cracks or pores present, limitations on temperature at which a material can be safely used, and the ability or inability to install and/or repair a particular refractory material in a cost effective manner or while the vessel is in service (i.e. at temperature).

Material Development Five process vessels from the targeted industries were selected for improvement through implementation of newly developed refractory material systems. These were rotary and reverberatory aluminum furnaces (aluminum industry), black liquor gasifiers (forest products), coal gasifiers (chemical industry), and lime kilns (forest products with possible extension to cement). Additionally, refractory systems were developed for two specific purposes. Initial efforts focused on materials designed for new (original) lining installations. These compositions were formulated with the intention of being direct replacements for currently available castable and brick formulations traditionally used in these furnace

All of the above mechanisms of attack will lead to drastic reductions in the energy efficiency of the furnace and can ultimately lead to complete failure of the refractory lining as described in further detail elsewhere [1]. The decrease in energy efficiency is further compounded by associated environmental impacts and related reductions in the economic viability of the

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applications. Subsequent efforts focused on refractories which were designed and specifically tailored for on-line (hot) maintenance or repair installations where the material could serve as a "patch" to extend the service life of a furnace between relines. A separate part of the project also focused on pursuing alternative application techniques and systems for the optimized installation of the newly developed refractory materials in an effort to maximize the properties of installed linings and to facilitate hot installation and repair. Several alumino-silicate, magnesia and spinel forming castable refractory systems were developed as shown in Table I. These materials were designed to either possess favorable phases (alumino-silicates or magnesia) when installed or to form favorable phases (spinel) during heating and drying as the furnace or process vessel was brought up to operating temperature. Additionally, the specific spinel composition chosen (aluminarich verses magnesia-rich) was selected based on the operating environment being acidic such as in aluminium furnaces and coal gasifiers (alumina-rich required) or basic such as in black liquor gasifiers and lime kilns (magnesia-rich needed) [2]. Table I. Materials Developed for Specific Industrial Furnace Applications Industry Application Material Aluminum AS (A) Primary lining AS(B) Repair material SF (A)* Repair material Black Liquor SF (B)+ Primary lining Repair material PBC Coal Gasification Repair Material SF (O* Insulation LWC Back-up lining SF(B)+ Primary lining Lime Kiln

Aluminum Through physical, chemical, and mechanical characterization of salvaged refractory materials from the aluminum industry a number of sources of refractory failure were previously identified [1, 3]. In summary these are poorly performing anti-wetting additives, degradation of the aggregates used in refractory castables, reaction of micro-silica in the refractory matrix with cement binder systems, and poor furnace maintenance practices leading to mechanical damage of refractory linings. The effectiveness of anti-wetting additives and the choice of aggregate and dispersant systems were investigated in this work. Three candidate materials were developed for aluminum applications (as shown in Table I). Two of these were aluminosilicate materials based on previous formulations produced by MinTeq, with one developed as a primary lining material and one developed with the intent of serving as a high-temperature repair material. Both materials are 70% alumina content pumpable formulas designed for shotcreting. The anti-wetting additive systems in these materials were redesigned to provide superior resistance to corrosion and the refractory matrix was modified to improve hot modulus of rupture (HMOR) at higher temperatures and to provide enhanced high temperature corrosion resistance. An alumina-rich spinel material which showed improved corrosion and erosion resistance was also developed for repair applications, but was not further pursued due to the substantially higher cost of this material compared to the developed aluminosilicate material which performed equally as well. Examples of cup testing results from the alumino-silicate materials developed for aluminum contact applications are shown in Figure 1. Both cast and shotcrete versions of the material were evaluated after being pre-fired at 871 and 1260°C (1600 and 2300°F) through testing at 815°C (1500°F) for 72 hours in aluminum alloy 7075 aluminum (magnesium-rich aluminum alloy).

AS = alumino-silicate material SF = spinel forming system PBC = phosphate bonded magnesia castable LWC = light weight castable A, B, and C in parenthesis designate sequential compositional designations * alumina-rich spinel + magnesia-rich spinel Although spinel-based materials were initially sought for all of these applications, it was found that in several cases they were not the best choice. Therefore alumino-silicate or magnesia-based systems were selected instead. Additionally, a light-weight backup refractory system was developed for use with the spinel forming systems to help offset the high thermal conductivity inherent in the spinel materials, as compared to traditional alumina-based or alumino-silicate materials. Shotcreting of monolithic refractory materials was selected as an alternative application method to traditional brick or castable linings that are often used for these applications. Thus, materials developed under this project were designed and tested for application by such a method. This method of applying refractory materials involves first mixing the refractory components with water and then adding an accelerator to the air supply of the nozzle system. Advantages of the method include rapid application, elimination of joints, and elimination of geometric constraints.

Figure 1. Refractory cup testing results for alumino-silicate materials develop for aluminum applications. Testing performed in 7075 aluminum alloy for (a) shotcrete sample pre-fired at 871°C, (b) shotcrete sample pre-fired at 1260°C, (c) cast sample pre-fired at 871°C, and (d) cast sample pre-fired at 1260°C. Such testing is qualitative and usually results in either a pass/fail or a ranking (excellent, satisfactory, poor) evaluation. The

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samples shown above were all rated as possessing "excellent" corrosion resistance due to the fact that the original cored geometry was maintained (sharp corners still present and no metal penetration was found). It was also determined that firing temperature did not seem to affect the corrosion behavior, nor was there a difference seen between cast and shotcrete samples (often a problem when quoting properties of a shotcrete material based on cast sample testing).

was veneered with the developed maintenance material. The temperature of the furnace was then raised to 927°C (1700°F) and held for four hours. After the hold, the crucible was cooled and sectioned to evaluate the adhesion of the repair material on the crucible wall. Results of this test are shown in Figure 4.

To further evaluate the performance of these materials, before moving to a full industrial trial, a rotary furnace simulation was conducted at the MinTeq research facilities in Easton, PA. Pictures of the constructed unit are shown in Figure 2. The furnace was lined with shotcrete panels which were exposed to molten aluminum metal heated by a gas burner. Pictures of installed panels of the alumino-silicate formulation before and after testing are shown in Figure 3. Based on the successful performance of these materials in this test, full industrial trials are currently being initiated.

Figure 4. Evaluation of alumino-silicate repair material through high-temperature adhesion test. A crucible containing solidified aluminum/dross/flux ( a) was heated and vennered with repair material (b) before being cooled and sectioned (c). Evaulation of the coating after cooling showed good adhesion of repair material to the degraded crucible wall. Black Liquor Gasification Black liquor gasification provides the pulp and paper industry with a technology which could potentially replace recovery boilers with equipment that could reduce emissions and, if used in a combined cycle system, increase the power production of the mill allowing it to be a net exporter of electrical power. In addition, rather than burning the syngas produced in a gasifier, this syngas could be used to produce higher value chemicals or fuels. However, problems with structural materials, and particularly the refractory lining of the reactor vessel, have caused unplanned shutdowns and resulted in component replacement much sooner than originally planned.

Figure 2. Rotary furnace simulation test system built at MinTeq in Easton, PA shown with gas burner assembly.

Figure 3. Installed refractory panels exposed to molten aluminum in rotary furnace simulation test (a) before and (b) after testing showing the relatively unchanged refractory surfaces after testing.

Previous work at ORNL resulted in the identification of fusioncast spinel brick materials which extended the lifetime of the hot face refractory lining of these vessels from months to years [4, 5]. Although these lining materials were highly successful, a lower cost option is still sought. It was thought that a non-brick material may offer such an alternative if the corrosion and wear resistance of the brick material could be maintained. With this in mind, two magnesia-containing materials were developed as shown in Table I. The first material was a magnesium-rich spinel forming shotcrete material designed for primary linings. The second material was a phosphate bonded magnesia castable designed as a hot repair material.

To further assess the performance of the repair material, a test was performed to evaluate the adhesion of this material on a candidate refractory wall consisting of degraded refractory material (alumino-silicate castable) and solidified aluminum. A spent crucible containing aluminum metal, dross, and flux was heated to 538°C (1000°F) in a laboratory furnace. Upon heating, the hot surface of the crucible bathed in the aluminum/dross/flux mixture

Both formulations were tested through laboratory immersion testing as described elsewhere [4, 5]. Testing was conducted for 100 hours at 1000°C (1832°F) using an in-house constructed laboratory immersion test system. This system has been demonstrated to successfully reproduce the corrosion products observed to form on refractories exposed in operating gasifiers using commercially generated smelt (nominal composition: 60-

1213

75% Na2C03, 20-38% Na2S04, 1-4% Na2S, and 1-4% Na2S203) [6]. Results of previous testing for the currently used fusion-cast spinel material were compared to the corresponding results obtained for the spinel shotcrete and magnesia castable materials as shown in Figure 5.

may be present. Therefore, there is a need for longer life refractory materials for this application that could improve gasifier reliability, availability, and affordability. Current state-of-the-art refractories for this application are highchrome containing alumina based refractories (60-95% Cr203). Typical lifetimes of these materials in this environment are four to eighteen months with replacement costs often being on the order of $1M and down time during replacement lasting as long as two to three weeks [7]. Improved materials are of interest which are longer lasting, less expensive, and non-chrome containing due to environmental and disposal issues associated with Chrome VI compounds. One approach to producing improved materials has been through the use of surface coatings [8]. These are low cost coatings using a colloidal approach for protection against corrosion attack of the refractory brick. This has been shown to be both valid in the laboratory and commercially successful through industrial trials. The approach taken in this project was to try and apply the spinel shotcrete technology developed for the black liquor application to this environment. Spinel-based materials were thought to be suitable for this application due to the high temperature and high alkali contents of the service environment. Yet, since this is an acidic environment, an alumina-rich as opposed to magenesia-rich spinel formulation was considered as shown in Table I. These materials would offer the advantages of being non-chrome containing and less expensive to produce and install.

Figure 5. Refractory test samples after exposure to molten smelt immersion test. Samples shown are (a) phosphate bonded magnesia castable repair material, (b) magnesium-rich spinel schotcrete for primary lining, and (c) currently used fusion-cast alumina. It was found that the spinel shotcrete material performed well when exposed to molten smelt for 100 hours at 1000°C (1832°F). Similar to the currently used fusion-cast spinel material, the spinel shotcrete sample retained its sharp edges and integrity even though it did show some signs of smelt penetration. A color change (darkening of the sample matrix) was seen in both materials, which was confirmed to be a chemical change due to penetration of sulfur and alkali. Yet, this chemical change did not result in the expansion of the sample structure as seen in previous high-alumina containing refractories used for this application, which led to spalling and failure of the hot face lining due to the formation of beta-alumina.

Evaluation of this material was performed through refractory cup testing of samples supplied by MinTeq. A reservoir was core drilled in a refractory block and filled with 35 grams of commercial slag. The sample was then heated to 1600°C (2912°F) at 5°C/min (41°F/min) under Argon and held for four hours before being cooled naturally back to room temperature. Upon cooling, the sample was sectioned for analysis. Results from the first round of samples tested are shown in Figure 6.

The phosphate bonded magnesia castable did not perform as well as the other two materials when subjected to the same conditions, but still maintained its integrity and only had slight dissolution of the exposed refractory surfaces. Less penetration of the smelt was evident in this sample as well. Based on these results it is still believed that this material would serve well as a repair material since it would not need to survive as long in the hostile molten smelt environment. Adhesion studies, similar to those performed for the candidate aluminum furnace repair material discussed previously, are still needed for this material. Additionally, industrial validation of both the spinel shotcrete primary lining material and the phosphate bonded magnesia castable repair material are desired. Coal Gasification

Figure 6. Refractory cup samples before and after exposure to molten coal slag. Alumina-rich spinel material before testing (a), after testing (b), and showing the interaction of molten smelt with the exposed refractory surface (c).

Similar to the black liquor gasifer, coal gasification systems involve high temperatures («1300-1600°C), aggressive chemical species (including sulfur, alkali salts, and heavy metals), and erosion/corrosion effects. Additionally, issues such as thermal cycling, variable environments, and elevated pressures (>400 psi)

1214

Cracking was seen in this material due to shrinkage issues during heating to 1600°C which are being investigated in future optimized compositions. Otherwise, these materials appeared to perform well with only minimum interaction between the molten slag and the refractory material. Visual observations were substantiated through X-ray diffraction data. At the sample interface between the melted smelt and the refractory, significant iron from the smelt (in the form of free magnetite - Fe304) and the formation of forsterite - (Mg,Fe)2Si04 were seen where it appears free magnesia is combining with iron at this interface. Yet, when the analysis was moved away from this area to the middle of the brick no free iron (magnetite) was seen and the amount of fosterite was greatly reduced. At the back side of the brick, low levels of iron and high levels of spinel (Mg2Al204) were seen. This would seem to indicate that although there is some interaction of the smelt with the brick at the interface, it is relatively confined to there.

was tested as developed through laboratory refractory cup testing. Refractory cubes (50x50x50 mm) were drilled using a 20 mm carbide tip masonry bit to a depth of 25 mm. The cups were filled with 6 grams of dried mud (nominal composition: 55-60% Na2S * 9H20, 4-5% K2C03, 0.5-1% NaCl, 35-37% Na2C03) before being heated to 1500°C (2732°F) in four hours. Samples were held for six hours at temperature, and then cooled over a twelve hour period before being sectioned for analysis. Samples were rated based on the cross-sectional area of penetration as shown in Figure 8. Previously tested materials exhibited large areas of penetration as shown in Figure 9.

To improve the performance of the spinel-based material, modifications are being made by MinTeq in the composition of this material. Additionally, surface coating technology similar to that currently used for this application, as discussed above, is being developed at ORNL for the spinel-based material system utilizing a unique ORNL coating process. A slurry coating process (as shown pictorially in Figure 7) is being considered since it is non-line-of-sight, can be used for complex shapes, will fill open porosity, requires no unique equipment, and can be automated. Once the coating has been developed and successfully applied to the current spinel-based refractory system, testing will be carried out through exposure to molten slag in refractory cup testing. Additionally, abrasion testing will be carried out on coated and uncoated spinel-based shotcrete materials. If successful, it is hoped to then pursue industrial testing of this material in actual gasifier applications.

Figure 8. Cross-sectional area of penetration used for evaluation of candidate lime kiln refractories.

Üüüi»l

•

" it

mmmg

Figure 9. Picture showing large cross-sectional areas of penetration for previously tested candidate lime kiln refractory materials.

Hi

Figure 7. Schematic representation of slurry based coating process, (used courtesy of Beth Armstrong, ORNL).

Static cup testing of the magnesia-rich spinel formulation in contact with industrially obtained lime mud is shown in Figure 10. This testing showed it to be highly resistant to attack at processing temperatures characteristic of those provided by industrial partners with no adherence of the mud to the refractory cup.

Lime Lime kilns are also high temperature and high alkali environments operating at temperatures as high as 1500-1600°C. In the various zones of the kilns high strength castables, high duty fireclay bricks, magnesia-based bricks and high-alumina bricks are utilized. This work focused on the burning zone of the kiln which is the hottest zone and where high-alumina or magnesia-based brick is traditionally used. Due to the rotating nature of the kiln, materials are subjected to both high temperature corrosion and mechanical abuse. Therefore, materials used for this application are desired to be resistant to chemical attack, abrasion resistant, spall resistant, thermally insulating, mechanically strong, and low cost.

Additionally, preliminary energy analyses have been performed that predict significant projected energy and economic savings can be achieved (as compared to use of current state-of-the-art materials) when materials such as these are used in lime kilns in conjunction with an insulating refractory back-up material like the one shown in Table I. This alumino-silicate material was developed under this project since the conductivity of the spinelbased material is substantially higher («5-8 W/mK) than that of a traditional alumino-silicate refractory («2-5 W/mK). Therefore, a traditional composite lining strategy can be used consisting of a more corrosion/erosion resistant and mechanically sound hot face lining teamed with a highly insulating back-up lining material. The advantage that the newly developed back-up lining material offers though is it still exhibits good mechanical properties

Based on the above list of desired properties and the use of lime kilns in the pulp and paper industry, the magnesia-rich spinel material developed for black liquor gasification applications was pursued as a candidate material for this application as well (as shown in Table I). This in-situ spinel forming shotcrete material

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(shown in Table II) despite having very low thermal conductivity («0.5 W/mK).

References 1. J.G Hemrick, K.M. Peters, and J. Damiano, "Energy Saving Strategies for the Use of Refractory Materials in Molten Material Contact," Energy Technology Perspectives: Conservation, Carbon Dioxide Reduction and Production from Alternative Sources, TMS, February (2009), 225-232. 2. Stephen C. Carniglia and Gordon L. Barna, Handbook of Industrial Refractories Technology - Principles, Types, Properties and Applications (Park Ridge, NY: Noyes Publications, 1992), 53-54.

Figure 10. Magnesia-rich spinel material after exposure to industrial lime mud in refractory cup test. Slug of lime solidified and fell out of cup after testing leaving a clean, unaltered refractory surfacewhere the lime mud was in contact with the refractory. Table II. Mechanical Properties of Insulation Back-Up Lining Material

Firing Temperature (°Q

Bulk Density (g/cm3)

110 815 1093 1260

1.15 1.09 1.11 1.11

Modulus of Rupture (MPa) 3.03 1.45 2.28 3.10

Cold Crushing Strength (MPa) 6.89 7.58 7.58 6.21

3. J.G. Hemrick, W.L. Headrick, and K.M. Peters, "Development and Application of Refractory Materials for Molten Aluminum Applications," International Journal of Applied Ceramic Technology, Vol. 5, No. 3, (2008), 265-277. 4. J.G. Hemrick, J.R. Keiser, and R.A. Meisner, "Material Characterization and Analysis for Selection of Refractories Used in Black Liquor Gasification," Materials Challenges in Alternative & Renewable Energy Proceedings, Cocoa Beach, Florida, February (2010). 5. J.G. Hemrick, J.R. Keiser, R.A. Peascoe-Meisner, J.P. Gorog, and W. Ray Leary, "Material Characterization and Analysis for Selection of Refractories Used in Black Liquor Gasification," Proceedings of the 44th Symposium on Refractories, St. Louis, Missouri, March (2008), 101-114.

Apparent Porosity

(%) 37 45 48 55

6. R.A. Peascoe, J.R. Keiser, C.R. Hubbard, J.P. Gorog, C.A. Brown, and B. Nilsson, "Comparison of Refractory Performance In Black Liquor Gasifiers and A Smelt Test System," Proceedings of the International Chemical Recovery Conference, Whistler, British Columbia, June 11-14, (2001), 297-300.

Acknowledgements

7. J.P. Bennett, K.S. Kwong, H. Thomas, and R. Krabbe, "PostMortem Analysis of High Chrome Oxide Refractory Liners from Slagging Gasifiers Using Analytical Tools/Thermodynamics," UNITECR' 09 Proceedings, Salvador, Brazil, October 13-16 (2009).

Research sponsored by the U.S. Department of Energy (DOE), Energy Efficiency and Renewable Energy-Industrial Technologies Program (EERE-ITP) under Award Number CPS Agreement #14954 with UT-Battelle, LLC. The authors wish to acknowledge Kelley O'Hara and Todd Sander of Missouri University of Science and Technology who contributed to the aluminum refractory testing, Hu Longmire and Adam Willoughby of ORNL who contributed to the black liquor gasification refractory testing and analysis, and Beth Armstrong of ORNL who is collaborating on the coal gasification refractory coatings work. The authors would also like to recognize the contributions of the late Fritz Henry of MinTeq for his refractory development efforts across this project. Finally, the authors would like to recognize Andrew Wereszczak and Fei Ren for their technical review of this manuscript.

8. J.P. Bennett, K.S. Kwong, A. Petty, C. Powell, H. Thomas, H.D. Prior, and M. Schnake, "Field trial results of an improved refractory material for slagging gasifiers," 23rd Annual International Pittsburgh Coal Conference, Pittsburgh, Pennsylvania, September (2006).

This submission was produced by a contractor of the United States Government under contract DE-AC05-00OR22725 with the United States Department of Energy. The United States Government retains, and the publisher, by accepting this submission for publication, acknowledges that the United States Government retains, a nonexclusive, paid-up, irrevocable, worldwide license to publish or reproduce the published form of this submission, or allow others to do so, for United States Government purposes.

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Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

STUDY ON THE ENERGY-SAVING TECHNOLOGY OF CHINESE SHAFT CALCINERS Guanghui LANG, Chongai BAO, Shoulei GAO, Ronald Lee Logan, Yan Li Sunstone Development Ltd, Beijing 100029, China

Keywords: Shaft Calciner, Energy-saving, Integrated Technology, Carbon Burning Loss When calcining, VCM discharged from the calcining pot is mixed with the airfromthe preheating air tunnel and burns in the first level flue. This process generates high temperature exhaust gas, which flows to the second level flue, then the third, and finally the last flue by the draft from exhausting system. After the exhaust gas passed through the last flue, it arrives at the heat accumulator room, where heat exchange happens between cold air and hot exhaust gas through bricks. The cold air can be heated up to 400-600°C, and then it goes to the first level flue to help VCM burning.

Abstract Shaft calcining technology is widely used in China for the calcination of green petroleum coke. It has the advantages of lower energy consumption and lower carbon burning loss compared to the rotary kilns. In this paper, a method of further reducing carbon burning loss using integrated technology is discussed. Introduction Rational utilization of resources is a basic state policy for China. During the production of carbon materials, calcination of GPC (Green petroleum coke) is a very important process, which is to heat up the GPC to a high temperature. When calcined, moisture and VCM (Volatile Combustion Matter) are removed from coke, crystal-lattice grows, real density increases, bulk shrinks, and mechanical strength enhanced. And, the performance of electrical conductivity, thermal conductivity, thermal shock resistance and oxidation resistance are all improved simultaneously. [1] Currently, more than half of the Chinese carbon plants use vertical shaft calciners as main method for petroleum coke calcination, and the current production capacity is about 3,500Kt/a.

The connection of level 1 to 8 flue are 'Z' type from the top to the end, and the green cokes are heated indirectly by the flue on both sides of the calciner. High temperature exhaust gas coming out from the last flue goes into a boiler to recycle the heat through the main flue. Fig. 2 shows the 8fluesof calciner.[2]

Vertical shaft calciner is an indirect heating calciner where GPC and flues are separated. The VCM from GPC when heated burns in the flue and heats up GPC inside the shaft. So it's somewhat a self-support system. Once it starts, you don't have to input fuels to maintain the calcining process. Fig. 1 is a cross sectional view of vertical shaft calciner.

1 - Air Damper; 2 - First Rue; 3 - VCM Damper; 4 - VCM Vertical Tunnel; 5 - Temporary Drying Furnace Device; 6 - Second Flue; 7 - Preheated Air Damper; 8 - Last Flue; 9 - Exhaust Gas Damper; 10 - Boiler Fig.2 Eightfluesof the calciner Energy available for us in the world is less day by day. Energy saving technology has been one of the most important fields to develop in. The advantage of shaft calciner - less carbon burnt loss and high quality calcining process - has made it very popular in carbon industry all over the world. In year 2008 and 2009, this technology was introduced to Brazil and Middle East countries. The advantage of energy-saving

Cooling Sleeves; 8 - Calciner Support Fig.l Cross sectional view of vertical shaft calciner

Now, there are three kinds of mainly used calcining equipment (rotary kiln, vertical shaft calciner and rotary hearth

1217

calciner). [3] A comparison of these three calcining equipment is shown in Table 1.

whole plant. Or, through a boiler, it can be used in Green Mill or daily lives. The energy balance diagram of heating and cooling zone of the shaft is shown in Fig.4 and Fig.5.

Table 1 calcining equipment comparison Item Technology Features Product Quality

Rotary Kiln Direct Heating

Productivity

6-12t/h

Carbon Loss Energy Consumption Recycle Heat Environment Protection Material Adaptability Investment Operation Cost Automatic Level

7-10%

80-100kg/h (per Shaft) 3-4%

High

Low

Low

Yes

Yes

Yes

Good

Good

Good

Wide

Common

Wide

Low

High

High

High

Low

High

High

Low

High

Shaft Calciner Indirect Heating Good

Fine

Volatile sensible heat 4.7% f

Rotary Hearth 1 Direct 1 Heating

Preheated air sensible heat 11 8% 83% VCM combustion heat

Raw Material sensible [ J \ heat 0.5%

Good 13-21t/h 7-8%

Ë. 2.7% Water \ \ ] evaporation heat 14.6% VCM incomplete combustion heat loss 24% Raw material heat content entering the cooling zone

5% Heat radiation

Fig.4 The heating zone energy balance diagram Preheated air sensible heat f

Carbon burnt lossbecause of air entrance 40.6% 59.1% Raw material heat content entering the cooling zone

We can see from table 1 that the advantages of vertical shaft calciner are less carbon burnt loss, less operation and maintenance cost, energy-saving and so on. It corresponds with the tendency of using resources in sustainable development way. Fig. 3 is theflowchart of energy circulation in calcining. 5.6% Calcined coke out of the heat

GPC

Power grid

Waste heat

Gas

generation

Steam

Calcination

39.8% Coolinh water out of the heat

Burning

Fig.5 The cooling zone energy balance diagram

Heating

From figure 4 we can tell that 83.1% of heat income in calcinations zone is from VCM, and the hot flue exhaust gas takes away 63.7% of heat. The proportion of the heat produced by carbon burnt loss and carried by the cooling water respectively is 40.6% and 39.8% in Figure 5. Therefore, reducing carbon burnt loss has the advantage of increasing product recovery efficiency and decreasing the quantity of the cooling water for the shaft.

m Plant

Heating medium

Power line

boiler

Heating water

Keep warm

Liquid

Green Anode

Ditch

Formation

Heat provided to formation

Given an example of 150000 ton/a pre-baked anodes production plant, it need to construct six sets of shaft calciners (28 shafts and 8 flues for each calciner), the output per shaft is 80-100kg/h, exhaust gas temperature around 850-900 °C, gas flow about 10000Nm3/h, temperature of exhaust gas from boiler is 400 °C. There is about 154 x 104Kcal/h heat by one shaft.

Heating

Fig.3 Flow chart of energy circulation in calcining There are 10 to 12 percent of VCM in GPC, which must be discharged during calcining. With these VCM, we don't have to input additional fuel. And, the GPC are heat treated under an air isolated condition, which allows low carbon burnt loss.

Totally, the production and daily life need about 584* 104Kcal/h. So the waste heat produced by one set of calciner can satisfy then, and other 5 sets of calciners can be used to generate electrical energy for the production and daily life of the whole plant. And, the temperature of hot gas from boiler is still about 400°C high. There is potential to recycle them further to save more energy.

Escaping VCM enter into the flue and burn. The heat produced is used for GPC calcination, and high temperature exhaust gas of 800-1000 °Ccan be used to generate electrical power, which can satisfy 80 percent electrical energy of the

1218

It's extremely important for a carbon plant to recycle the waste heat from calciner exhaust gas. It's high efficient, environment friendly, and economical. There is great potential for energy saving. In the recent years, heat recycling is widely applied in many Chinese carbon plants successfully. And, it will be introduced to more and more plants all over the world. So, how to improve the efficiency of heat recycling and decrease carbon burnt loss is a necessary step for energy saving. The direction of improvement Shaft calciner has very good sealing around the body. However, air may leak into it during discharging. Actually, this is the main cause of carbon burning. [4]

Cooling water

Air leaking from the discharging gate will react with the carbon material inside the shaft. It happens at the cooling zone, where carbon material temperature is about 300-500 °C, and generate C0 2 . C0 2 rise up, react with the carbon material and generate CO, which will burn in the flue finally. In this process, some carbon materials are burnt and generate additional greenhouse gas - CO2. On the other hand, it brings additional heat to the cooling zone so that the cooling water quantity has to be increased also. It's a negative process which leads to more cost and less environment protection. So, if we cut off the access of air at the discharging spot, we can reduce the carbon burnt loss. And therefore, we can reduce our cost, save energy and protect the environment with less CO2 emission.

I

W 1

f_

Cooling water ir extraction

CPC Discharge Fig. 6 The diagram of the draft sealing technology Conclusion (1) The Chinese vertical shaft calciners have big advantage in energy saving. They consume energy in a high efficient way with less carbon burnt loss and high CPC quality.

According to the actual structure of the discharging unit, we find a effective way to cut off the air, sealing this unit by draft. As Fig 6 shows. We fix an air extraction pipe at the CPC (Calcined Petroleum Coke) discharging unit and connect it with a draft generator. A negative pressure of about -10-0 millimeter of water inside the discharging pipe can cut off the access of air to the carbon materials inside the shaft.

(2) The draft sealing technology at the discharging unit can reduce the carbon burnt loss, and therefore reduce the quantity of cooling water and C0 2 emission. This technology can improve the performance of calciners' energy consumption. References

When 1000kg GPC are calcined, we can get about 750 kg CPC only. About 10% of moisture and 10-13% of VCM are lost or burnt. In the process of CPC discharging, there is air entering the calciner, which will burn about 3-4% of coke and generate additional heat. This additional heat needs about 40% more cooling water. However, if we seal the discharging unit with this technology, the air leaking in can be reduced about 70% and, carbon burnt loss reduced to 0.5-1% accordingly. We can save carbon burnt loss about 6000-7500 tons per year. That means 25,000 tons of C0 2 emission and 36,000 tons of circulation water per year can be saved also. This technology can realize our purpose of reducing carbon burnt loss and C0 2 emission.

1. Pingfu Wang, "Production Technology of Prebaked Carbon Anodes Used in the Aluminium Electrolytic Cells" (Metallurgical Industry Press, 2002), 130-169. 2. Pingfu Wang, "Analysis and Study on the Chinese Vertical Shaft Calcining technology for petroleum coke"(Carbon Technology, 2009), 41-45. 3. Lining Chen, "The Production Process of Rotary Kilns and its control characteristics" (Carbon Technology, 1999), 33-37. 4. Tao Xu, "Graphite Products Technologies" (Wuhan University of Technology Press, 1992), 11-15.

1219

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

AUTHOR INDEX Light Metals 2011 A

Aarhaug, T Abdullah, F Abreu, A Adamson, H Ahmed, A Ai,Z Akhtar,S Al-Dubaisi, A AlHosni,S Al Zarouni, A Alamdari,H Allard, B Alli,M.... Almeida Neto, L Ambarkutuk, A Ammar, H Andersen, D Antille,J Apisarov, A Araujo,J Argyropoulos, S Arkhipov,A Arkhipov,G Armoo, J Artola,J Aune,R Ayhan,T Azari,K Azevedo, C

B

Backhouse, N Bagcivan,N Bainbridge,1 Bakken,J Ballentine,F Ban, Y Banerjee, T Bao,C Bao,S Barbato, C Batista, E Baudino,M Baudrenghien, J Baxter, R Bayliss,C Bedel, M Beier, S Belt,C Beltramino, L Berkovich,A Berlioux, J Berntsen,H Bertherat, M Bertolo,J Bertran,G Bigot, J Billat,G Bilodeau,J Bin,C BJ0rnstad,H Bjarno,0 Bo, W Bobzin,K Boenigk,W Bojarevics, V

263, 339, 345 745 587, 907 81 449, 693, 803 443 763 23 859 333, 745, 1079 973,1111,1161 1091, 1097 97 621 821 1161 1009 569 563 907 809 1047 1047 461 847 763 621 1161 179

Boltersdorf, C Bouchard, J Bouchard, P Boufounos, D Bourgier, A Bratcher, D Broek,S Brooks, G Buchholz, A Buck, H Buehrig-Polaczek, A Bugge, M Bussmann, M

C

Caboussat, A Calixto,J Canada, M Canming,X Cannova,F Cantin,K Caratini,Y Carlberg, T Carlson, W Carvalho,R Chanda,A Chandler, J Chao,Z Charette,0 Chateeriji, S Chen, J Chen,W Chen,X Chιnard,S Choudhury,P Chunming, Z Clark, D Cobo, E Cockcroft, S Colavito,D Coleman, L Combeau,H Cooksey, M Coppack, J Correa, L Cotι,P Coulaud,C Coulombe,M Courchιe,R Courtenay, J Crosdale,M

865,931 705 687 763 107 443,455 69 1217 775 47 291,491 959 635 437 309 699 979 1205 847 865 635 57 657 351 351 847,859 1091 1041 497 1017 295 201, 221,241 705 889 549

D

Daglilar, S Dahle,A Damdah,L Dando, N Darmstadt, H Das, S Dasgupta Davis, J Dedyukhin,A Delcorde,X Dιry,S Desgroseilliers, B Desilets,M

1221

889 997 269 113 635 1189 361,531 575, 725, 751, 1085 1179 711 705 415 809

581 281 931,937 591 931,937 967 1041 675 185 57 269 859 145 525 693 369, 575, 599, 605, 979 227 309,357 ...1003 431 197 959 847 783 1211 325 699 543 663 57 739 871 1097 653 769 75

821 815 757 325,361 931 575 29 725 563 467 467 467 1035

Dιsuets, M Di Lisa, D Ding,K Ding,T Dion, M Dispinar, D Donaldson, D Dong, G Dong, H Dorreen, M Dou,Z Doutre,D Dubι, D Duchesne, C Duchesneau, L Duncan, A Dupuis,C Dupuis,M

E

Edwards, L Einarsrud, K Engh, T Erdegren, M Eskin,D Eufrasio,M

F

Fafard, M Fateh,B Faure, M Fehrenbach, F Fei,Z Felberbaum,M Fellner,P Feng,J Feng,N Feng, Y Fergusson, L Fiot,N Fischer, W Fleckenstein, J Folkers,A Forato, 0 Fortin, H Franca, S

G

Gaertner,H Gagne, J Gagnon, F Gao,B Gao,J Gao,S Gao,Y Gariιpy, R Garrido,F Gauvin,G Gendre, M Ghosh, K Ghuloom,K Gilbert, M Gill, J Gillner,A Girault,G Giri,U Grab, H

1041 875 985 251 931 731 171 201, 221 455 319, 369 1155 809 1111 407, 967 953 75 739 519

895,947 555 719, 775 675 669 881

537, 973, 979, 991, 1003, 1161 725 351 657 51 815 513 1053 1029,1155 543 81 871,881 1067 325 425 859 979, 1003 47

Grama, A Grandfοeld,J Groutso,T Gu,S Gudbrandsen, H Gudmundsson, H Guilin,L Guillot,M Guoyin, Z Guozhi,L Gusberti,V Guthrie,R

H

Haakonsen, H Haarberg, G Hacini,L Haiyan,Y Hakonsen,A Hamaguchi, M Han,H Han,M Hang,W Hannart,B Hansen, J Hartmann, C Hassan, H Hassan, J Hautus,M He, H He,J Heidari,H Hemrick,J Hennings, J Herbert, J Hial,S Hives, J Holm,K Holtkamp,J Hongliang, Z Hosler, B Hou,J Howard, S Hu,C Huang, K Huang, X Hui, F Hui-Lan,S Hui-lan,S Hyland, M

345 269 537 405 757 1217 301,319,605,1103 467 13 1003 865 803 449 319 797 705 351 69 647

I

Ing,D Inoue,T Instone, S Isac,M Ivanov, T

J

Jaouen, G Jarry,P Jensen, E Jensen, T Jentoftsen, T Jiang, G Jianping,P Jiao,K

1222

399 751 399 89,509 461 471 245 991 615 33, 145 853 797

415 461 1041 245,251 793 913 405 19 567 657 1017 705 449 803 605 1053 1155 1111 1211 657 1193 175 513 1061 705 145 483 1117 191 319 237 865 197 201 221,241 285, 399, 745,1079

647 913 769 797 705

641 699 1017 137 461 1103 205,1023 1103

Jie,L Jilai,X Jin,X Jing,Y... Johnsen, L Jonville,C Jouet-Pastrι, L Juan, D Jun,L Jungi, L

K

Kan, H Kandev,N Kano,K. Kapoor,A Karadeniz,0 Karhausen, K Katgerman, L Kato, C Keiser,B Keller, F Kennedy, M KertU Khaliq,A Khatun,A Khoshneviss, D Kildea,J Kirkpatrick, D Kiss, L Klett, C Kobbeltvedt, 0 Koch, H Koide,J Komatsu,N Kotsis,1 Kshatriya, N Kumar, A Kumar, C Kvithyld,A

L

Ladam,Y Lai, Y Lalpoor, M Lang, G Laplante,F Larouche,D Lauzon-Gauthier, J Lavoie,P Lebeuf,M Lewellyn, M Li,B Li,D Li,J Li, W Li,X Li, Y Liang, X Lifschitz,J Lin,X Lindner, F Lindsay, S Liu, F Liu,J Liu,K Liu,Q Liu, Y

1143 197 1143 205 959 859 657 33 1023 51

Liu,Z Logan, R Lorentsen, 0 Lossius,L Lowe, M Lu,D Lü,X Lu,Z Lubin,M Lv,G

M

503 1003 913 985 821 625 669 325, 587, 907 185 853 763 821 751 1035 57 185 959 581, 997 125, 151 415 711 913 913 113 29,103,255 333, 699 69 719,731,775

393 1117,1135 669 1217 953 991 967 369,1073 1097 107 599,1029 797 1117,1135 227, 309, 357 237 1217 319 315 319 889 163,361 509 405 1135 1053 1129,1149,1155

Ma,J Mackay,G Mackay,K Mahieu,P Maijer,D Maiwald,D Makhlouf,M Maltais, B Mannweiler, U Mao,J Marceau, D Marino, J Marks, J Martin, G Martin, L Martin, M Martin, 0 Massambi, S Mathieu, J Mathis,A Matsuki,T Matsuzaki,K Matthies,C McCulloch,R McNerney, J Medeiros, M Meher,S Meier, J Meier, M Meijer,M Melo,D Meng,L Metson,J Migchielsen, J Mimna,R Mingchun, L Mingliang, S Miotto, P Miranda, J Mishra,B Mishra, C Misra,C Missalla, M Miwa,K Mnikoleiski, P Moffatt,S Mohamed Al-Jallaf, M Mohapatra,B Mohapatra, S Molenaar, D Monteiro, A Montoro,E Moody, K Moraes, E Morales, F Moras, A Moreno, C Morsi, Y

1223

227 1217 393 841,941 725 443,455 1117 251 947 1155

1129 269 269 881 783 875 831 739 1067 319,455 997, 1041 947 309 803 191 1003 581 351 657 739 827 1199 711 437 41 13 231 185 853, 1067 421 281 497 151,285 647 185 121 51 291 281 93 93,97 131 125,157 827 875 107 333, 745, 1079 93 681 985 179 291 663 179 859 361 315 575

Mukhlis,R Mulder, A Mullett,M Murakoshi,Y

N

Naixiang, F Namboothiri, S Nanda,S Nele,M Neple,S Neumann, S Niedermair, F Niehoff, T ÍΟΟΓ,Α

Nordmark, A

O

Oakes, J Ofico,I Okuyama,N Ordronneau, F Osen,K Ostrem,0 0ye,H

P

Padhi,B Palchowdhury, A Pan,X Panchal,J Patel,P Patnaik,R Pattnaik,P Paul,W Paulsen, K Paulus, R Peel, A Perander, L Perron, A Perruchoud,R Petit, 1 Phillips, E Picard, D Pilote, S Pingin,V Pires, R Pisano, J Plikas,T Pluchon,C Poizeau, S Pomerleau, L Pradhan,R Prasad,A

Q

Qi,X Qian,J Qiang,D Qin,H Qin,J Qingsheng,L Qiu,S Qiuyue,Z Quiroga, G

1085 425 81 1199

205, 1023 369 681 47 881 625 1173 1185 693

731

1189 881 913 865 263,339 1061 941

231 681 251 925 1073 211 69 477 415 1091 1193 151, 285 581 1067 967 185 973,1161 1041 1047 437 269 531 657 387 865 69 687

R

R0dseth,J R0Tvik, S Raahauge,B Raghavan, P Raghavendra, K Ramana, Y Ranaud,G Rappaz, J Ratvik,A Redkin,A Reeb,B Reek,T Reilly,C Renaudier, S Rey Boero, J Reusch,F Rhamdhani, M Ribeiro, F Ribeiro, J Ribeiro Alves Filho, J Richard, D Riquet,J Robichaud,P Rodrigues-Schroer, A Rolle, J Rosefbrt,M Rosenkilde,C Rouleau, M Rout, A Roy, A Roy,P

1179 841,1061 137 29, 103,255 449 211 657 581 339,345,841 563 125 375 783 581 847 769 725,751,1085 907 157 57 525 657 739 1211 925 711 339 997 231 549 847

S

Sadoway,D Sagin,A Saha,R Sahin,Y Sakai,K Salles, M Sampaio, J Sandnes, E Santos, H Santos, J Sato, Y Schmidt, H Schmidt, T Schneider, A Schulz, R Schwarz, P Segatz,M Severo, D Seyed Ahmadi, M Shah,U Shaofeng,F Sharma,A Sharma,P Shaw,R Shi-Wen,B Shi-wen,B Shimizu,T Shishido,T Shiwen,B Shuchan,W Silva, F Simard,C Simard, S

455 405 615 319 443 197 309,357 145 315

1224

4

387 821 69 215 913 587 13 555 281 281 1073 125,157 647 525, 531 1111 543 393 853 809 531 1023 431 97 75 201 221,241 1199 913 245 145 13 467 467

Simpson, M Singh, R Singleton, V Skybakmoen, E Skyllas-Kazacos, M Smith, J Soffer,M Solheim, A Sommerseth, C Songyun, D Sorhuus,A Soucy,G Souza,L Spencer, B St-Arnaud,P St-Georges,L Stam, M Steele,T Steen,1 Stevens, D Stiegert,J Suarez, C Sulger,P Sun, S Sun,X Sylvain, L

T

Tabereaux, A Takahashi,K Tamura, T Tangstad,M Tao, Y Tarcy, G Taylor, J Taylor, K Taylor, M terWeer,P Tessier, J Theiss,S Thonstad,J Tian,Z Tiefenbach, P Tikasz,L Tin'ghaev,P Ting-an,Z Tjahyono,N Toptan,F Trempe, 0 Tu, G Turhan, S Turpain,C

U

Usta,B Uysal,0

V

Van Son, V Vιkony,K Verreault,J Vershenya, A Vitchus,B Viyyuri,S vonKaenel,R von Krüger, P Vroomen,U

Vry,B

175 175 1123 263, 339, 461, 1061 1123 1211 319 263, 339, 381, 393, 461, 1061 263, 339 1143 295 1097 291 941, 959 973 997 285,425 663 793 719 889 41 853 599 319 325

W

Wang,D Wang, F Wang, H Wang,J Wang,W Wang,X Wawrynink, K Waz,E Webb,J Wedde,G Wehrli,J Wei,C Welch, B Welshons,D Wenxin,H White, D Wijayaratne, H Williams, F Wilson, S Wimroither,G Wind, S Wischnewski, R Woloshyn,J Wong,D Wood, G Wood,R Wu,C Wu,J Wu,M Wu,Y Wyborney,M Wynn,A

319, 329 1199 827 775 497 329,407,483,491 725 81 301, 319, 369, 399,425, 605 5 407, 967 705 513 1135 1061 437 563 33, 145 301 821 991 599 215 467

X

Xenidis,A Xia,W Xianqing, L Xiao,J Xiaobing, L Xiaobing,Y Xiaoping, Y Xiaoxiao, W Xiquan,Q Xue, J Xue-zheng,Z

Y

215 215

325 581 467 531 937 1185 569 275 705

1225

Yadav,D Yan,J Yan, L Yan,X Yang, H Yang,J Yang,Q Yang,W Yanping,Q Yanqing, L Yao, G Yao,L Yaowu,W Yexiang, L Yin,J Yingfu,T Yingwen, B Yuezhong,D

907

405 1029 1155 509 1053 483,491,503 103,255 699 41 295 81 3135 333,745,1079,1123 41 205 1169 399 131 719 1173 137 157, 179 531 301,319 491 657 831 1217 137 121 925 663

113 19 63 405 615 497,591 63 33 1023 1053, 1103 221,241

97 405 33, 33,145 1085 509 309 455 575 121 1143 1129, 1149 783 205 1143 19 567 51 205

z

Zaikov,Y Zaloznik,M Zavatti,J Zeigler,D Zhai,X Zhan,L Zhang, H Zhang, L Zhang, N Zhang, S Zhang, T Zhang, W Zhang, X Zhang, Z Zhao,J Zhao, Q Zhao,X Zhihe,D Zhihui,W Zhu,H Zhu,J Ziegler,D Zografidis, C Zou,Z

563 699 315 369 599 1053 1117,1135 757 503 357,1029 1155 301, 319 1029, 1129,1149 1009, 1129, 1149 917 309, 357, 917, 917, 1155 405 33, 145 205 237 783,1053,1103 537,901,973,991,1161 113 1135

1226

Light Metals 2011 Edited by: Stephen J. Lindsay TMS (The Minerals, Metals & Materials Society), 2011

SUBJECT INDEX Light Metals 2011 3

300kAPotlines.

..405

5Cu / (10NiO-NiFe2O4) Cermet

1135

Abrasion Resistance 1067 ACD 509,737 ACD Model 567 Adsorption 251 Advanced Process Control 69 Air Emissions 361 Air Gap 991 Air Monitoring Network 315 Al-Si Alloys 815 Al-Ti-C Grain Refiners 821 Alcoa 587 AlFeSi 711 Alumina 5,23,151, 163, 185, 197,221,503 Alumina Calcination 137 Alumina Discharge 93 Alumina Feeding 449 Alumina Leaching 201, 241 Alumina Mixing 543 Alumina Production 63 Alumina Quality 285 Alumina Refinery 191 Alumina Technology 5 Aluminate Spinel 1085 Aluminium 345,415,431,461, 513, 519, 537, 675, 731, 775 Aluminium Electrolysis 319, 1123 Aluminium Electrolysis Cell 263, 549 Aluminium Electrolysis Pots 1091 Aluminium Process 339 Aluminium Reduction Cell 543, 591,1003, 1009 Aluminium Smelting 605 Aluminum 295, 325, 657, 663, 719, 757, 763, 821, 827 Aluminum Alloys 669,699,711 Aluminum Carbide Formation 1097 Aluminum Dross 197 Aluminum Electrolysis 567,581,1023, 1053, 1097, 1117 Aluminum Hydroxide 205 Aluminum Manufacturing 421 Aluminum Melting and Holding Furnaces 1169 Aluminum Processing 737 Aluminum Production Decrease 497 Aluminum Reduction 483, 509, 599 Aluminum Reduction Cell 405 Aluminum Reduction Cells 1029 Aluminum Reduction Pot 443,455 Aluminum Smelters 269 Aluminum Smelting 483 Aluminum Tri-hydroxide 121 Alunorte Performance 57 Amperage 415 Amperage Creep 525 Analysis Method 483 Ancillary Equipment 477 Anode 871, 881, 895, 913, 991 Anode Assembly 985 Anode Baking 853 Anode Baking Furnace 847, 859, 865, 875 Anode Carbon 985 Anode Connection 979,1003 Anode Cover 399 Anode Crust Cover 351

Anode Current Density Anode Effect Anode Effects Anode Gas Anode Interfaces Anode Optimization Anode Paste Anode Power Losses Anode Quality Anode Reaction Anode Spikes Anode Stub Contact Anode To Cathode Distance Anodes Anodic Gas Bubbles ANSYS AP18 Approach Asperity ASTM ASTMD4292 Attrition Attrition Index AzZabirah

333 281, 319, 325, 329, 333,467 309, 357 1155 997 889 1161 1003 967 901 471 1003 1009 471 555 519,1047 415 657 1009 941 959 163 163 23

B

Bake Furnace Design 853 Baking 871,881 Basicity 245 Bath 381 Bath Chemistry 491 Bauxite 5,23,93,113,211 Bayer Liquor 215 Bayer Process 51,63,81,103, 107, 121,131, 151,171,179,185, 255 Bayer Residue 89 Bend Strength 1129 Bifilm Index 731 Billet 641 Billet Casting 793 Bimetal Clad Sheet 615 Blaine 1143 Boride 751 Boron Treatment 751 Bubble Dynamics 575 Bubble Transport 549 Buildup 621 Bulk Density 947 Bulk Density Testing 941 Burner 1185 Butt 351

C2oA13M3S3 Calcination Calcined Coke Calcined Coke Size Distribution Calcined Petroleum Coke Calcining-leaching Process Calcium Aluminate Slag Calcium Fluoride Cancranite CaO Ratio CaO2Al 2 0 3 Carbon Carbon Anode Carbon Anodes Carbon Burning Loss

1227

201 131,151,157,895,931 931 937 925,959 197 201,221,241 563 255 201 205 865, 871, 881, 895, 947 973,1009,1143 841 1217

Carbon Cathodes 1053 Carbon Dioxide 901 Carbon Residue 719 Carbonate 51 Cast Iron 985, 991,1009 Castable 663 Casthouse 635, 641, 657, 693, 793 Cathode 519,537,1079 Cathode Cast Iron Rodding 1017 Cathode Design 569 Cathode Interfaces 997 Cathode Material 1103 Cathode Wear 1061, 1097 Cathodes 3017 Causticization 51 Cell Emissions 531 Cell Life 1047 Cell Stability 569 Central Indian Bauxite 29 Ceramic Foam Filters 763 CFD 1185 Chalco 509 Characteristics 1091 Characterization 345 Chemisorption 725 Chisel Bath Contact 467 Circulating Fluidized Bed 125,157 Clinker 97 Coal 185,913 Coal Ash 97 Coil 621,763 Coke 913,1067 Coke Specification 953 Cold Cracking 669 Cold Water Model 1155 Collector Bar 519 Collector Bar and Cathode Block Preheating 1017 Collector Bar As Resistance Heater 1017 Communication 871 Compensation 621 Composition 399 Comprehensive Energy Consumption 63 Compression Behavior 1161 Computational Fluid Dynamics 543 Computational Fluid Dynamics (CFD) 531 Computerized Tomography 973 Contact 1009 Contact Angle 775 Contact Pressure 985 Contact Resistance 985 Continuous PFC 309 Control 301 Cooling Rate 241 Cooling Way 227 Corrosion 1211 Corrosion Behavior 1085 Corrosion Resistance 1123,1129 CPC 917 Creep 625 Creep Deformation 1053 Critical ACD 567 Cryolite 513 Cryolite Melts 563 Crystallinity of Carbon 841 Cu-Ni-NiO-NiFe204 Composite Ceramic 1129 Current Density 461 Current Effciency 591 Current Efficiency 449,461 Cycle 647 Cyclone 769

D

DC DC Busbars DC Casting DC Consumption De-noising Decision Making Deformation Behavior Dendrite Arm Spacing Density Design Design of Experiment Desilication Desulfurization Diaspore Diaspore Bauxite Diastar Die Pick Up Differential Pulse Voltammetry Differential Thermal Analysis Diffusion Digesting Properties Digestion Studies Dilution Direct Chill Direct Chill Casting Discrete Simulation Dissolution Dissolution Kinetics Dry Hydrate Dry-scrubbing DSP Dual Chamber Furnace Dual Duct Durability Dusting Dynamic Simulator

E

693 525 669, 675, 687 509, 567 599 605 615 681 973 519 291 29 19 33 63 75 803 215 483 901 33 29 647 681 699 421 809 237 125 285 89 647 361 647 471,901 137

Economics 5 Electrical Conductivity 563 Electrical Contact Resistance 1003 Electrical Preheating 1041 Electricity Production 393 Electrochemical Wear 1073 Electroheating Balance 591 Electrolysis 387,483 Electrolysis Expansion 1117 Electrolytic Cell 997 Electromagnetic Field 1029 Electromagnetophoretic Force 549 Electroneutrality 537 Elemental Carbon 821 Emission 295 Emission Decrease of Equivalent Weight Carbon Dioxide 497 Emission Monitoring 269 Emissions 151, 185, 345, 1185 Energy 151, 171, 653, 1205 Energy Consumption 281, 393, 587, 859, 875, 1041 Energy Efficiency 125, 157, 179 Energy Models 375 Energy Price 375 Energy Saving 509, 569 Energy Saving At All Levels 591 Energy Saving Furnaces 1169 Energy Savings 1023 Energy-saving 1217 Environment 1091 Environmental Design Criteria 191

1228

Equipment Eutectic Modification Expansion Project Explosion Extraction

F

Fast Imaging FEM Ferro Silico Manganese Filtration Fines Finite Element Analysis Finite Element Method Finite Element Modeling Fire Operation Firing System Flatness Fleet Management Floatation-Bayer Process Flocculant Flocculants Flocculation Flue Wall Fluoride Emission Fluoride Emissions Fluoride Inmission Fluxing Forced Convection Formulation Fortin Bubble Fracture Furan Resin Furnace Furnace Efficiency Furnace Optimization

G

Garbage To Garbage Gas Analyzer Gas Collection Efficiency Gas Treatment Center Gibbsite Granule Granules Granulometry Graphite Graphitized Cathodes Green Coke Greenwood-Williamson GTC Guidelines Gypsum Gypsum-bonded Investment

H

Hall-Hιroult Hall-Hιroult Cell Hall-Hιroult Process Heat Capacity Heat Flow Heat Integration Heat Preservation Lining Heat Recovery Heat Recuperation Heat Release Rates Heat Transfer

793 815 57 657 197

725 1009 275 769 907 525 1041 387, 537 847 875 621,625 421 89 245 107 107 881 301 291, 351 315 647 809 1161 581 669 1117 635, 871, 881,1173, 1205, 1211 1169 853,1185

Heat Transfer Coefficient Heat Treatment Hematite Hertz HES HF HF Emissions HF Levels High Draft High Energy Efficiency Highspeed High Speed Photography High-Sulfur Bauxite High-Temperature Mixing Horizontal Single Belt Casting (HSBC) Human Factors Humid Atmospheres Hydro Cyclone Hydro-Process Hydrogarnet Hydrogen Fluoride Hydrolyzing Hygiene

I

Idle Power Consumption Image Analysis Improvement Improvement Process In-Line Treatment Incipient AE Logic Inclusions Industrial Ccale Measurements Inert Anode Inert Anodes Ingot Line Initial Stage Al Oxidation Innovation Cathode Structure Inserts Integrated Technology Intense Magnetic Field Interface Intermetallic Phases Interstage Coolers Investment Casting Ionic Equilibrium Iron Iron-Rich Phases ISO

647 269 361 163 211,285 1079 745 399 775 1067,1073 917 1009 163 301 97 705

J

Jarosite.

K

K2TiF6.. KPI 991 387,1041 151,575 1035 687 179 509 393 917 531 175, 381,1185

L

L-sub-c / Lc Crystallite Height Lab Scale Measurements Large Bubbles Laser Ablation Laser Scanning Layout Development Leaching

1229

387 835 255 1009 793 269,339 285 263 361 443,455 803 575 19 889 797 605 719 75 89 89 269 245 1091

497 503, 1103 621 477 737 333 745 555 1135,1149 1123 635 725 1029 569 1217 33 615 711 175 705 537 113 757 941

.255

.821 .693

841 555 581 725 1061 437 221

Leaching and Divalent Alkaline Earth Metal Oxide Lean Design Levers Lime-Bayer Process Limestone Liquid Layer Diffusion Lotus Leaf. Low Cell Voltage Low Temperature Aluminum Electrolyte LowToxicity Low-Median Grade

231 437 693 89 97 237 705 509 503 889 19

M

Macrosegregation 699 Magnesium Smelting 205 Magnetic Field 763 Magneto-hydrodynamic 569 Magnetohydrodynamics 549 Maintenance Cost 875 Management 301, 653 Marginal Bauxite 13 Material Balance 591 Mathematical Modeling 797, 1047 Ma'aden 23 Mechanical Properties 731 Mechanical Property 1135 Mechanical Stirring 145 Mechanical Vibration 827 Melting 1179 Membranes 81 Meniscus 763 Meniscus Behavior 797 Mercury 185 Metal Cleanliness 745 Metal Quality 731 Method Validation 315 MgO 201 Micro-casting 705 Microporosity 783 Microstructure 699, 827, 991 Microstructured Surface 705 Mineralogical and Chemical Characterization 13 Mixing 907 Modeling 519,641,835,865,1185 Modelling 635 Moisture 339 Molten Aluminum Stirring 1199 Monolithics 663 Mould-Wall 687 Multivariate Statistics 407

N

Nanofiltration Nanopowder NEUI400 NEUI500kA Family New Application Progess New Structure Cathode New Structure Cell Ni-Fe Superalloys NiFe 2 0 4 Non Homogeneous Novel Structure Cathodes Cells Noyes-Whitney Equation Nucleation Nucleation Site Distribution Numerical Modeling Numerical Simulation

Numerical Simulations

O

OEE Offline OPC Operation Operation Costs Operations Integrity Management Operator Training Optical Methods Optimization Organics Oxalate Oxidation Oxide Film Oxyfuel

P

P155 Technology Package Particle Breakage Particle Gradation Particle Shape Particle Size Particle Size Distribution Particulate Emissions Pencil Graphite Electrode Penetration Perfluorocarbon Periodically Attenuating Permanent Magnet Permeability Petroleum Coke PFC PFC Emission Phase Composition Physical Properties Pitch Pitch Burn Pitch Impregnation PIV Plant Management PLS Polyethylene Glycol Polymerized Aluminum-ferric Chloride Pore Distribution Pore Size Distribution Porosity Pot Control Pot Gas Treatment Pot Lifetime Pot Room Operation Potroom Potroom Electrical Safety Potroom Ventilation Power Cost Power Failure Power Generation Power Modulation Prebaked Anode Preheating Pressure Pressureless Liquid Phase Sintering Primary Aluminium Smelters Process Process and Operational Improvement Process Control

81 1149 443 455 443 1155 509 1123 1149 1009 1023 237 815 783 1179 145, 1029

1230

581

693 621 97 415 875 41 137 555 865 51 215 719 783 1185

467 477 125, 157 1149 937, 1161 1161 121 291 215 1053 325 121 1199 901 841,895,941,947 281,319,325 357 1135 1117 1117 875 1073 581 57 967 251 245 1103 285 925 491 295 1067 421 477 525 361, 531 591 405 917 369, 375 967 1047 131 1111 361 431 41 269, 407,425,605

Process Management Process Monitoring Process Safety Product Quality Productivity Program Project Development Protection

Q

Quenching Factor Analysis

R

Radiation Ramming Paste Ras AzZawr Recycle Recycling Red Mud Reduction Reduction Cells Reduction Slag Refinement Refractory Refractory Construction Refractory Performance Relative Density Remelting Residual Hydroxide Residue Resource Utilization Restart Reverberatory Furnace RMPCT Roasting Roasting Pretreatment Roll Bonding Roll Speed Ratio Rolling Rolling Ingot Rough Surface

S

Safety Salt Fluxing Sandy Alumina Saudi Arabia Schedule Optimization Scrap Sea Water Neutralization Sealing Seed Precipitation Tank Seeded Precipitation Segregation Self-Disintegrating Separation SGA Shaft Calciner Shale Shell Heat Exchanger Si-Al Alloys SiC Side Ledge Measurements Sideledge Sidewall Sidewall Heat Transfer Sidewall Lining Material

75, 587 407 41 125, 157 393,865 653 5 1079

835

687 1091 23 731 1173 81, 93, 97, 103, 107, 231, 237,245, 255 113 309, 357, 407 205 827 663, 1205, 1211 847 859 1129 731 285 113 19 405 1179 69 103 33 615 615 625 693 1009

Simulation Sintering Sintering Atmosphere Sintering Process SMART Centres Smelters Smelting Soderberg Sodium Aluminate Solution Sodium and Fluoride Penetration Sodium Carbonate Solid-liquid Phase Flow Solidification Solios Solvent Extraction SOP Spent Anode Spent Potlining Stability Start-up Statistical Analysis Status Breakup Steady State Steam Economy Steam Savings Stimulus Strip Casting Stub Deterioration Stub Diameter Stub Replacement Stub To Carbon Voltage Drop Super-Concentration Superheat Surface Defects

T

TAC Target Molar Ratio TBD Technology Development Temperature Temperature Measurement Temperature Varying Properties Thermal Conductivity Thermal Contact Conductance Thermal Diffusivity Thermo Gravimetry Thermo-Electro-Mechanical Thermo-Electro-Mechanical Simulation Thermocouple Thermomechanical Simulation Through Process Modelling TiB2 TiB2-C Cathodes Time Dependent Property Titanium Diboride (TiB2) Composite Tortuosity Traffic Transitions Metals Transport Transport Number Tri-Hydrate Alumina Tunable Diode Laser Spectroscopy Turbulent Flow Two-phase Flow

657,871 737 227 23 421 1173 81 881 145 227 675 241 647 163 1217 97 369 809 775 375 381,387 1085 369 1035

U

UBC.

1231

437,641 231 1135 89 449 309,357,431 113,325,491 907 121,251 1103 103 145 681,699,827 871,881 913 693 351 275 591 425 967 497 693 69 175 647 797 979 979 979 985 51 381 675

693 29 941 455 991 491 1035 399,1035 1003 1035 211 519 1041 991 669 625 745,1079 1117 775 1111 901 437 751 537 513 211 269 387 581

.1173

Ultrafine Powder Ultrasonic

V

Van Der Waals Forces VBD Vertical Drags Vertical Shaft Calciner. Vibrated Bulk Density Voltage Voltage Grade

W

Wall Induced Forces Waste Heat Recovery Waste Water Water Balance Wavelet Wear Wear Profiles Wettability Wettable Cathode Wetted Cathode Wifi Winding Wireless Work Practices Wrought Alloys

X

X-Ray Micro-Tomography XC Filter XPS XPS Imaging

Y

Yield Stress Yoke Bending

Z

Zeta Potential ZrO,

1143 ..221

47 931, 937, 941, 953, 959 675 917 925, 937, 953, 959 587 497

575 295 191 .....179 599 1211 1061 775 1111 509 871 625 871 301 711

783 769 1097 1097

47 979

47 1129