2nd International Symposium on High-Temperature Metallurgical Processing

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Author: Jiann-Yang J. Hwang | Jaroslaw Drelich | Jerome Downey | Tao Jiang | Mark Cooksey

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2nd International Symposium on

High-Temperature Metallurgical Processing

140th Annual Meeting & Exhibition 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 ProceedingsrVolume 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.

2nd International Symposium on

High-Temperature Metallurgical Processing Proceedings of a symposium sponsored by the Pyrometallurgy Committee of the Extraction and Processing Division and the Energy Committee of the Extraction and Processing Division and the Light Metals Division of TMS (The Minerals, Metals & Materials Society) Held during the TMS 2011 Annual Meeting & Exhibition San Diego, California, USA February 27-March 3, 2011 Edited by Jiann-Yang Hwang Jaroslaw Drelich Jerome Downey Tao Jiang Mark Cooksey

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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., 111 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 978-1-11802-938-1 Printed in the United States of America. 10 9 8 7 6 5 4 3 2 1

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A John Wiley & Sons, Inc., Publication

TIRÁIS

TABLE OF CONTENTS 2nd International Symposium on HighTemperature Metallurgical Processing Foreword Editors

xi xiii

2nd International Symposium on High-Temperature Metallurgical Processing Energy Efficient New Metal Production Technology Intrinsic Hydrogen Reduction Kinetics of Magnetite Concentrate Particles Relevant to a Novel Green Ironmaking Technology H. Wang, M. Choi, andH. Sohn

3

A Laboratory Investigation of the Reduction of the Iron Carbonate Bearing Ore to Iron Nugget by Means of the ITmk3 Technology 11 N. Panishev, R. Tahautdinov, A. Posazhennikov, and V. Bastrygin Behavior of Coal-Based Direct Reduction Reaction of Iron Oxide Pellets by Microwave Heating H. Zhu-cheng, W. Hua, H. Bing, P. Hit, andX. Guang-bin Carbothermal Reduction of Titanium Concentrate at High Temperature R. Huang, C. Bai, X. Lv, G. Qiu, andL. Lei

15 25

A Simulation Study on Flue Gas Circulating Sintering (FGCS) for Iron Ores...33 T. Jiang, Z. Fan, Y. Zhang, G. Li, andX. Fan Optimizations of Preparation for U 3 0 8 by Calcination from Ammonium Durante Using Response Surface Methodology B. Liu, J. Peng, and D. Huang

41

Microwave Field Attenuation Length and Half-Power Depth in Magnetic Materials Z. Peng, J. Hwang, X. Huang, M. Andriese, and W. Bell

51

v

Vanukov Furnace Technology: Application Experience for Processing Different Types of Raw Materials and General Development Trends 59 V. Bystrov, V. Paretsky, A. Vernigora, R. Kamkin, A. Mamaev, and A. Kuznetsov

Microwave Heating and Iron and Steel Production A Study of Coal-Based Direct Reduction of Composite Binder Magnetite Preheated Pellets D. Zhu, V. Mendes, T. Chun, J. Pan, andJ. Li Microwave Dielectric Properties of Pyrolyzed Carbon Z Peng, J. Hwang, W. Bell, M. Andriese, andS. Xie Fugitive Emissions Related to Oxidation of Liquid Silicon during Ladle Refining M. Ncess, G. Tranell, andN. Kamjjord Reduction Kinetics of Iron Oxide in CaO-Si02-Al203-FexO-C Mixtures Y. Zhang, and P. Masset Optimization of the Process Variables for Making Direct Reduced Iron by Microwave Heating Using Response Surface Methodology L. Dai, J. Peng, and H. Zhu

69 77

85 95

101

Study on Nucleation and Growth Mechanism of Iron Crystal Grain in CoalBased Shaft Furnace Direct Reduction Iron Pellets by Microwave Heating ....111 Z Huang, Z. Liao, B. Hu, L. Yi, and Y. Zhang Investigation on a Microwave High-Temperature Air Heat Exchanger J. Liu, Y. Li, L. Liu, J. Peng, L. Zhang, S. Guo, H Luo, H Wang, andG. Chen

119

Refractories, Slag and Recycling Study on Preparation of High-Quality Synthetic Rutile from Titanium Slag by Activation Roasting Followed by Acid Leaching 127 Y. Guo, S. Liu, T. Jiang, and G. Qiu Calculation of Phase Equilibria Relations in CaO-Si02-FeOx-MgO System ...137 N. Wang, C. Huang, X. Xin, Z. Zou, Z. Zhang, Y. Xiao, and Y. Yang

vi

Dissolution Behavior of Rhodium into Molten Slag C. Wiraseranee, T. Okabe, cmdK. Morita "One Step" Technology to Separate Copper, Zinc, Lead from Iron in Metallurgical Slag and Pyrite Cinder: Part 2 - Pilot Test D. Zhu, D. Chen, J. Pan, Y. Cui, and T. Chun

143

151

Effect of Oxygen to Alumina Ratio on the Viscosity of Aluminosihcate and Alumínate Systems 161 J. Xu, J. Zhang, C. Jie, F. Ruan, and K. Chou Blast Furnace Burdens Preparation from Metallurgical Dusts and Sludges with Composite Binder 169 K. Zhang, Y. Zhang, T. Jiang, G. Li, andZ. Huang Determination of FeO Containing Liquid Slag Surface Tensions Using the Sessile Drop Method C. Schmetterer, and P. Masset

177

Preparation of Partially Stabilized Zirconia and Interface Structure Analysis D. Li, S. Guo, L. Liu, J. Peng, L. Zhang, and C. He

185

Characteristic of Mineralization of Specularite Iron Ores during Composite Agglomeration Processing 191 H. Zhang, H. Yu, G. Li, Y. Zhang, Q. Li, and T. Jiang

Ferrous and Nonferrous Metals Enhancing the Pelletization of Brazilian Hematite by Adding Boron Bearing Additives 199 W. Yu, D. Zhu, T. Chun, andJ. Pan Study on Improving the Quality of Pellet Made from Vale Hematite Pellet Feed V. Mendes, D. Zhu, M. Emrich, J. Pan, and T. Chun

211

Decomposition and Oxidation of Bismuthinite in Nitrogen-Oxygen Atmospheres R. Padilla, R. Villa, andM. Ruiz

221

Pyrometallurgical Controls of Silver-Residue Smelting in a Short Rotary Furnace A. Nabei, and K. Yamaguchi

229

vu

A Study of Pelletization of Manganese Ore Fines D. Zhu, V. Mendes, T. Chun, andJ. Pan

237

Reduction of Carbon-Burdened Chromite Pellets in the Presence of Additives G. Li, J. Li, M. Rao, G. Bai, and T. Jiang

245

Production of Strontium Metal from Strontium Oxide Using Vacuum Aluminothermic Reduction Y. Dem iray, and O. Yücel

255

Treatment of Metals and Pellets Heats of Reaction in the Formation of TiB2 Reinforced Titanium Aluminide Composites 263 A. Baker, S. Kampe, and T. Zahrah Hot Workability of 1.2690 Ledeburitic Tool Steel and Development of Microstructure M Tercel), and G. Kugler

271

Effects of Binders on Oxidized Pellets Preparation from Vanadium/TitaniumBearing Magnetite 279 G. Han, Y. Zhang, Y. Huang, Z. Sun, G. Li, and T. Jiang Constituents and Porosity of Lead Concentrate Pellets Produced in the Trepce Plant 289 A. Haxhiaj, andJ. Drelich Oxidized Pellet Preparation from Refractory Specularite Concentrates Using Modified Humic Acid (MHA) Binders 299 G. Bai, D. Zhang, Y. Zhang, G. Han, andZ. Su

Raw Materials Processing An Innovative Process on Benefíciation of Superfine Low Grade Hematite Ore D. Zhu, Y. Xiao, T. Chun, andJ. Pan Calcination Behavior of Sivrihisar Latérite Ores of Turkey E. Keskinkilic, S. Pournaderi, A. Geveci, and Y. Topkay a

vm

309 319

Function of High Pressure Roll Grinding in Producing Magnetite Oxidized Pellets 327 Y. Guo, H. Hao, T. Jiang, andJ. Fan Improving the Pelletization of Fluxed Hematite Pellets by Hydrated Lime D. Zhu, W. Yu, T. Chun, andJ. Pan Research on the Ball Milling and Followed by Microwave Reduction of Panzhihua Low Grade Ilmenite Concentrate Y. Lei, Y. Li, J Peng, L. Zhang, S. Guo, and W. Li Study of Strengthen Pelletization of Nickel Latente J. Pan, X. Zhou, D. Zhu, and G. Zheng

335

345 355

Waste to Wealth: Production of Fe-Ni from Lateritic Ore/Chromite over Burden of Sukinda Deposits in Orissa, India 363 B. Bhoi, C. Mishra, and H. Mishra Mineralization Behavior of Fluxes during Iron Ore Sintering M. Gan, X. Fan, T. Jiang, Y. Wang, L. Hu, W. Li, Q. Wang, and Lu. Xie

371

Microwave Assisted Breakage of Metallic Sulfide Bearing Ore M. Andriese, J. Hwang, W. Bell, Z. Peng, A. Upadhyaya, andS. Borkar

379

Sintering and Synthesis Crystallization Behavior of Calcium Ferrite during Iron Ore Sintering 389 X. Fan, L. Hu, M. Gan, T. Jiang, W. Li, Q. Wang, L. Xie, andZ. Yu Enrichment Behavior of Phosphorous in CaO-Si02-FeOx-P205 Based Slag N. Wang, Y. Shen, Z. Tian, andM. Chen

397

Numeric Simulation of the Cooling Process of the Iron Ore Sinter J. Yin, X. Lv, C. Bai, and G. Qiu

403

Author Index

411

Subject Index

.415

ix

Foreword This volume collects selected papers presented at the 2nd International Symposium on High-Temperature Metallurgical Processing organized in conjunction with the 2011 TMS Annual Meeting in Sand Diego, CA, USA. As the title of symposium suggests it is on thermal processing of minerals, metals and materials and intends to promote physical and chemical transformations in the materials to enable recovery of valuable metals or produce products such as pure metals, intermediate compounds, alloys (including steel), or ceramics through various treatments. The symposium was open to participants from both industry and academia and focused on innovative high-temperature technologies including those based on non-traditional heating methods as well as their environmental aspects such as handling and treatment of emission gas and by-products. Since high-temperature processes require high energy input to sustain the temperature at which the processes take place, the symposium intends to address the needs for sustainable technologies with reduced energy consumption and reduced emission of pollutants. The symposium also welcomed contributions on thermodynamics and kinetics of chemical reactions and phase transformations that take place at elevated temperatures and papers on characterization of materials used or produced in high-temperature processing. Given the spread among numerous journals - not always easily accessible to many researchers - we decided to compile information on research actitivities in the area of metallurgy at elevated temeperature in an easily accessible source and this book is the result. The availability of focussed scientific information into a few accessible resources should be attractive and gratifying to many researchers. Over 250 authors have contributed to the symposium with a total of 73 presentations. After reviewing the submitted manuscripts, 49 papers were accepted for publication on this book. The book is divided into seven sections and each section has different focus. It includes: Energy Efficient New Metal Production Technology, Microwave Heating and Iron and Steel Production, Refractories, Slag and Recycling, Ferrous and Nonferrous Metals, Treatment of Metals and Pellets, Raw Materials Processing, and Sintering and Synthesis. We hope this book will serve as a reference for both new and current metallurgists, particularly those who are actively engaged in exploring innovative technologies and routes that lead to more energy efficient and environmentally sustainable solutions. To our knowledge, this is the first book exclusively dedicated to this important and burgeoning topic published in the 21 century. This book could not materialize without contributions from the authors of included papers, time and effort that reviewers dedicated to the manuscripts during the review process, and help received from the publisher. We thank them all! Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey December 2010 XI

Editors Jiann-Yang (Jim) Hwang is a Professor in the Department of Materials Science and Engineering and the Director of the Institute of Materials Processing at Michigan Technological University. He is also the Editor-in-Chief of the Journal of Minerals and Materials Characterization and Engineering. Dr. Hwang received his PhD degree from Purdue University in 1982. He has been a TMS member since 1985. His research interests include the characterization and processing of materials and their applications. He has been actively involved in the areas of separation technologies, pyrometallurgy, microwaves, hydrogen storages, ceramics, recycling, water treatment, environmental protection, biomaterials, and energy and fuels. He has more than 20 patents, published many papers, and founded several companies. He has chaired the Materials Characterization committee and the Pyrometallurgy committee in TMS and has organized several symposiums. Jaroslaw W. Drelich received his BS degree in chemistry and his MS degree in chemical technology from the Technical University of Gdansk (Poland) in 1983. Jaroslaw earned his Ph.D. degree in metallurgical engineering from the University of Utah in 1993. Dr. Drelich then worked at the University of Utah for four years as postdoctoral fellow and research assistant professor. During that time his work concentrated on fundamental and advanced concepts of surface chemistry that he applied to mineral processing, materials recycling and processing, oily soil remediation, and oil sands processing. Dr. Drelich came to Michigan Technological University (Michigan Tech) in 1997 and works there as associate professor. His main research interests are in applied surface chemistry and interfacial engineering for ore dressing and materials processing, materials recycling, characterization of materials surfaces, modification and testing of biomaterials, and formulation of antimicrobial materials. Some of Dr. Drelich's current research activities include formulation xiii

of superhydrophobic surfaces and coatings, measurements and modeling of colloidal forces between heterogeneous surfaces, biodégradation of implant materials, processing of minerals, synthesis of novel antimicrobial materials, and rehydroxylation studies on ceramic artifacts. Aside from teaching several courses at Michigan Tech and advising capstone senior design projects, Dr. Drelich has published over 130 technical papers, holds 8 patents and has more than 40 conference presentations to his credit. Prior to this symposium he has organized or co-organized three other international symposia as part of the annual ACS, SME and TMS meetings. Dr. Drelich also serves on the External Advisory Board for the Journal of Adhesion Science and Technology and edited several special issues of this journal containing invited papers on atomic force microscopy, adhesion force measurements, and wetting phenomena. Additionally, he organized and chaired several sessions on materials recycling, applied surface chemistry, wetting phenomena and application of atomic force microscopy during international symposia and domestic TMS and SME meetings. Dr. Drelich is the active member of the TMS EPD Materials Characterization and TMS EPD Pyrometallurgy Committees. Currently, he also serves as co-Chair of the TMS Energy Committee. Jerome P. (Jerry) Downey is an Associate Professor and Goldcorp Distinguished Professor of Metallurgical and Materials Engineering at Montana Tech of the University of Montana. Dr. Downey is a registered professional engineer with a doctorate degree in Metallurgical and Materials Engineering from Colorado School of Mines. His experience encompasses industrial operations, applied process research and development, and corporate management. He has specific technical expertise in chemical and metallurgical thermodynamics, thermal processing, materials synthesis and processing, and hazardous materials treatment. Dr. Downey's current research activities focus on the study of fundamental properties of slags, molten salts, and glasses; synthesis of ceramic materials for energy applications; and remediation of acid rock drainage. Prior to accepting his Montana Tech professorship in 2006, Dr. Downey was Vice President at Hazen Research, Inc., where he directed the Thermal Process Department. Dr. Downey has been a member of TMS since 1977; he is the current Chair of the Pyrometallurgy Committee, a co-organizer for the 2nd International Symposium on High-Temperature Metallurgical Processing (2011 TMS Annual Meeting), and a co-organizer for the International Smelting Technology Symposium (2012 TMS Annual Meeting). xiv

Tao Jiang received his BS degree in 1983 and MS degree in 1986 in metallurgical engineering from Central South University of Technology, China. He earned his Ph.D. degree in mineral processing engineering in 1990 from the same university. Then Dr. Jiang worked at the university for ten years as assistant professor and full professor (from 1992). During that time his research concentrated on the fundamentals and new technologies for sintering, pelletizing and direct reduction of iron ores, as well as the extraction of gold ores. Dr. Jiang was a Visiting Scholar to the Department of Metallurgical Engineering, the University of Utah in 2000 and worked there with Dr. Jan D. Miller. During this stay of three years, his research interests expanded to treatment of industrial wastewater, interfacial fundamental in agglomeration of ore fine, catalytical and non-cyanide leaching of gold ores. Dr. Jiang returned to China in 2003 and has worked there as a Professor in the School of Minerals Processing & Bioengineering, Central South University. Some of his current research activities include beneficiation, agglomeration, reduction and utilization of complex iron ores, and treatment of refractory gold ores. He has accomplished more than 30 projects from the government and industry, including National Science Fund for Distinguished Young Scholars program, sub-project of National Basic Research Program of China (973 Program). Dr. Jiang has published over 200 technical papers, 5 books, holds 22 patents and has more than 30 conference presentations. He has won one second-class national invention prize for the research and development of the direct reduction process of composite binder pellets, and 8 items of science and technology prizes of provincial and/or ministerial level. Currently, Dr. Jiang serves as Specially-appointed Professor of Chang Jiang Scholars Program of China and the Dean of the school. He also serves as viceChair of the TMS Prometallurgy Committee, member of Ironmaking Committee, Chinese Society for Metals.

xv

Mark Cooksey is the Research Group Leader of Process Engineering at CSIRO, where he leads a team of 35 to develop and deliver science and engineering expertise for the minerals and related process industries. He has a background in aluminium production, and has previously held research and development engineering positions at Rio Tinto Alean and General Electric. He still leads some aluminium research projects at CSIRO. Mark has Bachelor Degrees in Materials Engineering, Information Technology and Applied Mathematics from the University of Western Australia, and is undertaking a PhD in Chemical and Materials Engineering at the University of Auckland. He was a TMS Young Leader in 2007.

xvi

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

2nd International Symposium on

High-Temperature Metallurgical Processing

Energy Efficient New Metal Production Technology Session Chairs: Jiann-Vang Hwang Anton Vernigora

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

INTRINSIC HYDROGEN REDUCTION KINETICS OF MAGNETITE CONCENTRATE PARTICLES RELEVANT TO A NOVEL GREEN IRONMAKING TECHNOLOGY Haitao Wang, Moo Eob Choi, and H. Y. Sohn Department of Metallurgical Engineering, University of Utah, Salt Lake City, Utah 84112 Keywords: hydrogen reduction, magnetite concentrate, ironmaking, intrinsic kinetics, nucleation and growth kinetics Abstract A novel ironmaking technology is under development at the University of Utah. This technology produces iron directly from iron ore concentrate by gas-solid suspension reduction. Hydrogen is the main reducing agent for high reactivity and for the elimination of carbon dioxide release in this ironmaking process. The direct use of concentrates allows bypassing the problematic pelletization/sintering and cokemaking steps in the steel industry. Intrinsic kinetics of the suspension reduction of magnetite concentrate particles by hydrogen has been measured. Experiments were carried out in the temperature range of 1423-1673 K, other experimental variable being hydrogen partial pressure, the amount of excess hydrogen and particle size. Under most experimental conditions, 95% reduction was attained within several seconds, which presents sufficiently rapid kinetics for a suspension reduction process. The reaction kinetics followed the nucleation-and-growth equation, and a rate equation that contains all the effects of the experimental parameters has been obtained. Introduction With the rapid increase in steel production, it is critical to control CO2 emissions which is the main greenhouse gas. It is inevitable to release a huge amount of CO2 when coke and/or coal are used as the main reducing agent in producing iron. As one of the methods to reduce the C 0 2 emission from ironmaking, the utilization of the gaseous reducing agents and fuels, including hydrogen, has been investigated [1-6]. The use of such gases forms the basis for a new ironmaking technology under development for producing iron directly from fine concentrates by a gas-solid suspension technology. This technology would reduce energy consumption by about

3

40% of the amount required by the blast furnace and drastically lower environmental pollution, especially CO2 emission, from the steel industry [4]. This is accomplished by adopting H2-based reductant and bypassing pelletization/sintering and cokemaking steps. Unlike other gas-based alternative ironmaking processes using shaft furnaces or fluidized-bed reactors, the suspension reduction technology is a high-intensity process because it will not suffer from the problems that other processes do when operated at high temperatures mainly from the sticking and fusion of particles. Our previous work on the kinetics feasibility tests showed that iron oxide concentrate can be reduced to a high metallization degree within a few seconds of residence time typically available in a suspension process. Previous related work and the reasons for expectation of sufficient reduction rate have been described elsewhere [5-6]. Further investigation aimed at systematic measurement of the intrinsic reduction kinetics of fine iron oxide particles, including the effects of temperature, partial pressures of hydrogen and water vapor, and particle size, has been conducted in this laboratory. Experimental Raw Material Characterization The magnetite concentrate used in this work was provided by ArcelorMittal. The concentrate particles are irregularly shaped and angular. The particle shapes of the samples of different sizes were examined by taking SEM micrographs and shown in Figure 1. The inner cross-section microstructure of the samples was examined by an optical microscope. The microphotos were taken by an on-site camera and shown in Figure 2. It is observed that there are differences in porosity among the samples with three different sizes.

a. 20-25 um

b. 32-38 um Figure 1. SEM micrographs of unreduced samples

4

c. 45-53 urn

a. 20-25 um b. 32-38 \¡m c. 45-S3 um Figure 2. Inner cross-section micrographs of polished unreduced samples obtained by an optical microscope. 500 X

For accurate determination of the rate of individual iron oxide concentrate particles, a high temperature drop-tube reactor system was fabricated [2]. The furnace system was made up of a vertical split tube furnace with a maximum working temperature of 1823 K and a cylindrical alumina tube (5.6 cm ID, 193 cm long). The concentrate particles were fed through a tube of 0.16 cm ID carried by a mixture of H2 and N2 gases with a total flow rate of 200 NmL/min. The duration of reduction of the concentrate particles was determined by the residence time of particles in the isothermal zone. Results and Discussion Reaction Mechanism Analysis The oxygen and total iron contents in the ArcelorMittal concentrate yielded a removable O/Fe ratio of 1.30, indicating that essentially all the iron oxide existed as Fe3Ü4. The amount of removable oxygen was determined by the weight loss in the expriment of concentrate reduction with a large amount of H2 in the horizontal furnace. The fact that all iron oxide in the concentrate exists essentially as Fe3Û4 was further confirmed by the XRD pattern [6]. The nucleation and growth kinetics expression given by the following equation was found to describe the rates of reduction in this study: (1) where X = fractional reduction degree, I = reaction time, n = Avrami parameter, E = activation energy, T= reduction temperature, dp = particle size, ko = constant, gas partial pressures,^ (dp) = function of particle size.

5

= function of

In this study, the gas mixture of hydrogen and nitrogen was used, which formed water vapor as a result of the reduction. Its amount changes during the course of the reduction, which reduces the reduction potential of hydrogen. We let (2) When the particle size and reduction temperature are fixed, the relationship between k and should be linear. The following expression for pressure dependence is used for iron-oxide reduction, the last step (FeO + H2 = Fe + H2O) of which is equilibrium-limited: (3) where K is the equilibrium constant for the reduction of FeO. At different initial partial pressures of H2, a set of experiments were conducted with the sample of 20-25 um at 1573 K. The results are presented in Figure 3, according to Equation (1) with n = 2, which was previously determined to be the best-fit value of n.

Figure 3. The relationship between

with

20-25 um samples at 1573 K. The results in Figure 3 indicate that m = 1/2 gives the best fit of the exprimental data. This is consistent wim the reaction mechanism in which adsorbed H2 molecules dissociate into H atoms before reacting with the oxide. Intrinsic Kinetics Measurement In our previous work [6], the kinetics feasibility tests showed that the reduction rate was fast enough to obtain 90-99% reduction within 1-7 seconds at 1473-1673 K, depending on the

6

amount of excess hydrogen supplied with iron oxide. Since then, further systematic experiments were carried out in the temperature range of 1423 and 1673 K. with various other parameters including the amount of excess hydrogen, partial pressure of hydrogen, and particle size. The experimental conditions were adapted to get moderate reduction degrees scattered evenly along the whole curve of the plot between reduction degree and residence time. The residence time is changed by the variable total flowrates according to the feeding rate change of iron oxide. Since the amount of formed water vapor varies with the degree of reduction, the value of fp(p¡¡ ,p¡,0) is not constant. For m = 1/2, the arithmatic average of fp(jp,¡_,pH 0 ) is the appropriate average driving force when gaseous species concentration varies [7], and was thus used for further rate analysis. The experimental results obtained at different various parameters were plotted according to Equation 1, as shown in Figure 4.

Figure 4. Relationship between

and residence time with 200% excess hydrogen.

The results show that reduction rate increased with the particle size at a lower reduction temperature while the effect of particle size became less significant when temperature was higher than 1573 K. It could be explained by the following reasons: "Gas-solid reactions usually involve the adsorption of gaseous reactants at preferred sites on the solid surface and the

7

formation of neclei of the solid product. For small particles the period of the formation and growth of nuclei occupies essentially the entire conversion range" [8]. But the number of preferential nuclei forming sites per unit area of the surface depend on the properties of the exposed surface of the single particle including scratches, lattice defects, and the presence of the impurities [9]. The iron oxide concentrates with different particle sizes were obtained by grinding before it was concentrated by various mineral processes, which could possibly make the particles have different properties of the surface. And also the gangue content is different among them. What is more, the cracks tend to be formed in the big particles during the reduction more readily than in the samller particles, and the dimension of solid between cracks can be smaller than the size of the smaller particles. All of these are possible reasons why big particles have higher reduction rate at a lower temperature. Determination of Particle Size Function As described above, samples with different particle sizes reacted at different rates. The rate dependence on the particle size was assumed to be (dpf. In order to determine the s value, the reaction rate constant k is rearranged as follows:

Therefore, in those cases at different reduction temperatures, the s values are obtained from the slopes of the plots of

as shown in Figure 5.

The relationship between s and temperature obtained from the results is shown in Figure 6. The activation energy is obtained from the plot of fin ko - E/RT) and 104/T, as shown in Figure 7. The slope in Figure 7 leads to the value of E/R = 55790, which yields an activation energy value of 464 kj/mol in the termperature range between 1423 and 1673 K. The complete rate equation is now expressed as follows: (6)

(7)

8

c. 1573 K

d. 1673 K

Figure 5. Relationship between particle size, dp and kB ■ e"^" 1 " • /¿(d,,). (dp in um)

Figure 6. Relationship between

Figure 7. Relationship between (Ln ko - E/RT) and 104/T

Í and reduction temperature

Conclusion A novel technology for producing iron directly from fine concentrates by a gas-solid suspension reduction is under development at the University of Utah. Intrinsic kinetics of the suspension reduction of magnetite concentrate particles by hydrogen has been investigated. The reaction rate

9

follows the nucleation-and-growth kinetics. The obtained activation energy is 464 kj/mol. The reduction rate has half-order dependence with respect of H2. Samples with larger sizes have higher reduction rates when reduction temperature is in the range of 1423-1673 K. Particle size has a negligible effect on reduction rate above 1573 K. A rate equation that contains all the effects of the experimental parameters has been formulated. References 1. H.Y. Sohn, "Suspension Ironmaking Technology with Greatly Reduced Energy Requirement and CO2 Emission," Steel Times International, May/June (2007), 68-72. 2. M.E. Choi and H.Y. Sohn, "Development of a Green Suspension Ironmaking Technology Based on Hydrogen Reduction of Iron Oxide Concentrate: Rate Measurements," Ironmaking and Steelmaking, 37(2) (2010), 81-88. 3. H.Y. Sohn, M.E. Choi, Y. Zhang, and J.E. Ramos, "Suspension Reduction Technology for Ironmaking with Low C02 Emission and Energy Requirement," Iron & Steel Technology (AIST Trans.), 6 (6) (2009), 158-165. 4. H.Y. Sohn, "Suspension Hydrogen Reduction for Iron Ore Concentrate" (Final Report TRP 9953- NonPropFinalReport, U. S. Department of Energy, American Iron and Steel Institution, 31 March 2008) 5. H.Y. Sohn, M.E. Choi, Y. Zhang and J.E. Ramos, "Suspension Reduction Technology for Ironmaking with Low C02 Emission and Energy Requirement," AISTech 2009 Proceedings, Vol. /(Warrendale, PA: AIST, 2009) 187-199 6. M.E. Choi, "Suspension Hydrogen Reduction of Iron Ore Concentrate" (Ph.D. Dissertation, University of Utah, 2010) 7. H.Y. Sohn, unpublished work, University of Utah, 28 September 2010. 8. S. Seetharaman and H.Y. Sohn, "Chapter 7: The Kinetics of Metallurgical Reactions," Fundamentals of Metallurgy, ed. S. Seetharaman (Cambridge, ENGLAND: Woodhead Publishing Limited, 2005), 299-210. 9. Y.K. Rao, "Mechanism and Intrinsic Rates of Reduction of Metallic Oxides," Metallurgical Transactions B, 10B (June) (1979), 243-255.

10

2nd International Symposium on High-Temperature Metallurgical Processing Ediled by: Jiann- Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cooks TMS (The Minerals, Metals & Materials Society), 2011

A LABORATORY INVESTIGATION OF THE REDUCTION OF THE IRON CARBONATE BEARING ORE TO IRON NUGGET BY MEANS OF THE ITmk3 TECHNOLOGY Panishev N.V., Tahautdinov R.S., Posazhennikov A.N., Bastrygin V.V. Magnitogorsk Iron & Steel Works (MMK) Open Joint Stock Company 93 Kirova St, Magnitogorsk 455002, Russia Siderite, Rotary Hearth Furnace, ITmk3, Iron Nugget Abstract The Bakal (South Ural, Russia) deposit of iron ore bearing iron carbonate (siderite) with the capacity of more than 1 billion tones belongs to the MMK. This ore cannot be fully processed via blast furnace technology because of high content of MgO. According to the investigations carried out in the USA and Japan in 1999-2004 the ITmk3 (Ironmaking Technology mark three) RHF technology is a breakthrough in Ironmaking. Four iron ore types (hematite, magnetite, high and low AI2O3/SÍQ2) were tested. Reduction, melting and slag The main objective of the investigation is to removal can be achieved in just 10 min. establish optimum operation conditions for the production of iron nuggets from iron carbonate bearing ore via the ITmk3 by means of the lab scale testing. Samples of iron ore and carbon reductant were chemically and physically analyzed. Green pellets were processed via a lab tube furnace to simulate RHF conditions. The nuggets and slag were chemically analyzed. This preliminary test work provides valuable information which may be used for large-scale testing in a commercially sized RHF. Introduction The ITmk3 process upon which this paper focuses was developed by Kobe Steel in its research facilities in 1996 [1]. After pilot testing at Kobe Steel's Kakogava Works in 19992004 the pilot demonstration plant was built and operations were successfully carried out by Mesabi Nugget joint venture in 2002-2004. Steel Dynamics has committed to building the commercial plant at the Mesabi Nugget site in Hoyt Lakes (Minnesota). Only 4 iron ore types (hematite, magnetite, high AI2O3/SÍO2 and low AI2O3/SÍO2), as well as 4 coal types with range of volatility, ash and fixed carbon were tested. Results were consistent with nugget characteristics confirming ITmk3'sflexibilitywith respect to raw material inputs. At the same time, no research in the field of nugget from iron carbonate bearing ore with high content of MgO has been done. Such kind of iron ore is restricted for charging into the blast furnace due to contamination of MgO. The reason for restriction is the limitation of load of MgO in the blast furnace. If a lot of MgO enters the blast furnace, slag becomes viscous. The viscous slag is obstacle for stable operation of the blast furnace.

11

Objectives of Investigation The main objective is to find optimum operation conditions for the production of iron nuggets from the iron carbonate (siderite) bearing ores via the rotary hearth furnace by means of the laboratory testing. Methodology The first step in the investigation involved performing lab scale testing on the iron nugget components. Samples of iron ore, fluxes, binder (clay), coal fines and coke breeze (as reductant) were chemically and physically analyzed. Typical ore testing included particle size analysis, % S, % C0 2 , % metal components, % gangue components. Typical reductant testing included particle size analysis, % carbon, % sulfur, % volatiles, % ash and ash analysis for % metal and gangue components. A pellet blend developed from this testing, defining carbon reductant addition, binder, particle size. A blend was mixed followed by rolling green pellets. Green pellets were processed via a lab tube furnace or a chamber furnace. Using such kind of furnaces to simulate RHF conditions, was it possible to vary several parameters such as furnace retention time (10-15 min), temperature (1350-1450°C). After the allotted time in the hot furnace, the nuggets and slag were quenched, then analyzed for % C, % S, % Fe in iron nuggets, and % gangue components in slag. Results and Discussion Each raw material is different in distribution of the particle size and chemical compositions (See Table 1). Table l.Dry Chemical Composition of Raw Materials, wt. % Material

Fe

CaO

Si02

MgO

A1203

C0 2

C

CaF2

Bakal Iron Ore

34.0

2.2

3.2

9.8

1.2

29.0

-

-

Coke Breeze

0.9

0.7

8.0

0.3

3.1

-

87.0

-

Quartzite

0.5

0.4

95.6

-

0.6

-

-

-

Clay

1.8

0.4

51.0

0.8

35.3

-

-

-

-

1.0

36.8

0.5

0.6

-

-

52.5

Fluor-spar

Mainly considered are raw material preparation (weighing, blending, grinding), pelletizing, heating, reacting and melting control. In particular, material preparation is very important, since influence of the properties of raw materials on operational results (especially

12

the temperature of melting of gangue) is large. Thus a precise preparation work for raw materials is needed for desirable and stable operation. Raw materials were blended in a disc grinder at a predetermined mixing ratio to reach desirable temperature (1300-1400°C) of melting of gangue [2]. Particle size of-200 meshes for raw materials usually results in higher iron metallization. More intimate contact between carbon and iron oxide speeds reaction. Ability to pelletize determines maximum particle size. The mixing ratio was determined to give good conditions of palletizing and the reactions in the furnace. The ratio of fix carbon to iron oxides and other properties of raw materials were analyzed. Then the mixing ratio was determined (binder - 1-2%, coke breeze -18-20%, fluxes -12-18%, iron ore -60-69%) with this information. The mixture was fed on the palletizing disc, and made into green pellets. Diameter of green pellet is less than 20 millimeters. If the pellet is too large, heat transfer rate is too slow inside the pellet. Green pellets were dried in the chamber furnace (150-300°C). The dried pellets were fed in the roasting furnace. To reach conditions for better reactions and heat exchange entire pellet bed must be heated to 13501450°C for 40-50% of the pellet residence time. Nuggets contained 96-97% of iron and 2-3% of carbon. Slag composition is shown below (See Table 2). Table 2. Chemical Composition of Slag, wt.% CaO

MgO

A1203

Si0 2

5.0-8.5

20.0-25.0

5.0-7.5

60.0-65.0

Carbonates (siderite, magnesite, dolomite, etc.) in the pellets dissociate when temperature of the pellets is above 600°C. FeC03 = FeO + C0 2 MgC03 = MgO + C0 2 CaC03 = CaO+C0 2

(1) (2) (3)

Carbon and iron oxide react when temperature of the pellets is above 1100°C. C + C0 2 = 2CO FeO + CO = Fe + C0 2 FeO + C = Fe + CO

(4) (5) (6)

Formation of pig iron and melting of nuggets and slag from gangue take place when temperature of the pellets is above 1300°C. 3Fe + C = Fe3C 3Fe + 2CO = Fe3C + CO

(7) (8)

Thus, the RFH should have heating, reduction and melting zones. Temperature in each zone should be controlled to get a good performance that is required in each zone. In heating

13

zone, heat transfer for heating up the pellets and dissociation of carbonates should be taken into account. Oxidation of the gas is high here, because no reduction reaction takes place. In reduction zone, temperature and atmosphere of the gas should be controlled to get a high reaction rate. The reduction reactions in this zone are very fast. At last, in melting zone, the conditions should be controlled to promote the reduction reactions, formation of pig iron (iron nugget) and melting of pig iron and slag. Conclusions The possibility of production of iron nugget from iron carbonate bearing ore with high content of MgO via ITmk3 process has been established. The initial promise shown by the ITmk3 process in laboratory tests may be used for large-scale testing in a commercially sized RHF. ITmk3 technology is a simple process with a single-step furnace operation. Reduction, melting and slag separation completes within 10-15 minutes. Process temperature is 13501450°C. ITmk3 technology makes no harmful impact on environment since the process does not require coking and sintering plants. References 1. J.A. Hansen, "Mesabi Nugget-The New Age of Iron", AISTTech 2004 Proceedings. 2004, vol.1, 545-550. 2. N.L. Zhilo, Formation and Characteristics of Blast Furnace Slag (Moscow: Metallurgy, 1974), 120.

14

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooks TMS (The Minerals, Metals & Materials Society), 2011

BEHAVIOR OF COAL-BASED DIRECT REDUCTION REACTION OF IRON OXIDE PELLETS BY MICROWAVE HEATING

HUANG Zhu-cheng1, WANG Hua1, HU Bing1, PENG Hu2, XIA Guang-bin2 'School of Minerals Processing and Bioengineering, Central South University, Changsha, 410083, China 2 Changsha SYNO-THERM Co., Ltd; Changxing Road; Changsha, 410013, China Key words : Microwave heating, Coal-based direct reduction, Temperature rising characteristics, End point temperature, Porosity Abstract The temperature rising characteristics and reaction of the reduction of iron oxide pellets with anthracite fines were studied using a MW-L0316V microwave oven and Leica-DM-RXP polarizing microscope. The results show that iron oxide pellets and anthracite fines have good microwave absorbing properties, and high temperature and reducing atmosphere in a relatively short time for reduction of iron oxide pellets can be achieved. The reduction reaction, which follows the unreacted core model, is rapid and the metallization increases from 13.41% to 56.57% as the reduction end point temperature increases from 850°C to 950°C. The iron is formed, fine grain crystals of iron are transferred and the contraction of metal phase occurs firstly at the surface of pellets, which hinders the reduction reaction at the center of pellets. The rate of reduction reaction increases slowly as the reduction end point temperature increases from 950°C to 1050°C. Introduction Iron ore direct reduction can be divided into gas-based reduction and coal-based reduction. Owing to the effect of the resource and the related economic benefits [1, 2, 3, and 4], the development of gas-based direct reduction technology is restricted, but the development of coal-based direct reduction supported by the relevant industrial policy has very good prospects for application. As a source of clean energy, in China microwaves have been applied in many fields [5, 6, and 7]. Chen Jin studied the reduction of carbon pellets under microwaves, and has made some progress on reduction kinetics and reaction mechanisms. Other countries also have related research in this field [8 -15], for example, Kotaro Ishizaki and Kazuhiro Nagata have done research on the reduction reaction of carbon magnetite heated by microwaves. The research shows that Fe304 is deoxidized to FeO at about 800°C, and FeO is deoxidized to Fe at 1000°C-1025°C. Due to the very good heat absorbing properties of iron oxide pellets and anthracite fines, the selective and rapid heating of microwaves can rapidly raise the temperature of pelleta and anthracite, and the iron oxide pellets can be rapidly reduced. In

15

order to provide enough theoretical basis for direct reduction by microwave heating, this paper studies the temperature rising characteristics, behavior of the reduction reaction and the changes of the iron oxide pellets and anthracite in the microwave field. Experimental Basic Principles Materials Warming Theory In the microwave field, microwave energy absorbed by per unit time and per unit volume of materials is [16]: Q = 2 n f 8 0 s".E 2 = 2 71 f 808'tgô.E2

( 1 )

Where, £0—the permittivity of the material without electric field; f—microwave frequency, 2.45GHz;s"—the imaginary part of the complex permittivity, which shows the effective loss factor in the microwave; s'—the real part of the complex permittivity, equal to the dielectric constant of the material, reflects the ability of shackle charge; tg8—material dissipation factor; E—the electric field strength in the microwave heating cavity. The heating rate of the heat energy produced by the material which has absorbed microwave [16]. ^ _ r _„K,22 A _TI

AT

2nfe0s"E

At

pC p

(2)

Where, T. , p. , Cp. and t respectively represent the temperature, density, constant pressure mass specific heat and heating time. Coal-based Direct Reduction Reaction Mechanism The reduction of iron oxides proceeds step by step. When the temperature is below 570°C, because of the thermal instability of FeO, it can generate Fe and Fe3Û4 4 FeO —> Fe + Fe3Û4 So the reaction of anthracite to deoxidize iron oxide pellets can be divided into two steps: Fe2Û3 —* Fe304 —* Fe. and the reaction formulas are as follows: C + 3Fe203 -* 2Fe304 + CO 4CO + Fe304 -» 3Fe +4 C0 2

AA

0

=120000-218.46T J/mol

(3)

ArGm" = -39328+34.32T J/mol

(4)

When the temperature is above 570°C, the reaction of anthracite to deoxidize iron oxide pellets can be divided into three steps: Fe 2 0 3 —» Fe 3 0 4 —<• FeO —» Fe, and the reaction formulas are as follows: C + 3 Fe203 -► 2Fe304 + CO CO + Fe304 — 3FeO + C0 2

A TGm" = 120000-218.46T J/mol ö

A,Gm =35380-40.16T J/mol

(3) ( 5)

CO + FeO -> Fe + C0 2 ArGm^ =-18150+21.29 T J/mol (6) When the temperature rises sufficiently, carbon gasification will occur (the Boudouard Reaction).

16

C + C0 2 -»2C0 ArGm"- =166550-171T J/mol ( 7) In the process of reduction, the energy needed by the reaction in the microwave oven is supplied by the microwave energy absorbed by iron oxide pellets and anthracite. Different reduction end point temperatures may lead to different gas compositions in the heating cavity of the microwave oven, and the different atmosphere can directly affect the process of the reduction reaction of the pellets. Experiments Experimental Materials The iron oxide pellets were supplied from a ball mill. Its main chemical components and size composition are showed in Table 1 and Table 2. The reducing agent used in the experiment is a kind of anthracite taken from Hunan province. Its industrial analysis and size composition are showed in Table 3 and Table 4. Table 1.Chemical Analysis of Iron Cxide Pellets. Chemical Components /% TFe

FeO

Fe203

Si02

A1203

MgO

CaO

CuO

V2Os

64.24

0.24

91.21

5.31

1.55

0.61

0.56

0.05

0.02

+16mm 5.24 Fcad 81.14 +5mm 17.94

Table 2.Size Composition of Iron Oxide Pellets. Size Composition /% bulk density /kg/m3 -5mm -10~+5mm -16+ 10mm 94.32 0.44 0 1774.42 Table 3. Industrial Analysis of Anthracite /%. Vad Aad Mad 4.96 10.47 3.43 Table 4. Size Composition of Anthracite /%. -5 - +3 mm -3 - +0.5mm -0.5+0.18mm 19.28 39.72 1.6

Compressive strength /N/a 2973 Porosity /% 10.22 Total 100 -0.18mm 21.23

Experimental Methods And Equipment In this materials warming experiment, 100g of anthracite and iron oxide pellets were separately heated by the microwave oven at a power of 1300W. In the reduction experiment, 10-16mm pellets and 0.18-5mm anthracite were prepared, and according to the Fcad of anthracite and Fe content of iron oxide pellets, and from the formula: C +

17

Fe2C>3= Fe + CO, it can be concluded that, theoretically, 31.38g of anthracite reacts fully with 120g of iron oxide pellets. Considering the inevitable loss occurring during the experiment, 43.93g of anthracite was used, that is to say 1.4 times more than the theoretical amount.

Fig.l Schematic diagram of microwave heating cavity.

Fig.2 Schematic diagram of microwave oven.

The prepared materials were placed into the mullite crucible, and then the crucible was placed into the heating cavity shown in Fig.l. The crucible and cavity were filled with insulator so as not to lose heat. The heating cavity was placed into the MW L0316V microwave oven as shown in Fig.2, and nitrogen was supplied as the protective gas. When the temperature reached target, the microwave oven was switched off, die crucible was removed, and nitrogen was used to cool the materials. Results And Discussion Temperature Rising Characteristics Of Iron Oxide Pellets And Anthracite The microwave generator was switched off when the temperature reached 1080°C, producing the heating curve shown in Fig.3. It can be concluded that the heating time of both iron oxide pellets and anthracite is less than one hour, so they both have good temperature rising characteristics. Also, the anthracite takes a shorter time to reach a high temperature when compared with the iron oxide pellets. From formulas (1) and (2), the rate of temperature increase of iron oxide pellets is obviously lower than anthracite of the same quality, which is caused by their different electromagnetic properties.

18

1200 1000

ü

800

'S =3

600

Q- 400

E

200

0

0

10

20

30

40

50

60

time/min

Fig.3 Heating curves of iron oxide pellets and anthracite heated by microwave. It can be also concluded from Fig.3 that the heating process consists of two stages: a rapid heating stage and a slow heating stage, which we can call the first stage and the second stage. In the first stage, because pellets and anthracitic contain some water, and water is a type of polar molecule with high permittivity, and the more water material contains, the more heat a material absorbs per unit time, the rate of temperature increase is high. As the water content in the material reduces because of evaporation, the rate of temperature increase in the second stage sharply decreases, which makes the whole process slower. Table 5 shows the rate of temperature increase of iron oxide pellets and anthracite under same qualities and power. Table 5.The Rate of Rise of Temperature of Various Stages of Iron Oxide Pellets and Anthracite Heated by Microwave. Stage 2 rate Average rate Stage 1 rate Material /°C/min /°C/min /°C/min /100g 25.32 11.11 16.6 Iron oxide pellets 43.19 10.74 22.81 Anthracite Influence Of End Point Temperature On The Metallization Of Reduced Pellets The reduced pellets were assayed and the results are shown in Fig.5. It can be concluded that, as the end point temperature increases, the metal iron content, total iron content and the metallization of the reduced pellets all increase. When the temperature rises from 850°C to 950°C, the metal iron content and the metallization increase significantly, while the total iron content increases slightly. When the temperature is low, the reduction reactions will not proceed to completion. Reactions (3) and (5) proceed, but reactions (4) and (6) occur to only a very small extent, so the content of iron is very low, which makes the metallization low.

19

Temperature/ "C

Fig.5 The influence of end point temperature on the metal iron, total iron and metallization of reduced pellets. When the temperature is 850°C, the metal iron content is only 9.56% and the metallization is only 13.41%. When the temperature rises to 950°C, reactions (4) and (6) proceed to completion, and the iron content rapidly increase to 45.07% and the metallization is 56.57%. Also, the total iron content rises from 71.29% to 79.27%. When the temperature reaches 1050°C, the reactions continue. However, compared to at low temperature, the iron content is almost unchanged, mainly because the reduction reaction follows the unreacted core model. The compact metal shell on the surface prevents the reduction from progressing further. Microstructure Of Reduced Pellets At Different End Point Temperatures The microstructure of the reduced pellets at different temperature(450°C, 650°C, 850°C and 1050°C) can be seen using the Leica-DM-RXP polarizing microscope. They are shown in Fig.6 and Fig.7. The bright white is Fe, the off-white is FeO, the dark grey is Fe304; the dark is voids.

20

Fig.6 The microstructure of the edge of reduced pellets at different end point temperatures ( 50xl0times,reflex ) . From Fig.6 and Fig.7, we can conclude that, with an increase in the end point temperature, the degree of reduction increases. When the temperature is 450°C, reduction does not occur at the surface or the center. Iron oxides inside the pellets are basically Fe 2 0 3 and Fe3Ü4 as shown in pictures Ai and A2. When the temperature rises to 650°C, the pellets rapidly reduce, and the reduction products are mainly FeO as shown in pictures Bi and B2. Increasing the temperature to 850°C, the pellets in the surface and center are largely reduced, and the FeO at the surface starts to be deoxidized to iron, and there is some iron phase in the center, shown in pictures Cl and C2, When the temperature reaches 1050°C, the generated iron phase starts to move and grow to a contractive shell, as seen in pictures Dl and D2.

Fig.7 The microstructure of the center of reduced pellets under different end point temperatures ( 50x 10times,refiex ) . Compared the microstructure in the surface and center, the reduction at the surface

21

proceeds faster than that at the center. When the reduction reaction starts, the surface of the pellets are firstly reduced with the anthracite around it, then generated CO penetrates the pellets to deoxidize the iron oxide through flaws in the surface. The process of reduction of the whole pellet is gradual. When the reduction at the surface is complete, compact iron is formed at the surface. The generated CO will penetrate the pellets through flaws formed by reduction expansion. However, the iron phase will contract and the crystal lattice will move, the metal iron phase formed later will move to the metal iron phase formed early at the surface, so the reduced pellets are compact and contractive, and there are some voids closed or connected in the center, as shown in picture D2. Influence Of End Point Temperature On The Porosity Of Reduced Pellets The porosity of the pellets at different end point temperatures can be measured and calculated using the Leica-DM-RXP polarizing microscope. The results are shown in Fig.8. It can be concluded that, with the temperature rising from 450°C to 850°C, the porosity of the reduced pellets increases. When Fe203 is deoxidized step by step from Fe304 to FeO to Fe, the crystal structure of the pellets is changing. Inside Fe203, oxygen atoms are arranged in a tetragonal structure, while inside Fe 3 0 4 and FeO, oxygen atoms form a face-centered cubic structure. With the end point temperature rising, the iron oxides are gradually reduced, simultaneously, oxygen atoms experience a drastic re-adjustment and rearrangement, which makes the porosity at 850°C 1.5 times higher than that at 450°C. When the temperature rises to 950°C, the oxygen atoms remain unchanged, while iron atoms spread into the flaws inside the iron lattice; additionally, owing to the growing of iron phase and iron grains, the area of iron phase increases. For the interaction of the two results, the porosity of the pellets at 950°C is 1.38 times higher than at 450°C, which is smaller than at 850°C. When the temperature rises to 1050°C, for the continuing contraction, the porosity of the reduction pellets is 1.5 times higher than iron oxide pellets.

400

500

600

700

800

900

1000

1100

temperature/ °C

Fig.8 The porosity of reduced pellets at different end point temperatures. Conclusion Through the analysis and discussion above, the following conclusions can be reached: (1) Both iron oxide pellets and anthracite have excellent microwave absorbing properties,

22

and the anthracite has a faster temperature rising rate than the same quality pellets. Also, the whole process can be divided into two stages. (2) As the end point temperature rises from 850°C to 950°C, the reaction is so rapid that the iron oxides are largely reduced; the metal iron, total iron and metallization of the reduction pellets sharply increase. (3) Because of the metal shell formed at the surface of pellets, the metallization is little changed as the end point temperature rises from 950°C to 1050°C. The whole process of reduction follows the unreacted core model. (4) The porosity of the pellets at 950°C is higher than at 450°C, but smaller than at 850°C. References 1. ZHANG Q.C., and XIAO Q., The Development of Production of Direct Reduction and Prospect in China (Sintering and Pelletizing, 1997),31-35. 2. GAO W.X., et al., The Development State of Coal-Based Direct Reduction and the Technology of Rotary Hearth Furnace (Iron and Steel, 2008),68-74. 3. O scar D. Sidio AREX, Direct Reduction (Iron and Steel Maker, 1992)10-11. 4. KEM,T.hrum," Development of Reduction Processes for the Steel Production," ISIJ International, 55(1), (2006),1-10. 5. TONG Z.F., BI S.W., YANG Y.H.," The Research State of Heating by Microwave in the Field of Metallurgy ¿'Materials and Metallurgy, 3(2), (2004),115-119. 6. CHEN J., "The Applied Basic Research on the Self-Fluxed Pellet Containing Coal by Microwave Reduction" (Ph. D. Thesis. Beijing: Iron and Steel Research Institute, 2003), 15-19. 7. CHEN J., et al.," Experimental Research of Microwave Heating on Iron Ore Concentrates Containing Coal and Lime," Iron and Steel, 39(6), (2004), 6-9. 8. David E C, Diane C F, and Jon K W, "Processing Materials with Microwave Eergy", Mat. Sei. Eng. A-Struct, 287(2), (2000), 153-158. 9. Standish N, Pramusanto." Reduction of Microwave Irradiated Iron Ore Particles in CO," ISIJ International, 31(1), (1991), 11-16. 10. Donald MALMBERG, Par. HAHLIN,Emil NILSSON," Microwave Technology in Steel and Metal Industry, an Overview," ISIJ international, 47(4), (2007), 533-538. 11. Koki NISHIOKA, Takeshi TANIGUCHI. "Gasification and Reduction Behavior of Plastics and Iron Ore Mixtures by Microwave Heating , "ISIJ International, 47(4), (2007), 602-607. 12. Kotaro ISHIZAKI, "Kazuhiro NAGATA. Localized Heating and Reduction of Magnetite Ore with Coal in Composite Pellets Using Microwave Irradiation," ISIJ International, 47(6), (2007), 817-822. 13. Kingman S W, Vorster W, Rowson N A, "The Influence of Mineralogy on Microwave Assisted Grinding, "Minerals Engineering, 13(3), (2000), 313-327. 14. Jones D A, et al, "Understanding Microwave Assisted Breakage", Minerals Engineering, 18(2), (2005), 659-669. 15. Standish N, Worner H K, Gupta G, "Temperature Distribution in Microwave Heated Iron Ore-Carbon Composites", Microwave Power and Electromagnetic Energy, 25(2),

23

(1990), 75-80. 16. JIN Q.H., DAI S.S., HUANG Ka-ma. Microwave Chemistry (Beijing, BJ:Science Press, 1999), 8-15.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

CARBOTHERMAL REDUCTION OF ILMENITE CONCENTRATE AT HIGH TEMPERATURE

Run Huang, Chenguang Bai, Xuewei Lv, Guibao Qiu, Lei Lei College of Materials Science & Engineering, Chongqing University, Chongqing 400044, China Emails: [email protected]. [email protected] Keywords: Ilmenite, Carbothermal reduction, Titanium dioxide Abstract The TiCh-rich slag and pig iron can be produced from ilmenite concentrate by the electric furnace process. In this study, ilmenite concentrate was reduced in a vertical furnace equipped with a weighting data acquisition system. The effects of reducing agent (carbon) amount on the iron reduction were studied. The phase transformation, morphology and chemical compositions of reduced samples were investigated by X-ray diffractometry (XRD), scanning electron microscopy (SEM) and energy dispersive spectroscopy (EDS), respectively. It was found that the reductive degree of ilmenite concentrate increased with the increase of reducing time. The phases of reduced samples were mostly iron, rutile, TÍ3O5 and Fe2TiOs. The mass loss percentage increased with increasing the carbon amount from 8% to 12%. When the carbon amount exceeded 12%, the mass loss percentage decreased with the increasing of carbon amount. Introduction Ilmenite and rutile are the main raw materials for producing titania pigment and titanium metal. With the development of the titanium industry, huge amounts of titanium resources are needed. As sources of high-grade titania mineral decrease worldwide, more attention should be paid to the processes involving low grade mineral such as ilmenite ore. In ilmenite, MgO and CaO are the gangue parts. Therefore, the ilmenite refining becomes more and more popular. In the previous studies, the methods of ilmenite refining are electric furnace smelting, direct acid leaching, selective chlorination and thermal reduction [1]. Electric furnace smelting method is widely used in the titanium industry, however, which consumes a great deal of energies and has a high CO2 emission. Its essence is to remove iron oxides and obtain a high titania slag. So the thermal reduction of ilmenite concentrate is an important step in titanium industry, which can decrease carbon reagent consumption and save energy. Over the past several decades, many studies have been carried out on the mechanism and kinetics of the reduction of different ilmenite ores, which are as follows: The carbothermal reduction of Panzhihua ilmenite initiated at about 860 °C. In the temperatures ranging from 900 °C to 1400 °C, the reaction rate was changed like a shape of "M". That was to say that the

25

reaction rate increased with the growth of temperature at first, then decreased, afterward increased again and at last the reaction was arrested [2]. The Mohammad A.R group [3, 4] showed that the carbothermal reduction of primary ilmenite concentrate was faster in hydrogen and initiated at a lower reduced temperature than in argon and helium. The reduction in argon and helium had about the same rate and extent. Titanium oxycarbide started at 1000°C in hydrogen. The conversion of primary ilmenite concentrate to metallic iron and titanium oxycarbide at 1200 °C was completed in 300 minutes. Monoxide as the reductant to reduce ilmenite had been studied intensively by a few research groups [5, 6]. The rate and the degree of reduction depended on the formation of a metallic shell of iron which inhibited the transfer of CO to the reaction zones; titanium had a strong effect on the mechanism and rate of reduction of iron oxides. Wang and Yuan [7] studied the reductive kinetics of the reaction between a natural ilmenite and carbon. At temperature below 1100°C, chemical reaction was the rate controlling step, at 1100 to 1250°C mixed controlling was rate controlling step, above 1250°C diffusion through the product layer controlled the reaction. Zhang [8] investigated the reduction of natural, sintered and preoxidized ilmenite concentrates in hydrogen methane mixtures. Preoxidation and sintering of ilmenite concentrates changed their phase composition and caused a sharp decrease in specific surface area of raw materials and an increase in carburization temperature. Sintering also increased the metallization temperature of iron. However, there are few reports about the effect of reducing agent (carbon) amount on the iron reduction. This study examined the effect of carbon amount on the reduction rate and degree, the changes of phase composition and the morphology of the ilmenite concentrate. Experimental Methods Materials and Apparatus The ilmenite concentrate and coke, whose chemical compositions are shown in Table 1, are used in this study. Combining the chemical analysis with the XRD patterns of the ilmenite concentrate, the main phases are shown in Fig.l. The schematic diagram of experimental apparatus is shown in Fig.2. It consists of a shaft furnace, a temperature controller, electronic balance and a gas shielded system.

26

Fig. 1 XRD patterns of ilmenite concentrate

Fig. 2 Schematic of the apparatus

Table 1 Chemical analysis of ilmenite concentrate and coke (wt.%) Ilmenite

T¡0 2

FeO

Fe 2 0,

CaO

MnO

MgO

Si0 2

A1203

v2o5

concentrate

45.64

36.45

6.53

1.12

0.86

3.22

3.65

1.02

<0.10

S

P

Ash

volatile

carbon

0.65

0.12

14.12

2.22

83.66

-

-

-

-

Coke

Procedure The weighing ilmenite concentrate and coke, which were sieved to particles with a size under 74um and were thoroughly mixed, then a certain amount of adhesive and distilled water were added. Later, the mixture was pressed into the pancake pellets with a pressure of 20 MPa, each was about 30g in weight, 30mm in diameter and 5mm in height. According to the experience from the Panzhihua Iron and Steel Group and the previous studies [9, 10], the levels of carbon addition were 8%, 10%, 12%, 14% and 16% in this study. Experiments were conducted in a vertical furnace, which was equipped an accurate temperature control system, an electronic balance to record the mass change with respect to time and temperature, and a gas manometer system for the atmosphere controlling. The temperature was automatically controlled by a computer within ±5 °C. A Ni-Fe-Cr wire attached to the electronic balance was used to hanging the specimen crucible in the reaction zone. The system was sealed and the nitrogen gas was purged to provide an inert atmosphere in the reaction zone. The phase and reduction rate development during reduction of ilmenite were studied by temperature programmed. When the temperature was 1000 °C, the pallets were put into the deoxidation furnace and the ramping furnace temperature was varied from 1000 to 1300 °C at 3 "C/min. When the temperature reached 1300 °C and after the predetermined time, the samples were cooled in the graphite crucible to the room temperature. The phases were detected by XRD. The electron images of the samples were taken by SEM and the analysis of elements were obtained by EDS. Results and Discussion The experimental results are presented in terms of the percent mass loss versus time and recorded by me weighting data acquisition system. The percent mass loss is defined as: AW=

W -W ° 'xl00% W0

(1)

where, Wo is the initial mass of pallet sample; Wt is mass of sample at time t; AW is mass loss of sample from starting to time t. All the experiments were investigated under the same temperature range and heating rate. The effect of carbon addition on reduction rate and degree,

27

the phases and the morphology will be discussed in the next paragraphs. Reduction Degree

Temperature ( °C)

Carbon amount (%)

Fig.3 Percent mass loss of ilmenite concentrate with the temperature

Fig.4 Percent mass loss of the theoretical calculation

Carbon addition has a significant effect on reduction degree, as shown in Fig.3. (The mass loss contains carbon loss and CO emission.) The theoretical mass loss percentage was got according to that all the iron oxides were reduced to iron and titania was reduced to TÍ3O5, as shown in Fig.4. Obviously, the reduction degree increased with increasing the temperature. The reduction rate increased sharply near 1300 °C, which was the same with the Yuan's study [2]. It was observed that the percent mass loss increased with increasing the carbon amount from 8% to 12%. With the increase of carbon amount, the number of iron reduced from ferrous oxides also increased, which enlarged the interface between ilmenite and carbon. When the carbon amount rose from 12% to 16%, the mass loss decreased. The reduction degree was approached to linear increasing with the temperature increasing. Percent mass loss of the theoretical calculation was shown in Fig.4, of which percent mass loss change tendency was the same with experimental data in Fig.3. Experimental data is lower than the theoretical calculation, because the reaction didn't reach the equilibrium in this experiment and there might be an experimental error. When the carbon amount is 12%, compared with the theoretical calculation and experimental data, the percent mass loss were both higher than others. Because the total mass of the pallets with different carbon amount were the same, so in which the more carbon amount, the less ilmenite concentrate were, then, which diminished the efficient contact area between ilmenite and carbon. Phase Analysis of the Reduced Samples The phases of the reduced ilmenite pellet at different temperatures for 100 minutes were analyzed by XRD, as shown in Fig.5. The reduced samples with different carbon addition contained the same phases Fe, (Fe, Mg)TÍ20s, Fe2TiOs, TÍ3O5 and TÍO2. Owing to the ilmenite concentrate with a high content of magnesium, there was a (Fe, Mg)Ti2Os phase, which was

28

different from previous study [10, 11]. Fig.3 shows the mass loss rose sharply near 1300 °C, so the reaction didn't reach the equilibrium in this study. The reduction proceeded through the stages of ferric to ferrous iron and ferrous iron to metallic iron [11]. Iron oxides couldn't be reduced completely in the initial reaction. FeO, MgO and TÍO2 form a solid solution in the reduced sample. Moreover, the intensity of pseudobrookite decreases with the increasing of carbon amount evidently, which suggested that pseudobrookite was reduced and decomposed. When the carbon amount is 10%, the intensity of iron peak was much higher than others.

Fig. 5 XRD patterns of reduced ilmenite concentrate with different carbon amount SEM-EDS Analysis

29

Fig.6 SEM patterns of crossing section of the samples reduced with different carbon addition The reduced samples with different carbon addition at the same experimental condition were observed by the SEM and EDS, as shown in Fig.6 and Fig.7. The amplification of a, b column

30

are 500 and 3000 in Fig.6. The morphology of the reduced samples with different carbon is almost resembled, where the magnification is 500 times. However, when the carbon amount is 8%, the burr region is iron phase; the gray region is titanium oxides in b column. When the carbon amount is 10%, the burr region disappeared gradually and the coalescence of iron was agglomerated evidently compared to 8%. The XRD intensity of carbon amount 10% also indicated the amount of iron was more relatively. Because ofthat die iron grains are instable and moved to agglomerate. When the carbon amount is higher than 10%, the surface is more and more compact and leveling. As a result, CO couldn't escape from the reactive layer fast, which influenced the reaction balance or the equilibrium constant. The X-ray Ka mappings of iron, titanium and oxygen are shown in Fig.7. It could be seen that the bright regions, which were marked by wire-frames, were iron phase, and gray region were titanium oxides. The carbon addition has an important effect on the carbothermal reduction of ilmenite concentrate. Through analyzed this study, the optimum carbon amount is between 10% and 12%. And according to the chemical compositions of the ilmenite concentrate, the best carbon addition amount are necessary to reduce all hematite and ferrous to iron and all titania to TÍ2O3.

16%

Ti

Fe

O

Fig.7 SEM and the Ka X-ray mapping images of the reduced sample with 16% carbon Conclusions The reduction degree, the phase composition and the morphology of ilmenite concentrate in the carbothermal reduction were investigated from 1000 to 1300 °C. The results can be summarized as follows: (1) The mass loss percentage increases with increasing the carbon amount from 8% to 12%. When the carbon amount exceeds 12%, percent mass loss decreases with increasing of carbon amount. (2) The reduced samples with different carbon addition contain the same phases Fe, (Fe, Mg)Ti 2 0 5 , Fe 2 Ti0 5 , Ti 3 0 5 and Ti0 2 . (3) When the carbon amount is 8%, the burr region is iron phase and the gray region is titanium oxides. When the carbon amount is higher than 10%, the surface is becoming smooth gradually.

31

References [I] Mo W, Deng G. Z, Luo F. C. Metallurgy of Titanium [M]. Beijing, China, Metallurgical Industry Press, 2006, 133-198. [2] Li, W.B., Yuan, Z.F, Xu, C. Effect of temperature on carbothermic reduction of ilmenite. J. Iron and Steel Res., Int. 12(4), 2005, 1-5. [3] Wang Y. M et al. Reduction Extraction Kinetics of Titania and Iron from an Ilmenite by H2-Ar Gas Mixtures. ISIJ International, Vol. 49 (2009), No. 2, pp. 164-170. [4] Mohammad A. R. D, Zhang G. Q and Oleg O. Carbothermal Reduction of a Primary Ilmenite Concentrate in Different Gas Atmospheres. Metallurgical and Materials Transactions B Volume 4IB, February 2010, 182-192. [5] Zhao, Y, Shadman, F. Kinetics and mechanism of ilmenite reduction with carbon monoxide. AIChE Journal, v 36, 1990, 1433-1438. [6] Zhang G. Q and O. Ostrovski. Reduction of Ilmenite Ores by Methane Containing Gas. Part I. Effects of Ore composition, temperature and gas composition. Canad. Metall. Q. 40(3), 2001,317-326. [7] Wang Y. M, Yuan Z. F. Reductive kinetics of the reaction between a natural ilmenite and carbon. Int. J. Miner. Process. 81 (2006) 133-140. [8] Zhang G. Q and O. Ostrovski. Reduction of Ilmenite Concentrates by Methane Containing Gas, Part II: Effects of Preoxidation and Sintering. Canad. Metall. Q. 40(4), 2001,489-497. [9] J. Pesl, R.H. Eric. High temperature carbothermic reduction of Fe203-TiOr-MxOy oxide mixtures. Minerals Engineering, 15,2002, 971-984. [10] M. Pourabdoli, Sh. Raygan, H. Abdizadeh and K. Hanaei. Production of high titania slag by Electro-Slag Crucible Melting (ESCM) process. Int. J. Miner. Process. 178, 2006, 175-181. [II] C.S. Kucukkaragoz, R.H. Eric. Solid state reduction of a natural ilmenite. Minerals Engineering 19 (2006) 334-337.

32

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

A SIMULATION STUDY ON FLUE GAS CIRCULATING SINTERING (FGCS) FOR IRON ORES Tao Jiang, Zhenyu Fan, Yuanbo Zhang *, Guanghui Li, Xiaohui Fan School of Minerals Processing & Bioengineering Central South University, Changsha, Hunan, 410083, China Keywords: iron ore sintering; flue gas circulation; desulphurization Abstract Iron ore sintering process is the main source of SO2 generated in the steel industry, of which the discharge amount of SO2 emission accounts for about 60% of the total. Aiming at the features of sintering flue gases and existing problems during the flue gas desulphurization, a technical route of flue gas circulation sintering (FGCS) for iron ores has been put forward. Under the simulated experimental conditions, effects of FGCS process on the main output and quality indexes and SO2 emission rule are researched using simulating flue gases. Compared the novel FGCS with conventional sintering process, SO2 in the final sintering flue gas is obviously enriched and the total exhaust gases can be reduced evidently. The decreasing of oxygen content in circulating gases has bad effect on the sintering indexes. Increasing the oxygen potential is beneficial to obtain high quality sinters. This investigation shows that the FGCS process is promising. Introduction Steel is a kind of essential material for modern society, but steel-making process is characterized by high energy-consumption and emission as well as heavy environmental pollution. Chinese steel industry has developed rapidly in the last decade, and the environmental problem caused by ironmaking and steelmaking have also aroused the greatest concern, especially the atmospheric pollution caused by SO2, CO2, etc. Iron ore sintering process is the main source of SO2 generated in the steel industry, of which the discharge amount of SO2 emission accounts for about 60 % of the total. The development of the desulfurization process for sintering flue gases to reduce SO2 emission in sintering process is not only the requirements of sustainable development but also the most important environmental tasks for steel industry [1-5]. Sintering flue gases (SFG) have the characteristic of low SO2 content, complex composition and poor stability unlike the flue gases from coal-fired power plants [6-8]. The equipment investment and operation cost for SFG desulfurization are relatively high. According to preliminary estimation, the investment on SFG desulfurization equipment accounts for about 20 %~30 % of the total costs of sintering plant. Additionally, lower desulfurization efficiency, complex by-products and greater technique risks are common drawbacks of the existing SFG Corresponding author: Dr. Yuanbo Zhang, E-mail: [email protected].

33

desulfurization processes. SFG desulfurization technology are developed slowly [5, 9-11]. In the 1970s, in order to utilize the surplus heat in sintering exhaust gas, a technical idea of flue gas circulating sintering was put forward in Japan. Emission optimizing sintering process has been also applied in industrial production, which has brought effect of energy saving. But there is no research on the behaviors of SO2 in circulating gases and sulphur enrichment in the sinters [12-14], In this paper, behaviors of SO2 emission are firstly studied using simulating flue gases. Based on the features of SO2 emission in sintering gases, including generation, absorption and discharge, a technical route of flue gas circulation sintering (FGCS) for iron ores has been put forward in order to enrich SO2 in the final flue gases. Compared the main output and quality indexes and the metallurgical properties of sinters by the new FGCS and traditional sintering process, the feasibility of the flue gas circulating sintering process is verified. Experimental Raw Materials Raw materials used in this study are taken from a domestic sintering plant. The chemical compositions of raw materials are given in Table 1. Industrial analysis of coke breeze is listed in Table 2. As shown in Table 1 and Table 2, the primary source of sulfur in raw materials is coke breeze, which is 0.74% (wt.)

Raw materials Bedded ore fine Return fine Limestone Dolomite Burnt lime

Table 1. Chemical composition of raw materials /wt% CaO MgO TFe FeO Si0 2 AI2O3 60.51 5.47 4.17 2.28 2.62 0.49 56.13 7.35 5.22 9.22 2.88 2.15 1.58 1.42 48.85 0.76 3.62 18.25 0.18 1.19 33.15 0.68 1.41 0.37 3.08 67.88 0.96 -

S 0.082 0.024 0.040 0.030 0.130

LOI 4.61 0.82 42.45 45.71 24.65

Table 2. Industrial analyses of coke breeze and chemical compositions of ash Avt% primary chemical composition of ash Volatile matter Fixed carbon ash S MgO S TFe Si02 CaO AI2O3 0.21 1.52 84.58 13.90 0.74 1.42 6.01 0.50 4.19 0.051 Methods Sintering test included proportioning, blending, granulating, feeding, igniting, sintering, cooling and sieving. Drum mixer of 600mm><250 mm was used for blending. Granulated mixtures were filled in a 0100x500 mm sintering pot and the sintering bed height was fixed at 500 mm. After ignition, simulated flue gases were sucked into the sintering mixtures, whose composition

34

was determined based on the composition of traditional sintering flue gases. Schematic of main equipment for flue gases circulating sintering experiment is shown in Fig. 1. Main sintering parameters are including: ignition time of 1.5min, ignition temperature of 1150 °C±50 °C, ignition negative pressure of 5 KPa and sintering negative pressure of 10 KPa. After sintering, the sintering cake was cooled for 3min and cooling negative pressure was controlled at 5 KPa. DELTA 65-3 Gas Analyzer made by German MRU Company was used for detecting the composition and temperature of sintering exhaust gases.

sampling point upper ( 100-150mm )

gas tank

middle ( 250-300mm ) gas tank bottom ( 350-400mm )

Figure 1. Schematic of the equipment for flue gas circulating sintering experiment Results and Discussion Conventional Sintering Conventional sintering test was carried out and optimized sintering parameters were determined as following: the moisture content of mixtures of 7.8 wt%, coke breeze dosage of 4.5 wt%, the basicity of 2.2 and burnt lime dosage of 5.5 wt%. Under these optimal conditions, ISO tumbling index of the product of 50.26 % (wt), qualified sinter of 79.24 wt %, productivity of 1.544 t'm"2«h"' and vertical sintering speed of 21.13 mm-min"1 were obtained By analyzing the residual sulphur content in thefinishedsinter it can be found that about 0.026 wt% sulphur was evenly distributed in sinter (shown in Table 3). The change of the main composition in the sintering exhaust gas is given in Fig.2. Seenfromthis figure, O2 level in the exhaust gases was 9-12 %, CO2 was stabled at 13-14 % and CO was about 1~2 % during the sintering.

35

Table 3. Results of residual sulphur content in the finished sinter by conventional process Sampling point

Residual sulphur content (wt%)

upper middle bottom

0.025 0.026 0.026

21 18

§? 15

~ 12 c CD

c

9

O

° 6 to to o> 3

0

0

3

6

9

12

15

18

21

24

27

si nt er t i me (rri n) Figure 2. Change of O2, CO2, CO content in the exhaust gas from conventional sintering process Simulated Flue Gas Circulating Sintering Based on the laboratory results, O2 content in the sintering gaseous medium has more effects on the sintering process [14], and the main output and quality indexes become obviously worse when O2 content in sintering gas medium is less than 18%. Accordingly, simulated sintering mixed gas medium was set to contain 18 % O2 and 500 ppm SO2 Combined with the emission rules of traditional sintering flue gases, the main gas composition of simulated sintering gas is determined and given in Table 4. The temperature of the mixed gas was controlled at 150 °C. Table 4. Main gas composition of simulated sintering gas medium Composition CO Water vapour C0 2 S0 2 N2 O2 1 Content / % 18 5 500 ppm 70.95 5 Effects on the Sintering Indexes Sintering parameters was set as moisture of 7.8 wt%, basicity of 2.2, coke breeze dosage of 4.5 wt% and burnt lime dosage of 5.5 wt%. Each test had been conducted twice and the results are shown in Table 5. Compared with the conventional sintering results, FGCS indexes slightly decline. Tumbling

36

index (TI) is 48.78 %, qualified sinter (QS) is 77.78 %, productivity (P) is 1.482 t«m"2'h"' and vertical sintering velocity (VSV) is 20.58 mm-min"1. Table 5. Results of simulated FGCS with 18 % O2 in the sintering gas medium No. 1 2 Average

VSV /mm-min"1 20.50 20.66 20.58

QS /% 77.65 77.90 77.78

TI /% 48.60 48.96 48.78

P /t-m"2-h"' 1.480 1.484 1.482

Optimization sintering has been done by increasing oxygen content to 21 %, and the results are shown in Table 6. Table 6. Results of simulated FGCS with 21 % O2 in the sintering gas medium Oxygen content /% VSV /mm-min"1 TI /% P /t-m"2-h"' QS /% 21 51.43 1.529 20.98 78.70 As seen from Table 5 and Table 6, when oxygen content in flue gas is increased from 18 % to 21 %, all sintering indexes were obviously improved, approximate to the results by the traditional sintering process. Therefore, enhancing the oxygen content of the sintering gas medium is useful to ensure sintering indexes in future industrial production of FGCS. Effects on the Residual Sulphur Content of Sinter In order to find out the residual sulphur contents of sinter, the upper, middle and bottom layer sinters are sampled respectively. The final analyses are presented in Table 7.

No. 1

2

Table 7. Results of residual sulphur content in different layer sinters Sampling point Residual sulphur content (wt%) upper 0.030 middle 0.032 bottom 0.027 upper 0.029 middle 0.031 bottom 0.028

According to the data in Table 7, residual sulphur content in different layer sinter changes between 0.029 wt% and 0.032 wt%, which is a little higher than it in conventional sinter (0.026 wt%). The results indicate the novel FGCS process will lead to sulphur enrichment in sinter to a certain extent. Therefore, it should be paid much attention that the residual sulphur content in sinter will be further increased if SO2 content in the circulating flue gases is increased.

37

Effects on the Sulphur Content in the Final Sintering Flue Gases MRU Gas analyzer was used to detect the changes of SO2 content in conventional and FGCS exhaust gases and the curves are plotted in Fig. 3. 2100

1500

c
w

600 300

0

0

3

6

9

12

15

18

21

24

27

si nt er t i me ( ni n) Figure 3. Changes of SO2 content in conventional and FGCS exhaust gases Seen from Fig. 3, the emission rules of SO2 in the two processes are similar. After the ignition, SO2 content in flue gas maintains a low level. With the disappearing of wet mixture zone in the sintering bed, SO2 content rises rapidly in the exhaust gases until the peak value emerges. Comparatively, final flue gas from FGCS process has a higher value of SO2 content (2006 ppm) than that in conventional sintering flue gases(1256 ppm ) The investigation shows that most of SO2 in the circulating flue gas is enriched in the final exhaust gases. Comparison of Metallurgical Properties of the Products The metallurgical properties of sinters by the traditional sintering process and FGCS process were tested. The traditional sinter is marked sample A and the FGCS sinter is sample B. Table 8 Reducibility and RDI of sample A and B /% Samp e

RDI ( + 3.15)

A B

62.49 63.32

RI 88.99 85.69

Table 9 Soft melting properties of different sinters /°C Sample

Ta

Ts

Tm

Td

( Ts-Ta )

( Td-Tm )

A B

1063 1070

1180 1178

1310 1305

1441 1427

117 108

131 122

38

As seen from Table 8 and Table 9, RDI of two samples are very close, indicating that FGCS has little effect on the RDI of finished products. The reducibility of two samples is higher than 85 % in spite of a little lower value for FGCS. Moreover, two samples have good soft melting performances. The softening zone and melting zone are relatively narrow. On the whole, the metallurgical properties of final sinter are still kept good even if FGCS process is used. Conclusion In this paper, simulated flue gas circulating sintering experiments has been done. Investigations show that the decrease of oxygen content in circulating gas will cause the sintering indexes to go down. If the oxygen content in the circulating flue gas is increased to 21%, high quality sinter can be obtained. The most of SO7 in the circulating flue gases will be enriched in the final exhaust gas and the sulfur content of finished sinter by FGCS is increased slightly. The behavior of sulfur during FGCS enables the amount of final flue gas to reduce greatly, which is much beneficial for the following desulfurization process. Acknowledgements The authors would like to express their thanks to National Science Fund for Distinguished Young Scholars (No.50725416) and the Graduate Degree Thesis Innovation Foundation of Central South University (No.2009ssxt245) for financial support of this investigation. References [1] S.J. Hao, et al. "Methods of Decreasing S02 Pollution in Sintering," Journal of Hebei Institute of Technology, 2(2006), 14-17. [2] W.M. Song. "Approach to Controlling of S02 Pollution at Integrated Iron and Steel Companies," Iron and Steel, 7(1999), 66-69. [3] X. Zhang, Z.C. Guo. "On Energy Consumption and Atmospheric Pollutants of China's Iron and Steel Industry," Iron and Steel, 1(2000), 63-68. [4] S.H. Shan, CF. Li. "Speed up the implementation of flue gas desulfurization sintering and promoting regional environmental improvement". China steel, 4(2006), 16-20 [5] X. Gao. "Situation of Sintering and Pelletizing Processes Desulfurization in Our Country and Count measures to Decrease SO2 Emission". Sintering and Pelletizing, 6(2008), 1-5. [6] Changsha black metallurgical and mining design institute. Sintering design manual.( Beijing: Metallurgical Industry Press, 1999), 127-131 [7] J.I. Mou. "In-Plant De-SOX, De -NOX in Iron Ore Sintering Process," Kuangye, 2,(2000), 41 -48. [8] E.O. Wigman. Siderology. (Beijing, Metallurgical Industry Press,1993), 40 [9] J.Y. Liao, et al. "An Overview on Desulfurization of the flue gas from sintering, " Southern Metals,2,( 2006), 1-3.

39

[10] L. Wang. "Comparison and Analysis of Desulfurization Technique for Sintering Smoke ". Hebei Metallurgy. 2,(2008), 53-56. [11] H. Wang, et al. "New Development of Controlling SO2 Pollution from Flue Gas," Power System Engineering.il,(2006), 5-7. [12] Y. Hosotani, et al. "Influence of Oxygen and Humidity of Inlet Gas on the Sintering Reaction," iron and steel, 83,(1997),293-298. [13] T. Shoho, et al. Proceedings of The Sixth International Iron and Steel Congress, (Nagoya:2 (1990), 171.) [14] H. Werz, J. Otto and J. Rengersem. "Source Pollution control at iron ore sintering plants through waste gas recycling," Stahl und Eisen, 115(1995), 37. [15] K.H. Chen. "Generation Mechanism of SO2 in Sintering Process and Its Emission Control," Sintering and Pelletizing, 8(2007), 13-16.

40

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

Optimization of Preparation of U 3 0 8 by Calcination from Ammonium Diuranate Using Response Surface Methodology Bingguo Liu ' 2, Jinhui Peng ' '2*, Daifu Huang' 1. Faculty of Materials and Metallurgical Engineering, Kunming University of Science and Technology, Kunming Yunnan, 650093, PR China 2. Key Laboratory of Unconventional Metallurgy, Ministry of Education, University of Science and Technology, Kunming Yunnan, 650093, PR China 3. No.272 Nuclear industry Factory, China National Nuclear Corporation, Hengyang, 421002, PR China Response surface methodology; Ammonium diuranate; Calcination; U3O8; Total uranium; C/4* Abstract The conditions to prepare U3O8 by calcination from ammonium diuranate were optimized, and the response surface design method was applied to analyze the influence on the total uranium and U'* of calcination temperature, calcination time and mass of sample. A quadratic equation model for the total uranium and t/ 4+ of ^Ogwas built and the effects of main factors and their corresponding relationships were obtained. The analysis of variance shows that calcination temperature and calcination time significantly affected the value of total uranium and U'*oî U3CV The optimal calcination conditions were as follows: calcination temperature 931.83 K, calcination time 24.32 min and 43.89 g. Under these conditions the value of total uranium and C/4*of U3O8 was 84.78% and 28.02%, respectively. The validity of the model was confirmed experimentally and the results were satisfactory. Introduction Nuclear energy is clean, effective, and a sustainable from of energy. It plays an important role in overcoming the impending global energy crisis, and in reducing energy vulnerability. In addition, nuclear energy is a possible cheaper and cleaner source of energy compared to fossil fuels, with no not really, mining of uranium produces CO2 contribution to green house gases [1]. Uranium dioxide, one of the most important fuel pellets used as nuclear fuel in light water reactors, is manufactured by calcining ammonium diuranate to triuranium octaoxide, followed by its reduction to uranium dioxide using cracked ammonia at 700°C [2]. Earlier attempts to optimize the important calcination process through several common techniques have not yielded desirable results [3]. These techniques either rely on the classical one parameter at a time approach that ignores the combined interactions between physicochemical parameters, or are theoretical in nature. Furthermore, these techniques also require large number of experimental data to be generated [4]. In order to produce triuranium ' Corresponding author: Jinhui Peng Tel.: +86 871 5192076; fax: +86 871 5191046. E-mail address: [email protected].

41

octaoxide in an optimized manner, which considers the interaction of different process parameters [5], there is a need to adopt a multivariate statistical technique. Response surface methodology (RSM) is a collection of mathematical and statistical techniques useful for analyzing the effects of several independent variables [6]. The main advantage of RSM is the reduced number of experimental trials needed to evaluate multiple parameters and their interactions [7], It deals with multivariate experimental design strategy, statistical modeling and process optimization [8]. Several previous researchers have shown that RSM is a powerful statistical tool in process optimization; it has been successfully applied to optimize the process parameters for biosorption of metals [9], and producing dyes [10] from synthetic solutions. This work aims to extend the RSM technique to calcination of ammonium diuranate. In this study the effects of calcination temperature, calcination time, and mass of sample on the value of total uranium and £/4+of triuranium octaoxide were investigated by means of central composite design (CCD, a part of RSM package). RSM was used to determine the optimal conditions and an empirical model correlating the decomposition rate to the three variables was then developed. Experimental Materials and Thermal Decomposition Behavior The ammonium diuranate used in the study was obtained from No.272 Nuclear industry Factory, China National Nuclear Corporation, and the particle size was less than 50 urn. The thermal behavior of ammonium diuranate has been reported in the literature. As can be seen, ammonium diuranate was decomposed into uranium trioxide at 623K, followed by its reduction to triuranium octaoxide using cracked ammonia. On the basis of these results, the thermal decomposition mechanism of ammonium diuranate could be shown as [11]: (M/„) 2 (/ 2 0 7 =2JV7/ 3 +2U03+H10

(1)

3UO,+H2=U,Os+H20

(2)

Calcination Experiments The calcination experiments were carried out by varying temperature, duration of calcination times, and mass of sample, using a conventional tube furnace. Initially, the furnace was preheated at 25 K/min until the desired temperature was reached. The ammonium diuranate was weighed and placed in a ceramic crucible which was located approximately in the centre of the tubular electric oven. During the reaction, the temperature was monitored by a PID (proportional-integral-derivative) temperature controller system. The product was removed from the tubular reactor and immediately put into a drier once the experiment was complete and allowed to naturally cool to room temperature. The value of total uranium and U** of triuranium octaoxide was determined using chemical analysis methods. Design of Experiment Using Response Surface Methodology On the basis of the initial decomposition results, RSM was employed to optimize the calcination

42

conditions in order to obtain maximum value of total uranium andi/4* of triuranium octaoxide, and CCD was employed to design the experiments. In this study, the effects of three independent variables, %* (calcination temperature), & (calcination time), and ^ (mass of sample), at two levels were investigated using central composite design (Table 1). Table 1. Independent variables and their levels used for center composite rotatable design Coded variable levels

Independent variables

Symbol

-1

+1

Calcination temperature (K)

X\

673

1073

Calcination time (min)

%2

20

50

Mass of sample (g)

X%

30

60

The value of the total uranium and U"* of triuranium octaoxide was taken as the two responses of the designed experiments. A total of 20 experiments consisting of 8 factorial points, 6 axial points, and 6 replicates at the central points were performed. The experimental data obtained from the designed experiment were analyzed by the response surface regression procedure using the following second-order polynomial equation [12, 13]:

r=A + £ ß,x, + £ ß„zf + £ ß,z,Zj. M

i-1

(3)

.(/

where y is the predicated response, /?„ is a constant, /?, is the i-th linear coefficient, ßä is the i-th quadratic coefficient, ßn is ij-th interaction coefficient, and %u s are independent variables. Results and Discussion Data Analysis and Evaluation of The Model by RSM The experiments were conducted based on the design matrix under the defined conditions and the response obtained from the experimental runs are shown in Table 2. Table 2. Experimental design matrix and results Run 1 2 3 4 5 6 7 8 9 10

Calcination temperature

r, (K)

673.00 1073.00 673.00 1073.00 673.00 1073.00 673.00 1073.00 536.64 1209.36

Calcination time 7 , (min)

Mass of sample

Total uranium

7,
y, (%)

20.00 20.00 50.00 50.00

30.00

30.00 30,00 30.00 60.00 60.00 60.00 60,00 45.00 45.00

81.67 84.58 82.91 84.62 78.37 84.57 83.80 84.87 77.84 84.66

20.00 20.00 50.00 50.00 35.00 35,00

43

i/4+

r,(%) 15.00 28.30 18.75 28.18 13.58 28.20 19.21 28.45 2.69 28.70

873.00 873.00 873,00 873.00 873,00 873.00 873.00 873.00 873.00 873.00

9.77 60.23 35,00 35.00 35.00 35.00 35.00 35.00 35.00 35,00

45.00 45.00. 19.77 70.23 45,00 45,00 45.00 45.00 45.00 45.00

84.30 84.40 84,33 84,41 84,48 84,66 84,64 84.52 84.60 84.78

18.77 27.22 26.74 27.30 28.25 28.11 28.66 28.07 28,10 28.73

According to the sequential model, the sum of squares can be obtained, and the models were selected based on the highest order polynomial where the additional terms were significant and the models were not aliased [14]. The responses of the value of the total uranium and l/ 4+ of triuranium octaoxide were considered in studying the effect of process variables The response of the value of total uranium and U'* of triuranium octaoxide and the independent variables were used to develop an empirical model after excluding the insignificant terms, which is presented by Eq. (4) and (5), respectively: y, =84.62+ \.1\X\ + 0.52 Xi .0.15-^3- 0.79 X\Xi + 0.33-ft Xi + 0.56#2 Xi - 1.2 x] - 0.11 xl~0l0Xl y2 =28.29+6.61 ^i+1.74-&-0.01 ^3-1.16-fi Xi+o.uXi

Xi +0.28#2 #3-4.26Xi ~ l-6&X¡-°-26X¡

(4) (5)

The quality of the model developed was evaluated based on the correlation coefficient [15]. The R2 value for Eq. (4) and Eq. (5) was 0.928 and 0.977, respectively, which indicated that 92.8% and 97.7% variability of the total variation in the total uranium and £/4+of triuranium octaoxide was attributed to the experimental variables studied. The R2 of 0.928 and 0.977 for Eq. (4) and Eq. (5) was considered relatively high, indicating that there was good agreement between the experimental value of the total uranium and U'* of triuranium octaoxide and the predicted one from this model. Furthermore, anlysis of variance (ANOVA, also a part of RSM) was also carried out to justify the adequacy of the model. The ANOVA for the quadratic model for the value of the total uranium and U,f of triuranium octaoxide is presented in Table 3. The model's adequacy was tested through the lack of fit F-test, in which the residual error was compared to the pure error. According to the software analysis, "Lack of fit F-value" of the total uranium of triuranium octaoxide of 97.5 implies that the lack of fit was significat. The "Model F-value" of the total uranium of triuranium octaoxide of 14.41 implies that the model was significant. Values of "Prob > F" of the total uranium of triuranium octaoxide less than 0.05 indicates that the model terms are significant, whereas the values greater than 0. lOare not significant. The same, "Lack of fit F-value" of the U4*of triuranium octaoxide of 47.9 implies that the lack of fit was significat. The "Model F-value" of the I/4* of triuranium octaoxide of 47.97 implies that the model was significant and there was only 0.01% chance that a "Model F-value" this large could occur due to noise. Values of "Prob> F" of the U"* of triuranium octaoxide is 0.0003, which is less than 0.05, which is indicates that the model terms are significant.

44

Table 3. Analysis of variance (ANOVA) for response surface quadratic model for the value of total uranium andO 4 of triuranium octaoxide Sum of squares

Mean square

F-value

Source

Df

Model

9

73.25

936.49

8.14

104.05

Residual

10

5,65

21.69

0.56

2.17

Lack of fit

5

5.59

21.25

1.12

4.25

Pure error

5

0.057

0.44

0.011

0.089

19

78.90

958.18

Cor total Y.

u

2

R =0.928,

■■

U

2>

t/ 4+

u"+

I"

Prob>F

U<+

Zf

u4+

5>

14.41

47.97

0.0001

<0.0001

97.5

47,9

O.0001

0,0003

ÄL.=0.864;adequateprecision=13.58>4

R =0.977, Ä ,. =0.957;adequateprecision=23.63>4

The checking of model adequacy is an important part of the data analysis procedure, since it would give poor or misleading results if it is an inadequate fit [16]. Multivariable linear regression was used to calculate the coefficients of the second-order polynomial equation and the regression coefficients, whose significance was determined using the P-value, summarized in Table 4. In this case,-^ 1 ,-^ 2 ,^ 1 and the interaction terms (%' %2) were significant to the total uranium of triuranium octaoxide, and *>, -*2, *' , -*-2 were significant to U"* of triuranium octaoxide, whereas the Y

Y

Y

Y

Y Y

Y

interaction terms ( « î * ! / ! * ! ^ ! * ! / ! ) were insignificant to the response. Table 4.Regression coefficient of polynomial function of response surface of the value of total uranium and t/ 4+ of triuranium octaoxide Regression Term

Df

coefficient

Standard error

Z"

0.31

0.60

6.61

0.20

0.40

1.74

0.20

0.40

-0,15

0.011

0.20

0.40

x,x2 x,x,

-0.79

-1.16

0.27

0.33

0.14

X¡X, !

0.56

Intercept

I"

I"

u

Z"

u"+

P-value

Z«

u4*

29.63

0.0001

O.0001

1.26

5.76

2.16

7.50

<0,0001

O.0001

0.072

0.85

0.98

2.62

0.0274

0.0014

-0.60

-0.88

0.30

0.90

0.4783

0.9783

0.52

-1,38

-2.32

-0.20

4.007

0.0140

0.0507

0.27

0.52

-0.26

-1.02

0.92

1.30

0.2393

0.7917

0.28

0.27

0.52

-0.035

-0.88

1.15

1.44

0.0623

0.6009

-1.20

-4.26

0.20

0.39

-1.64

-5.13

-0.76

-3,40

0.0001

O.OOOl

-0.11

-1.68

0.20

0,39

-0.55

-2.55

0.33

-0.82

0.5977

0,0015

-0.10

-0.26

0.20

0.39

-0.54

-1.12

0.34

0.60

0.6216

0.5175

X,

1.71

X,

0.52

X,

x,

limites

,+

85.30

28.29

x;2

Upper confidence

limites

26.95

84.62

x,

Lower confidence

83.93

Response Surface Analyses Total Uranium of Triuranium Octaoxide The best way to visualize the influence of the independent variables on the response is to draw surface response plots of the model [17]. The three-dimensional response surfaces constructed to

45

show the effects of the calcination variables on the value of triuranium octaoxide using the fitted quadratic polynomial equation obtained from regression analysis are shown in Fig. 1 and Fig. 2.

Fig. 1 Three-dimensional plot of the response surface

Fig. 2 Three-dimensional plot of the response surface

for the total uranium of triuranium octaoxide ( y )

for the total uranium of triuranium octaoxide ( y, )

as related to temperature ( ^ ) and time ( ^ 2 )

as related to temperature ( ^ )and mass of sample ( ^ 3 )

Fig. 1 shows the effect of calcination temperature and calcination time on the value of total uranium of triuranium octaoxide with the mass of sample at 40g. Fig. 2 shows the effect of calcination temperature and mass of sample on the value of the total uranium of triuranium octaoxide, with the calcination time set at 35 min. It was observed that the value of total triuranium octaoxide significantly increased with increasing calcination temperature and calcination time, while the value of total triuranium octaoxide is seen to increase with a decrease in mass of sample within the experimental range studied. This is because thermal decomposition of ammonium diuranate is an endothermic reaction and would accelerate with increasing temperature [11]. Moreover, a longer calcination time leads to a more complete calcination reaction. So, the total uranium of triuranium octaoxide increased with increasing calcination temperature and calcination time. lf+ of Triuranium Octaoxide The three-dimensional display of the response surface plot of the if* of triuranium octaoxide as a function of the calcination temperature, calcination time, and mass of sample are shown in Fig. 3 and Fig. 4.

46

Fig. 3 Three-dimensional plot of the response surface for [ A

Fig. 4 Three-dimensional plot of the response surface for(/**

of triuranium octaoxide ( y ) as related to temperature

of triuranium octaoxide ( y ) as related to temperature

(%} ) and mass of sample ( ^ 3 )

( % t ) and time (%2 )

Fig. 3 shows the effect of calcination temperature and calcination time on the value of W* of triuranium octaoxide (mass of sample was fixed at 40 g), and Fig. 4 shows the effect of calcination temperature and mass of sample on the value of U'* of triuranium octaoxide (calcination time was fixed at 35 min). As seen from Fig. 3 and Fig. 4, the value of C4+ of triuranium octaoxide increased abruptly with increasing calcination temperature and then increased smoothly after 973 K, while the value of U'* of triuranium octaoxide decreased with increasing mass of sample within the experimental range studied. This is attributed to the reduction reaction of uranium trioxide. As a result of endothermic reaction, the reduction of uranium trioxide would accelerate with increasing temperature [11]. In addition, the reduction of uranium trioxide is a gas-solid phase reaction. The diffusion of gas becomes more difficult with increasing mass of sample, resulting in a decrease in concentration of V* in triuranium octaoxide. Optimal Conditions and Verification of the Model Thus, based on the above model, the optimal conditions for maximizing the total uranium and U'* of triuranium octaoxide were 931.83 K, 24.32 min, and 43.89 g, and the values of total uranium and u'* of triuranium octaoxide were 84.78% and 28.02%, respectively. In order to confirm the optimized conditions, the accuracy of the model was validated with experiments under the optimized conditions. An experiment was carried out with parameters as suggested by the model. The conditions used in the confirmatory experiment were as follows: calcination temperature 932 K, calcination time 24 min, and mass of sample 44 g, giving values of total uranium and [/4+of triuranium octaoxide of 84.62% and 28.16%, respectively, which concurred with the model prediction. The model, therefore, can be considered to fit the experimental data very well in these experimental conditions; with an error margin of only 0.19% and 0.5%, respectively. Therefore, the model is acceptably valid.

47

Conclusions The effects of calcination temperature, calcination time, and mass of sample were optimized using RSM. A quadratic model was developed to correlate the calcination variables with the total uranium and U"* of triuranium octaoxide. This study showed that response surface methodology was an appropriate approach to optimize conditions for achieving suitable values of total uranium and U'* of triuranium octaoxide. The experimental and predicted values were very close, which reflected the correctness and applicability of RSM. The value of the adjusted determination coefficient was 0.864 and 0.957, respectively, showing a relatively high significance. Using RSM to optimize the conditions, the optimal conditions were found to be 931.83 K, 24.32 min, and 43.89 g. Under these conditions, the predicted value of total uranium and U** of triuranium octaoxide of 84.78% and 28.02% was in good agreement with the actual experimental values (84.62% and 29.18%). Acknowledgements Financial support for this work from the National Natural Science Foundation Council of China (50734007), Technology Project in Yunnan Province (2007GA002) and Analysis and Testing Foundation of Kunming University of Science and Technology (2008-16) are gratefully acknowledged. References [1] XU, Y.C., "Contested regimes," Energy, 33 (2008), 1197-1205. [2] Song, K. W., Kim, K.S. and Kang, K. W., "Grain size control of U 0 2 pellets by adding heat-treated U 3 0 8 particles to U0 2 powder," J. Nucl. Mater., 317 ( 2003), 204-211. [3] Öberg, T.G. and Deming, S.N., "Bacterial leaching of nickel latérites using chemolithotrophic microorganisms: Process optimisation using response surface methodology and central composite rotatable design," Measurement and control. 98(2009), 241-246. [4] Shweta, S., Anushree, M. and Santosh, S., "Application of response surface methodology (RSM) for optimization of nutrient supplementation for Cr (VI) removal by Aspergillus lentulus AML05," J. Hazard. Mater., 164 (2009), 1198-1204. [5] Gratuito, M. K. B. et al., "Production of activated carbon from coconut shell: Optimization using response surface methodology," Bioresour. Techno!., 99 (2008), 4887-4895. [6] Myers, R. H. and Montgomery, D.C., "Response Surface Methodology" John Wiley & Sons, Inc., USA, 2002. [7] Chen, M. J., Chen, K. N. and Lin. C.W., "Optimization on response surface models for the optimal manufacturing conditions of dairy tofii," J. Food Eng., 68 (2005), 471-480. [8] Fu, J. F., Zhao, Y. Q. and Wu, Q. L., "Response surface optimization of phosphorus species adsorption onto powdered alum sludge,"/ Hazard. Mater, 144 (2007), 499-505. [9] Kiran, B., Kaushik, A.and Kaushik, C.P., "Response surface methodological approach for optimizing removal of Cr(VI) from aqueous solution using immobilized cyanobacterium," Chem. Eng. J. 126 (2007), 147-153. [10] Mohana, S., et al., "Response surface methodology for optimization of medium for

48

decolourization of textile dye direct Black 22 by a novel bacterial consortium," Bioresour. Technol., 99 (2008), 562-569. [11] HU Zheng-jie, "Defiuorination problem with the converting process of ADU to UO2 powder," Atomic Energy Science and Technology, 32 (1998), 352-357. [12] Bezerra, M. A., et al., "Response surface methodology (RSM) as a tool for optimization in analytical chemistry," Talanta, 76 (2008), 965-977. [13] Rodriguze-Nogales, et al., "Experimental design and response surface modeling applied for the optimisation of pectin hydrolysis by enzymes from A. niger CECT 2088," Food Chem., 2007, 101 (2007), 634-642. [14] Tan, I. A. W., et al., "Optimization of preparation conditions for activated carbons from coconut husk using response surface methodology," Chem. Eng. J., 137 (2008), 462-470. [15] Ahn, J.H., et al., "Optimization of microencapsulation of seed oil by response surface methodology," Food Chem., 107 (2008), 98-105. [16] Korbahti, B. K. and Rauf, M. A., "Determination of optimum operating conditions of carmine decoloration by UV/H2O2 using response surface methodology," J. Hazard. Mater., 161 (2009), 281-286. [17] Raymond, et al., "Response surface methodology," Tecimometrics, 312 (1989), 432-438.

49

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

MICROWAVE FIELD ATTENUATION LENGTH AND HALF-POWER DEPTH IN MAGNETIC MATERIALS Zhiwei Peng1, Jiann-Yang Hwang1, Xiaodi Huang1, Matthew Andriese1, Wayne Bell1, 'Department of Materials Science and Engineering, Michigan Technological University, Houghton, MI 49931, USA Keywords: Microwave, Field Attenuation, Half-power Depth, Dielectric Parameters Abstract The equations for determining microwave field attenuation length and half-power depth in magnetic materials were derived from Maxwell's equations. The microwave field attenuation length and the half-power depth for a magnetite concentrate were calculated and the temperature dependence of them was determined. It is demonstrated the microwave field attenuation length and the half-power depth highly depend on temperature and decrease with increasing temperature up to 900 °C. The variations of field attenuation length and the half-power depth with temperature indicate the microwave power attenuates much faster than the field strength in materials. The evaluations of field attenuation length and microwave half-power depth can be used to characterize the microwave dissipation behaviors in the sample and optimize the dimensions of the material in microwave heating. Introduction Microwave has been applied in materials processing for its distinguished advantages including rapid heat transfer, volumetric and selective heating [1,2]. However, most research work involving microwave heating are confined to non-magnetic materials (e.g. ceramics), with limited attention paid to the heating of mixed property materials (e.g. magnetite and other ferrites) under microwave irradiation [3-5]. Since many magnetic materials exhibit good and even better microwave absorption capability than non-magnetic counterparts, it is quite necessary and deserved to take advantage of this property and then improve the heating efficiency. To achieve this aim, further understanding on the microwave heating mechanism, especially the microwave propagation/dissipation behaviors in magnetic materials is required. The dissipation behavior of microwave in materials has been identified as one of the key factors that affect the efficiency of heating under microwave irradiation and this behavior can be characterized by several parameters. Apart from traditional penetration depth (TEM power penetration depth, Dp = l/(2*rx), where a is the TEM field attenuation factor), two other parameters, field attenuation length (Df) and half-power depth (Dh), are also applied to indicate the microwave dissipation in materials [6],The accurate determination of such parameters can help in optimizing the dimensions of the load in a microwave furnace, producing more uniform heating under microwave irradiation and avoiding thermal runaway. In this paper, two simple equations were developed to quantitatively determine the microwave field attenuation length and half-power depth in magnetic materials. A typical magnetic material,

51

magnetite concentrate, was taken as an example to show the temperature dependence of field attenuation length and half-power depth from room temperature to as high as 900 °C at 915 MHz and 2450MHz. Derivation of Field Attenuation Length and Half-power Depth Equations In a homogeneous medium and without convection or external currents, the differential Maxwell's equations in lossy dielectrics are given by VxE = -ja)ßH

(1)

WxH = jweE

(2)

V £ =0

(3)

V// =0

(4)

where E is the electric field intensity (V/m), H is the magnetic field intensity (A/m), j is the imaginary unit (j2= -1), w is the angular frequency of microwave (rad/s), and u and e are complex relative permeability and permittivity, respectively. Considering vector multiplication, for Maxwell's equations, we have VxVxE

= V(V-E)-V2E

= -ja)ßVxH

= ar:fi£E = -fE

(5)

= jmNxE

=
(6)

VxVxH=V(VH)-V2H Thus, we obtain W2E-fE V2H-fH

=0 =0

(7) (8)

The two equations above are also known as Heimholte equations. Assuming a TEM plane wave traveling in the z direction with x component of electric field intensity, Heimholte equation would yield d2E ^-rEx=0 az

(9)

The general solution to this wave equation is Ex(z) = C^-"!+C2er'

=Cie-ae-J0'+C2eazeiflz

52

(10)

where Ci and C2 are constants. In time domain, the representation of E is Ex(zj) = Re{Ex(z)eJ'"} = C,e-'" cos(ox-ßz)+C2ea!

cos(at + ßz)

(11)

Here, the propagation constant ( y) is a complex number and can be expressed as (12)

r=a+jß

where a is the field attenuation constant (Np/m), and ß is the phase constant (rad/m). From the definition, we have 7 = J<*>(fJ£)"2 = [-o)2e„Mo (er'Mr '- £, "ßr ")+J»?£<Jk ( ^ X "+ er "Mr ')]" 2

(13)

where so is the permittivity of free space (8.854 x 10"12 F/m), uo is the permeability of free space (4it xlO"7 H/m), er' is the real part of complex relative permittivity, er" is the imaginary part of complex relative permittivity, ur' is the real part of complex relative permeability, and ur" is the imaginary part of complex relative permeability. By separating the real part and imaginary part of Eq. (13) and equating the real part with a, we have a = ^ { f , "ßr "- e, 'ßr '+ [(*, 'ßr i + (e, "ßr "f + {e, ' ß , ")2 + (ßr 'e, "f ]" 2 }'

(14)

The wave is attenuated as it traverses the medium and therefore the power is dissipated. According to the definition of the field attenuation length (skin depth), Df, we obtain D,

= ¿ = ¿ {

f

' "^

"~£r

'V,

1+

[(£r

'Mr 'f+{S,

"ßr "f

+ {€, 'ßr "f+(ßr

'£r " t ^

^

or, in terms of loss tangent, D, =

4

—[(l + tan 2 S £ tan 2 ^ +tan2Ô„ +tan 2 S c )" 2 +tan<í t a n ¿ H - i l

(16)

where tan5E and tan5^ are dielectric loss tangent and magnetic loss tangent, respectively. As to microwave half-power depth, Dh, it is defined as the distance from the surface into the dielectric at which the traveling wave power drops to 1/2 from its value at the surface. From this definition, we have

53

or, in terms of loss tangent, =

ln(2H

T/j + tan2 4 tan2 Su + tan2 Su + tan2 £ f + tan .5; tan Su - l T " '

( 18)

Determination of Field Attenuation Length and Half-power Depth To show the variations of microwave field attenuation length and half-power depth with increasing temperature in magnetic materials, a typical American magnetite concentrate was used to do permittivity and permeability measurements and then determine the temperature dependence of field attenuation length and half-power depth at two most practical microwave frequencies, 915 MHz and 2450 MHz. The phase compositions were determined using a Scintag XDS2000 powder x-ray dif&actometer with a graphite monochromator and Cu Ka radiation and the X-ray diffraction pattern is shown in Figure 1. The analysis shows the sample contains 3 phases, mainly magnetite (Fe3Û4) with minor amount of quartz (SÍO2) and siderite (FeCCy.

29 (Degree) Figure 1. X-ray diffraction pattern of magnetite concentrate. Both the complex relative permittivity and permeability were measured using "cavity perturbation technique" [6,7], This method measures the difference in the microwave cavity response between a cavity with an empty sample-holder and the same cavity with a sampleholder plus the sample, and uses this to calculate either the permittivity or permeability, depending on the field type (electric or magnetic) in the region of the cavity in which the sample is placed. The technique measures the frequency shift and the change in loaded Q (quality factor) of a resonant mode of the cavity caused by inserting a sample. The details of the experiment and

54

the measured values of the real and imaginary parts of complex relative permittivities and permeabilities of magnetite concentrate at the given microwave frequencies (915 MHz and 2450 MHz) can be found in the published literature [6]. The experimental data of complex relative permittvities and permeabilities were used to calculate thefieldattenuation length and the half-power depth as a function of temperature at the specified frequencies, as shown in Figure 2 and Figure 3. It can be seen that the variations of microwave field attenuation length and the half-power depth follow the same "trend" in the whole temperature range and decrease with increasing temperature. Notice that there is only scale difference between the curves in Figure 2 and Figure 3. At the same temperature and microwave frequency, the half-power depth is found to be much smaller than the field attenuation length. It is associated with the Poynting vector, which indicates the microwave power attenuates much faster than the field strength in materials. Also, as indicated in both figures, the values at 915 MHz are much larger than those at 2450 MHz in the temperature range below ~ 550 °C and only small difference is found at higher temperatures. This phenomenon is believed to be closely related to dramatic increase of permittivity with temperature (not shown in this paper, see literature [6]) and the Curie point effect of magnetite (Tc= 585 °C). At high temperature (e.g. > 600 °C), the large permittivity presents dominating effect on the scales offieldattenuation length and microwave half-power depth. Furthermore, the loss of magnetism indicates the influence of permeability on the microwave dissipation in materials can be neglected in this temperature range. Hence, the evaluations offieldattenuation length and microwave half-power depth can be used to characterize the microwave dissipation behaviors in the sample, which may help in optimizing the dimensions of the material in microwave heating.

Figure 2. Variation of microwave field attenuation length (Df, meter) as a function of temperature.

55

Figure 3. Variation of microwave half-power depth (Dh, meter) as a function of temperature. Conclusions Based on Maxwell's equations, two simplified equations for determining microwave field attenuation length and half-power depth in magnetic materials were derived. The microwave field attenuation length and the half-power depth for a magnetite concentrate were calculated and the temperature dependence of them was determined. The variations of field attenuation length and the half-power depth with temperature at 915 MHz and 2450 MHz indicate both parameters are highly dependent on temperature and decreased with increasing temperature up to 900 °C. At the same temperature and microwave frequency, the half-power depth is found to be much smaller than the field attenuation length, indicating much faster power dissipation than field attenuation in materials. The determinations of field attenuation length and the half-power depth can help in optimizing the dimensions of the material in microwave heating. Acknowledgements The authors gratefully acknowledge the financial support from Michigan Public Service Commission, U.P. Steel, and the United States Department of Energy (DOE). References [1] D. E. Clark, D. C. Folz, and J. K. West, "Processing Materials with Microwave Energy," Mater. Sei. Eng. A- Struct, 287 (2000), 153-158. [2] Y. V. Bykov, K. I. Rybakov, and V. E. Semenov, "High-temperature microwave processing of materials," J Phys D: ApplPhys., 34 (2001), R55-75.

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[3] H. Fukushima, T. Yamanaka, and M. Matsui, "Microwave Heating of Ceramics and Its Application to Joining,"/. Mater. Res., 5 (1990), 397-405. [4] G. Xu, I. K. Lloyd, Y. Carmel, T. Olorunyolemi, and O. C. Wilson Jr, " Microwave Sintering of ZnO at Ultra High Heating Rates,"./. Mater. Res., 16 (2001), 2850-2858. [5] E. B. Kulumbaev, V. E. Semenov, and K. I. Rybakov, "Stability of Microwave Heating of Ceramic Materials in a Cylindrical Cavity," J. Phys. D: Appl. Phys., 40 (2007), 6809-6817. [6] Z. Peng, J. Y. Hwang, J. Mouris, R. Hutcheon, and X. Huang, "Microwave Penetration Depth of Materials with Non-zero Susceptibility," ISIJ Int., 50 (2010), 1590-1596. [7] C. A. Pickeles, J. Mouris, and R. M. Hutcheon, "High-temperature Dielectric Properties of Goethitefrom400 to 3000 MHz," J. Mater. Res. 20 (2005), 18-29.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooks TMS (The Minerals, Metals & Materials Society), 2011

VANUKOV FURNACE TECHNOLOGY: APPLICATION EXPERIENCE FOR PROCESSING DIFFERENT TYPES OF RAW MATERIALS AND GENERAL DEVELOPMENT TRENDS V.P Bystrov1, V.M. Paretskiy2, A.S. Vernigora1, R.I. Kamkin1, A.Y. Mamaev1, A.V. Kuznetsov1 'National University of Science and Technology "MISIS" Leninsky prospect, 4, Moscow, 119049, Russia 2 State Research Center of Russia Federation "Gintsvetmet" Instate Academika Koroleva Street, 13, Moscow, 129515, Russia Keywords: Vanukov furnace, Pyrometallurgy, Injection, Non-ferrous metals, Processing Abstract Vanukov furnace technology is an efficient, proven pyrometallurgical injection technology, used extensively in Russia and Kazakhstan for a number of different applications. The technology was most widely adopted for processing copper sulfide concentrates for matte in smelters in Norilsk, Revda (Russia) and Balkhash (Kazakhstan). The following applications also were developed and tested at industrial scale: treatment of sulfide lead and lead-zinc concentrates, latérite nickel ore, production of cast iron, treatment of antimony goldcontaining ores, and municipal solid waste. In this paper, current experience of Vanukov furnace application to these technologies is described with a number of general development trends and new areas of application. Introduction At the present time, the Vanukov process, developed at the Moscow State Institute of Steel and Alloys at the Department of metallurgy of non-ferrous, rare and noble metals with collaboration with other research organizations, is a highly efficient technological solution for processing different kinds of raw materials. The process was invented by Professor Andrey Vladimirovich Vanukov, whose research work began in 1949. From 1956 to 1975 test runs of the technology were carried out at the smelters in Norilsk and Balkhash. In 1976 a pilot Vanukov furnace was built in Ryazan. Furnace designs was modified to suit the various test programs. A major objective of each experiment was the determination of optimal working conditions for processing different types of raw materials such as copper, copper-nickel, copper-zinc, lead or pyrite concentrates, and materials containing secondary lead and zinc. After the death of Andrey Vladimirovich Vanukov in 1986, this technology and furnaces based on the technology were named after him: "Vanukov process" and "Vanukov furnace". Today, the following furnaces based on Vanukov technology are either working or were tested at semi-industrial scale: three in Norilsk, two in Revda (Russia) and two in Balkhash (Kazakhstan) for treatment of copper sulfide concentrates for matte, one in Orsk (Russia) for processing latérite nickel ore, one in the Buruktal plant (Russia) for antimony gold-containing ores for antimony oxide sublimates and gold-containing metallic antimony, and one in China for processing sulfide lead and lead-zinc concentrates for production of black lead and zinccontaining sublimates.

59

Process Overview The Vanukov process is a versatile pyrometallurgical technology for processing a wide range of raw materials into different types of products. A key principle of the process is the side injection of the blow into the melt via the tuyeres [1, 2]. A highly enriched (up to 90% oxygen) blow is normally used as an oxidizing reagent. Coal, oil or natural gas can be used as fuel. The ratio of charged fuel to blow creates the necessary oxidizing or reducing atmosphere in the reaction zone, depending on smelting requirements. Separation walls are empolyed to create separate zones within the furnace with different oxidizing-reducing atmospheres. Level of the blow injection divides the melt bath into 2 functional zones. The tuyeres are evenly distributed along the length of side walls and are located on the level of 0.5 - 1.0 m below the surface of the melt. The upper, or super tuyere zone, is a zone where all major chemical interactions, mass and heat transfer processes take place. The specific blow discharge is about 800 - 1200 m3/h per square meter of the horizontal section of the furnace. The combination of mentioned conditions is characterized by turbulent agitation of the melt with an energy of agitation of 50-80 kW/m3. Vigorous turbulent agitation of the melt and injection of fuel oxidizing reagent provides efficient heat and rapid chemical reactions, massheat transfer processes, coalescence of fine-dispersed suspended phases and formation of products. After formation products settle down to the under tuyere zone for settling and separation by gravity. After separation the products exit the zone through siphon skim holes located at opposite ends of the furnace. Process gases leave the furnace through an uptake located over the shaft of the furnace. The reaction shaft of the furnace is fully constructed from water-cooled elements without any brick lining in front of them. An accretion of frozen slag serves as a protection layer for the cooling elements during the furnace campaign [1, 2]. Vanukov furnace has the following advantages: • High productivity of the furnace. The technology can process up to 100-120 tons per day per square meter of the furnace • Efficient consumption of heat. The exothermic reaction of oxidation of combustible components from raw materials and fuel produce heat inside an agitated slag, effectively maintaining its temperature. • There are no strict requirements for feed preparation. Fineness and moisture content in feed can varied widely. The technology can accommodate moisture up to 6-8% and not be finely ground. • Low dust emission is typical. During smelting, because of low offgas velocity (1.5 - 2.0 m/s), dust emission is about 1.0 - 1.5%. • Campaign length. Current, confident campaign length of the furnace, applied for copper sulfide concentrates processing is about 2 years. There are possibilities for extending campain duration. • A wide range of raw materials can be processed using Vanukov furnace technology which leads to high flexibility and controllability of the process. Applications of Vanukov Furnace Technology

60

Aforementioned qualities permitted the design and develop a number of highly economic and environmentally efficient technologies for the treatment various kinds of raw materials based on the Vanukov furnace. The initial application of the technology was for processing of copper sulfide concentrates for matte. Today this application continues to be the major application for the technology. Copper Sulfide Concentrates Processing Processing of copper sulfide ores and concentrates using a Vanukov furnace began at the Norilsk Copper Plant in 1978. Today, smelters in Russia and Kazakhstan produce about 800 thousands tones of copper in matte from copper, copper-nickel and copper-zinc concentrates. Furnaces are installed in Norilsk Copper Plant (3 furnaces), SUMZ Copper Plant in Revda, Russia (2 furnaces) and Balkhash Copper Smelter in Balkhash, Kazakhstan (2 furnaces). Furnace Design An example of the Vanukov furnace for the treatment of copper sulfide concentrates is the Vanukov furnace .Nsl of SUMZ Plant (Revda, Russia). The furnace has an area of36m 2 (15 m long and 2.4 m wide). Longitudinal and crossal views are included in Figure 1.

Figure 1. Sections of Vanukov furnace. 1 - feed ports; 2 - matte siphon; 3 - uptake; 4 - shaft constructed from cooling elements; 5 - tuyeres; 6 - slag siphonsiphon; 7 - tuyeres for afterburning The reaction shaft of the furnace is fully constructed from copper cooling elements. Side wall cooling elements are arranged in three horizontal rows and 25 vertical groups. The first row is vertical while the second and third are sloped. Under the uptake there is a fourth row of cooling elements. At the slag end, 'beam' copper coolers are used in to form a partition wall. At the matte end, in its lower part, "beam' elements are used in the melt, and at the level of the sidewall in the second and third rows the same cooling elements as in sidewalls are mounted. The roof is also constructed from copper cooling elements. A refractory lining is not used in front of any of the coolers.

61

Feed ports are located at the matte end of the furnace, at the center of the roof. Blow is uniformly distributed along the length of the furnace, and discharged into the furance via horizontal tuyeres, installed in the first row of cooling elements. The enriched blow (about 90% 0 2 ) is injected into the agitating slag-matte emulsion. If needed, a gas burner can be inserted in the tuyere to compensate the lack of heat in the furnace. There are also additional tuyeres in the uptake, right over the fourth row of cooling elements. They are used to oxidize gases coming from the melt bath. Major slag and matte-forming processes take place in the super tuyere zone. The top surface of the agitated bath is located at the level of the center of the second row cooling elements. After formation, droplets of matte and slag filter down and form two layers on the hearth of the furnace, namely amatte layer on the bottom and slag layer above it. Liquid products exit the furnace through the skim holes under the partition walls. Skim holes are placed on different levels to prevent slag from exiting furnace at the matte skim hole. Operating Conditions General operating conditions of the furnace are listed in Table I. Table I - Parameters of Vanukov furnace work Parameter Value Feed rate 90-105 ton/h Blow rate 15900-16000 m3/h Oxygen content in blow 85% Natural gas rate -1350 m3/h Matte temperature 1170-1200°C Slag temperature 1250-1270°C Average product compositions are given in Tables II and III. Component Cu S Fe S¡02 CaO

Table II - Composition of liquid products Average content in slag (wt %) Average content in matte (wt %) 0.7 50 0.6 17 40.3 23.8 33.3 0.8 Table III - Composi tion of offgas Component Average content in offgas (vol %) 42.3 S0 2 7.9 C0 2 2.8 02

Vanukov furnace .N»l in SUMZ was launched in 1994. Common furnace campaign is about two years.

62

Sulfide Lead and Lead-Zinc Concentrates Processing For treatment of sulfide lead concentrates and lead-containing semi-products, a two-zone Vanukov furnace is employed. It consists of two chambers, one for smelting and one for reducing, with a common hearth. Both shafts are constructed with cooling elements, have their own uptakes, feed ports, injection and afterburning tuyeres. The furnace is also equipped with separate siphon for slag and metal (Fig. 2)

Figure 2. Two-zone Vanukov furnace for smelting sulfide lead concentrates and lead-containing semi-products Two operating zones are connected by a skim hole which provides directional flow of leadcontaining slag from the smelting zone to the reducing zone. Bottom cooling elements in both zones are penetrated by tuyeres for injection of blow into the slag melt. Top cooling elements have tuyeres for offgas afterburning. The outside wall of the reducing zone is has a siphon for continuous discharge of waste slag and is connected to the operating zone via a channel. For continuous discharge of the metal lead and slag from the siphons, tap holes placed at different levels. The two-zone Vanukov furnace permits realization of complete oxidation of sulphide lead concentrates in the smelting zone and reduction to produce a lead-rich slag in the reducing zone. Melting of the concentrate is carried out at temperatures of 1100 - 1150°C and reducion can cause slag to reach 1270°C. Metallic lead contains 0.5 % sulfur and almost all of the noble metals (90 %). Waste slag contains about 1 - 3 % Pb which responds to dissolved losses. The offgas of the first chamber rich with SO2 (20 - 25 %) is cooled, cleaned in electrostatic precipitator and sent to the sulfuric acid plant. Offgas from the second chamber contains combustible components which are burned using afterburning tuyeres. This enables some of the heat to be returned back to the melt. Fumes of metal zinc and lead are gathered from both the melting and reducing steps and sent back to the oxidation zone [3].

63

Latente Nickel Ore Processing The Vanukov process was used in OJSC "Ujuralnikel" Orsk city for processing latérite nickel ores to nickel matte. A two-zone Vanukov furnace was constructed for this purpose. The first zone was low-oxidizing where ore was smelted and the second reducing zone was used for sulfidizing the melt and dividing it to slag and matte. Basic coal and natural gas were used for fuel. Processing latérite ores this manner is more economic and envionronmentally efficient than shaft furnace technology [4, 5]. Antimony Gold-Containing Ore Processing The Vanukov process can be used for processing antimony gold-containing ores. The SakhaUral Antimony Plant (Svetly settlement, Orenburg region) for gold-antimony sulfide concentrate for the production of commercial grade antimony metal began operation in 1996. The Vanyukov process provides sublimation of volatile components, collection of precious metals in small amount of the metal alloy, and production of free-running slag with minimum content of precious elements. It represented a fundamentally new process flowsheet for the production of antimony. Concentrate melting is carried out at temperatures 1250 - 1300°C. Slag compositoin was: 21 -25 % Fe, 36 - 42 % Si0 2 , 7 - 1 2 % CaO. Sublimates contain: 92 - 96 % Sb203, 0,2 - 0,3 % Fe203, 0,2—0,8 % S, 0,6—1,0 % As, 0,5 - 1,5 % Si02. Reduction melting of the sublimates for rough antimony production and its subsequent fire refining is carried out according to well-proven technology. This type of Vanukov furnace (Fig. 3) has two feed ports: the first one in the roof and the second one in the frontal end of the furnace. The second feed port forms a type fore-chamber which allows reduction of entrained dust through the hole made particularly for removal of sublimates in the roof. On the opposite side of the fore-chamber is a skim hole which connects the Vanukov furnace to an electric settler where the melt is divided to slag, matte and metal smelt.

64

Figure 3 - Vanukov furnace for working antimony goldcontaining ores Cooling elements of the first row are made of copper, upper rows are made from steel. There are eight tuyeres in the side walls of the first row and eight afterburning tuyeres in the second row. Charred coal, steaming coal and oil can be used as a fuel. Discharge of the slag is via a tap hole in electric settler. Periodical tapping of matte and the metal smelt is made through the tap holes [6]. Iron Ore Processing (Slag Processing) The Vanukov process was also developed and widely tested for ferrous metallurgy at the Novolipetckij plant. This was called the "Romelt" process. It is used for producing cast iron without coke from different types of ores and also from ores containing significant amounts of zinc and lead. A Romelt furnace looks like a typical Vanukov furnace for smelting copper sulphide concentrates. The main difference is that in the Romelt process there is a reducing atmosphere in slag bath. The reducing atmosphere is generated by burning coal to produce CO in the bath and the afterburning CO to CO2 in the upper part of the furnace. This enables most of the chemical energy to be liberated as heat and causes the slag to reach a temperature ofaboutl500°C[7]. Coal Gasification in Slag Melt Coal gasification is a process of carbon oxidation to produce gases containing variable concentrations of CO and CO2. Coal gasification in a slag melt permits an environmentall safe and effective method for the producing of thermal energy from coal. In the Vanukov process for coal gasification a mixture of oxygen and air are blown into slag bath through the side tuyeres. Melt over tuyeres is stirred by bubbles which intensifies the agitation. Coal is loaded into the agitated bath. Coal particles are quickly oxidized in the melt to produce CO and CO2. The ratio of CO/C02 can be regulated by the composition of the blown mix [8]. Municipal Solid Waste Treatment At this moment a zero waste discharge scheme of solid waste treatment has been developed [9]. Requirements for the safe processing of dioxins and handling are provided by melting solid waste in the slag bath of a Vanukov furnace. This environmentally clean scheme was tested successfully during the work of an experimental-industrial plant in Ryazan in the 1990's. Contents of dioxins and harmful compounds were measured in the offgases while testing. These measurements showed a near zero discharge of of primary (1300°C) and secondary (450°C) toxic materials in dust [9]. Development Trends

65

There are several new technological developments underwayemploying the Vanukov process. One of them is continuous converting and slag reducing in copper production. This is a process for obtaining either rich copper matte or blister copper in one step in the first oxidating zone of the furnace, while rich copper slags are reduced in the second reductive zone. A second new application is for the production of construction materials. Slags from the Vanukov furnaces of different processes are chemically stable can be used as a base material for producing construction materials such as cement clinker, thermal insulating materials, cast stone materials and abrasive materials. Anotherway to improve Vanukov technology is to extend furnace campaigns. This is very important develpment which is now economically possible. Conclusion The Vanukov process is both a well explored and developed process. Different variantions of its use demonstrate its versatility and potential for processing various materials. This technology has been used for the production of commercial products from almost all feed materials (ore, flux and fuel). The Vanukov process has been demonstrated to be a stable, productive and ecologically safe metallurgical process, proven at many smelters. References 1. A.V. Vanukov, N.I. Utkin, Complex Processing of Copper and Nickel Ores (Chelabinsk, Russia: Metallurgy Publishing Company, 1988), 292-304. 2. A.V. Vanukov, ed., Smelting in Slag Bath (Moscow, Russia: Metallurgy Publishing Company, 1988). 3. U.P. Romanteev et al., Metallurgy of Lead (Moscow, Russia, MSISA, 2005), 68-84. 4. A.N. Fedorov, A.A. Komkov, V.N. Bruek, "Vanukov process development for latérite nickel ore processing in OJSC "Ujuralnikel"," Nonferrous Metals, Russian edition, 12 (2007), 3337. 5. V.l. Kostin, "Experience of Vanukov Furnace exploitation in OJSC "Ujuralnikel"," Nonferrous Metals, Russian edition, 11 (2008), 45-48. 6. A.N. Fedorov, D.K. Donskih, V.Y. Zajtcev, "Organization of antimony processing in SakhaUral Antimony Plant," Nonferrous Metallurgy Transaction. News of Higher Schools, 4 (1999), 16-19. 7. V.A. Romenets, Romelt Process (Moscow, Russia, MSISA, 2005). 8. A.V. Balasanov, A.V. Leherzak, V.A. Romenets, Gasification of coal in slag melt (Moscow, Russia, Stalproekt Institute, 2008). 9. V.M. Paretskij et al. "Waste Burning in Slag Melt," Municipal Solid Waste Journal, 9 (2009), 34-38.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

2nd International Symposium on

High-Temperature Metallurgical Processing

Microwave Heating and Iron and Steel Production Session Chairs: Chenguang Bai Jerome Downey

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cook TMS (The Minerals, Metals & Materials Society), 2011 A STUDY OF COAL-BASED DIRECT REDUCTION OF COMPOSITE BINDER MAGNETITE PREHEATED PELLETS Deqing ZHU1, Tiejun CHUN1, Vinicius Mendes u , Jian PAN1, Jian LI 3 'School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, Hunan, P.R. China; 2 Vale, Rio de Janeiro 20030, Brazil; 'Research Institute, Baoshan Iron & Steel Co., LTD., Shanghai, 201900,PR China Keywords: magnetite concentrate; direct reduction; preheated pellets; fired oxide pellets; coal-based grate-rotary kiln Abstract A study of the coal-based direct reduction behaviors of composite binder magnetite pellets was carried out in a simulating coal-based grate-rotary kiln process. It is shown that preheated pellets possess much better reducibility than fired oxide pellets: 40 min are required for preheated pellets to reach over 90% metallization degree compared to 100 min for fired oxide pellets. The compressive strength of preheated pellets decreases dramatically at the earlier stage of reduction, climbs quickly after reducing for 30min and achieves a high value at the end of reduction. However, the compressive strength of metallized pellets from reducing of fired pellets is much lower, more cracks and fractures being formed. In the preheated pellets, crystallite of metal iron is coarse. In the oxide pellet, a large number of interconnected crystals of metal iron are formed inside pellets, and pellet structure is very compact. Introduction Directly reduced iron is an excellent feed in electric arc furnace (EAF) steelmaking, and possesses low tramp element content and steady component. It is generally preferred in the production of high quality steel [1-3]. Nearly 64.44 million tons of DRI were produced in 2009 over the world, over 80% of which were produced by gas-based processes, the left by coal-based processes. However, the DRI capacity was only 600 thousand tons in China in 2009. The EAF steel production has exceeded 40 million tons annually, and most of EAF operate on hot metal as feed [4]. Extensive researches were investigated on DRI processes [5-7], including gas-based process and coal-based direct reduction of composite binder pellets. Due to lack of high grade lump iron ores and natural gas, more attentions have been paid to coal-based direct reduction of pellets processes in China. However, there are some disadvantages for the traditional coal-based rotary kiln process using fired oxide pellets as burden, especially the reduction degradation of fired pellets during reduction because of phase transferring from hematite to magnetite, leading to reduction swelling and kiln accretion. One DRI plant utilizing the latter process with an annual output of 150 thousand DRI has been put into operation since 2007, where grate-rotary kiln process was utilized and the kiln

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accretion was eliminated in the process of coal-based direct reduction of composite binder pellets [8]. Based on the previous researches of parameters optimization of coal-based direct reduction of composite binder pellets [9], the behaviors of oxide pellets and composite binder preheated pellets in coal-based direct reduction were conducted in this paper. Experimental Raw Materials The raw materials include magnetite concentrate, bentonite, and noncoking coal, which are supplied from Xinjiang. Composite binder FH (major components of sodium húmate) is produced according to the invention patent held by Central South University [10], Chemical analysis of the magnetite concentrate, bentonite and coal used in the experiments are shown in Tables I, 2 and 3, respectively. FC(total)

FeO

Table 1. Chemistry of magnetite concentrates (wt%). Si0 2 A1203 CaO MgO K,0 Na 2 0 Cu Pb Zn

69.21 27.01 2.67 FeÄ

0.46

0.68

LOI

0.45 0.011 0.012 0.00t 0.0086 0.021 0.042

Table 2 Chemical analyses of reduction coal and LOI (wf%). MgO P S CaO Si0 2 A1203 0.26

1.79

0.56

Binders

Fe 2 0 3

Bentonite

7.92

59.09

16.38

1.08

FH agent

236

27.92

13.01

0.38

0.27

1.11

0.013

0.37

0.32

LOI 95.51

Table 3. Chemical analyses of binders 'wi%). K:0 CaO MgO SiO : A1 : 0,

Na : 0

LOI

2.42

0.13

3.16

8.75

0.27

1.70

4.62

49.19

The fineness of magnetite concentrate is 55,5wt% and 80.1wt% passing 0.044 mm and 0.074 mm, respectively. The morphology of magnetite particles under SEM is shown in Fig.l.

( SEM*500 > (a)

( SEM*1O00 ) (b)

Figure 1. Morphology of magnetite concentrates particles under SEM.

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The coal samples produced from Xinjiang is of bituminous type. Its chemical analysis, industrial analysis and ash softening and melting properties were measure and tabulated in Tables 4 and 5, respectively. It can be seen that the Xinjiang coal is suitable for direct reduction. FC, ad/% 58.50

Table 4. Analysis of coal sample. V, daf/% M, ad/% A, ad/% 7.62

31.55

7.48

SI 3

*Footnotes: FC, ad=fixed carbon on air dry basis; M, ad=moisture on air dry basis; A, ad=ash on air dry basis; V, daf= volatile matter on dry ash free; SI=sticking index of coal slag. DT 1100

Table 5. Softening and melting properties of reduction coal (°C) ST HT FT 1170

1190

1260

*Footnotes: DT = distortional temperature, ST = softening temperature, HT = half global temperature, FT = flowing temperature Experimental Procedure The experimental flowsheet to simulate the innovative one-step process include proportioning pellet feeds, mixing, balling, drying, pre-heating of green balls in oxidizing atmosphere and reducing hot preheated pellets by using coal as reductant. Green balls were made of 1.5wt% composite binder FH and Xinjiang magnetite concentrate in a disc pelletizer (diameter 1000 mm, rim height 300 mm, angle of inclination 45°and rotating 28 rpm). The finished green pellets were dried in the oven at 105°C for 4h for further experiment. Preheating and reduction of pellets were carried out in stainless steel pot (diameter 65 mm, depth 100 mm) using an electrically heated tube furnace of 800 mm long and inside diameter of 80 mm . After stainless steel pot was heated at 900°C for 15 min, dry pellets were put in. When preheating of pellets was finished, reduction coals (mass ratio of coal to pellets is 2 in order to keep enough reducing atmosphere) were loaded into the stainless steel pot. The pot was moved to the highest temperature area of shaft tube furnace for reduction of hot preheated pellets to take place. The tests flowsheet to simulate the two-step process, i.e. the reduction of the fired oxide pellets is a little different from the above. Green balls were made of 1.5wt% bentonite and magnetite concentrate, and then dried for further preheating and firing in tube furnace to prepare fired oxide pellets for the following reduction tests. Green balls were hardened by preheating at 800°C for 10min, firing at 1150°C for 20 min and cooling to produce fired oxide pellets. The cooled oxide pellets would be loaded into the shaft tube furnace for reduction. Results and Discussions Reducibilitv of Pellets Reduction index against reduction time is shown in Fig.2. Reduction index

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increases dramatically with an extension of reduction time. Then the reduction index remains steady when reducing for about 60 min and 100 min for preheated pellets and fired oxide pellets, respectively. It can be concluded that preheated pellets possess better reducibility than fired oxide pellets.

0

Z)

60

«

80

1C0

Reduction time/nin

Figure 2. Reduction index against reduction time (reduction at 1050°C). Fig. 3 shows the effect of reduction time on metallization degree of reduced pellets. Metallization degree of both types of pellets is very similar to each other before reducing for 15 min. However, metallization degree of preheated pellets increases dramatically while metallization degree of oxide pellets increases slowly after reducing for 20 min. The metallization degree of preheated and oxide pellets amounts to 82.1% and 61.3%, respectively when reducing for 40 min. Especially the time for preheated pellets to reach the metallization degree of over 90% is only 50 min whereas over 80 min for oxide pellets. This agrees with that of Fig. 2, preheated pellets possess better reducibility.

Figure 3. Metallization degree vs reduction time (reduction at 1050°C). Compressive Strength of Pellets Fig.4 reveals the variations of the compressive strength of pellets with reduction duration. The compressive strength of both preheated pellets and fired oxide pellets decreases dramatically and drops to 50 and 210 N/pellet when reduced for 20 min, respectively. Compressive strength of preheated pellet drops from 540 N/pellet to 50 N/pellet and the minus gain is 90.3% while fired oxide pellet decreases from 3340 N/pellet to 210 N/pellet and the minus gain is 93.6%. However, the compressive strength of preheated pellet soars compared to a slow increase for oxide pellets after reduced for 40 min. The final

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compressive strength of metallized pellets from preheated pellets and oxide pellets after reducing for 100 min is 2590 N/pellet and 850 N/pellet, respectively. This agrees with Fig. 2 due to superior reducibility of preheated pellets. Also, there is a close relationship between metallization degree and compressive strength of pellets during reduction. There are other factors affecting the compressive strength of reduced pellets because there is a huge difference in compressive strength of the two types of reduced pellets.

0

20

40

60

80

100

Reduction time/min

Figure 4. Compressive strength vs reduction time (reduction at 1050°C). Unbroken Pellets Proportion The dramatic decrease in the compressive strength of reduced pellets is partly ascribed to some breakage and cracks of pellets forming during reduction. As shown in Fig.5, there is little difference in the unbroken pellets proportion before reducing for 20 min, then the unbroken pellet percentage of preheated and fired oxide pellets decreases with an extension of reduction time and drops to 67% and 55% for about 30 min, respectively. However, the unbroken proportion of preheated pellets reaches 100%, which means cracks of preheated pellets disappeared at the end of reduction. Compared to reduced pellets from preheated pellets, more reduced pellets from oxide pellets are broken, as shown in Fig. 6. More cracks and broken pellets mean that higher degradation occurs due to the reduction of hematite into magnetite, leading to much higher risk in kiln accretion. This probably can explain why the kiln accretion can not be eliminated in the traditional two step direct reduction process, where oxide pellets are used as burden.

0

20

40

60

80

100

Rotation timefain

Figure 5. Comparison of unbroken pellet percentage between preheated pellets and fired oxide pellets (reduction at 1050°C).

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(a)

(b)

Figure 6. Photos of reduced pellets after reduction at 1050°C for 100min, (a) reduced product of preheated pellets; (b) reduced product of oxide pellets. Metal Iron Grains Growth Metal Iron grains in the preheated and oxide pellets reduced at 1050°C for different time were examined by using optical microscopy and depicted in Fig.7. When reduction time reaches 20 min, there is some metal iron in preheated pellets and 40min are required for the oxide pellets. In the preheated pellets, crystallite of metal iron is coarse. In the oxide pellet, a large number of interconnected crystals of metal iron formed inside pellets, pellet structure is very compact.

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Figure 7 Photomicrograph of preheated and oxide pellet reduced with non-coking coal (white dots are metal iron) Conclusions Preheated pellets possess much better reducibility than fired oxide pellets. 60 min are required for preheated pellets to reach over 90% metallization degree compared to 100 min for fired oxide pellets. As fired pellets during reducing at the earlier stage, the compressive strength of preheated pellets drops dramatically. However, compressive strength of preheated pellets climbs quickly and achieves a high value at the end of reduction. In the preheated pellets, crystallite of metal iron is coarse. In the oxide pellet, a large number of interconnected crystals of metal iron formed inside pellets, pellet structure is very compact. References 1. Yang. J, Tomoyuki Mori, and Mamoru Kuwabara, "Mechanism of Carbothermic

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Reduction of Hematite in Hematite-Carbon Composite Pellets," ¡SUM., 47(2007), 1394. 2. Qiu. GZ, Jiang T, and Xu. J.C, Direct reduction of cold-bonded pellets, (Changsha, Central South University Press, 2001), 125. 3. Kotaro Ishizaki and Kazuhiro Nagata, "Production of Pig Iron from Magnetite Ore-Coal Composite Pellets by Microwave Heating," ¡SUM., 46 (2006), 1405. 4. Xu. K.D and Yin. R.Y, Proceedings of Chinese EAF Fowsheet and Engineering, (Beijing, Metallurgy Industrial Press, 2005), 234. 5. Duan. D.P, Wan. T.J, and Ren. D.N, "New Process for Coal-based Direct Reduction of General Grade Iron Ores," Iron & Steel, 8(8) (2001), 7-9. 6. Li. Y.Q, Chen H, and Zhou. Y.S, "Study and Development of BL Direct Reduction Process," Iron & Steel, 34(2) (1999), 6. 7. Tao. J, et al, Direct Reduction of Composite Binder Pellets and Use of DRI, (Ahmedabad, India, Electrotherm Press, 2007), 345. 8. Zhu D.Q et al., "One-step process for direct reduction of Xinjiang magnetite concentrate," J. Cent. South Univ. Sei. Technol, 38(3) (2007), 42. 9. Zhu. D.Q, Xu. X.F, and Ou. Y.Q, "One-Step Direct Reduction of Damp Milled Magnetite Concentrate Pellets," Iron & Steel, 42(1) (2007), 6. 10. Xu.J.C, et al., Composite Binder of Pellets Coal-Based Direct Reduction, Patent No: CN1088618, Central South University, (1994).

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

MICROWAVE DIELECTRIC PROPERTIES OF PYROLYZED CARBON Zhiwei Peng1, Jiann-Yang Hwang1, Wayne Bell1, Matthew Andriese1, Shuqian Xie1'2 'Department of Materials Science and Engineering, Michigan Technological University, Houghton, MI 49931, USA department of Metallurgical Physical Chemistry, Northeastern University, Shenyang, Liaoning 110004, P. R.China Keywords: Activated Carbon, Permittivity, Dielectric Polarization, Relaxation Time Abstract Pyrolyzed carbons are generally known as good microwave absorbers and their dielectric properties still remain to be fully explored. In the present study the dielectric properties and dielectric polarization-relaxation phenomenon of a typical activated carbon was investigated. The experimental results indicate the complex permittivity is highly dependent on temperature and frequency. The decrease of permittivity with increasing temperature from room temperature to -100 °C is probably ascribed to the release of the water vapor adsorbed on the surface of activated carbon, and the variations of permittivity with temperatures between 100 °C and 450 °C in the frequency range of 300 MHz to 3000 MHz are mainly attributed to the decreased relaxation time of dielectric polarization. Introduction Pyrolyzed carbons have been identified as excellent microwave absorbers owing to their high dielectric and electrical loss under microwave irradiations. Among them, activated carbon continues to attract more and more attention from researchers. Due to its special physical properties, the production, modification and regeneration of activated carbon with the assistance of microwave have been widely studied over the last few decades [1-6]. However, only few attempts have been applied to explore the dielectric properties of such materials due to the complexity and variability of their organic precursors and activation processes [7,8]. In order to shed a light on this problem, a typical sample, Norit® A Supra USP powdered activated carbon, was adopted to perform dielectric characterization, which may provide useful information for its wide application in the fields of purification, energy (hydrogen) storage and so on. For general dielectric materials, susceptibility to microwave heating is governed by the frequency and temperature dependent complex permittivity [9]: e{(o) = £(a>)-je(a>) = j£0{er((o)-jer(m)\

(1)

where £o is the permittivity of free space (8.86><10"12 F/m), e,' is the real component of the complex relative permittivity, e" is the imaginary component of the complex relative permittivity, and j= (-1)"2. According to Debye theory, the real part of permittivity and imaginary part of permittivity are highly dependent on the frequency and can be described as

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£'(ß>) = £. + ,'

2~2

£"(«) = H—ri— l + WT

(2)

(3)

where &» is permittivity at the high frequency limit, es is the static permittivity and x represents the polarization relaxation time, which is the inverse of the frequency of maximum dielectric loss [10]. Above equations do not give the obvious relationship between the permittivity and temperature. However, from dielectric theory, it is expected that the relaxation time (t) may be significantly affected by the temperature variations, depending on the type of materials. In this work, the activated carbon was taken as an example to make a quantitative examination of such temperature effect. Experimental The high surface area carbon samples, which were activated through steam, were purchased from Norit Americas Inc (USA). The microstructure was imaged by Hitachi S-4700 field-emission scanning electron microscope (FE-SEM). The BET surface area and pore distribution were characterized by using a Micromeritics ASAP 2000 instrument. The complex permittivity of the activated carbon was measured by the cavity perturbation technique, which is based on measuring the difference in the microwave cavity response between a cavity with an empty sample-holder and a cavity with a sample-holder plus the sample [11]. The high temperature measurements (from room temperature in -50 °C steps to -450 °C) were performed by moving the hot naturally packed sample under stagnant air atmosphere and its holder rapidly from a conventional furnace into the high electric field (central) region of a thickwalled, well-cooled TMono cavity with connection of a Hewlett-Packard 8753 vector network analyzer, which was used to transmit electromagnetic wave transmission and the detection of frequency shift resulting from the insertion of samples. The cavity modes used in this experiment include six frequencies (397, 912, 1429, 1948, 2466, and 2986 MHz) covering the approximate range between 300 MHz and 3000 MHz. The detailed descriptions of the measurement technique and the apparatus are found in the published literature [12]. Results and Discussion The porous structure of the activated carbon is shown in Figure 1. As revealed by the images, enormous pores distribute in carbon particles, whose sizes are ranging from 10 um to 90 urn. The BET surface area is 1836 m2/g, which contributes to the superior adsorption capability for gas and liquid phases.

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Figure 1. FE-SEM images of the activated carbon, (a) *2,500 magnification, (b) x 10,000 magnification. The temperature and frequency dependent complex permittivity was determined using cavity perturbation technique. The experimental results indicate, as expected, the permittivity including both its real and imaginary parts highly relies on the temperature and frequency (Figures 2 and 3). For the real component of complex relative permittivity over the temperature range measured and frequencies covered, er', it remains the maximum values at room temperature. The effect of temperature can be illustrated according to the analyses of the two different regimes: stage I, room temperature to -100 °C; and stage n , 100 to -450 °C. This feature can also be observed through the plot of sr" versus £r' (Figure 4). In stage I, the values of er' ands r " decrease with increasing temperature, which is probably due to the release of the water vapor adsorbed by the activated carbon from atmosphere. As discussed before, the activated carbon exhibits excellent adsorption capability on gas and liquid phases. Although the amount of moisture is very small compared to that of activated carbon (no apparent weight loss was detected in the permittivity measurement), the effect of water due to its intense interaction with microwave on the permittivity of carbon should be considered. In stage n , both er- and er show few variations (small decrease for both parameters) during the temperature range. There are two factors that may account for this phenomenon. The primary factor should be associated with the completed removal of water in this range, which contributes partly to the stable permittivity. The other factor may be related to the decreased density presented during the measurement process. Since the sample was naturally packed in the sample holder, it is possible that the carbon expands at high temperature for its much higher temperature expansion coefficient than the holder made from quartz. This effect was confirmed by the increased length of the bulk of the sample after the permittivity measurement. As to the effect of frequency, it is obvious that the permittivity decreases with increasing frequency. This should be attributable to the increase of angular frequency (
79

Figure 2. Real component of complex relative permittivity as a function of temperature and frequency.

Figure 3. Imaginary component of complex relative permittivity as a function of temperature and frequency.

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Figure 4. Plot of sr" versus Er'. To provide further understanding of the dielectric properties of the activated carbon, a corresponding analysis of dielectric polarization-relaxation has been performed. From the equations discussed previously, the characterization of dielectric polarization-relaxation can be indicated by the relationship between real and imaginary components of complex permittivity. As shown in Figure 5, a plot of s r ' versus (er"/co) was depicted. According to this plot, the e„ (the intercept of plot) and T (the inverse of the slope of the plot) were determined. The static permittivity, ES, was calculated. The variations of these parameters along with temperature are summarized in Table 1. It indicates the relaxation time mainly decreases with increasing temperature until ~250 °C. At even higher temperatures the values of calculated relaxation time become fluctuated, indicating the weak temperature dependency of dielectric relaxation of activated carbon at temperatures between about 250 °C and 450 °C.

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Figure 5. Plot of £r' versus (sr"/co). Table 1. Dielectric Polarization-relaxation Parameters of the Activated Carbon T(°C) T(10"VS) Es/Eo Eoo/ £o 9.6782 24 87.75381 1.26997 167.9073 7.7323 96 3.36542 407.499 7.4802 147 7.98977 7.2274 196 174.2775 3.21906 114.9507 6.8939 246 1.98043 52.2905 6.4283 298 0.76664 346 94.88393 7.3059 1.39303 396 0.98717 73.62006 6.6111 86.73684 443 1.21890 7.0016 Conclusions The permittivity of activated carbon was determined from room temperature to about 450 °C and from about 300 MHz to 3000 MHz. The variation of complex permittivity below 100 °C (stage I) is probably attributed to the release of water vapor adsorbed by the carbon and above 100 °C (stage n), the sluggish decrease of permittivity with increasing temperature is partly ascribed to the decreased density of the bulk of the sample. It is also found that the permittivity decreases with increasing frequency over the whole temperature range and the imaginary permittivity is more strongly dependent on frequency than the real permittivity. The analysis of dielectric polarization-relaxation shows the effect of water vapor dominates the relaxation behaviors at first (stage I), followed by a decrease of relaxation time with increasing temperature until ~250 °C.

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There is only a weak temperature dependency of dielectric relaxation of activated carbon at even higher temperatures (250- 450 °C). Acknowledgements The authors would like to thank Michigan Public Service Commission, U.P. Steel, and the United States Department of Energy (DOE) for their financial support and Microwave Properties North, Canada, for the help in permittivity measurement. References [I] L. M. Norman and C. Y. Cha, "Production of Activated Carbon from Coal Chars Using Microwave Energy," Chem. Eng. Comm., 140 (1996), 87-110. [2] J. A. Menéndez, E. M. Menéndez, M. J. Iglesias, A. Garcia, and J. J. Pis, "Modification of the Surface Chemistry of Active Carbons by Means of Microwave-induced Treatments," Carbon, 37 (1999), 1115-1121. [3] J. A. Menéndez, E. M. Menéndez, A. Garcia, J. B. Parra, and J. J. Pis, "Thermal Treatment of Active Carbons: A Comparison between Microwave and Electrical Heating," J. Micron: Power Electromagn. Energy, 34 (1999), 137-143. [4] J. M. V. Nabais, P. J. M. Carrot, M. M. L. R. Carrot, and J. A. Menéndez, "Preparation and Modification of Activated Carbon Fibers by Microwave Heating," Carbon, 42 (2004), 13091314. [5] H. M. Williams and G. M. B. Parkes, "Activation of A Phenolic Resin-derived Carbon in Air Using Microwave Thermogravimetry," Carbon, 46 (2008), 1169-1172. [6] F. K. Yuen and B. H. Hameed, "Recent Developments in the Preparation and Regeneration of Activated Carbons by Microwaves," Adv. Colloid Interface Sei. 149(2009), 19-27. [7] J. E. Atwater and R. R. Wheeler Jr, "Complex Permittivities and Dielectric Relaxation of Granular Activated Carbons at Microwave Frequencies between 0.2 and 26 GHz," Carbon, 41 (2003), 1801-1807. [8] J. E. Atwater and R. R. Wheeler Jr, "Micorwave Permittivity and Dielectric Relaxation of a High Surface Area Activated Carbon," Appl. Phys. A, 79 (2004), 125-129. [9] A. Von Hippel, Dieletrics and Waves (New York, Wiley Press, 1954). [10] P. Debye, Polar Molecules (Pennsylvania, Lancaster, Lancaster Press, 1929). [II] C. A. Pickeles, J. Mouris, and R. M. Hutcheon, "High-temperature Dielectric Properties of Goethite from 400 to 3000 MHz," J. Mater. Res. 20 (2005), 18-29. [12] Z. Peng, J. Y. Hwang, J. Mouris, R. Hutcheon, and X. Huang, "Microwave Penetration Depth of Materials with Non-zero Susceptibility," ISIJ Int., 50 (2010), 1590-1596.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

FUGITIVE EMISSIONS RELATED TO OXIDATION OF LIQUID SILICON DURING LADLE REFINING Mari K. Nsss, Gabriella M. Tranell, Nils Eivind Kamfjord Norwegian University of Science and Technology (NTNU); Alfred Getz vei 2, NO-7491 Trondheim, Norway. Keywords: Liquid silicon, oxidation, ladle refining, condensed silica fume, boundary layer Abstract In oxidative ladle refining (OLR) of silicon, the metal surface is oxidized resulting in the formation of a condensed silica fume (SÍO2). In the current work, industrial measurement campaigns were performed aiming to measure the fume generation during OLR. A thorough discussion of the possible mechanisms has been included in order to improve our understanding of the Si(i)-02(g) system. The measurement campaigns were performed at the Elkem Salten plant in Norway. In addition to fume generation from OLR, the metal temperatures and ladle purge gas amount were recorded. The results of this work suggest that fume generation during OLR results from splashing of the metal and/or oxidation of the metal surface, with oxygen transport to the metal surface being the limiting factor. Other mechanisms of SÍO2 formation were investigated, however insignificant.

Introduction One of the main environmental and economical challenges facing the metallurgical industry is fugitive emissions of both materials and energy. In the production of metallurgical grade silicon (MG-Si), refining of the silicon generally takes place through an oxidative ladle treatment (purging with an air-Û2 mixture). In the refining process, exposure of silicon to air results in the formation of condensed silica fumes (SÍO2) - one of the main sources of fugitive emissions in silicon production plants. The oxidation of liquid silicon, as opposed to solid silicon, has not been widely studied and described in the literature [1], thus a thorough investigation of the oxidation mechanisms, and factors affecting the oxidation reaction rate are essential in order to deal with the problem. In the current work, the rate of liquid silicon oxidation has been studied through industrial measurements of fume generation from refining

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ladles, where the purge gas rate and composition have been varied. The results from these measurements were discussed in light of existing models describing the liquid silicon oxidation kinetics.

Theory The oxidation of liquid silicon is widely accepted as a process consisting of two steps: formation of volatile SiO gas, followed by combustion in air to form SÍO2 [2]:

SiOig)+iOKg)^Si02(l)

(2)

The heterogeneous reaction between oxygen and liquid metal takes place at a phase boundary. As illustrated by Figure 1, the reaction involves mass transport of gaseous or liquid species from the bulk through a boundary layer to the interface, where the reaction takes place. The chemical reaction is generally considered to be very fast due to the high temperature, thus the rate limiting step for the oxidation is mass transfer of reactants to the interface and/or reaction product from the interface [3]. The process is thus referred to as diffusion controlled. In addition, there will be a counter flux of carrier gas (nitrogen in the case of oxidation in air), and also a condensed phase of SÍO2 may be present. All of these species constitute the boundary layer through which the oxygen must diffuse to reach the reaction interface.

Figure 1 : A sketch of the mass transport mechanisms in the system. The C¡" are the bulk concentrations. The C¡ are the concentrations at the interface, and 8 is the thickness of the boundary layer.

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The diffusion of Si(i) and 02
pSi0(eq),

from the reaction:

the surface of the silicon will stay uncovered by SÍO2 and SiO will continue to form. This is the state of active oxidation, where the rate of attack is increasing with increasing oxygen partial pressure. When the partial pressure of SiO reaches the pSICKeq),

a layer of SÍO2 may be formed at

the surface, and there is a transition from active to passive oxidation. Wagner makes an important assumption prior to the derivation of the equations for the rate of oxidation; the reaction between Si and O2 has a low activation energy, thus the reaction is rapid and the partial pressure of oxygen at the surface, p'0 is low, - much lower than the bulk partial pressure, p'0 . The derived equation for the consumption of silicon in gram-atoms per unit area per unit time,/», during active oxidation is according to Wagner

/•. =

Js>

2p„ °Dn ' ' S0iRT

(4)

where D0 is the diffusion coefficient of oxygen through the boundary layer with thickness S0 , R is the gas constant and T is the absolute temperature. The rate of oxidation is determined by the supply of oxygen for the formation of SiO. When p°0 becomes larger than p¿(max), the rate of attack becomes

"

SSl0RT

(p0^Ssl0RT

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where Dsto is the diffusion coefficient of SiO molecules, Sso is the effective thickness of the boundary layer for diffusion of SiO, p'slo

is the SiO pressure at the surface and K is the

equilibrium constant for the reaction in equation (2). The partial pressure of SiO at the surface will increase linearly withp¿ , and is equal to the equilibrium partial pressure (1.5*10 2 atm, at T = 1410°C) when the oxygen pressure reaches maximum (6.1* 10"3 atm, at T = 1410°C). Here, the SiO pressure drops to a value eight orders of magnitude lower. As seen in equation (5), the steady state rate of attack is proportional to the surface partial pressure of SiO, thus the rate will drop when the SiO pressure drops [5]. Wagner does not discuss the fate of the SiO gas except when the partial pressure of oxygen reaches the critical point of formation of SÍO2. It might be crucial to take into consideration the counter flux of N2 and SiO inside the boundary layer, and also the possible presence of SÍO2, which would be a fog of condensed phase - oxidized SiO gas. Hence, the system is probably more complex than pictured by Wagner, and the flux of oxygen is most likely altered by these influences. Ratto et al. [1] extended Wagner's theory and made it more general. With extensive derivations they took into account the counter flux of SiO-gas and also derived a transport regime with a heterogeneous boundary layer, where condensed SÍO2 was present. They considered the two limiting cases of null reaction and instantaneous reaction in the boundary layer, with local equilibriums. They compared their results to Wagner's results, and found that the fundamental concept is in complete agreement; however, the fluxes of each chemical species in the layer will be entirely different, due to the reactions in the boundary layer [1]. This discrepancy is shown schematically in Figure 2.

Figure 2: Pressure profiles of the species in a heterogeneous boundary layer. Ç is the nondimensional distance from the surface; z/S. Dotted lines = no reactions (Wagner), solid lines = instantaneous reactions in the boundary layer. Modified from Ratto et al.[l].

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In the heterogeneous layer with instant reactivity case, i. e. the SiO oxidizes to SÍO2 instantly as it meets oxygen on its way out of the layer, Ratto et al. came to an important conclusion: the oxygen concentration at the surface of the molten silicon will be very low and does not change with increased bulk concentration of oxygen. An increase in the bulk oxygen partial pressure will only increase the oxidation/combustion of the SiO gas. This implies that saturation will never occur and active oxidation will just go on without passivation [1]. The relative pressure profiles of oxygen, SiO and SÍO2 in the heterogeneous system with null and instant reactions can be seen in Figure 2. In Wagner's approach (null reaction, dotted lines) the concentration profiles are linear, and in the case of instantaneous reaction they are strongly non-linear. In the industrial ladle refining case, more complexity is introduced to the system as an oxygen/air mixture is purged through the metal from the bottom of the ladle. This generates a turbulent environment for the oxidation, combined with the continuous exposure of new silicon surface to the ambient air.

Experimental Two industrial measurement campaigns were carried out at Elkem Salten in April and August, 2010. The tapping of the silicon was discontinuous during the measurements, in order to record only the fume formed during refining, rather than the fume from the tapping. The rate and composition of the purging gas were varied in order to see whether this would affect the rate of the fume formation. Videotapes of the refining process, in addition to measurement of the amount of SÍO2 formed, were recorded. The equipment used to measure SÍO2 amounts in the off gas channel was a LaserDust MP with process parameters 200°C, 1013 mbar and 1.2 m measuring path. The apparatus had a relative error of 2 % in the measurements, however it was not calibrated for this kind of fume particles. It is a reasonable assumption that the measurements have an absolute error of 30% [6]. At the time of the measurements, the temperature in the metal was in the range of 14501500°C. Gas temperature measured at -0.5 m above the metal was typically in the range of 370500°C. Figure 3 shows the location of the LaserDust MP, and how a ladle top/silicon surface typically looks like during refining.

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Figure 3: Left: A photo of the top of a ladle with silicon being refined, showing the dynamic nature of the silicon surface and the air above. Right: Photo of the LaserDust MP installed at the off-gas channel. There is a layer of slag and solidified silicon along the perimeter of die silicon, and the exposed area of liquid silicon was typically in the range 0.3-0.5 m2, when we take into account that the surface is curved up convexly from the stirring of the refining bubbles.

Results The average amounts of silica fume measured in the off-gas are shown in Table 1, and plotted against the amount of oxygen and total gas flow in Figure 4. For the varying total gas flows, the measured amounts correspond to 31-71 mg/m2s of SÍO2 (0.051-0.119 moles/m2s of SÍO2). It indicates that the rate of oxidation is increased with increasing total gas flow; however, it is neither significantly affected by the concentration of oxygen in the gas nor of the oxygen flow rate. Table 1 : The numerical values from the silica fume measurements, with calculated standard deviations in parentheses. Amount of oxygen and total gas flow rate is also included. Measurement M 1 2 3 4 S 6_ Si02, mg/Nm3 111 (20) 126(15) 71.9(9.9) 127 (8.0) 96.9 (8.2) 65.9(10) 2 S¡02, kg/m h 14.6(2.7) 25.6(3.1) 14.6(2.0) 21.3(1.3) 16.2(1.4) 11.0(1.7) Oxygen flow, m3/h 5 10 5 4 18 4 Total gas flow, m3/h 15 24 8 21 22 8

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Figure 4: Amount of oxygen and total amount of gas in the refining, plotted against the measured amount of silica fume (kg/m2h). If we use these values of fume amount, and a diffusion coefficient for oxygen of £>0=3.84 cm2/s from Ratto et al. [1], we can calculate the thickness of the boundary layer at 1500°C using Wagner's equation for active oxidation (eq. (4)). The thickness is then in the range 0.93-2.2 cm, assuming that the temperature does not change significantly within the boundary layer (which would affect the diffusion coefficient). The thickness of the boundary layer decreases with higher total gas flow. In an initial CFD study, it was found that the supply of silicon to the surface was 10 kg/s. This is several orders of magnitude larger than the measured amounts of fume, thus it may be concluded that the availability of silicon is not limiting in this case [7]. Discussion In the discussion of the formation of condensed silica fume during OLR, it is also important to consider the different paths in which the silica fume formation in the industry may occur. Three different oxidation paths may be envisaged, as outlined in Figure 5. The first scenario is that the SiO-gas formed in the refining bubbles is oxidized as the gas is released at the top. The second possibility, referred to as splashing, is where the bubbles from the purging gas will drag silicon up in the air, in which the droplets will oxidize to SÍO2. The last possibility is the surface oxidation, where the continuous exposed surface of liquid silicon is oxidized by the oxygen in the air above. As the rate of formation of silica fume is independent on the amount of oxygen in the refining gas, it is probably the dynamic mechanisms in Figure 5b and c which are most dominant.

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Figure 5: Simple sketch of the possible macroscopic mechanisms in the ladle: a) oxidation of SiO in the bubbles from the refining, b) oxidation from splash of liquid silicon due to the drag of the purge gas (PG) in the bubbles, and c) surface oxidation when in contact with oxygen in air. In addition to the results from the industrial measuring, the amount of SÍO2 formed due to SiO from the bubbles was estimated using Tang's equation for the SiO-pressure in the bubbles [3]: exp(11.13 + 39464/r 7exp(-23.66 + 113623/r where T is the absolute temperature. The estimated amount from equation (6) was orders of magnitude smaller than the measured amounts of fume, thus found insignificant. This supports our result that the amount of silica fume formation is not significantly affected by the oxygen amount in the purging gas in the present set-up. Even though we can exclude SiO from the refining bubbles as a dominating macroscopic mechanism, we cannot decide which of the other two dynamic mechanisms is dominating. The splash of metal caused by bubble drag will increase with increasing gas flow rate, however, the surface area of liquid silicon available for oxidation will also increase. Thus we cannot say with certainty which of the two dynamic mechanisms is dominating, and both may contribute to the silica fume formation.

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A problem with comparing the models of Wagner and Ratto et al. with the system in a real situation in the ladle is that the silicon surface is turbulent and moving, which leads to active oxidation even though there should have been a passivation, according to Wagner. The complexity of the boundary layer most likely slow down/hinder the oxygen transport to a large extent, making the supply of oxygen to the surface low enough to get active oxidation. Although Ratto et al. operates with much lower bulk oxygen pressures than ambient air in their derivations, it might be that their theory about the heterogeneous layer with instant reactions applies in our system. The oxygen in the air which is transported through the boundary layer meets a flow of SiO and is instantly consumed to such an extent that the true oxygen partial pressure at the silicon surface is in the active oxidation range.

Conclusions Results from industrial measuring campaigns where oxidation rate have been measured, show that the rate of oxidation of liquid silicon increases with increasing gas flow in the refining process. Theory and preliminary CFD modeling suggests that the rate limiting step in the surface oxidation is the oxygen transport to the surface, however, the dominating dynamic mechanism is notfinalfrom these results. Further evaluation of this system should be considered by doing modeling with CFD and more advanced theoretical calculations. Examination of collected silica fume from the off-gas channel with SEM is planned, in order to investigate the morphology and thus the dynamic mechanisms. In the future we intend to perform small scale laboratory experiments, in order to investigate the system in a more controlled environment.

Acknowledgements The financing of this work was provided by the Norwegian Research Council and FFF (Ferro Alloys Industries Research Association) through the FUME project (Fugitive emissions of Materials and Energy). We would like to express our appreciation to SINTEF for their contribution to this work, and Elkem for allowing us to carry out the measurements at their plant in Salten. Also we would like to thank Norsk Elektro Optikk AS (NEO) for using their equipment, and for assisting in the technical part of the measurements.

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References 1

Ratto, M., Ricci, E., Arato, E. & Costa, P. (2001) Oxidation of metals with highly reactive vapors: Extension of Wagner theory Metallurgical and Materials Transactions B 32, 903911.

2

Gelain, C , Cassuto, A. & Le Goff, P. (1971) Kinetics and mechanism of low-pressure, high-temperature oxidation of silicon, Oxidation of metals 3, 139-151.

3

Bongiorno, A. & Pasquarello A. (2005) Atomic-scale modelling of kinetic processes occuring during silicon oxidation, Journal of Physics: Condensed Matter 17, 2051-2063.

4

Tang, K. (2008) Thermodynamic analysis of oxidative ladle refining of silicon melt including models of thermophysical properties of the silicon melt and SiOi-Al]P}-CaO slag, SINTEF Report

5

Wagner, C. (1958) Passivity during the oxidation of silicon at elevated temperatures, Journal of Applied Physics 29, 1295-1297.

6

Personal communication with Norsk Elektro Optikk AS (NEO), (2010).

7

Olsen, J. E. , (2010), MEMO: CFD perspective on dusting from Si-ladles, (Unpublished Work), SINTEF

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooks TMS (The Minerals, Metals & Materials Society), 2011 REDUCTION KINETICS OF IRON OXIDE IN CaO-Si02-Al 2 0 3 -Fe x O-C MIXTURES Yuanyuan Zhang and Patrick J. Masset1 Freiberg University of Mining and Technology Centre for Innovation Competence VIRTUHCON, Group "Multiphase Systems" Fuchsmühlenweg 9, Reiche Zeche, D-09596 Freiberg, Germany Key Words: FexO reduction, kinetics, slag Abstract The reduction of FexO in slag is one of the fundamental reactions in iron and steelmaking technology, including the blast furnace process (BF). Experiments on the reduction reaction of iron oxide using CaO-Si02-Al203-FexO-C mixtures were carried out using thermogravimetric analysis (TGA) under reducing conditions at temperatures up to 1100°C. The composition of the primary slag was chosen so that the FexO concentration varied between 0 and 75 wt % and the CaO/SiC>2 ratio being kept constant to 0.90. From the TGA curves recorded isothermally at different temperatures the kinetic parameters of the iron oxide reduction were determined. Introduction In blast furnace the iron ores and coke are charged in layers from the top. The reduction of iron oxide by coke is one of the fundamental reactions in the blast furnace process. Although a great amount of work has been done on the topic of reduction of composites [1, 2, 3], gasification of carbon [4, 5] and reduction of iron oxides by CO [4, 5, 6], the mechanism of reduction in iron oxide - carbon powder mixtures is not fully understood. Fruehan [1] and Rao [2] studied the reaction rate of the reduction of iron oxide by different types of carbon and concluded that the reaction mechanism is controlled by the oxidation of carbon at temperatures up to 1200°C. Rao [2] studied the reaction of hematite with graphite as carbon source up to 1080°C. An activation energy close to 300 kJ/mol was calculated within the temperature range of 950°C to 1087°C. In his study, Fruehan [1] determined the rate of reduction of Fe2Û3 and FeO by coal, coke and coconut charcoal as carbon sources within the 900°C to 1200°C temperature range. He found that the reduction of Fe203 by carbon occurs in two stages and that the reduction of Fe203 to FeO is faster than that of FeO to Fe. This is due to the fact that the CO2/CO equilibrium ratio is higher, and hence, the rate of oxidation of carbon is faster. Fortini and Fruehan [3] have studied the reduction reaction of wustite by graphite and wood charcoal. They found that the wustite reduction can have a significant effect on the overall rate of reduction in composites at high temperatures or in the presence of a large excess of carbon. The calculated activation energy of the reduction reaction by wood charcoal was close to 368 kJ/mol, while the activation energy of the reduction reaction by graphite was 444 kJ/mol. From literature [1] it has been shown that for the direct reduction of iron oxide by carbon (Eq. 1) the overall mechanism of reduction generally consists of two elementary steps (Eq. 2 and 3) 1

Corresponding author: [email protected] Tel/Fax: +49 3731 39 4810/4555

95

and the gasification of carbon (Eq. 3) is considered to be the rate controlling step in the reduction of the composite. FexOy(s) + C(s) -> F C A H , (S) + CO (g)

(Eq. 1)

FexOy(s) + CO (g) -» FexO^i) (s) + C0 2 (g) (Eq. 2) (Eq.3) C(s) + C0 2 (g) = 2C0(g) where x=l, 2, 3, and y=l, 2, 3 Until now less attention has been paid to the kinetics of the reduction of iron oxide contained in slag by carbon. Most of the previous work has been done using pure Fe203 or FeO and a certain kind of carbon source (coconut charcoal, coal char and metallurgical coke). Therefore, in the present work the rate of reduction of FeO alone and of FeO containing slag (FeOAl203-CaO-Si02) was investigated at different temperatures under inert atmosphere. Experimental CaO (99.99%, Sigma Aldrich), A1203 (99.995%, Alfa Aesar) and Si02 (99.995% Sigma Aldrich) powders were mixed and molten under Ar gas for one hour at 1400°C in a BN (boron nitride) crucible using a conventional tubular furnace. The samples were ground in a ball mill. FeO (99.9% Sigma Aldrich) was added in well defined amounts and the mixtures were remelted under Ar - 2%02 gas mixture. After cooling, the samples were ground and the graphite powder was mixed to the prepared mixture. Three FeO based slags were prepared with the compositions shown in Table 1. Sample ID 1 2 3

Table 1. Composition of the samples (in mass %) A1203 Si02 CaO FeO 26.7 7.0 12.3 24.0 11 21.3 23.7 14.0 ___——-—~"~ __—-—-~~~~ 75.0

graphite 30.0 30.0 25.0

A Netzsch DTA/TGA analyser (STA 449C) was used to measure the mass change of the samples under isothermal conditions at temperatures of 980CC, 1030°C and 1080°C. The sample was contained in a BN crucible and a continuous flow of Ar gas (50ml/min) was used to prevent the accumulation of gaseous products. For the data analysis the conversion degree was defined as follows (Eq. 4): m0-mm Am, .„ ., a(t) = —= '-— (Eq. 4) m0 -mf m0 -mf where w0is the initial mass, m{l) the mass at the instant t and mf represents the final mass. Although the heating rate was kept high to reach the investigation temperature, the begin of the reduction reaction can not be completely avoided. The final mass corresponds to the expected mass change for the complete reduction of FeO in the slag. The reaction rate daldt is given by Eq. 5 Aa _aM(tM)-al(tl) (Eq. 5) daldt **'■ Ai tM -1, The relationship between the reaction rate and the conversion degree is given by Eq. 6: daldt = k{T)t" (Eq. 6) where k(T) the kinetic constant and n the exponent for time dependence. k(T) is a function of the temperature as shown in Eq. 7

96

(Eq.7)

k(T) = Aexp(--^-) Kl

1

where A is the pre-exponential factor in min" (if n=0), T the absolute temperature in K, R the perfect gas constant in J-mof'-K"1 and Ea the activation energy in kJ-mol"1 Results and Discussion Influence of the Temperature Fig. 1 and 2 show typical TGA curves of the FeO + graphite mixture (Fig. 1) and FeO containing slag + graphite (Fig. 2) specimens. The curves were composed of two parts. The first part is ascribed to the reduction of FeO by graphite for which FeO is directly accessible by graphite and exhibits a linear behaviour with time. The second part of the curves is the continuation of the reduction process but for which FeO becomes less accessible due to the formation of a reaction layer at the FeO surface. Then the curves exhibit a more complex shape which depends on the different processes occurring in the formed layer (diffusion, restructuration). The kinetic constants were determined using the first part of each curve. They are compiled in Table 2. In a first approximation the kinetics of the reduction reaction of FeO by C was considered with previous investigations [1, 2].

Sample ID

1 2 3 *) not determined

Table2. Values of rate constants and activation energy Ea (kJ/mol) k (min"1) T(K) 0.0034 1253 203 0.0058 1303 0.0145 1353 0.0059 1303 n.d.* 0.0161 1353 0.0174 1253 140 0.0205 1303 0.0476 1353

1.0 0.9 0.8 0.7 ö

S 0.6 CD

o 0.5 ¡2 ¡0.4 8

0.3 0.2 0.1 0.0 0

50

100

150

200 Time (min)

250

300

350

400

Fig.l Dependence of temperature on the conversion rate for the sample 3

97

Time (min)

Fig.2 Dependence of temperature on the conversion rate for the sample 1 The results shown in Figs.l, 2 and Table 2 indicate that at a given temperature the reduction of pure iron oxide by graphite is faster than in the slag system. Fig. 3 shows the temperature dependence of the rate constants measured for the reduction of iron oxides by graphite. The values measured in this work differ from the published values of Fortini and Fruehan [1]. However, the values are of same order of magnitude compared with the data from Rao [2]. From the temperature dependence of the rate constants the activation energy was calculated according to Arrhenius low.

10*x1/T(1/K) Fig.3 Arrhenius plot ln(k)=f(l/T) - evaluation of activation energy of the reduction reaction of FeO by C Influence of FeO Contents on Reduction Rate Fig. 4 depicts the evolution of the mass of specimens containing different amounts of FeO. The slope of the curves is independent of the FeO content in the slag. The time of the reduction of FeO was found to be proportional to the FeO content. Therefore the activation

98

energy of the reaction is thought to be independent of the FeO content. The activation energy of the reduction of FeO by graphite is 140 kJ/mol, while the activation energy of the reduction reaction of the FeO containing slag system and graphite is 203 kJ/mol. It appears that the components of the system have significant influence on the reduction kinetics. The surrounding slag may act as a physical barrier which limits the evolution of the produced gas at the outer surface of the FeO particles. 100.0 99.5 99.0 98.5

^

98.0

I 97.5 97.0 96.5 96.0 95.5

0

50

100

150

200

250

300

350

400

Time (min)

Fig. 4 Dependence of FeO contents on the reduction rate at 1030°C Conclusions The results of the present work indicate that the reduction reaction of FeO by C follows a first order kinetics. The mass loss rate and the total mass loss increase with increasing temperature. The FeO content in the slag has no influence on the kinetics of the reduction reaction. The reduction of pure iron oxide by graphite is faster than in slag system, hence the matrix in which FeO is contained modifies the reduction kinetics. Acknowledgments This research work was supported by the Federal Ministry of Education and Research. The authors gratefully acknowledge the help and encouragement of Dr. Schmetterer and Ms. Starke throughout this study. References [1] R. J. Fruehan, Metall. Trans. B, 8B, (1977) 279-286 [2] Y. K. Rao, Metall. Trans., 2 (1971) 1439-1447 [3] O. M. Fortini and R. J. Fruehan, Metall, and Mat. Trans. B, 36B (2005) 865-872 [4] K. Watanabe, S. Ueda, R. Inoue and T. Ariyama, ISU international, 50(4) (2010), 524-530 [5] S. Ueda, K. Yanagiay, K. Watanabe, T. Murakami, R. Inoue and T. Ariyama, ISU International, 49(6) (2009) 827-836 [6] E. T. Turkdogan and J. V, Vinters, Metall. Trans., 3 (1972) 1561-1574

99

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cook TMS (The Minerals, Metals & Materials Society), 2011

Optimization of the Process Variables for Making Direct Reduced Iron by Microwave Heating using Response Surface Methodology Linqing Dai1,2, Jinhui Peng1,2 , Hongbo Zhu1'2 'Key Laboratory of Unconventional Metallurgial (Kunming University of Science Technology ),Ministry of Education, Kunming, 650093, China 2 Faculty of Metallurgical and Energy Engineering, KUST, Kunming, 650093, China Keywords: direct reduced iron (DRI), microwave heating, response surface methodology (RSM) Abstract To optimize the process variables for making direct reduced iron (DRI) by microwave heating, the reduction temperature, reduction time and the ratio of coal to material were studied with the central composite design (CCD) and their interactions on the Metallization rate were also investigated. The predictive polynomial quadratic equations model was analyzed by ANO VA. Optimal conditions of making DRI can be concluded as follows: 28 min at 1139°C, the ratio of coal to material is 20.95%. Under these conditions, the metallization rate is up to 97.06%. 1. Introduction There are some technical problems in making direct reduced iron (DRI) or sponge iron using iron ore concentrates containing coal and lime. In reduction of the iron ore concentrates containing coal with microwave heating, the strength of pellets needn't to be considered. It reduces the costs and improve quality of materials, but also eliminates the need for pellet preparation process, simplifying the process and achieving a cleaner production. Microwave energy is a clean energy, and has characteristic of selective heating. The speed of heating is also fast. It researched by Chen Jin that when the mole mass of C: O is 0.627 and the microwave power is 15KW, it only spends 15min from 15°C to 1021.6°C with 1 kg iron ore containing coal with microwave heating [1]. There are many factors affecting metallization rate. In order to find a suitable condition for reduction process, it is necessary to design a series of suitable experiments. Response surface methodology (RSM) is one of the relevant multivariate techniques

101

which can deal with multivariant experimental design strategy, statistical modeling and process optimization [2-5]. It is used to examine the relationship between one or more response variables and a set of quantitative experimental variables or factors. This method is often employed after the identification of the vital controllable factors to find the factor settings that optimize the response. Designs of this type are usually chosen when a curvature in the response surface is suspected. The process optimization of the microwave reduced iron ore was seldom reported in literature. Hence the present work intends to assess the effects of variables such as reduction time, reduction temperature and the ratio of coal to identify the optimum conditions using a central composite design (CCD). The characteristic of product was assessed using scanning electron microscope (SEM). 2. Materials and method 2.1 Materials and apparatus The materials were obtained from Kunming Iron & Steel Group Co., Ltd. in Yunnan province, China. The compositions of ferrous materials are presented in Table l.The compositions of coal and flux are presented in Table 2 and Table 3. Prior to mixing, the mill scale and coal were milled with ball mill. The mixture was heated with microwave to given temperature using a self-made microwave oven. The content of metallic iron and Fe in DRI were analyzed by Technical Center of Kunming Iron & Steel Co.,Ltd.[6j.

Magnetite Mill scale Table 1 Magnetite Mill scale

Table 1 Compositions of ferrous materials (weight percent,%) CaO MgO MnO TFe FeO A1203 Si0 2 0.04 0.43 0.81 62.98 24.86 6.86 1.36 1.44 0.62 53.60 2.84 0.64 0.63 69.18 P 0.033 0.070

TiC-2 0.58 0.15

Pb 0.004 0.006

Zn 0.009 0.016

K20 0.056 0.020

S 0.024 0.020

Moisture 5.80 1.40

Na 2 0 0.260 0.085

Table 2 Proximate analyis of coal (weight percent,%)

CaO 41.10

c

Ash

Volatile

S

P

Moisture

70.42

22.40

9.25

0.9

0.014

1.69

Table 3 Composition of flux (weight percent,%) MgO Si0 2 A1203 S 26.39 <0.5 0.038 0.016

102

P <0.005

Table 4 Independent variables and their levels used for central composite rotatable design

Ranige and levels Variables

Symbol

Temperature(°C) Time(min) Ratio of coal (%)

- 1

- 0.595

0

+ 0.595

+1

Xi

900

961

1050

1139

1200

X2

20 5

28 9.05

40 15

52 20.95

60 25

2.2 Experimental methods The mixture was mixed by hand. A total of 20 experiments were conducted with 75 g of sample in a batch mode using a special reactor with the volume of 500 ml. The surfaces of mixture were well covered with 5 g of coal, so as to prevent oxidation of the DRI. The reduced product was analyzed for the Fe and metallic iron content using titration analysis. The metallization rate was estimated using Eq. (1). metallization rate = ^ - x 100% m0

(1)

where mo and mi correspond to Fe and the analyzed metallic iron content of product. 2.3 Experimental design RSM helps to optimize the process which was influenced by number of operating parameters with a minimum number of experiments as well as to analyze the interaction between the parameters. Reduction temperature (xi), reduction time (X2) and the ratio of coal (X3) were chosen as the independent variables with their levels and ranges shown in Table 4. Table 5 shows the actual values of the independent variables at which the experiments were conducted to estimate the response variable the metallization rate (Y). The chosen independent variables used in process optimization were coded according to Eq. (2). =

2k_l2i
(2)

where %¡ is the dimensionless coded value of the ¡th independent variable, Xo is the value of Xi at the center point and Ax is the step change value. The metallization rate was the response variable of the experimental conditions in the design of experiments. The reduction temperature was varied from 900°Cto 1200°C The reduction time was varied from 20 min to 60 min and the ratio of coal was varied from 5% to 25%. A total of 20 experiments consisting of 8 factorial points, 6 axial points and 6 replicates at the central

103

points were performed [13]. Experimental results obtained from the CCD model were described in the form as given in Eq. (3),

y — ßo + ^

k

;=1

k

ßix> + 2 , ßiCXi i=l

2

Jc

+ 5 L ßijxocj + £

(3)

i-<j

where ßo is the value for the fixed response at the central point of the experiment; and ß„ ßj and ßy are the linear, quadratic and cross product coefficients, respectively. The analyses of variance (ANOVA) and response surfaces were performed using the Design-Expert software (version 7.1.5) from Stat-Ease Inc., USA. The optimized reduction conditions were estimated using the software's numerical and graphical optimization tools. 3. Results and discussion 3.1 Response analysis and interpretation The results of experiments were shown in Table 5. It is found that the metallization rate varies from 27.58% to 97.68% in response to the variation in the experimental conditions. According to the sequential model sum of squares were selected based on the highest order polynomials, where the additional terms were significant. The ANOVA of quadratic model is presented in Table 6 which proves the validity of the model. The reduction temperature has the greatest effect on metallization rate with the highest F-value of 47.92 whereas the ratio of coal and the reduction time were found to be less significant. The model F-value of 10.11 implies the significance of the model. The validation of model was an important part of the data analysis procedure since an inadequate model could lead to misleading results. For the fixed model, adequate precision can be ensured with a signal to noise ratio greater than 4. An adequate precision ratio of 10.998 indicates the ability of model to precisely navigate through the design space. Not all the effects of parameters on metallization rate were significant, while values of Prob > F less than 0.05 indicate that the model terms were significant. In this case, xi, X2> 3te and the interaction terms X32 was significant model terms whereas the others were insignificant. In order to enhance the effect of significant parameters, the insignificant parameters were eliminated. The final equation in terms of coded factors is shown in Eq. (4) as, y,= - 81.86+16.23x1 + 5.47x2+11.99x3 - 1.24JCIX2 - 1.88xiX3 - 2.42x2X3 - 5.02X,2 + 1.09X22 - 5.40X32

(4)

2

The R value for Eq. (4) was found to be 0.9010 close to unity, indicating the good agreement between the experimental and the predicted metallization rate. It does not show any significant non-linear pattern (S-shaped curve) indicating non-normality in the error term. Fig. 1 shows an approximate linearity which confirms normality of the data.

104

3.2 Process optimization The experimental and predicted metallization rate is shown in Fig.2.The figure shows a close proximity of the model prediction with the experimental data signifying the validity of the regression model. Fig. 3 shows three-dimensional plot of the reduction temperature and time on metallization rate. It shows that the metallization rate increases significantly with increasing temperature. An increase of the temperature from 900°C to 1200°C increased the metallization rate remarkably from 27.58 to 96.58%. Fig. 4 shows the combined effect of temperature and ratio of coal on the metallization rate. It can be observed that an increase in both coal ratio and the temperature increases the metallization rate. Fig. 5 shows that the metallization rate increases significantly with increasing the coal ratio. It can be seen from Fig. 3 to 6 that the effect of temperature and coal ratio on the metallization rate were remarkable, while longer time show only a marginal effect. With the objective to maximize the metallization rate at the suitable temperature, reduction time and coal ratio, the optimum reduction conditions were identified using the Design-Expert software. It reports the optimum conditions to be a temperature of 1139°C, reduction time of 28 min and coal ratio of 20.95%, which gives metallization rate of 97.06% (Table 7). Table 5 Experimental design matrix and results

Std 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15

Y, rcn

961

1139

961 1139

961 1139

961 1139

900 1200 1050 1050 1050 1050 1050

X2<min)

*(%)

y,(%)

28 28 52 52 28 28 52 52 40 40 20 60 40 40 40

9.05 9.05 9.05 9.05 20.95 20.95 20.95 20.95

44.19 76.70 61.79 89.56 69.51 94.73 77.63 97.68 27.58 96.58 69.51 89.20 32.33 89.68 81.03

105

15 15 15 15 5 25 15

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a

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o o o o o Lft LA L/l L/« (_ o o o o o

Fig. 1. Normal % probability versus internally studentized residuals

Fig.2. Comparison of model prediction with the experimental data

Fig.3. Effect of temperature and time on metallization rate (ratio of coal: 15%)

107

Fig.4. Effect of temperature and ratio of coal on metallization rate (time:40min)

Fig.5. Effect of time and ratio of coal on metallization rate (temperature: 1050°C) Table 7 Optimum reduction conditions with'model validation metallization rate/% Time/min Ration of coal/% Temperature/°C Predicted Experimental 28 20.95 97.06 94.73 1139 The SEM analysis of final product was shown in Fig.6. It clearly indicates the sponge morphology of the reduced iron.

108

Fig.6. SEM ¡mage of DRI in Optimum reduction conditions 4. Conclusion It is efficient and environmental friendly to make DRI with microwave heating. The present study was aimed to explore the effects of reduction temperature, time and coal ratio on the metallization rate by optimizing the process conditions using RSM. The proposed quadratic model agrees well with the experimental data, with correlation coefficients (R2) of 0.9010. The temperature and coal ratio were found to have significant effect on the metallization rate while reduction time showed little effect. Based on "Design-Expert" software, the optimum conditions were identified to be a temperature of 1139°C, a reduction time of 28 min and a coal ratio of 20.95%, with the metallization rate of 97.06%. Furthermore the SEM analysis of product indicated that the spongy iron was produced. Reference [1] Cheng Jin, Liu Liu, Zeng Jiaqing, et al. Experimental Research of Microwave Heating on Iron Ore Concentrates Containing Coal and Line. Iron and Steel. 2004, 39(6): 1-5. [2] M.A. Bezerra, R.E. Santelli, E.P. Oliveira, L.S. Villar, L.A. Escaleira, Response surface methodology (RSM) as a tool for optimization in analytical chemistry, Talanta 76 (2008) 965-977. [3] Ranjana Yadav, Archana Devi, Garima Tripathi, Deepak Srivastava, Optimization of the process variables for the synthesis of cardanol-based nivolac phenolic resin using response surface methodology, European Polymer Journal, 43 (2007)3531-3537. [4] Malihe Amini, Habibollah Younesi, Nader Bahramifar, et al, Application of response surface methodology for optimization of lead biosorption in an aqueous solution by Aspergillus niger, Journal of Hazardous Materials. 154 (2008)694-702. [5] B.H. Hameed, I.A.W. Tan, A.L. Ahmad, Preparation of oil palm empty fruit

109

bunchbased activated carbon for removal of 2,4,6-trichlorophenol: optimization using response surface methodology, Journal of Hazardous Materials. 164 (2009) 1316-1324. [6] Jun Tao, Ling Zheng. Determination of metallic iron in direct reduced iron by potassium dichromate titration after decomposition of sample by ferric chloride. Metallurgical Analysis 29(2009)65-68.

110

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

Study on Nucleation and Growth Mechanism of Iron Crystal Grain in Coalbased Shaft Furnace Direct Reduction Iron Pellets by Microwave Heating Zhu-cheng Huang, Zhen-yuan Liao, Bing Hu, Ling-yun Yi, Yuan-bo Zhang (School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, China) Keywords: oxidized pellet, microwave heating, direct reduction, nucleation and growth of iron crystal grain, etc. Abstract Nucleation and growth mechanism of iron crystal grain in coal-based shaft furnace direct reduction process by microwave heating were investigated by optical microscope, scanning electron microscope and EDX. The results indicate that microwave has selectivity on various minerals which absorb microwaves. The high temperature in certain area of pellets and thermal stress on the interface of minerals are formed, results in the wustite grain smashed along the joints, and iron crystal grains turned out a spherical shape obviously. The iron crystal nucleus is formed firstly at the edge of wustite, on the interface between grains and in the holes in pellets, and grows gradually while reduction carries on from surface layer to inner core. Microwave heating promotes the rate of transportation and accumulation of the iron crystal grain, increases growth rate of iron crystal grain and contributes to formation of dense iron-jointed crystal texture eventually.

Introduction Microwave energy as a new energy, which selectively heating materials through internal energy dissipation, has the advantages of heating uniformity, high thermal efficiency, fast heating, clean and pollution-free compared to conventional heating motheds. Microwave energy also has obvious superiority in increasing productivity. Thus, it has been widely used in many areas, including electronics, food processing, chemical, medicine, environmental protection, and family life etc [1], As a rapidly developed green metallurgical method, microwave heating also has been widely recognized by areas of grinding, pretreatment, pre-reduction, drying, furnacing, metal extraction, and waste materials processing and utilization. Some research achievements are gradually putting into practice, showing great potential in chemical and metallurgical industry [23]. Researchers have paid many attentions to mineralogical microstructure of pellets for conventional heating [4-7], Conventional heating is external heating, which would cause many problems, along with the processing of carbothermic reduction. The reaction between ore and coal absorbs a large amount of heat, easy to form "cold-center", and slow down the rate of material reduction. If enhancing external temperature further, temperature gradient between

111

inside and the surface of the material increases, pellets will be cracked to powder. For microwave heating of pellets containing carbon, research shows that mineral particles and coal particles all have the characteristics of endogenous heating, which can not only provide good conditions for reduction kinetics, satisfy the self-reduction properties of iron ore containing carbon, increase the metallization of materials, but also decrease the activation energy, lower reaction temperature and shorten reduction time [8-12]. However, there are rare researches about microwave heating on the coal-based direct reduction of oxidation pellets. This paper focuses on effects of microwave heating on mineralogical structure of reduced pellets. Materials and Experimental Procedure Iron-ore oxidized pellets produced by Ezhou pellet plant of Wuhan Iron and Steel Company were used as raw material; the chemical composition is listed in Table 1. Coke is the reducing agent, its industry analysis and granularity composition is respectively shown in Table 2 and Table 3. Table 1. The chemical composition of oxide pellets S P Mi;!s Compositions TFe FeO FeA SiO¡ Hfh MgO CaO K20 Na20 CuO Vft Content0/. 63.99 0.24 91.21 4.52 2.30 0.49 0.56 0.091 0.075 0.05 0.02 0.047 0.030 0.036 Table 2. The industrial analysis of coke Compositions

fixed carbon

Contenf/o

75.75

ash content 20.24

S

volatile content 3.24

-

water 0.77

Table 3. Size distribution of coke particle size/mm Content%

>5

-5~+3

-3-+0.5

-0.5-+-0.18

-0.18

17.94

1.60

39.72

1928

21.23

The experiment applies external matching carbon method, according to matching carbon ratio (carbon iron ratio in clumps) 0.354, mix pellets and coke together, put them into the industrial microwave shaft furnace [13] and heating them. Along with the increasing of material temperature, oxide pellets and coke mainly occurs the following reduction reaction: FexOy+yCO =xFe +yC0 2 C + C0 2 =2CO

(1) (2)

After natural cooling, cutting, grinding, polishing, and drying of the tested samples, we obtain microsection for microstructure studies. Test conditions and chemical analysis are shown in Table 4; procedure of the test is shown in Figure 1. The microstructure and composition of oxide pellets in different temperatures were studied by optical microscope and scanning electric microscope energy spectrum analyzer.

112

Table 4, Test condition and chemical composition Test samples

microwave power

test condition ratio of iron to Reduction carbon Temperature K

Reduction time

chemical composition /% metallized TFe MFe rate

LI

7.5kw

0.354

850

40min

66.72

0

0

L2

9kw

0.354

950

40min

81.44

35.75

44.05

L3

10.5kw

0.354

1000

40min

82.26

45.43

55.23

L4

13kw

0.354

1050

40min

86.93

71.24

81.95

oxide pellet

coal

mix I •* microwave heating direct reduction

i

cooling screening magnetic separation

I

'

'

I

DRI

I

metalized rate

I

non-magnetics

Microstructural analyses

Figure 1. Experimental procedure Results and Discussion Nucleation Mechanism of Iron Crystal Grain Fe2Û3 is prone to be reduced into Fe3C>4, microwave heating oxide pellet, when the temperature reaches at 850°Q pellets and coal initiate the reduction reaction, the hematite is reduced into magnetite, and magnetite crystals mostly are tabular (Figure 2). By 40 minutes, in the outer layer of pellets, you can see a few round-ball magnetite grains (area "A" in Figure 2). This phenomenon is due to magnetite is a very good microwave absorbing material, microwave heating enables the internal magnetite rapid warming, internal-heat promotes magnetite particles expand fragmentation, this fragmentation is helpful for carbothermic reductionfreaction (2)), and improves the rate of the reduction reaction. Formation of magnetite effectively enhances the pellet's capacity of microwave absorption, and

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the local high temperature formed around magnetite, promotes hematite reduced into magnetite, and then into wustite. When the temperature reaches 950°Q the majority of magnetite is reduced into spherical wustite (Figure 3), this spherical structure is due to that microwave can heat mostly useful minerals, but not heat gangue minerals. Thus, obvious local temperature gap is formed between wustite and gangue minerals, thus resulting in thermal stress. When the thermal stress is large to a certain extent, cracks would appear in the grain boundary, and lead to wustite appears fragmentation. In addition, under the effect of microwave, harmonic oscillator of the medium molecules make strong resonance absorption, the material for absorbing capacity of microwave energy is caused by the dielectric body, while the dielectric body is comprised of polar molecules and polarizable molecules. Under the effect of external electric field, there are two movements of material dielectric: one is polarizable molecules make relative displacement, and forms two polars, which is the so-called dielectric polarization, and the other is reverse of the polar molecules. In the microwave electromagnetic field, orbit of polarizable molecules would vibrate relative to the nuclear at the frequency of electric field, polar molecules would rotate along with electric field back and forth[14]. Because of this special function of microwave, wustite grain smashed along the boundary and turned into a spherical shape.

Figure 2. Tabular and rounded wustite crystals at the surface layer of a pellet reduced at 850°C (SEM) grey—wustite; dull gray—gangue; black— hole

Figure 3. Rounded grains of wustite in the center of a pellet reduced at 950°C ( SEM ) brilliant white—iron crystal grain; light gray—wustite; dull gray—gangue; b l a c k hole

As the surface of spherical wustite grain is not smooth (Figure 4), according to adsorption theory [15], uneven surface and impurities can be "active center", resulting in gas molecules to adsorb on it. Therefore, CO firstly adsorbs at the edge of wustite, on the interface between the boundaries of grains and in the holes of pellets. In the microwave field, non-symmetry and polarization distortion of CO molecule orbital, which leads to the energy of molecular orbital enhanced and causes the instability of molecular orbital, at a certain temperature, prone to dissociation adsorption on the surface of wustite, the following reaction would take place: FeO+CO=Fe+C0 2

(3)

Thus, wustite grain begins to form iron grain crystal and further to form nuclei, iron crystal nucleus sporadic scattered in the pellet (Figure 5). According to the above, when temperature reaches above 850°Q reduction reaction of iron oxide is carried out in this order: Fe2Û3- Fe304-FeO-Fe step by step.

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Figure 4. Rounded grains of wustite in a pellet reduced at 950°C (SEM) brilliant white—iron crystal grain ; light gray—wustite; dull gray—gangue; b l a c k hole

Figure 5. Iron crystal nucleus sporadic scattered in a pellet reduced at 950°C ( x200 reflected light) brilliant white—iron crystal grain; light gray—wustite; dull gray—gangue; b l a c k hole

Figure 6. Iron crystal nucleus encased gangue structure in a pellet reduced at 950°C (SEM) brilliant white—iron crystal grain ; gray— gangue; black—hole

Figure 7. Gangue at the center of a pellet reduced at 1000°C (SEM) brilliant white—iron crystal grain ; light gray—wustite; gray("A")— gangue; black—hole

Growth Mechanism of Iron Crystal Grain Once the iron crystal nucleus has been formed, it becomes a new "active center", As temperature increasing, the number of "active center" sharply increases, As the reaction gradually from the wustite's surface layer to the core, the new iron grain crystal moves to the "active center" and accumulates with them, This makes iron crystal nucleus continuously grows up, as far as all the wustites are completely reduced into iron, ring-like iron crystal nucleus structure will form in the pellet (Figure 6). As the oxidized pellet contains a little gangue materials, mainly in the form of complex silicate minerals, they contain variable valence of iron, which can produce a high microwave loss, and can also form some heat. Reduced reaction by microwave heating at 1000°C for 40min, partial wustite in the gangue will be reduced, contrast the X-ray scanning results of gangue (see Table 5) for the center and the edge of the reduced pellet (see Figure 7 and 8 respectively), we can know that, the edge of the pellet contains more FeO than the center, which shows that majority of FeO have been separated from the gangue. Then the isolated FeO are reduced into metal iron, and

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move together to form the ring-like iron crystal nucleus structure, a dense worm-like iron-jointed crystal texture eventually (Figure 9).

numbered list "A" in Figure 7 "+" in Figure 8

Table 5. SEM X—ray microanalysis of gangue Compositions, wt, % FeO MgO A1203 Si0 2 CaO 56.64 6.66 0.38 34.02 1.30 17.39 7.12 13.48 53.69 6.04

Mn0 2 1.00 2.28

total 100 100

Changes of the Inner Structure of the Pellets From the view of kinetics of process metallurg, because the oxidized pellets after calcined by furnacing are comparatively dense, so the reduction reaction are affected by the gas diffusion, due to the concentration of reducing agent CO gradually decreases from the edge to the center of the pellet, which induces the reaction carries on from surface layer to inner core (Figure 10).

Figure 8. Gangue at the edge of a pellet reduced at 1000°C(SEM) brilliant white—iron crystal grain ; gray ("+") —gangue; black— hole

Figure 9. Worm-like iron-jointed crystal texture in a pellet reduced at 1000°C(SEM) brilliant white—iron crystal grain; gray— gangue; black—hole

When temperature reaches at 850°Ç the pellet is divided into two layers: the outer layer is wustite layer, and the inner layer is unreacted layer —hematite layer (Figure 10a), this due to the CO gas density inside the pellet is low. However, when the temperature rises to above 950°Ç the pellet is divided into three layers, the outer layer is the dense iron-jointed crystal texture, middle layer is ring-like iron crystal nucleus layer, and the inner layer is wustite layer(Figure 10b ~ lOd). This is because of the concentration of CO in outer layer is the highest, iron oxide has been basically reduced to metal iron, middle layer mainly takes place reaction (3), so ring-like iron crystal nucleus layer is formed. Concentration of CO in the inner layer is lowest, the reaction occurs in this layer is mainly the formation of wustite. Along with the temperature gradually increases, concentration of CO inside the pellet increases, the edge of the inner layer would gradually advance towards to center, shown as in Figure 10. Comparing Figure 10b with Figure 10c, we can know that the dense iron-jointed crystal texture and ring-like iron crystal nucleus structure in Figure 10c is nearer to the center than Figure 10b. Relationship between Figure 10c and Figure lOd is also the same.

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a-850°C; b—950°C; c-1000°C; c-1050°C Figure 10. The inner structure of the pellet undergone gradually changing temperature In addition, the boundaries between iron-jointed crystal texture layer and ring-like iron crystal nucleus layer are not obvious, but ring-like iron crystal nucleus layer and wustite layer would form significant boundaries—macroporous ring-like iron crystal nucleus structure, which is die result of transportation and accumulation of the iron crystal grain. In microwave field, the transportation rate of grains is greater than the accumulation rate, the number of crystallization centers reduces, which is helpful for the growth of iron crystal nucleus. Conclusions 1) Microwave selectively heats metallic useful minerals, and puts less effect on gangue minerals, thus the high temperature difference in certain area between wustite and gangue phase would be formed and results in thermal stress. When the thermal stress is great to a certain extent, crackes would be appear at the joints. Microwave can make dielectric medium polarize and the polar molecules reverse, orbit of polarizable molecules would vibrate relative to the nuclear at the frequency of electric field, polar molecules rotate back and forth with the electric field, polar molecules such as vibration, making wustite takes the shape of spherical. 2) Reduction reaction process in coal-based shaft furnace by microwave heating is carried out in the order: Fe2C>3-Fe3C>4-FeO-Fe step by step, thereinto, magnetite is a good microwave absorbing material, microwave heating can make the internal magnetite rapid heating, and internal-heat induces wustite grain smashed along the joints. The iron crystal nucleus is formed firstly at some "active center" such as at the edge of wustite, on the interface between grains and in the holes of pellets, and then wustite is reduced to iron crystal grain from die surface layer to the core, iron crystal nucleus grow up, and microwave promotes the rate of transportation and accumulation of the iron crystal grain, which can make iron crystal grain grows rapidly, and form dense iron-jointed crystal texture eventually.

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3) By microwave heating, wustite can been separated from the gangue and reduced into metal iron, then the isolated wustite are reduced into metal iron, and move together to iron crystal uncleus structure, this will makes iron crystal nucleus grows larger and larger. References [l]David E.C., Diane CF., Jon K.W., Processing Materials with Microwave Energy[J],Mat.Sei.Eng.AStruct.,2000,287(2): 153-158. [2]A1-Harahsheh M, Kingman S W, Microwave-assisted Leaching-A Review[J],Hydrometallurgy,2004,73(3-4):189203. [3]Haque KE, Microwave Energy for Mineral Treatment Processes-A Brief Review[J], Int.J. Miner. Process, 1999,57(l):l-24. [4]T.I.Otsuka,D.Kunji.j.Chen.Eng.Jap.,1969,2(l):46-50 [5]Y.K.Yao.Metall.Trans.ISIJ,1983,23(6):190-196 [6]R.J.D.Carvalho,P.GO.Netto,J.C. [7]Abreu.Can.Metall.Q., 1994,33(3): 229-2354 [8]Jin Chen, Liu Liu, and Jia-qing Zeng, Study on the Microwave Heating of Pellets Containing Coal[J]. Journal of Chinese Electron Microscopy Society,2005,24 (2) : 114—119 [9]Jin Chen, Liu Liu, and Jia-qing Zeng, Experimental Research of Microwave Heating on Iron Ore Concentrates Containing Coal and Lime [JJ.Iron and Steel,2004,39(6):l-5 [10]Qing Lv, Wei-dong Liu, and Li-guo Zhao, The observation of the reduction of the pelletcontaining carbon through microscope [J]. Journal of Hebei Institute of Technology (Natural Science Edition),2001,23(4):19-24. [ll]Jin Chen, Liu Liu, and Jia-qing Zeng, Study on the Microwave Heating of Pellets Containing Coal[J]. Industrial heating,2001,6:8-10 [12]Qin-han Jin, Shu-shan Dai, and Ka-ma Huang, Microwave Chemistry [M], Science Press, Beijing,1999. [13]Wang Xia,Huang Zhu-cheng, Jiang Tao,Peng Hu,and Xia Guang-bin."Coal-based direct reduction of iron concentrate pellets by microwave heating".The Minerals, Metals & Materials Society,2010,1,363-371. [I4]Jin Chen, Wan-ming Lin, and Jin Zhao, Smelt Technology Use Noncoke Coal, Beijing: Chemical Industry Press,2007. [15]Yue-xun Shi,, Blast Furnace gas - solid reaction kinetics, Beijing: Metallurgical Industry Press,1996.

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2nd International Symposium on High-Temperature Metallurgical Processing Ediled by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

INVESTIGATION ON A MICROWAVE HIGH-TEMPERATURE AIR HEAT EXCHANGER Jianhua Liu ''2, Yingwei Li ' ,2 , Lijun Liu ' ,2 , Jinhui Peng 1 ' 2 * Libo Z h a n g u , Shenghui Guo 1 2 , Huilong Luo 2 , Hongpo Wang1'2 ,Guo Chen1,2 1. Faculty of Materials and Metallurgical Engineering, Kunming University of Science and Technology, Kunming Yunnan, 650093, PR China 2. Key Laboratory of Unconventional Metallurgy, Ministry of Education, University of Science and Technology, Kunming Yunnan, 650093, PR China Keywords : Air heat exchanger, Microwave, High-temperature, Microwave absorbing material Abstract As essential equipment in the metallurgical industry, current air heat exchangers cannot meet the requirements of high temperature heat transfer. In this paper, an energy efficient air heat exchanger, based on accumulation of the heat generated by microwave absorbing materials, is presented according to heat transfer theory and principles of the microwave field. A tubular shaped heat exchanger was designed and built up through temperature-rising curve, and experimental research of microwave high-temperature air heat exchangers was carried out. It was found from experiments that the tubular shaped heat exchanger fabricated using material B could be used for high temperature heat transfer system, and equilibrium temperatures could reach up to 4 5 7 ^ 8 5 ° C Introduction Heating or cooling of air by heat exchangers is commonly practiced in a wide range of industries including construction, chemical, machinery, textile, printing, and dyeing [1-3]. Generally such heating or cooling of air is achieved in pipe heat exchangers. There are several drawbacks for such heat exchangers, namely, (i) the limitation of temperature to which the pipes can be heated, (ii) tendency to oxidize at high temperature, and (iii) poor strength of metallic pipes at high temperature. These limitations do not allow pipe heat exchangers to be operated at high temperatures [4-6]. Microwave heating of inert solids, which are stable at high temperature, is a possible solution to achieve high temperature of the medium in air heat exchangers. It is well-known that microwave heating is an efficient, clean, safe, easy-to-control heating method. Microwave heating utilizes the dielectric and magnetic properties of a material. When microwaves penetrate and spread into the dielectric material, electrons, ions, etc move in the internal electromagnetic field, while the elasticity of inertia and friction impede these movements. This results in a significant thermal * Corresponding author: Jinhui Peng Tel.: +86 871 5192076; fax: +86 871 5191046. E-mail address: [email protected].

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effect, increasing the surface temperature rapidly [7-8]. This paper studies mainly design and experimental research of microwave high-temperature air heat exchanger. Design of Microwave Air Heat Exchanger System The microwave high-temperature air heat exchanger is based on generation of heat in a microwave field. In such a heat exchanger, the important components requiring special attention for design are the regenerator, heat transfer devices, materials of construction and design of the microwave cavity. An enhanced heat transfer device in the microwave cavity is a core component in a microwave heating system. Whether it is reasonable and optimal will directly influence and determine the efficiency of microwave air-heating system. Enhanced heat transfer devices in microwave cavities have an important role in converting microwave energy into heat energy and maximizing the heat transfer to air. Material Selection An enhanced heat transfer device in microwave cavity absorbs continuously microwave energy and converts it into heat energy. It is very important to select a suitable microwave absorbing material for the enhanced heat transfer device, which will: (1) be a strong absorbing material, i.e. the material must have sufficiently large dielectric loss factor; (2) have a high chemical stability, must be able to withstand the high temperature; (3) have sufficiently high strength, structural and dimensional stability at high temperatures; (4) have low coefficient of expansion, strong high temperature corrosion resistance; (5) have high thermal conductivity. So, two kinds of microwave absorbing material A and B were selected. A susceptor is a material utilized for its property to absorb electromagnetic energy. The term "susceptor" is derived from "susceptance", representing the tendency of the material to convert electromagnetic energy into heat [9,11]. The heat exchanger materials presented here are made up of a microwave absorbing material. The microwave absorbing materials A and B can ensure temperatures of 1200 - 1300 can be reached in a short period of time. In view of the above requirements, we selected microwave absorbing materials A and B for enhanced heat transfer devices in a microwave cavity and made comparative experiments. Design of Structure The structural design of heat exchanger includes the choice of flow channel layout and flow pattern. The enhanced heat transfer device has special features, different from ordinary heat exchangers, which continuously absorb microwave energy and convert it into heat, which is passed to the air by way of convective heat transfer. Therefore, the only fluid flowing through enhanced heat transfer devices is air. The layout of the flow channels is relatively simple. Also, the flow pattern is only an air flow pattern, not a choice of two kinds of fluid flow pattern. Considering the characteristics of microwave heating as well as the characteristics of the two materials themselves, the structure of two kinds of heat exchangers are shown in Fig. 1 and Fig.2.

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Fig. 1 The structure of material A Fig. 2 The structure of material B Air Heat Exchanger Test System The high-temperature air heat exchanger test system is shown in Fig.3. The system mainly consists of the following components: (i) Air supply system, the high-temperature fan providing air for heat exchanger, (ii) Power Systems, consist of the power and 150-250V adjustable transformer, air speed is controlled through the voltage-controlled, (iii) Test System, Imports of air speed is measured by the hot-wire anemometer. The outlet and inlet temperature are measured through the thermocouple. The heat exchanger temperature is measured in the center of the heat exchanger material.

Fig. 3 Air heat exchanger test system Results and Discussion Effect of Microwave Power The outlet air temperature curve of the heat exchanger is presented in Fig.4. At constant wind velocity, the outlet air temperature increases with increasing microwave power. The result is that the outlet air temperature is near 430°C when the microwave power is 33kW. The reason could be that the heat energy storage of the microwave heating body increased with the increasing power.

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Fig.4 Effect of microwave power on outlet air temperature Effect of Different Structures When the internal temperature of the heat exchanger rose to around 1150, the air blower was switched on. Fig.5 and Fig.6 show the comparison of the internal and outlet air temperatures of the heat exchangers with different interior structures. The internal temperature using Material A dropped much faster than Material B. The heat storage capacity of Material B is better than Material A. After the air blower was switched on, the trend of the outlet air temperature for the two heat exchangers is basically the same. Obviously, the outlet air temperature of Material B heat exchanger is about 200 higher than Material A. Therefore, Material B is a better heat exchanger structure.

Fig.5 Inner air temperature of two heat exchangers Fig.6 Outlet air temperature of two heat exchangers Effect of Air Speed By applying different voltages in the air blower (160V, 190V, 220V and 250V), we get different air speeds in the two kinds of heat exchangers, as shown in Table 1[10] Table. 1 Air blower speed for different voltages

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Air blower voltage/V

160

190

220

250

Air blower speed/m/s

1.81

2.18

2.33

2.45

The results are shown in Fig.7 and Fig.8. Air speed indeed has an effect on the outlet air temperature of the heat exchangers. However, the air speed is not proportional to the temperature of the microwave heating body; there is an optimal air speed that fully mixes hot and cold air and produces the best heat transfer.

Fig.7 The outlet temperature before air blower switch on Fig.8 The internal temperature after air blower switch on Heating Curve of Tubular Heat Exchanger The heating curve of the tubular heat exchanger is shown in Fig.9 and Fig. 10. When the microwave power is 3kW, the heat exchanger can be heated to 1000 in 90 minutes.

Fig.9 heating curve of tubular heat exchanger

Fig. 10 outlet and inlet temperature of tubular heat exchanger

Conclusion (1) The heat exchanger presented in this paper is a simple structure* corrosion resistant,

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high-temperature stability in microwave field and also environmentally clean (non-polluting). Also, it provides an economical and simple choice for the food, chemical, timber and other industries which need heating by hot air. (2) The results obtained experimentally indicate that the tubular shape of Material B can be used for a microwave-absorbing heating transfer system. (3) The experimental data show the rate of temperature decease of the heat exchangers is different, when they reach equilibrium heat transfer. The internal temperature of the tubular heat exchanger was 457—485°C. The temperature of the tubular heat exchanger is minimized when the air speed is about 2.18 m/s. So this is the appropriate air speed that is able to effectively extract the internal energy of the heat exchanger and convert it into heat. Acknowledgements The research is supported by National Basic Research Program of China (No: 2007CB613606) and National Natural Science Foundation of China (No: 50734007). References [I] Francesco Marra, Maria Valeria De Bonis, Gianpaolo Ruocco, "Combined microwaves and convection heating: A conjugate approach," Journal of Food Engineering, 97(2010) 31-39. [2] Jin mingcong, Cheng shangmo, Zhao yongxiang, Heat Exchanger (ChongQing,CQ: University Press, 1990), 156-160. [3] Jiao ming, Xu hong, "Development and Research Direction of new high efficiency Heat exchanger," Chemical Engineering Design Communications, 33 (2007), 49-53. [4] Peng Jinhui, and Yang Xianwan, The New Applications of Microwave Power (YunNan, KM: Science and Technology Press, 1997), 25. [5] Dumitru Niculae, "Air heating system utilizing microwave susceptor ceramic materials," (Global congress on microwave energy applications, GCMEA2008), 173-177. [6] Flavio C.C. Galeazzoa, Raquel Y. Miura, Jorge A.W. Gut "Experimental and numerical heat transfer in a plate heat exchanger," Chemical Engineering Science, 61 (2006), 7133-7138. [7] Metin, "Mathematical model of finned tube heat exchangers for thermal simulation software of air conditioners," Heat Mass Transfer, 29(2002), 547-557. [8] G.Hed, Bellander, "Mathematical modelling of PCM air heat exchanger," Energy and Buildings, 38(2006), 82-89. [9] M.K.Ghosal, GN.Tiwari, Srivastava, "Thermal modeling of a greenhouse with an integrated earth to air heat exchanger: an experimental validation," Energy and Buildings, 36(2004), 219-227. [10] M.De Paepe, A.Janssens, "Thermo-hydraulic design of earth-air heat exchangers," Energy and Buildings, 35(2003), 389-397. [II] NEI Xin, ZHOU Jun-hu, Lv Ming, CEN Ke-fa, "Test study on new type of high-temperature air heater," Thermal power generation, 37 (2008), 58-67.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

2nd International Symposium on

High-Temperature Metallurgical Processing

Refractories, Slag and Recycling Session Chairs: Patrick Masset Gabriella Tranell

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

STUDY ON PREPARATION OF HIGH-QUALITY SYNTHETIC RUTILE FROM TITANIUM SLAG BY ACTIVATION ROASTING F O L L O W E D BY ACID LEACHING Yu-feng Guo1, Shui-shi Liu, Tao Jiang, Guan-zhou Qiu School of Minerals Processing and Bioengineering, Central South University Changsha, Hunan 410083, P R China Key words: titanium slag, activation roasting, acid leaching, synthetic rutile Abstract The preparation of high-quality synthetic ruitle from electric furnace titanium slag was systematically investigated by activation roasting followed by acid leaching process. The results showed that impurities such as Ca and Mg in titanium slag can be effectively removed and Ti0 2 grade of the product was markedly improved by the process. High-quality synthetic rutile containing 88.54% Ti0 2 , 0.42% (CaO+MgO) was obtained under the conditions that roasted with 7.5%(wt) H 3 P0 4 at 1000 °C for 2h, followed by two-stage leaching in boiling dilute sulfuric acid for 2h. Mechanism studies showed that anosovite solid solution and silicate minerals were reacted with phosphoric acid in the roasting process. As a result, titanium components in titanium slag turned into rutile (TÍO2) and impurities were transformed into acid-soluble phosphate and quartz. Introduction With the gradual depletion of natural rutile and the rapid development of titanium industry in China, the natural rutile reserves cannot satisfy the requirement for production of titanium pigment and titanium sponge. Ilmenite is, therefore, becoming main raw material for titanium industry. However, because of its low TÍO2 grade, it generally requires preconcentration to titanium-rich material such as titanium slag or synthetic rutile [1-3]. Ilmenite resources are extremely rich in Pan-xi area of China, which account for 90.54% of total ilmenite reserves in China [4]. The titanium slag produced by electric smelting of ilmenite concentrate has high contents of Ca and Mg, and is not suitable as a raw material for chloride process and titanium sponge production. Therefore, upgrading titanium slag

1

Corresponding Author: Dr. Yufeng Guo, Email: [email protected], Tel: +86-731-88830542

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to prepare raw material suitable for chloride process and titanium sponge production has important significance for the development of China's titanium industry. Among the study of preparation of synthetic rutile from titanium slag [5-12], Canada QIT Company has successfully produced high-quality titanium-rich material from titanium slag, and achieved commercial production, but the kernel technology is unknown to outside. Many other processes, however, demand oxidation and reduction at a high temperature pretreatment. Furthermore, in the acid leaching process, they generally require high temperature and high pressure. High energy consumption, environmental pollution and high requirements for equipment limit the industrial application of those processes. In this paper, the process of activation roasting with phosphoric acid followed by acid leaching process has been studied, which can effectively remove the impurities in titanium slag such as Ca and Mg. The product obtained is suitable as a raw material for chloride process and titanium sponge production. Material Properties and Research Methods Material Properties Titanium slag used in this study was provided by Panzhihua Iron and Steel Company. The main chemical composition of the sample is shown in Table 1. Phosphoric acid is analytically pure regent. Table 1. Main chemical composition of titanium slag / wt% Composition Content Composition Content

Ti0 2 72.42 A1203 2.65

MFe 1.27 Cr 2 0 3 0.037

TFe 7.02

v2o5 0.14

Fe 2 0 3 0.10 MnO 0.74

FeO 7.31 P 0.0035

Si0 2 5.37 S 0.045

MgO 8.21

CaO 1.36

Ig 0.10

Research Methods Titanium slag was thoroughly mixed with phosphoric acid and pelletized; the pellets were roasted at various reduction temperatures and time in a tube furnace. The roasted products were grinded to 100% -74um. Specified amounts of preheated sulfuric acid of known concentration and the roasted products were loaded into the glass reactor. The stirring speed was maintained constant by means of a digital controlled stirrer. At the end of the runs, the slurry was filtered under vacuum. The resulting leached residues were washed with water and then calcined at 900 °C for 30min. The final product was analyzed chemically; the mineralogical analysis was carried out by X-ray diffraction.

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Results and Analysis Effects of Phosphoric Acid Dosage Effects of phosphoric acid dosage are given in Figure 2 and Figure 3. Parameters, which were kept constant during tests were at 1000 °C for 2h for roasting, and 2h with spend H 2 S0 4 containing 25% H 2 S0 4 for the first stage leaching, 2h with spend H 2 S0 4 containing 40% H 2 S0 4 for the second leaching.

Figure 1. Effects of phosphoric acid dosage on TÍO2 grade and recovery

Figure 2. Effects of phosphoric acid dosage on leaching of impurities

As shown in Figure 1 and Figure 2, T1O2 grade of the product was rapid increased as the phosphoric acid dosage increased from 0 to 10%. The leaching of impurities such as MgO, CaO, AI2O3, and Fe increased substantially at the same time. However, the leaching of silicon varies unconspicuously. The results suggested that roasting with phosphoric acid damaged the structure of anosovite in titanium slag. Impurities, which are dissolved in anosovite, are more likely to be leached by acid. Effects of Roasting Temperature Effects of roasting temperature are given in Figure 3 and Figure 4. Parameters which were kept constant during the tests were 2h for roasting, and 2h with spend H2SO4 containing 25% H2SO4 for the first stage leaching, 2h with spend H2SO4 containing 40% H2SO4 for the second leaching.

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Figure 3. Effects of roasting temperature on TÍO2 grade and recovery (H3PO4 content 7.5 wt %)

Figure 4. Effects of roasting temperature on leaching of impurities (H3PO4 content 7.5 wt %)

From Figure 3 and Figure 4, as the roasting temperature increased from 900 °C to 1200 °C, TÍO2 grade of the product was increased, leaching rate of impurities such as MgO, CaO and AI2O3 varied unconspicuously, while the leaching of Fe and Si0 2 decreased. The reason may be that silicate glass phase and iron phase in titanium slag were changed into new phases as the roasting temperature increased, the new phases, however, were not easily leached by acid. Accordingly, the optimal roasting temperature was selected to be 1000 °C. Effects of Roasting Time Effects of roasting time are given in Figure 5 and Figure 6. Other parameters kept constant were at 1000°C roasting, with spend H2SO4 containing 25% H2SO4 for the first stage leaching, 2h with spend H2SO4 containing 40% H2SO4 for the second leaching.

Figure 5. Effects of roasting time on Ti0 2 grade and recovery (H3PO4 content 7.5 wt %)

Figure 6. Effects of roasting time on leaching of impurities (H3PO4 content 7.5 wt %)

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From Figure 5 and Figure 6, as the roasting time increased from lh to 2h, Ti0 2 grade of the product was increased unconspicuously, which always stay in about 88%. Leaching of impurities such as MgO, CaO, A1203, Fe and SÍO2 varied unconspicuously. The results showed that roasting time has little effect on the experiment, but in order to make the titanium slag and phosphoric acid react completely, the roasting time was selected to be 2h. The Final Product Table 2 showed the main chemical composition of the product which prepared under the optimal condition. As shown in Table 2, the product, prepared by activation roasting followed by acid leaching from Panzhihua electric titanium slag, contains 88.54% Ti(>2, and 0.42% (CaO+MgO). It is a high-quality synthetic rutile product which is capable to meet the requirements of chloride process and titanium sponge production. Table2. Main chemical composition of the product/wt% Composition Content

Ti0 2 88.54

MgO 0.31

Si0 2 7.73

TFe 2.15

MnO 0.13

AI2O3

0.64

CaO 0.11

S 0.02

Mechanism of Activation Roasting Studies have shown impurity elements such as Mg, Mn and Fe were solid dissolved in anosovite, and stable compounds, MgTi205, MnTi205 and FeTi205 will be generated when smelting the titanium concentrates [13]. This so-called M 3 0 5 solid solution made the anosovite extremely stable, which was insoluble in HC1, NaOH, H3PO4 and cold H2SO4 solution. Therefore, it was difficult to upgrade the titanium slag by direct acid leaching. In this study, the slag phase was changed as the high-temperature and oxygen potential increased when titanium slag roasting with phosphoric acid. Due to the bonding force of PO/'and MesOî is greater than that of Ti„03n+i2(n+1)", the impurity elements were released from the anosovite lattice by formation a phosphate glass phase. This reaction undermines the stable structure of M3O5 and. On the other hand, oxidizes Ti components in titanium slag to Ti0 2 (ruitle) [14]. X-ray diffraction analyses of titanium slag samples roasted with different H3PO4 dosages and at different temperatures were analyzed to identify the mechanism of activation roasting.

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Figure 7. XRD analysis of titanium slag roasting with different dosage H3PO4 (Roasting temperature 1000 °C, roasting time 2h) (■-Me305 , ♦-Ti0 2 (ruitle) , A-Si0 2 , A-Mg3(P04)2 , * - Silicate minerals) As shown in Figure 7, the major phases of the roasted product without phosphoric acid were Ti0 2 (rutile), Me3Os and a small amount of silicate phase. While for the product roasted with 7.5wt% phosphoric acid, the diffraction peak of the Me 3 0 5 phase is weakened and the diffraction peak of Ti0 2 (rutile) strengthened. With diffraction angle 20 between 20° and 25°, there were two diffraction peaks of Mg3(PC>4)2 and a diffraction peak of Si0 2 . This indicated that anosovite solution structure and silicate minerals structure were damaged by roasting with phosphoric acid. As a result, impurities in anosovite were transformed into a phosphate, and titanium phases were transformed into rutile phase Ti0 2 , silicon phases into Si0 2 .

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Figure 8. X-ray diffraction analysis of titanium slag roasting at different temperature (H3PO4 content 7.5 wt%, roasting time 2h) (■-Me305 , ♦-Ti02(rutile) , A-Si0 2 , A-Mg3(P04)2 , *-Silicate minerals) Figure 8 shows the major phases of the product roasted at 400 °C were Me3C>5 and a small amount of silicate phase. As the roasting temperature rose to 1000 °C, the main phases of the roasted product were rutile Ti0 2 and Me3Os. The diffraction peaks of Me3Os, however, were weakened or even disappeared. The results indicated that temperature is important to the formation and aggregation of rutile phase. Conclusions Impurities such as Ca and Mg in titanium slag were effectively removed by phosphoric acid activated roasting followed by acid leaching. Ti0 2 grade of the product was significant improved. High-quality synthetic rutile containing 88.54% Ti0 2 with 0.42% (CaO+MgO) was obtained under the conditions of roasting at 1000 °C for 2h with 7.5wt %H3PC>4 dosage, the first stage leaching of 2h with spend H 2 S0 4 containing 25% H2SO4, the second stage of 2h with spend H 2 S0 4 containing 40% H 2 S0 4 .

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Phosphoric acid can react with anosovite solid solution and silicate minerals in the roasting process. As a result, titanium components in titanium slag were transformed into rutile Ti0 2 while impurities were transformed into acid-soluble phosphate and quartz. The activation roasting destroyed the structures of anosovite and silicate minerals. Acknowledgement The authors want to express their appreciation to The National Basic Research Program of China (973 Program) (No.2007CB613606) for financial support of this research. References [I] [2] [3]

[4] [5] [6] [7] [8]

[9]

[10]

Yang Yan-hua, Lei ting, Mi jiaRong. "Titania feedstock preparation methods and development proposals," Yunnan Metallurgy, 35 (1) (2006), 41 - 44 Yang Shao-li, Sheng Ji-Fu. Technology of smelting Ilmenite and iron (Beijing: Metallurgical Industry Press, 2006), 42 - 44 Liu bang-yu, Wang Ning, Chen Juan et al. "Ilmenite processing and comprehensive utilization of resources with high production of titanium slag," Minerals Engineering, 2007, supplement, 388 - 389 Wu Xian, Zhang Jian. "China's resource distribution and characteristics of titanium." Titanium Industry Progress, 23 (6) (2006), 8-12 Knizysztof borowiec. Method of Upgrade Titania Slag and Resulting Product.US Patent:5830420, 1998 QIT-Fertitane lnc.UGS process, http://www.qit.com Jacobus philippus Van Dyk, et al. Benefication of Titania slag by oxidation and reduction treatment. United States Patent, 6803024, Oct, 12, 2004 Michel Gueguin, Tracy, Canada. Method of preparing a synthetic rutile from titaniferous slag sontaining magnesium values. United States Patent 4933153, Jun.12, 1990 Michel Gueguin, Tracy, Canada. Method of preparing a synthetic rutile from a titaniferous slag containing alkaline earth metals. United States Patent 5389355, Feb.14, 1995 Michel Gueguin, Tracy, Canada. Method of preparing a synthetic rutile from a titaniferous slag containing magnesium values. United States Patent 5063032, Nov.5, 1997.

[II] J.P.VAN DYK , P.C.PISTORKJS. "Evaluation of a process that uses phosphate additions to upgrade titania slag," Metallurgical and Materials Transactions, 30B (4) (1999), 823-826 [12] Gerald W. Elger, Ruth A. Holmes, both of Albany, Oreg. Purifying titanium-bearing

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slag by promoted sulfation. United States Patent 4362557,Dec.7, 1982 [13] Lei Ting, Mi Jia-rong, Zhou Lin et al. Furnace slag composition and cleaning mechanism. The Fifth National Symposium on Rare Metals, 11(2006), 141 - 147 [14] Zhang Li, Li Guang-qiang, Sui Zhi-tong. "Prepared by the modified high titanium slag leaching of Ti-rich materials," Mineral Resources, 6(2002), 6 - 9

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

Calculation of Phase Equilibria Relations in CaO-Si0 2 -FeO s -MgO System Nan Wang1, Cuihuan Huang1, Xueqian Xin1, Zongshu Zou1, Z. Zhang2, Y. Xiao2, Y. Yang2 1

School of Materials and Metallurgy, Northeastern University Shenyang, China 2 Delft University of Technology, Delft, the Netherlands

Keywords: Bottom ash, Vitrification, CaO-Si02-FeOx-MgO system, Equilibrium phase relation, CALPHAD Abstract Vitrification is considered as an attractive procedure for municipal solid waste incineration ash, which can realize the volume reduction and innocent treatment of incineration ash, as well as reuse of bottom ash. The main oxide components of molten ash slag are CaO, AI2O3, SÍO2, Fe2C>3, MgO, and Na2Û, accounting for more than 90% of the total mass. The phase relations and thermodynamic properties of the CaO-Al203-Si02-FeOx-MgO-Na20 oxide system is highly required, and is of great scientific significance for optimization of vitrification processing and sustainable resource utilization of melting slag. A study on the equilibria phase relations of CaOSi02-MgO-FeOx system was done by using CALPHAD method based on FactSage in this paper. The modeling parameters for the thermodynamic properties of liquid phase in CaO-SiCh-MgOFeOx system were established and the equilibrium phase relations of MgO-Si02-FeOx and CaOSi02-FeOx-MgO systems were calculated for different temperatures and oxygen partial pressures. Introduction The main oxide components of vitrified bottom ash slag are CaO, AI2O3, SÍO2, Fe2Û3, MgO, and Na 2 0, accounting for more than 90% of the bottom ash by weight [1]. During slag processing and modification, the knowledge of thermodynamic properties and liquidus temperature in the related slag system is highly required [2-5], which is not yet available. Through literature survey [6-9], it is clear that the systematic information on the concerned complex slag system is scarce. Most binary subsystems except FeOx-Na20 have been investigated well. Among ternary subsystems, few phase relations of the Na20-MgO-Si02 ternary system are available. For the phase relations of the system Na20-Mg0-Al203, all the studies are focused on the Al203-rich region. Investigation on the CaO-Al203-FeOx system is limited to Fe203-containing system with relatively higher oxygen potential. Most quaternary systems are not studied well except the CaOMgO-Al203-Si02 system, mainly due to their metallurgical importance. Figure 1 demonstrates the research status of the vitrified bottom ash slag with almost zero effort [10]. The effort here is defined based on the published literature by considering the number of publications, the investigated composition ranges, and the research details. So far the engineers in the field have to make assumptions of the thermodynamic properties of the simplified oxide system, without taking into consideration the interactions of the different oxide components, which hinders the development of the slag valorization and utilization. In order to provide thermodynamic properties for the application of the vitrified bottom ash, the six-component oxide system (CaOMgO-Al203-Si02-FeOx-Na20) will be investigated through both experimental approach and computer modeling. In order to interpret and further to predict thermodynamic properties CaO-MgO-Al203-Si02FeOx-Na2Û multi-component oxide system, the experimental work is not enough because the experimental determination of phase diagrams and phase relations is a time-consuming and

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costly task. Moreover, in some oxide system, it is very difficult to do experimental investigation over wide temperature and composition ranges under the current experimental conditions. Therefore, CALPHAD approach will be employed to calculate the thermodynamic properties and phase relations in these oxide systems [11]. The calculation of phase diagrams reduces the effort required to determine equilibrium conditions in a multi-component system. A preliminary phase diagram of the CaO-MgO-Al203-SiC>2-FeOx-Na20 system can be obtained from extrapolation of the thermodynamic functions of constituent subsystems such as CaO-Si0 2 FeOx-MgO, CaO-Si02-FeOx-Al2C>3, CaO-Si02-Na20-MgO and CaO-Si02-Na2 0-Al203 . Thus, in the present work, the CaO-Si02-MgO-FeOx system, one subsystem of the complex CaO-MgOAl203-Si02-FeOx-Na20 system was investigated based on CALPHAD approach, and the modeling parameters for the description of liquid phase in CaO-Si02-FeOx-MgO system were established. The equilibrium phase relations and liquid region of MgO-SiC>2-FeOx and CaOSiCh-FeOx-MgO systems were calculated for different temperatures and oxygen partial pressures.

I Slag type

Fig. 1 Research status of the related oxide systems CALPHAD Theory and Thermodynamic Model Thermodynamic Description and Model For the calculation of phase equilibria in a multi-component system, it is necessary to minimize the total Gibbs energy, G, of all the phases that take part in this equilibrium, as shown in equation (1). Where n is the numbers of component, and q> is the numbers of phase. ° =ÉÈG'=G..

(« = l,2,-,n;# = a,Â-.f»)

(l)

A thermodynamic description of a system requires the assignment of thermodynamic functions for each phase. The CALPHAD method has to employ a variety models to describe the temperature, pressure and concentration dependencies of the free-energy functions of the various phases. In this work, in the case of molten silicate systems, a modified quasi-chemical model for short-range ordering is employed. The modified quasi-chemical theory is a mathematical formalism which has the advantages of simplicity and generality and which appears to have the characteristics required for relatively reliable interpolations and extrapolations of data into unmeasured regions and for extensions which can be used for multi-component systems. In this CaO-Si02-FeOx-MgO system, a modified quasi-chemical model is used for the liquid phase and compound energy model is used for the solid solutions. The polynomial model is used for extrapolation from subsystems to the high-order ones. Equilibrium Phase Relations of CaO-SiCh-FeOj-MgO System Model Parameters for the Thermodynamic Property of Liquid Phase

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According to the literatures [12-14], the modified quasi-chemical model was used for the liquid phase in CaO-Si02-FeOx-MgO system. The compound energy model was selected to describe the thermodynamic properties of spinel solid solution, olivine solid solution, proto-pyroxene solution, ortho-pyroxene solution and low clino-pyroxene solution. The simple polynomial models was used for wollastonite solution, a-Ca2SiC>4 solution and a'-Ca2SiC>4 solution, and the Köhler polynomial models was used to describe the thermodynamic properties of simple oxide solid solutions. According to the modified qusai-chemical model, the model parameters for the liquid phase in CaO-Si02-FeOx system were assessed by employing the experimental data [15], and then the equilibrium phase relations of MgO-Si02-FeOx and CaO-Si02-FeOx-MgO system were calculated based on both the assessed model parameters and those from literatures [14], with the Equili and Phase diagram modules of the FACT oxide solution databases in FactSage program. The assessed model parameters are as the followings.

Calculated Equilibrium Phase Relations in MeO-SiCb-FeCX System The equilibrium phase relations of the MgO-SiC>2-FeOx system for different temperatures and oxygen partial pressures were calculated based on both the assessed model parameters for liquid phases and those from literatures to describe the solid solutions. The calculated results are shown in Figure 2.

Fig. 2 Calculated equilibrium phase relations of MgO-Si02-FeOx system

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It is shown in Figure 2, in MgO-Si0 2 -FeO x system, under the condition of 1673K and low oxygen partial pressure of 10"8 arm, the liquid region emerges in the high FeOx zone closing to the SiCVFeOx boundary, and the liquid region shows the decreasing tendency as the temperature decreases to 1573K, while all the two-phase and three-phase regions containing liquid tend to expand without new phase. The phase relations for MgO are very simple, and only olivine and simple oxides (mainly MgO solid solution) can be found. Spinel solid solution appears with the increasing oxygen partial pressure and the phase relations in high-ferrous oxide region are changed to a large extent. The liquid field shrinks toward the low-ferrous oxide region. With the temperature decreasing to 1573K, the liquid region becomes smaller and disappears eventually. Calculated Equilibrium Phase Relations of CaO-SiO^-MgO-FeOv System Compared to the main components of vitrified bottom ash slag such as CaO, SÍO2 and FeOx, the amount of MgO is relatively small. Thus, the equilibrium phase relations of CaO-Si02-FeOx system at 1873K and with the oxygen partial pressure of 10"8 atm was calculated first, and the equilibrium phase relations of CaO-Si02-MgO-FeOx system with 2.1wt% MgO were calculated according to the relative proportion of Si0 2 , CaO, Fe 2 0 3 and MgO in the vitrified bottom ash slag. The effect with the addition of MgO on the equilibrium phase relations of CaO-Si02-FeOx system was discussed.

Fig. 3 Calculated equilibrium phase relations of CaO-Si02-FeOx system

Fig. 4 Calculated equilibrium phase relations of CaO-Si02-MgO-FeOx system with 2.1wt% MgO It can be found in Figure 3, there exists a large liquid phase region in CaO-Si02-FeOx system, under the condition of 1873K and an oxygen partial pressure of 10"8 atm. and the high-ferrous oxide region is composed of complete liquid phase. Compound CasSiOs emerges in the highCaO region closing to the CaO-Si02 boundary. The stable region of Ca3SiÛ5 is relatively small

140

which changes significantly with temperature. The effect of MgO addition on the Hquidus of CaO-SiC>2-FeOx system is shown in Figure 4. With addition of 2.1wt% MgO, pure MgO solid phase appears in the high CaO region, and Ca3SiOs disappears simultaneously. Moreover, "L+M" (primary phase region of simple single oxide solid solution) extends rapidly to the highFe oxide region and SÍO2 primary phase field reduces which contributes to the extension of liquid phase field to the high-SiU2 region. Conclusions Thermodynamic calculation has been applied to make an investigation on the equilibrium phase relations of CaO-Si02-MgO-FeOx system. The model parameters for the calculation of liquid phase in CaO-Si02-MgO-FeOx system were established and the equilibrium phase relations of MgO-Si02-FeOx and CaO-MgO-Si02-FeOx systems were calculated for different temperatures and oxygen partial pressures. The conclusions are as follows. (1) The model parameters for the calculation of liquid phase in CaO-Si02-MgO-FeOx system are assessed based on the literature data. (2) Based on the calculated equilibrium phase relations of MgO-Si02-FeOx, the liquid phase region emerges in the high-ferrous oxide region closing to the Si02-FeOx boundary in MgOSi02-FeOx system at low oxygen partial pressure (10 8 atm). Spinel solid solution appears with increasing oxygen partial pressure which changing the phase relations in high-ferrous oxide region to a larger extent. The liquid field shifts to low-ferrous oxide region rapidly. (3) As for the CaO-Si02-MgO-FeOx system with 2.1wt% MgO, the phase relations are very complex in high-CaO region, and MgO solid phase appears simultaneously. Complete liquid phase can be found in high-ferrous oxide region. Acknowledgements This research work was supported by the National Natural Science Foundation of China (Grant No. 50974034) and the Royal Netherlands Academy of Arts and Science (KNAW) (Project 08CDP026). References [1] Y. Xiao et al., "Vitrification of Bottom Ash from a Municipal Solid Waste Incinerator," Waste Management, 28 (6X2008), 1020-1026. [2] S. Y. Kim, T. Matsuto, and N. Tanaka, "Evaluation of Pre-treatment Methods for Landfill Disposal of Residues from Municipal Solid Waste Incineration," Waste Manage Res, 21 (2003), 416-423. [3] K. L. Lin, C. T. Chang, "Leaching Characteristics of Slag from the Melting Treatment of Municipal Solid Waste Incinerator Ash," Journal Hazardous Materials, B135(2006), 292302. [4] M. Suzuki, T. Tanaka, "Prediction of Phase Separation in Silicate Glass for the Creation of Value Added Materials from Waste Slag," ISIJInternational, 46(10) (2006), 1391-1395. [5] F. Andreola et al., "Reuse of Incinerator Bottom and Fly Ashes to Obtain Glassy Materials," Journal of Hazardous Materials, 153(3)(2008), 1270-1274. [6] G. Eriksson et al., "Critical Evaluation and Optimization of the Thermodynamic Properties and Phase Diagrams of the CaO-Al 2 0 3 , Al 2 0 3 -Si0 2 and CaO-Al 2 0 3 -Si0 2 Systems," Metall. Trans. B, 24 (50)(1993), 807-816. [7] L. Barbieri, A. Corradi, and I. Lancellotti, "Critical Thermodynamic Evaluation and Optimization of the MgO-Al 2 0 3 , CaO-MgO-Al 2 0 3 and MgO-Al 2 0 3 -Si0 2 Systems," Journal of the European Ceramic Society, 20(2000): 1637-1643.

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[8] E. Schiirmann, G. Kraume, "Effect of Oxygen Partial Pressure on Liquidus for the CaOSi0 2 -FeO x System at 1573K," Arch. Eisenhuilenwes, 47(6)(1976), 327-331. [9] E. F. Osborn, A. Muan, "Experimental Investigation and Optimization of Thermodynamic Properties and Phase Diagrams in the Systems CaO-Si02-FeOx," Ceramic Foundation, 1960, 57(178)(1960), 62-70. [10] Z. Zhang, "Literature Review of the CaO-Al 2 0 3 -Si0 2 -FeO x -MgO-Na20 System of Vitrified Bottom Ash Slag from Municipal Solid Waste Incinerators"(Internal Report, TU Delft, 2007). [11] A. D. Pelton, "Thermodynamic Database Development Modelling and Phase Diagram Calculations in Oxide Systems," Rare Metals, 25(5)(2006), 473-480. [12] H. Jung, S. A. Decterov, and A. D. Pelton, "Critical Thermodynamic Evaluation and Optimization of the CaO-MgO-Si02 System," Journal of the European Ceramic Society, 2005,25(2005), 313-333. [13] P. Wu, G. Eriksson, A. D. Pelton, "Critical Evaluation and Optimization of the Thermodynamic Properties and Phase Diagrams of the Calcia-iron(Il) Oxide, Calciamagnesia, Calcia-manganese(II) Oxide, Iron(II) Oxide-Magnesia, Iron(II) Oxidemanganese(II) Oxide, and Magnesia-manganese(II) Oxide Systems," Journal of the American Ceramic Society, 76(1993), 2065-2075. [14] H. Jung, S. A. Decterov, and A. D. Pelton, "Critical Thermodynamic Evaluation and Optimization of the Fe-Mg-O System," Journal of Physics and Chemistry of Solids, 65(2004), 1683-1695. [15] T. Ogura et al., "Activity Determinator for the Automatic Measurements of the Chemical Potentials of FeO in Metallurgical Slags," Metallurgical Transactions B, 23B(1992), 459465.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

DISSOLUTION BEHAVIOR OF RHODIUM INTO MOLTEN SLAG Chompunoot Wiraseranee1, Toru H. Okabe2, Kazuki Monta2 'Graduate School of Engineering, The University of Tokyo; 7-3-1 Hongo, Bunkyo-ku; Tokyo 113-8656, Japan institute of Industrial Science, The University of Tokyo; 4-6-1 Komaba, Meguro-ku; Tokyo 153-8505, Japan Keywords: Rhodium, Molten Slag, Platinum Group Metals, Recycling Abstract Aiming at the development of more effective recycling processes, dissolution behavior of rhodium into molten slag was investigated. Using a pure rhodium crucible and Na 2 0-Si0 2 slag system, equilibration experiments were carried out for 18 h at 1423-1623 K under various oxygen partial pressures of 0.005-1 atm with varied slag basicity. Rhodium solubility increased with increasing oxygen potential, slag basicity, and temperature. Rhodium dissolved into slag as rhodate ion ([Rh02]") by the endothermic reaction: Rh(s) + 3/40 2 (g) + l/202"(in slag) = [Rh0 2 ]" (in slag). In the dissolution reaction, the "rhodate capacity" of slag is newly-defined. The correlation between rhodate capacity and optical basicity is reported. Introduction Due to its high chemical resistance and outstanding catalytic properties, rhodium (Rh), one of the platinum group metals (PGMs), is utilized in various applications such as catalysts, molten glass vessels and containers, and electronic devices. Owing to the rarity of Rh resources in nature, recycled Rh from scraps such as spent automobile catalytic converters is currently considered as an important source of Rh supply [1]. According to the recycling practices, Rh is currently mainly recovered from scraps along with other PGMs such as platinum (Pt) and palladium (Pd) by a dry process (e.g., smelting), a wet process (e.g., leaching and subsequent separation of PGMs from aqueous solution), or a combination of both processes [2]-[4]. Among these processes, smelting is commonly applied because PGMs can be effectively separated from a relatively large amount of various impurities coexisting in scraps. In the smelting process, Rh and other PGMs are collected in the collector metal (e.g., iron and copper), and impurities in scraps are removed into slag [4]-[6], The process may be carried out at a temperature range of 1573-1973 K, in either an oxidizing or reducing atmosphere. Due to the high oxidation resistance at high temperatures, Rh is expected to be collected efficiently in the metal collector rather than being oxidized and forming RhOx in a molten slag. However, Rh loss into slag does occur in this process [7], The dissolution behavior of precious metals, including Pt, ruthenium (Ru), gold (Au), and silver (Ag) into slag has been reported. Nakamura et al. [8]-[9] investigated the dissolution behavior of Pt at 1373 K and 1873 K using various oxide melts such as Na 2 0-Si0 2 , CaO-Si0 2 and CaOAI2O3. It was suggested that Pt may exist as Pt cation (Pt2+) in an acidic slag and as a platínate ion ([Pt02]2") in a basic slag. Shuto et al. [10] investigated the dissolution behavior of Ru into molten slag at 1373-1873 K using Na 2 0-Si0 2 , CaO-Si0 2 , and Na 2 0-Si0 2 -Al 2 0 3 slag systems. It was suggested that Ru dissolved into slag as an acidic oxide ([Ru02]"). The "ruthenate capacity" of slag was defined, and the loss of Ru into slag in the actual process was estimated by applying

143

the theoretical optical basicity of slag. Swinbourne [11] investigated the solubility of Au and Ag in PbO-SiC>2 slag, and reported that the solubility of both Au and Ag increased with increasing oxygen partial pressure and content of PbO in the PbO-SiCh slag system. Even small amounts of the losses of Rh in recycling processes are taken into account seriously due to its scarcity and high price. Until recent years, the actual influences of factors such as temperature, oxygen partial pressure, and slag basicity on the dissolution behavior of Rh into slag have not been yet clarified. Therefore, the estimation of Rh loss into slag in a particular process is difficult. Hence, to develop an effective recycling process, the dissolution behavior of Rh into molten slag was investigated. In this study, the dependence of Rh solubility on oxygen partial pressure, temperature and slag composition were determined. The dissolution mechanism into molten slag is then discussed. An alternative method to estimate Rh loss into slag in a real process using theoretical optical basicity is also proposed. Experimental Principle Rh may dissolve into molten slag as metallic Rh, Rh oxide, or rhodate ion. The generic dissolution reaction can be written as follow: Rh (s) + m 0 2 (g) + no 2 " (in slag) = [Rh0 2m+n ] 2n - (in slag)

(1)

The equilibrium constant of the dissolution reaction (1), K(¡), is written as:

<1)

"^-(Po2)m-K2-)"

()

'

where a¡ and p, are the activity of a component ; and the partial pressure of a component i, respectively. According to the reaction (1), the dissolution of Rh into slag depends on temperature, oxygen partial pressure (p0 ) and oxide anions (O 2 ) in slag. Thus, the dissolution behavior can be determined from the dependence of Rh solubility on those parameters. Meanwhile, the dissolution of Rh as a cation results in the changing of Rh solubility upon varying the oxygen partial pressure. The dissolution as rhodate ion results in Rh solubility depending on both the oxygen partial pressure and activity of oxide ions(a 0 2 .). To determine the dissolution mechanism of Rh, a pure Rh crucible was equilibrated with molten Na2Ü- SÍO2 slag. Therefore, the activity of Rh is unity. The relationship between the Rh solubility, oxygen partial pressure, and the activity of oxide ions can be expressed as: log (mass% [Rh02m+n]2n") = m log p0 + n log aQ2. + log Km - log /

2n.

(3),

where // is the activity coefficient of a component ; in the Henrian 1 mass% standard state. Experimental procedures Na20-Si02 slag was prepared by melting a mixture of Na 2 C0 3 and Si0 2 (reagent grade) at 1473 K for at least 1 h. Equilibrium experiments were carried out in a SiC electric furnace using a pure

144

Rh crucible (18 mm inner diameter, 20 mm outer diameter, 21 mm in depth) and Na20-SiC>2slag (4 g). A schematic diagram of the experimental setup is shown in Figure 1. The oxygen partial pressure of gas flowing in the system was controlled using dried air or a mixture of pure O2 and Ar gas. Moisture in all gases was removed by passing the gases though silica gel. CO2 in Ar and O2 gas was removed using soda-lime and magnesium perchlorate. The temperature in the furnace was controlled within ±1 K using a proportional-integral-derivative (PID) controller with a Pt/6%Rh-Pt/30%Rh thermocouple. After equilibration for 18 h, the sample was removed from the furnace and quenched to room temperature using Ar gas. All samples were stored in a desiccator to shield them from moisture.

Figure 1. A schematic diagram of the experimental setup. Analysis Slag samples powder (0.1-1 g) were digested to acid-soluble form by alkali fusion using 0.8 g K2CO3 and 0.2 g LÍ2B4O7 fluxes in a Pt crucible. The fiised samples were then dissolved in HC1 (25-50% by volume) at room temperature. The Rh and sodium contents in the sample solution were determined by inductively coupled plasma-atomic emission spectrometry (ICP-AES). The Si0 2 content in slag was determined by the gravimetric method. Results and Discussion Dependence of Rh Solubility on Oxygen Partial Pressure and Temperature The dependence of Rh solubility on the oxygen partial pressure into 50(mass%)Na2O-50SiO2 slag was investigated at 1473 K. Over the oxygen partial pressure range of 0.005-0.21 atm, Rh solubility increased with increasing oxygen partial pressure (Figure 2).

145

-2.5

-2.0

-1.5 -1.0 log/?o2(atm)

-0.5

0.0

Figure 2. Oxygen partial pressure dependence of Rh solubility in 50(mass%)Na2O-50SiO2 slag at 1473 K. According to equations (1) and (3), the valence state of Rh dissolving into slag can be determined from the slope of the plot between the Rh solubility and oxygen partial pressure in logarithmic form. Considering the dependence of oxygen partial pressure at a constant temperature and a fixed slag composition, the slope "m" can be determined from the plot, which is close to 3/4. Therefore, it can be claimed that Rh dissolves into slag as Rh0 3 / 2 . At this point, the dissolution reaction can be written as shown in equation (4). Considering that the common oxidation state of Rh is Rh3+, the dissolution of Rh as RJ1O3/2 is reasonable. Rh (s) + - 0 2 (g) = RhOs« (in slag)

(4)

The dependence of Rh solubility on temperature in 50(mass%)Na2O-50SiO2 slag in air over 1423-1623 K is shown in Figure 3. The solubility increases with increasing temperature, indicating that the dissolution reaction is endothermic. Applying the van't Hoff equation (5), the enthalpy change of Rh dissolution into slag is calculated as +49 kj/mol, from the slope of the plot between the Rh solubility and the reciprocal temperature (1/7), as shown in equation (6).

6.0

6.2

6.4 6.6 6.8 7.0 7.2 KH/rCK"1) Figure 3. Temperature dependence of Rh solubility in 50(mass%)Na2O-50SiO2 slag in air.

146

ô \nK

AH°= - 2.303 R

AH" R

d log (mass% Rhin slag)

8(j)

(5)

(6)

Provided that Rh dissolved into slag as Rh0 3 / 2 , the dissolution can be divided into two main steps: the oxidation of Rh to RI1O3/2 and the dissolution of RI1O3/2 into slag. The enthalpy change of Rh oxidation to RI1O3/2 is calculated as -167 kj/mol, from the thermodynamic data of Rh2C>3at 298-1300 K [12]. Consequently, the enthalpy change of the dissolution of Rh0 3 / 2 into slag is calculated as +216 kJ/mol. According to a previous study, Pt also dissolved into slag by the endothermic reaction [8], However, Ru dissolved into slag by the exothermic reaction in which the enthalpy change of dissolution was determined as -130±20 kJ/mol [10]. Dependence of Rh Solubility on Basic Oxide Content The dependence of Rh solubility on slag composition in the Na20-SiC>2 slag system at 1473 K in air is shown in Figure 4(A).

1000 I 800 S ^g>600 J2 400 S S 200

♦

♦

♦

♦

o '—■—'—■—'—■—«—■—«

2.0

30

40 50 60 70 -5 -4 -3 -2 -1 log Cs2Na 2 0 content in slag/ mass% (A) (B) Figure 4. (A) Dependence of Rh solubility on slag basic oxide content in Na 2 0-Si02 slag system in air at 1473 K. (B) Relationship between Rh solubility and sulfide capacity in Na 2 0-Si0 2 slag system in air at 1473 K. Rh solubility increases with increasing basic oxide, i.e., Na20, content. Therefore, O2" ions affects the dissolution of Rh into slag. From equations (1) and (3), a number of moles of O2", "n", which reacts with a mole of Rh, can be determined from the linear relationship between logarithm of Rh solubility and aQ2. at constant temperature and oxygen partial pressure. However, aQl_ cannot be directly measured. To replace aQ2. by measurable parameters, sulfide capacity, C s 2 ., is applied [13]. Sulfide capacity is the capacity of a particular slag to hold S2" ions, defined as:

147

S2+02-=S2-+-02 2~¿ " " ' 2 c

(7)

_ (mass% S2- in slag) • (p0l ) 2 _ aQl. = K s2-= i m ~i /s2 (Ps 2 ) 2 -

(8)

The relationship between sulfide capacity and Rh solubility is obtained, as shown in equation (9). Although the activity coefficients of sulfide and rhodate ions are dependent on slag composition, the last term of equation (9) is considered to be constant if these coefficients have the same dependency on the slag composition. As a result, it is expected that Rh solubility and sulfide capacity will have a linear relationship where n can be estimated from the slope of the plot between Rh solubility and sulfide capacity. K log (mass%[Rh02m+n]2n-) = n log C s2 . + m log p0i + log —^-

(f )" - log - — -

K

( 0)>

/[Rh02 m + n ] 2 »-

(9)

The sulfur partition between Na20-SiÛ2 slag systems and carbon-saturated iron melts at 15231623 K was investigated [14]. Using the correlation of sulfide capacity with the slag composition of the Na2Û-Si02 slag system, the relationship between Rh solubility and sulfide capacity is plotted in Figure 4(B). From the plot, n is estimated to be 1/2. Although the slope is lower than it is estimated, because of the obvious increasing trend of the Rh solubility with increasing basic oxide content, the n value is assigned to the nearest possible value. From the present data, Rh dissolves into slag as rhodate ions, and the Rh dissolution reaction into slag can be written as: Rh (s) + ^ 0 2 (g) + ^-O2- (in slag) = [Rh0 2 ]" (in slag)

(10)

The existence of [RhOJ" ions is supported by research on the formation of ternary compounds such as MgRh 2 0 4 [15] and CaRh 2 0 4 [16]. Rhodate Capacity According to the definition of the capacity of slag, "rhodate capacity" is defined regarding the dissolution reaction of Rh into slag as rhodate ions in equation (10). The rhodate capacity, the capacity of slag to hold Rh as rhodate ions, can be expressed as: C

[Rho2l-

_ (mass% [Rh0 2 ]- in slag) _ (aQ2. ) 5 " "1 -Km~f (Po 2 ) 4 f l -

(11)

optical basicity is one of the basicity measures of slag, proposed by Duffy and Ingram [17]-[18]. It is expressed in terms of the electron donor power of the oxide ion in each oxide. The optical basicities of a few oxides are listed in Table I.

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Table I. Optical basicity of some oxides [18]

Optical basicity 1.15 0.48 1.00 1.00 0.73-0.81 0.65

Oxides Na 2 0 Si0 2 CaO FeO Fe 2 0 3 Cu 2 0

The rhodate capacity of a slag tends to increase with increasing slag basicity. The relationship of rhodate capacity and optical basicity, An,, is shown in Figure 5. Hence, the correlation between rhodate capacity and optical basicity, using Na20-Si02 slag within the range of Na20 content from 45 to 65 mass%, can be written as: logC[Rh02]. = 2.2/(th-2.6

(12)

-0.2 , -0.4 S

1-0.6 cm a

-0.8 -1.0 0.75

0.80

0.85

0.90

0.95

Figure 5. The correlation between rhodate capacity and optical basicity. Provided that the optical basicity of a particular slag system is known, the maximum Rh loss into slag can be estimated from equation (12). Conclusion The dissolution behavior of Rh into molten slag using a Na20-Si02 slag system was investigated. It was revealed that the Rh solubility increases with increasing temperature at 1423-1623 K, oxygen partial pressure at 0.005-1 arm, and basic oxide content of 45-65(mass%)Na20. Rh dissolves into slag as rhodate ions, [Rh02]\ by the endothermic reaction: Rh(s) + 3/402(g) + l/202"(in slag) = [Rh02]"(in slag). The rhodate capacity of slag was defined and its correlation with the optical basicity was determined. References 1. D. Jollie, Platimun 2010 (England: Johnson Matthey Public Limited Company, 2010). 2. J.S. Yoo, "Metal Recovery and Rejuvenation of Metal-loaded Spent Catalysts," Catal. Today., 44 (1998) 27-46.

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3. C. Hagelüken, "Recycling of Electronic Scrap at Umicore's Integrated Metals Smelter and Refinery" (Paper presented at the European Metallurgical Conference EMC 2005, Dresden, 18-21 September 2005), 152-161. 4. T.H. Okabe, H. Nakada, and K. Morita, "Recovery Technology of Platinum Group Metals," Journal of the Surface Science Society of Japan, 29 (10) (2008), 592-600. 5. A. Gibbon, J.E. Harry, and D. Hodge, "The Plasma Process for the Recovery of the Platinum Metals from Autocatalyst" (Paper presented at 8th International Symposium on Plasma Chemistry ISPC-8, Tokyo, 1987), 1862-1867. 6. M. Benson et al., "The Recovery Mechanism of Platinum Group Metals from Catalytic Converters in Spent Automotive Exhaust Systems," Resour. Conserv. Recy., 31 (2000) 1-7. 7. J.W. Matousek and J.G. Whellock, "Precious Metals Losses in Pyrometallurgical Processing," CM Bulletin, 83 (933) (1990), 97-101. 8. S. Nakamura and N. Sano, "Solubility of Platinum in Molten Fluxes as a Measure of Basicity," Metall. Mater. Trans. &, 28B (1997), 103-108. 9. S. Nakamura et al., "The Influence of Basicity on the Solubility of Platinum in Oxide Melts," Metall. Mater. Trans. B., 29B (1998), 411-414. 10. H. Shuto, T.H. Okabe, and K. Morita, "Ru Solubility and Dissolution Behavior in Molten Slag," Mater. Trans., submitted. 11. D. Swinbourne, "Solubility of Precious Metals in Slags" (Paper presented at the European Metallurgical Conference EMC 2005, Dresden, 18-21 September 2005), 152-161. 12.1. Barin, Thermochemical Data of Pure Substances (Weinheim, Germany: VCH Verlagsgesellschaft, mbH, 1989). 13. F.D. Richardson and C.J.B. Fincham, "Sulphur in Silicate and Alumínate Slags," J. Iron. Steel. I., (September 1954), 4-15. 14. R. Inoue and H. Suito, "Sulfur Partitions between Carbon-saturated Iron Melt and Na20-SiÛ2 Slags," Transactions ISIJ, 22 (1982), 514-523. 15. J. Nell and H. O'Neill, "The Gibbs Free Energy of Formation and Heat Capacity of ß-Rh2U3 and MgRh2Û4, the MgO-Rh-O Phase Diagram, and Constraints on the Stability of Mg 2 Rh 4+ 0 4 ," Geochim. Cosmochim. Ac, 61 (19) (1997), 4159-4171. 16. K.T. Jacob and Y. Waseda, "Solid State Cells with Buffer Electrodes for Measurement of Chemical Potentials and Gibbs Energies of Formation: System Ca-Rh-O," J. Solid. State. Chem., 150 (2000), 213-220. 17. J. A. Duffy and M. Ingram, "An Interpretation of Glass Chemistry in Terms of the Optical Basicity Concept", J. Non-Cryst. Solids, 21 (1976), 373-410. 18. J. A. Duffy, "A Review of Optical Basicity and Its Application to Oxidic Systems," Geochim. Cosmochim. Ac, 57 (1993) 3961-3970.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

"ONE STEP" TECHNOLOGY TO SEPARATE COPPER, ZINC, LEAD FORM IRON IN METALLURGICAL SLAG AND PYRITE CINDER: PART 2 - PILOT TEST De-qing Zhu, Dong Chen, Jian Pan, Yu Cui, Tie-jun Chun School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, China Keywords: Pyrite Cinder, Metallurgical Slag, Strand Grate - Rotary Kiln, Chlorination, Prerduction, Prereduced Pellets Abstract A study of processing pyrite cinder and metallurgical slag using "one step" technology in strand grate - rotary kiln is presented, which aims to efficiently recover the valuable metals of iron, lead, zinc and copper. The effect of chloridizing roasting parameters including roasting time and temperature on the removal rates of nonferrous metals was investigated. It is shown that copper and zinc are difficult to remove and lead is removed more easily. The preheated pellets were reduced in a rotary kiln, where zinc and lead can be removed easily and high iron grade and good metallurgical performance prereduced pellets can be obtained. Furthermore, the chlorination and the reduction mechanisms of copper, lead and zinc are discussed. It is revealed that the phases of copper, lead and zinc in slag are complex and ferrite and silicate of metals are difficult to remove. Introduction As two kinds of significant secondary resources, metallurgical slag and pyrite cinder contain not only abundant iron but also considerable copper, lead, zinc and so on. Several tens of millions of tons of metallurgical slag and pyrite cinder are discharged in nonferrous metal and chemical industries annually in China [1, 2]. In addition, the accumulated waste slag in storage is more than one hundred million tons, which not only occupies much land but also pollutes the environment [3]. However, because the metallurgical slag and pyrite cinder are treated by high temperature processing, the properties are distinctly different from the properties of natural ores, and iron is in close association with copper, lead and zinc. So, metallurgical slag and pyrite cinder are difficult to efficiently utilize and only a small amount is used as additives in cement, paving, brickmaking and auxiliary additives [4-7]. Now, more and more metallurgical slag and pyrite cinder are used as blast furnace burden by sintering and pelletizing, if they contain low nonferrous metals and high grade of iron [8-10]. However, there are some disadvantages by sintering and pelletizing: first, the grade of iron should be high, especially oxidized pellets (the grade of iron should be more than 63%); secondly, the nonferrous metals must be very low because they cannot be removed by these technologies. According to the literature [11-13], high temperature chloridizing roasting technology has been successfully applied to recover iron and valuable nonferrous metals. But the conventional high temperature chloridizing roasting is mainly used for processing high iron grade ores and heavily erodes equipment, especially rotary kilns. Therefore, it is necessary to explore a feasible process to recover iron and nonferrous metals

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from metallurgical slag and pyrite cinder which contain low iron grade and high content of nonferrous metals. In the previous study, the "one step" grate-kiln process including high temperature chloridizing roasting and prereduction at a laboratory scale was successfully tested. The process was investigated at a pilot scale in the present paper. In addition, the chlorination and the reduction mechanisms of copper, lead and zinc in grate chain and rotary kilns are discussed. Raw Materials

Raw Materials and Experimental

The chemical composition of the pellet feed is listed in Table 1. Pellet feed containing 61.24% Fe is too low for oxide pellets. And the high content of nonferrous metals is not suitable for ironmaking. So, copper, zinc and lead will be removed and the grade of iron will be further upgraded in this research. In addition, the content of sulfur in the pellet feed is slightly high, which may lead to high content of sulfur in the preheated pellets. 6.43% of SÍO2 in the pellet feed is unfavorable to ironmaking. But SÍO2 favors chlorination by promoting the formation of hydrogen chloride [13]. Table 1. Chemical composition of pellet feed* (mass %) Fetctai FeO Fe 2 0 3 SiQ2 AI2O3 CaO MgO P S Cu 61.24 19.73 65.60 6.43 1.21 0.51 0.41 0.050 0.27 0.218 *Pellet feed is composed of metallurgy slag and pyrite cinder.

Pb 1.788

Zn 0.199

The industrial analysis of the bituminous coal is given in Table 2. The bituminous coal which contains 49.74% of fixed carbon, 11.14% of ash and 32.85% of volatile matter is suitable to reduce pellets [14]. 0.42% of total sulfur in coal is slightly high. Table 2. Industrial analysis of bituminous coal Fcad/mass% Coking index Vdaf7mass% Ad/mass% Mad/mass% 49.74 11.14 6.27 32.85 2 *St—total sulfur in coal.

St/mass% 0.42

Experimental Methods Chloridizing roasting and Preheating of Pellets. Part removal of nonferrous metals and hardening of green balls will be achieved by preheating green balls containing chlorides in strand grate and imparting enough strength to preheated pellet to withstand reduction in a rotary kiln. Figure 1 shows the set-up for preheating and chlorination. At the beginning of preheating, a batch of green balls (12kg) was charge into a preheating cup. A combustion chamber, filled with natural gas and air, was used to supply heat for preheating. The combustion chamber was heated to the desired treatment temperature by adjusting the natural and air flow rates. The green balls were dried by blasting and drafting heated air at 250-300°C at an appropriate wind speed. Then the temperature of the preheating cup was raised to the set temperature and kept constant for the set time at 2.4m/s wind speed. The compressive strength of the preheated pellets was measured by using an intelligent compressive strength instrument (type: ZQYC) according to the standard of ISO4700. The removal rate of nonferrous metals was calculated by Equation (1) based on the chemistry of the preheated pellets and the dry pellets.

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Rm=(M0xm{l-Mlxm,)/(M0xmll)x\OQ

( 1)

where : Rm — Removal rate, %; Mo — The content of nonferrous metals in dry green balls, %; m 0 — Weight of dry green balls, g; M, — The content of nonferrous metals in preheated or prereduced pellets, %; mt — Weight of preheated or prereduced pellets, g;

Figure 1. General overview of grate chain Prereduction of Preheated Pellets, The prereduction tests were performed in a rotary kiln (see Figure 2). Natural gas and air, fired in the rotary kiln, were used to supply heat for reduction. At the beginning of reduction, the temperature of kiln was heated to 1000°C and a batch of hot preheated pellets (about 24kg) was charged into kiln with a size of 1000mm in diameter and 500mm width at Irad/min and preheated for 10min. Then 20% of the coal was charged into the kiln by a self-conveyor feed. When the temperature reached the target value, the remaining coal was uniformly charged into the kiln and coal filling was finished five minutes earlier than the end of reduction. When the reduction finished, the hot prereduced pellets and the remaining coal which was not combusted completely were loaded into a stainless steel pot, which was cooled by nitrogen gas to prevent the prereduced pellets from oxidizing. Finally, the cool prereduced pellets and the remaining coal were dissociated by magnetic separation and then the compressive strength and the chemical composition of the prereduced pellets were measured. The calculation of the removal rates of nonferrous metals during reduction is presented in Equation ( 1 ).

Figure 2. General overview of rotary kiln

153

Results and Discussion Preheating and Chlorination in Grate Chain The effect of preheating temperature on the compressive strength of the preheated pellets and the removal rates of nonferrous metals are illustrated in Figure 3. It can seen that the compressive strength of preheated pellets dramatically increases and the removal rates of copper and zinc slightly increase with an increase in preheating temperature, whereas the removal rate of lead clearly increases. However, the removal rates of copper, zinc and lead begin to decline when the preheating temperature is beyond 1050°C. The highest removal rates of copper, lead and zinc are 13.64%, 76.37% and 25.00%, respectively, when the pellets are preheated at 1050°C. Preheating at higher temperatures accelerates the chlorination reactions and the volatilization of metal chlorides, which favors the removal of nonferrous metals. But it is obvious that the higher the temperature, the more volatilization of CaCb and HCl, which results in less HCl to take part in the chlorination reactions. The porosity of the preheated pellets decreases with increasing comprehensive strength, which is unfavorable to the diffusion of HCl and the volatilization of metal chlorides. So the removal rates of nonferrous metals decrease when the preheating temperature exceeds 1050°C in the chain grate. After the optimization of preheating temperatures, the contents of nonferrous metals of preheated pellets were still too high, so the preheating time should be further extended.

Figure 3. The effect of preheating temperature on the compressive strength of the preheated pellets and the removal rates of nonferrous metals (2% CaCl2 and preheating for 10min with 2.4m/s wind speed, a-compressive strength, b-removal rate of nonferrous metals) Preheating for a longer time favors the solid phase diffusions and reactions fully occurring inside pellets [15], which results in higher compressive strength of the preheated pellets (see Figure 4a). The removal rate of copper slightly increases and the removal rate of zinc is constant with an extension of preheating time, while the removal rate of lead visibly increases within 10min (see Figure 4b). It is clear that the preheating time must be more than 10min.

154

Figure 4. The effect of preheating time on the compressive strength of the preheated pellets and the removal rates of nonferrous metals (2% CaCh and preheating at 1050°C with 2.4m/s wind speed, a-compressive strength, b-removal rate of nonferrous metals) In the pellet feed, the zinc, lead and copper mainly exist in the form of zinc oxide, zinc ferrite, lead oxide, lead silicate and copper oxide. Figure 5 shows that lead oxide and lead silicate are very easily removed, copper oxide is semi-removed and zinc oxide and zinc ferrite are the most difficult to remove. The removal rate of zinc is generally consistent with the test result in grate chain. In contrast, the removal rates of copper and lead in the grate chain (see Figure 3b) are obviously less than that in the tube furnace (see Figure 5). In the grate chain, the high temperature and the high flow rate of gas (2.4m/s) were used to heat pellets. The high flow rate of gas increases the volatilization of CaCb, and the dosage of calcium chloride is very important to the removal of lead and copper (see Figure 6), so the removal effects of lead and copper in the grate chain is far inferior to the removal effects in tube furnace. Thus it can be seen that the dosage and the volatilization of calcium chloride have a great influence on the removal of lead and copper.

Figure 5. Removal rates of nonferrous metal phases contained in pellets (2% CaCl2 and preheating at 1050°C for 10min in a tube furnace)

figure 6. The effect of dosage of calcium ™nfen™s metal phases contained in pellets (preheating at 1050°C for 10min in a tube furnace)

c h l o n d e on removal rates of

Reduction of Preheated Pellets

155

The tube furnace tests indicated that the higher the temperature, the higher the quality of prereduced pellets within the temperature range of 1000-1100°C. But the ash content in coal will easily react with iron bearing materials and materials of low melting point will be produced at high temperatures, which will easily result in the formation of accretion inside a rotary kiln. Therefore the operating temperature of rotary kiln must be 100-150°C less than the softening temperature of coal [ 14].In this pilot test, the reduction temperature was set at 1075°C. Figure 7 depicts the effect of the ratio of C/Fe on the metallization rate and the compressive strength of the prereduced pellets and the removal rates of nonferrous metals. With the ratio of C/Fe increasing, the compressive strength and the metallization rate significantly increases, the removal rate of zinc slightly increases and the removal rate of lead is consistently higher than the removal rate of copper. The concentration of CO increases with increasing ratio of C/Fe, which promotes the reduction of iron oxides and nonferrous metals. Also, metallization is favoured because metallic iron granules interconnect and metallic bonds are being fully developed. All of these contribute to the improvement of compressive strength of prereduced pellets. When the ratio of C/Fe reaches 0.7, the metallization rate, the compressive strength and the removal rates of zinc, lead and copper reach 73.2%, 535N per pellet and 84.5%, 95.27% and 13.64%, respectively. After optimization of the ratios of C/Fe, the compressive strength of the prereduced pellets was too low, so the reduction time should be further prolonged.

Figure 7. The effect of the ratio of C/Fe on the metallization rate and the compressive strength of the prereduced pellets and the removal rate of nonferrous metals (2% CaCl2; preheating at 1050°Cfor 10min; reducing at 1075°Cfor 70min; a- the metallization rate of prereduced pellets and removal rate of nonferrous metals, b- the compressive strength of prereduced pellets) Figure 8 shows the effect of reduction time on the metallization rate and the compressive strength of the prereduced pellets and the removal rates of nonferrous metals. In contrast to the removal rate of copper being kept steady, the removal rates of zinc and lead slightly increase and the compressive strength of the prereduced pellets significantly increases while the metallization rate falls with increasing reduction time. The reducing atmosphere attenuated in the whole period of reduction when the reduction time was increased to 130min, leading to a fall in metallization rate. After the optimization of C/Fe ratio and reduction times, the compressive strength of the prereduced pellets was also too low, so the comprehensive strength of preheated pellets needs to be further enhanced.

156

Figure 8. Effect of the reduction time on the metallization rate and the compressive strength of the prereduced pellets and the removal rate of nonferrous metals (2% CaCh; preheating at 1050°C for 10min; reducing at 1075°C with a C/Fe ratio of 0.7; a- the metallization rate of prereduced pellets and removal rate of nonferrous metals, b- the compressive strength of prereduced pellets) With the comprehensive strength of preheated pellets increasing from 760 N pre pellet to 1461 N pre pellet, the comprehensive strength of prereduced pellets significantly increases from 660 N to 1069 N, and the metallization rate of prereduced pellets and the removal rates of nonferrous metals drop (see Figure 9). FeîCh granules are more easily interconnected and the Fe2Û3 granules become large with higher preheating temperature, which results in higher compressive strength of preheated pellets. The larger Fe2Û3 granules and the inter-connected FezCh granules favor the higher compressive strength of prereduced pellets because the larger Fe2Ü3 granules and the inter-connected Fe2Ch granules can promote the metallic iron granules interconnection and metallic bonds being fully developed. However, the porosities decrease with the increasing comprehensive strength of preheated pellet, preventing CO diffusion inside pellets and reducing the reaction surface area [14]. So the metallization rate of the prereduced pellets and the removal rates of nonferrous metals drop.

Figure 9. The effect of the compressive strength of preheated pellet on the metallization rate and the compressive strength of the prereduced pellets and the removal rate of nonferrous metals (2% CaCl2; reducing at 1075°C for 130min with a C/Fe ratio of 0.7; a- the metallization rate of prereduced pellets and removal rate of nonferrous metals, b- the compressive strength of prereduced pellets) The reduction status of iron in prereduced pellets is shown in Figure 10, where metallic iron

157

exists in the form of graininess and jointed. The metallic iron granules interconnect and form reticular bridges of metallic bonds, which contribute to the high comprehensive strength of prereduced pellets. But in this test, the pellet feed contains high impurity content (especially 6.34% SÍO2 and 1.21% AI2O3). The impurities prevent the metallic iron from interconnecting. On the other hand, the metallization rate of prereduced pellets is low and the centre of prereduced pellets was reduced to ferriferrous oxide. So the comprehensive strength of prereduced pellets is obviously lower than that of direct reduced pellets. The centre of prereduced pellets was reduced to ferriferrous oxide and the reduction of lead silicate to lead must have ferrous oxide to take part [16], so part of the lead cannot be removed. The concentration of CO in the centre is too low to reduce the zinc oxide, which results in low removal rate of zinc.

Figure 10. Reduction status of iron in the prereduced pellets Reflection 200 times (MFe—bright white, graininess, jointed; FeO—white, spotted state; Fe304—Daffodil yellow, massive and granular; Slag phase—charcoal grey, embedded in centre and massive and granular in outer layer; Hole—black, irregular) Chemical compositions of the prereduced pellets that were prepared by the pilot test with the optimal parameters of preheating and reduction are presented in Table 3.The results show that the high degree of iron, low content of phosphorus, sulfur and zinc and slightly high content of copper and lead, can basically meet the quality requirement of blast furnace burden. Table 3. Chemical compositions of prereduced pellets (mass %) TFe Fez+ Fe J+ MFe Si0 2 A1203 CaO MgO Cu Pb Zn S P 73.17 32.64 6.74 33.79 7.78 1.99 1.65 0.87 0.18 0.16 0.065 0.017 0.039 The metallurgical performances of the prereduced pellets that were prepared by the pilot test are presented in Table 4. Table 3 shows that the iron exists almost in the form of metallic iron and Fe2+ in the prereduced pellets and only 6.74% iron exists in the form of Fe3+, resulting in the reducibility index of the prereduced pellets is 30.12% lower than that of the oxidized pellets. However, most of Fe3+ exists in the form of Fe 3 0 4 . So the expansion step of the reduction of Fe203 to Fe304 did not almost happen in the test of metallurgical performance and the reduction swelling index is -3.7%. The prereduced pellets have high indexes of reduction strength (RDI+ÍJ) and reduction disintegration (RDI+315) , low abrasion index (RDI.05) and high compressive strength after the assay of metallurgical performance. So, the prereduced pellet is likely to have good metallurgical performance. It is proved that the prereduced pellets have not only good cold strength but also good hot performance.

158

Table 4. Metallurgical performances of prereduced pellets ReducMity index(RI)/%

J ^ g ^ E f f ^ , Reduction swelling mdex(RSI)/% - ^ , 3 m n f + 3 . 1 5 m m -0.5mm

Compressive strength ^ ^

30.12 -3/7 97.47 98.83 0.37 894 Compressive strength refers to the strength of the prereduced pellet that was assayed after metallurgical performance test. Conclusions 1) High temperature chloridizing roasting in grate chain can remove the most of lead (78%) and only 20% Zn and 22.72% Cu. The high flow rate gas dramatically reduces the removal rate of lead and copper. Thus, more work will be dong to restrain the volatilization of calcium chloride and remove copper in grate chain. 2) Prereduction in rotary kiln is an effective way to not only further remove lead and zinc but also dramatically heighten the grade of iron. When preheated pellets were reduced at 1075°C for 130min with 0.7 ratio of C/Fe, the metallization rate and the compressive strength of the prereduced pellets and the removal rates of zinc, lead and copper are 46.18%, 1069 N per pellet, 67.5%, 91.21% and 18.18%, respectively. The test reveals that the high removal rates of lead and zinc need high metallization rate. And high enough comprehensive strength of prereduced pellets demands high comprehensive strength of preheated pellets. 3) The "one step" grate-kiln technique is a feasible way to remove nonferrous metals and heighten the grade of iron. The quality of prereduced pellets can cover the base needs of blast furnace. Especially, the metallurgical performances of prereduced pellets are very good. References 1 Guo Z.H. et al., "Mineralogical Characteristics and Environmental Availability of Non-ferrous Slag," Journal of Central South University (Science and Technology), 38(2007), 1000-1005. 2 Liu X.Z. et al., "Comprehensive Utilization of Burned Slag of Sulphuric Acid Making," Met. Mme, 315(2002), 51-54. 3 Chen Y.H. et al., "Migration of Fe, Zn, Cu in Slag of Sulfuric Acid and Latent Effect on Environment," Journal of Guangzhou University, 15(2) (2001), 78-81. 4 PENG Z.J., Wei G, "Exploitation and Comprehensive Utilization of Vitriolic Dregs and M etallurgical Dregs," Journal of Wuhan University of Science and TechnologyfNatural Science Edition), 25(2)(2002), 114-116. 5 Zheng Y.J. et al., "New Technology of Copperas Preparation from Pyrite Cinders," CHEMICAL ENGINEERING (CHINA), 33(4) (2005), 51-55. 6 C.J. Shi, C. Meyer and A. Behnood, "Utilization of Copper Slag in Cement and Concrete," Resources, Conservation and Recycling, 52(2008), 1115-1120. 7 Yang S.P., Li S.J. and Liu X.M., "Reutilization and Perspectives of the Metallurgical Waste Slag," Steelmaking, 24(3) (2008), 59-62. 8 Zhu D.Q. et al., "Preparation of High Quality Magnetite Concentrate from Pyrite Cinder by Composite Pellet Reduction-Roasting and Magnetic-Separation," The Chinese Journal of Nonferrous Metals, 17(4) (2007), 650 ~ 656. 9 Chen T.J., Zhang Y.M., "Experiments of Acid Pellet Using Treated Pyrite Slag and Its Commercial Application," Research on Iron and Steel, 142(2005), 1-4. 10 Nurcan Tugrul, Emek Moroydor Derun and Mehmet Piskin, "Utilization of Pyrite Ash Wastes

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by Pelletization Process," Powder Technology, 176 (2007), 72-76. 11 Li S.C, "High Temperature Reducing-Chlorination Process of Polymetallic Symbiotic Poor Tin by Rotary Kiln in Yunxi," Science and Technology ofYnnxi, 19 ( 1992 ) , 20-34. 12 Eduardo A. Brocchi, Francisco J. Moura, "Chlorination Methods Applied to Recover Refractory Metals from Tin Slags," Minerals Engineering, 21(2008), 150-156. 13 Metallurgical Laboratory of Central South College of Mining and Metallurgy, Chlorination Metallurgy (Beijing, Metallurgy Industry Press, 1978). 14 Qiu G.Z. et al., Direct reduction of cold-bonded pellets (Changsha, Centre South University Press, 2001), 176-181. 15 Fu J.Y., Jiang T, and Zhu D.Q., Principles of sintering and pelletizing (Changsha, Centre South University Press, 1995). 16 Peng R.Q. et al., Metallurgy of lead and zinc (Beijing, Science Technology Press, 2003).

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

EFFECT OF OXYGEN TO ALUMINA RATIO ON THE VISCOSITY OF ALUMINOSILICATE AND ALUMÍNATE SYSTEMS Jifang Xu', Jieyu Zhang1, Chang Jie1, Fei Rúan1, Kuochih Chou 12 1

Shanghai Key Laboratory of Modem Metallurgy and Material Processing, Shanghai University, Shanghai 200072, P. R. China; 2 Metallurgical and Ecological Engineering School, University of Science and Technology Beijing, Beijing 100083, P. R. China;

Keywords: viscosity, aluminosilicate and alumínate systems, oxygen to alumina ratio Abstract Viscosity is an important physical parameter of slag in metallurgical processes. Most studies of aluminosilicate and alumínate systems assume that aluminum occurs in tetrahedral coordination. In this case, the O/Al value was utilized to describe the slag structure in terms of the network character of aluminosilicate and alumínate systems. Composition and viscosity data from the literature and the trends were analyzed according to the O/Al value and temperature. It is shown that the O/Al value explains the trends in more than ten systems, containing more than 200 data points. In aluminosilicate system, the trend that viscosity value changed with the O/Al value which implies that alumina works as network modifier such as basic oxides or works as network former. In aluminate system, irrespective of the system and temperature, viscosity decreased steadily when O/Al value increased. Introduction Metallurgical slags, which are made of oxides and many also contain fluorides, play an important role in metallurgical processes such as iron making, steel making and secondary refining processes. Aluminosilicate slags are applied in a wide range in metallurgical processes. A polymerizing effect of the addition of aluminum and an association of aluminum with low-field strength cations is further supported by the physical variations in aluminosilicate and aluminate systems. The behavior of alumina in silicate melts has been studied by various authors [1-6]. It is generally accepted that AI2O3 is amphoteric [7] because Al 3+ ions can behave in profoundly different ways in a silicate melt. When added to a pure silica melt, AI2O3 acted as a network modifier, breaking the bridging oxygen of the pure silica network, thereby decreasing the viscosity in a similar manner to other network modifiers, such as CaO, MgO etc. to form [l/2Ca (AIO4)]4" complexes. However, when added with alkali or alkali-earth metal oxides, some of the Al3+ ions assumed tetrahedral coordination and replaced Si in the liquid network, the missing charge was compensated by Me+ or Me2+ ions. Therefore AI2O3 acted as network former [8, 9], As low viscosity is beneficial for mass transfer in the slag phase, viscosity is an important physical property that influences the slag performance. However, the measurement of viscosity at high temperature is very difficult and time-consuming. Several models have been developed for predicting viscosity from chemical compositions and temperature. The viscosity estimation of AbOs-containing silicate melts and aluminate melts is difficult to perform because viscosity is also sensitive to the structural changes of silicate slags and many structural units have to be considered. Generally, conventional basicity indices such as the lime to silica ratio provide an indication of the degree of polymerization, which is sensitive to the physical and chemical

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properties of slag [10]. Each MexO is considered to break a bond of a three dimensional network of tetrahedral units of Si0 4 4 " by supplying an additional oxygen and charge compensating the electron at the broken bond with the cations. At higher MexO contents, the network breaks down further to form rings and then to discrete units of silica compounds. The choice of basicity of the aluminosilicate and alumínate systems has its own limitations. When the content of AI2O3 is low, the conventional basicity can be characterized to determine the slag metallurgical, physical and chemical properties (including viscosity, desulphurization performance, etc.). When the AI2O3 content is high, the correlation between the conventional basicity and viscosity of slag is not obvious, and the conventional basicity is not suitable to evaluate the metallurgical and physical and chemical properties of the slag. The use of O/Si ratio for assessment of slags has been evaluated [11, 12] and the effect of the O/Al value on viscosity of the aluminosilicate and alumínate systems is discussed in this paper. Calculation Of The O/Al Value Molten slags are made up of tetrahedral and many other ionic species. The overall structure and behavior of slags is influenced by the relative proportions of the different ions. The O/Al value can account for the presence of different oxide species and hence will be a superior indication of slag composition than conventional basicity. 01 Al ratio = -i

(1)

Where, x¡ is the mole fraction of the /th oxide component, n¡ is the number of oxygen ions in the rth oxide, XA is the mole fraction of alumina, nA is the number of aluminum atoms in alumina. For pure Alumina, O/Al ration is equal to 1.5. The addition of alkali or alkaline earth oxides to alumínate systems increases the O/Al ratio to a value greater than 1.5 and breaks up the three-dimensional network with the formation of singly bonded oxygen, which does not participate in die network. The O/Al ratio gives an idea about the extent of depolymerization in the slag. The O/Al ratio is easy to calculate and reflects on the overall structure of slag. The outlines of the O/Al ratio are: (i) the addition of alkali or alkali-earth metal oxides to aluminosilicate and alumínate systems leads to increasing the O/Al value; (ii) increasing the O/Al value results in the progressive breakdown of the aluminosilicate and alumínate systems structure into smaller units (the smaller units will have relatively higher mobility and this will result in an decrease in viscosity of slag); (iii) the presence of different oxides can be accounted for in this approach and the main lacuna in with respect to CaF2, which could not be accommodated and for which alternate theoretical approaches are being explored, in the present work in progress, the F" is equal to O2"; (¡v) iron oxide has been considered as FeO only. In this paper, the O/Al value was utilized to describe the slag structure in terms of the network character of aluminosilicate and alumínate systems. Composition and viscosity were taken from the literature and the trends have been analyzed in terms of the O/Al value and temperature. The following systems, covering a range of temperatures, were analyzed: SÍO2-AI2O3, CaO-Al 2 0 3 , FeO-Al 2 0 3 , CaO-Si0 2 -Al 2 0 3 , MgO-Si0 2 -Al 2 0 3 , CaO-MgO-Si0 2 -Al 2 0 3 and mould fluxes. Many observations can be made from the present investigation to explain the trends effects of O/Al value on the aluminosilicate viscosity in different systems, covering more than 200 data points. Typically, the authors cited experimental procedures as well as significant trends; rigorous

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interpretation of the trends (with respect to structure) has not been provided in much of the literature. Chemical compositions were used to calculate the oxygen to alumina ratio values and then used for analyzing the trends in viscosity for a few slag systems, covering a wide range of temperature. The literature accounts of slag viscosity, in general, have been based on viscosity trends in specific systems/composition ranges. Results And Discussion Viscosity Of The Binary Aluminosilicate And Aluminate Systems Viscosity values of many slags in the Si02-Al203 [13], CaO-Al203 [14] and FeO-Al203 [15] binary system, with the O/Al value between 1.50 and 22.10, are shown in Figure 1. As expected, viscosity decreases with increasing temperature. As shown in Figure 1, the viscosity of the Si02-Al2C>3 binary system increased with increasing value of O/Al. This is due to that alumina works as network modifier such as a basic oxide, when it is added to a pure silica melt. The O/Al value decreased, breaking the bridging oxygen of the pure silica network, and thereby decreasing the viscosity in a similar manner to other network modifiers. In CaO-Al203 and FeO-Al203 binary system, if the system and the temperature are not considered, the viscosity was found to decrease steadily with increasing the O/Al value. The results indicted that in aluminate system, alumina works as network former, which is similar to Si02 in silicate system, and some of the Al3+ ions assume tetrahedral coordination. Higher concentrations of CaO and FeO decrease the O/Al value; the viscosity is reduced as the Al tetrahedral network is broken up.

Figure 1 Viscosity of the binary aluminosilicate and aluminate systems as a function of the O/Al ratio Viscosity Of The Multi-Component Aluminosilicate Systems The chemical composition of the Si02-CaO-Al203 system [14, 16] varies over a wide range for the three components, 10.0-65.0 mass% A1203, 10.0-60.0 mass% CaO, 10.0-70.0 mass% Si02, and the O/Al value varies from 2.11 to 14.43. The relationship between viscosity and the O/Al value at 1773K and 2073K are shown in Figure 2. In each case, the viscosity decreases gradually with increasing O/Al value, which implies that the network is being increasingly broken up. And the trend is identical for different author's data and different temperature. If Si02 content is not changed, then CaO will be added gradually. The result shows that it will increase the O/Al value, which breaks up the bridging oxygen of network, reduces the degree of polymerization, and decreases the viscosity. With increasing Si02 content, the viscosity increases, as expected. And the trend is identical for different level of the Si02 content. The overall trend confirms the suitability of O/Al ratio for evaluating viscosity in aluminosilicate melts.

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Figure 2 Viscosity of Si02-CaO-Al2C>3 slag system as a function of the O/Al ratio The chemical composition of the Si02-MgO-Al203 system [17] varies over a wide range for the three components, 10.0-25.0 mass% A1203, 10.0-30.0 mass% MgO, 60.0-65.0 mass% Si0 2 , and the O/Al value varies from 6.35 to 15.53. The relationship between viscosity and the O/Al value, with the temperature range from 1923K. to 1993K, is shown in Figure 3. The polynomial second order correlation coefficients of the two sub-sets of slags (Si0 2 at 60 mass% and 65 mass%) are similar. The overall trend confirms that the viscosity decreases gradually as the O/Al value increases. The tendency to form more extreme anionic units is ranked in terms of the parameter in the hierarchy Mg2+>A13+ and the cations with smaller radii and higher valence favor the formation of the more depolymerized and polymerized anionic units. Thereby, in the Si02-MgO-Al203 system, the gradual addition of MgO lead to the progressive breaking of these oxygen bonds with the formation of non-bridging oxygen (NBO), and the increases O/Al value results in the complex polymers of melts tetrahedral breaking down into smaller units, which leads to lower slag viscosity.

Figure 3 Viscosity of Si02-MgO-Al203 and Si02-MgO-CaO-Al203 systems as a function of the O/Al ratio The Si02-MgO-CaO-Al203 system [17-19] with the chemical composition varies over a wide range for the three components: 5.0-30.0 mass% AI2O3, 5.0-30.0 mass% MgO, 5.0-45.0 mass% CaO, 35.0-65.0 mass% Si0 2 , and the O/Al value varies from 4.61 to 30.46. The relationship between the viscosity and the O/Al value at 1773K was shown in Figure 3. In all cases, the viscosity gradually decreased with increasing the O/Al value. When Si0 2 content constant, and the network-breaking cations (e.g. Ca2+, Mg2+) is gradually increased, the O/Al value will

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increases. The higher O/Al value leads to the progressive breaking of these oxygen bonds and decreasing of viscosity. With increasing SÍO2 content, the viscosity increased as expected. The trend is identical for slags with different SÍO2 contents. The observations justify the choice of the O/Al as the chemical parameter for following the trends in viscosity. Viscosity Of The Multi-Component Aluminate Systems The viscosity of the CaO-Al203-MgO system [20] with the O/Al value in the range of 2.57 to 2.86 at 1845K was shown in Figure 4. In addition, the polynomial second order correlation coefficients may be formed. It is found that the viscosity decreases with increasing the O/Al value. The viscosity of CaO-AhOs-FeO slags system [21] with the O/Al value between 1.97 and 2.64 at 1673K, also as shown in Figure 4. The results indicated that the relationship between the logarithm of viscosity and O/Al value follows by a linear trend. The viscosity also decreases with increasing the O/Al value. This implies that, in these slag systems, AI2O3 does act as a network former. This interpretation is supported by the fact that the O/Al value increases steadily with increasing the content of CaO, MgO and FeO; and increasing the O/Al value imply that the network is being increasingly broken up. Therefore, the choice of O/Al ratio as a chemical parameter is logical and useful.

Figure 4 Viscosity of multi-component aluminate systems as a function of the O/Al ratio Viscosity Of The Mould Fluxes The viscosity of mould flux in the absence of CaF2 [22, 23] with the composition in the range of 51.10-62.00 mass% A1203, 18.23-36.40 mass% CaO, 2.50-12.00 mass% MgO, 3.70-6.00 mass% Na20, 2.78-5.50 mass% Cr203, 1.39-3.50 mass% Si02, and 1.30-3.50 mass% FeO, with the temperature between 1923K and 1993K, is shown in Figure 5. The O/Al value was 2.10 to 2.41. The viscosity decreases with increasing temperature, as expected. When the O/Al value increases, the viscosity decreases gradually. The trend is identical for different temperatures. The relationship between the viscosity of the Al203-CaO-MgO-Na20-Cr203-Si02-FeO slag system with the small amount of SÍO2 and the O/Al value was the same as that of the aluminate systems.

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Figure 5 Viscosity of mould flux without CaF2 as a function of the O/Al ratio In the CaO-Si02-Al203-MgO-Na20-Li20-MnO-CaF2 system [24], the Al 2 0 3 /Si0 2 molar ratio ranges from 0.06 to 2.14, in the composition range 5.0-40.0 mass% A1203, 11.0-46.0 mass% Si0 2 , and constant amounts of 10.0 mass% CaO, 12.0 mass% Na 2 0, 2.0 mass Li 2 0, 1.5 mass MgO, 1.5 mass% MnO, 22.0 mass% CaF2 was shown in Fig 6 with the O/Al value range from 2.60 to 22.22. The Al203-CaO-MgO-Na20-Cr203-Si02-FeO-CaF2 system [23] included twenty different slag compositions, with the composition range 50.00-62.00 mass% AI2O3, 16.57-28.64 mass% CaO, 1.85-2.65 mass% MgO, 3.70-6.00 mass% Na 2 0,4.10-5.50 mass% Cr 2 0 3 , 1.85-3.50 mass% Si0 2 , 1.30-3.50 mass% FeO and 2.00-10.00 mass% CaF2, was also shown in Fig 6. The O/Al value varies from 2.15 to 2.53. The general trend is the same in two cases, the viscosity decreases with increasing temperature as well as with increasing the O/Al value. With addition of F" ions, the O/Al value increases steadily. The F" ions play great role in the degree and type of polymerization of the slag structure, because it results in breakdown of large and low mobility aggregate complex alumínate anions to smaller sizes, the viscosity of the slag decreased. This observation is in agreement with the relationship between the O/Al value and the viscosity of mould flux with CaF2. This finding confirms that the suitability of the chemical parameter for studying the trends in viscosity.

Figure 6 Viscosity of mould flux with CaF2 as a function of the O/Al ratio Different systems, spanning a wide range of composition and temperature, have been considered in the present work. The correlation between viscosity and the O/Al ratio is consistent in different temperature ranges, slag compositions and slag systems. These observations clearly

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indicate that the correlation between the O/Al ratio and melt viscosity is valid, irrespective of the compositional constraints. This is so because the O/Al ratio represents the overall slag structure rather than just the slag composition. The slag structure in also influenced by temperature, which is reflected by the gradual reduction in viscosity with increasing temperature as clearly shown in Figure 6. Therefore, the O/Al ratio can be applied regardless of composition range and temperature. In this context, the O/Al ratio has good potential since it is not constrained by composition/system. Further efforts will be required for consideration of a larger database on viscosity and the application of the O/Al ratio to the proposed parameters of basicity (such as the lime to silica ratio) has already been demonstrated-in the context of crystallization of mold slags. There is a growing need to have a generic model for viscosity as function of chemical composition and temperature-irrespective of whether the slags are from iron making process or steel making process or continuous casting process. This is being examined in further detail, taking into consideration viscosity data for a wide range of slag compositions and temperatures Conclusions In order to improve the control of metallurgical process, the viscosities of aluminosilicate and alumínate systems were investigated. Data of compositions and viscosity were analyzed more than ten binary and multi-component aluminosilicate and alumínate systems from the literature. The trends were analyzed in terms of the O/Al value and temperature. The following observations were made: (1) The viscosity of the slag system decreased with increasing temperature, as expected. The trend that the viscosity of the slag system changed with the O/Al value is identical at different temperatures. (2) The effects of the O/Al value on the viscosity of aluminosilicate system indicated that alumina is amphoteric and can work as a network modifier or as a network former in aluminosilicate systems. In alumínate system, irrespective of the system and the temperature, the viscosity was found to decrease steadily with increasing the O/Al value, which is due to the increased breaking of the Al tetrahedral network. (3) The O/Al value has been successfully used to explain the trends in different systems, which contain more than 200 data points. The observations confirm that the choice of the O/Al value is logical and useful for studying the trends in viscosity and for describing the slag structure in terms of the network character of aluminosilicate and alumínate systems Acknowledgments The authors express their thanks to the NSFC under Contract No.50874072 for their kind financial support of this research. We also wish to acknowledge the support from the Program for Changjiang Scholars and Innovative Research Team in University (Grant No.IRT0739) and the Innovation Fund for Graduate Students of Shanghai University (SHUCX101054). References [1] E.T Turkdogan and P.M. Bills, "A Critical Review of Viscosity of CaO-MgO-Al203-Si02 Melts," Am. Ceram. Soc. Bull., 39(1960), 682-687 [2] E.F. Riebling, "Structure of Sodium Aluminosilicate Melts Containing at Least 50 mole % Si02 at 1500°Q" J. Chem. Phys., 44(1966), 2857-2865 [3] B. O. Mysen, D. Virgo and I. Kushiro, "The Structural Role of Aluminum in Silicate Melts-a Raman Spectroscopic Study at 1 Atmosphere," Amer. Mineral., 66(1981), 678-701

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[4] [5] [6]

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[21] [22] [23] [24]

J. B. Murdoch, J. F. Stebbins and I. S. E. Carmichael, "High-resolution 29Si NMR Study of Silicate and Aluminosilicate Glasses: the Effect of Network-modifying Cations" Amer. Mineral, 70(1985), 332-343 B. Mysen, "Aluminosilicate Melts: Structure, Composition and Temperature," Contrib Mineral Petrol, 127(1997), 104-118. M.J. Toplis, and D.B. Dingwell, "Shear viscosities of CaO-Al 2 0 3 -Si0 2 and MgO-Al 2 0 3 -Si0 2 liquids: Implications for the Structural role of Aluminium and the Degree of Polymerisation of Synthetic and Natural Aluminosilicate Melts," Geochim. Cosmochim. Acta, 68(2004), 5169-5188. T. Kou, K. Mizoguchi and Y. Suginohara, "The Effect of AI2O3 on the Viscosity of Silicate Melts" J. Jpn. Inst. Met., 41(1978), 775-781 A. Kondratiev and E. Jak, "A quasi-chemical viscosity model for fully liquid slags in the Al 2 0 3 -CaO-'FeO'-Si0 2 system," Metall. Mater. Trans. B, 36(2005), 623-638. Q.F. Shu, "A Viscosity Estimation Model for Molten Slags in Al 2 0 3 -CaO-MgO-Si0 2 System" Steel Research Int., 80(2009), 107-113 Frederick D. Richardson, Physical Chemistry of Melts in Metallurgy (Volume 1), (New York, NY:Academic press, 1974), 81 S.R. Sankaranarayanan: "Crystallization and related phenomena in continuous casting mold powders" (Ph.D. thesis, Drexel University, 1992), 80-82 K. Santhy, T. Sowmya, and S.R. Sankaranarayanan, "Effect of Oxygen to Silicon Ratio on the Viscosity of Metallurgical Slags," ISIJ Int., 45(2005), 1014-1018. G Urbain, Y. Bottinga and P. Richet, "Viscosity of Liquid Silica, Silicates and Alumino-silicates," Geochim Cosmochin Acta, 46(1982), 1061-1072 Paul Kozakevitch. "Viscosity of lime-Alumina-Silica Melts between 1600 and 2100°Q" Physical chemistry of Process Metallurgy, ed. George R. St. Pierre (New York, NY: Interscience Publishers, 1961,), 97-116 Verein Deutscher Eisenhuttenleute ed., Slag Atlas (2nd Edition) (Duesseldorf: Verlag Stahleisen, 1995), 356 J.S. Machin and T.B. Yee, "Viscosity Studies of System CaO-MgO-Al 2 0 3 -Si0 2 : U, CaO-Al 2 0 3 -Si0 2 ," J. Am. Ceram. Soc, 31(1948), 200-204 J.S. Machin and T.B. Yee, "Viscosity Studies of System CaO-MgO-Al203-Si02:IV, 60, and 65% Si0 2 ," J. Am. Ceram. Soc, 37(1954), 177-186 J.S. Machin and D.L. Hanna, "Viscosity Studies of System CaO-MgO-Al 2 0 3 -Si0 2 : I, 40% Si02,"J. Am. Ceram. Soc, 28(1945), 310-316 J.S. Machin, T.B. Yee and D.L. Hanna, "Viscosity Studies of System CaO-MgO-Al 2 0 3 -Si0 2 : m, 35,45, and 50% Si0 2 ," J. Am. Ceram. Soc, 35(1952), 322-325 A. Kondratiev, P.C. Hayes and E. Jak, "Development of a Quasi-chemical Viscosity Model for Fully Liquid Slags in the Al 2 0 3 -CaO-'FeO'-MgO-Si0 2 System. Part 2. A Review of the Experimental Data and the Model Predictions for the Al 2 0 3 -CaO-MgO, CaO-MgO-Si0 2 andAl 2 0 3 -MgO-Si0 2 Systems," ISIJ Int., 46(2006), 368-374. B. Vidacak, S.C. Du, and S. Seetharaman, "An Experimental Study of the Viscosities of Al 2 0 3 -CaO-'FeO' Slags," Metall Mater. Trans. B, 32(2001), 679-684. R.C. Behera, and U.K. Mohanty, "Effect of Oxide Fluxes on the Viscosity of Molten Aluminothermic Ferro-chrome Slags," ISIJ Int., 41(2001), 827-833. R.C. Behera, and U.K. Mohanty, "Viscosity of molten Al 2 0 3 -Cr 2 0 3 -CaO-CaF 2 slags at various Al 2 0 3 /CaO ratios," ISIJ Int., 41(2001), 834-843. Z. Zhang, G Wen, P. Tang, and S. Sridhar, "The Influence of Al 2 0 3 /Si0 2 Ratio on the Viscosity of Mold Fluxes," ISIJ Int., 48(2008). 739-746

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

BLAST FURNACE BURDENS PREPARATION FROM METALLURGICAL DUSTS AND SLUDGES WITH COMPOSITE BINDER Kecheng Zhang, Yuanbo Zhang *, Tao Jiang, Guanghui Li, Zhucheng Huang (School of Minerals Processing & Bioengineering, Central South University, Changsha, Hunan, 410083 , China) Key words: Metallurgical dusts & Sludges; comprehensive utilization; blast furnace burden; composite binder Abstract With rapid development of Iron and Steel industry, plenty of metallurgical dusts and sludges are generated, which have not been effectively utilized because of their poor hydrophilicity and bad ballability. Agglomeration and roasting of metallurgical dusts & sludges with a new-type composite binder are studied in this paper. The results indicate that the dosage of the new binders can be decreased to 2.5 wt% from 3.5 wt% when using the wet-grinding process. Under the optimal conditions of briquetting pressure 2070 N/cm2, briquetting moisture content 13 wt%, roasting temperature 1000 °C and roasting time 40 min, the qualities of the green, dry and roasted briquettes meet the requirements of the industry production. The finished products have the compression strength of more than 2000 N/cm2, which can be used as high-quality burdens for blast furnaces. Introduction With rapid development of Chinese Iron & Steel industry, there are more and more serious resource shortage and environmental pollution existing. Converter sludges, containing 30% ~ 40 wt% moistures, are typically difficult for recycling [1]. Foreign processing technologies for sludges treatment are relatively advanced including cold-bonded agglomerates for blast furnace, RHF reduction for the dusts and sludges with high content of Zn and Pb, smelting and reduction. Domestic iron and steel companies use a small amount of sludges in sintering. However, the sludges have the characteristics of high moisture and containing harmful impurities, which directly affect the sintering mixture permeability and product quality [2-5]. There are also other technologies such as production of metallized pellets, cooling agents for steelmaking, etc [6]. All of these deficiencies and applied are not in large-scale. In order to find out a rational way to utilize the converter sludges effectively, a new type of composite binder is used for agglomerates preparation for the blast furnace burdens. In this paper, * Corresponding author: Dr. Yuanbo Zhang, E-mail: [email protected].

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the optimal technological parameters are obtained, which can provide theoretical guidance for the industrial application of converter sludges. Experimental Materials The iron-bearing materials used in this study are converter sludge and fine ores, the densities of which are 3.668 g/cm3 and 4.317 g/ cm3, respectively. The chemical composition is listed in Table 1. Particle size distribution is given in Table 2 and Table 3. Table 1. Chemical compositions of sludges and fine ores / wt% Species TFe P FeO MFe Si0 2 A1203 CaO MgO MnO S C Sludge 60.34 19.07 1.64 2.58 0.80 0.63 0.021 0.30 0.18 7.05 ores 46.15 37.73 2.53 2.25 0.29 14.21 6.91 0.67 0.072 0.17 2.89

Fractions (mm) Content / wt%

Fractions (meshes) Content/ wt%

Table 2. Particle size distribution of fine ores / wt% +12 -5+3.2 -3.2+0.9 -0.9+0.4 -12+8 -8+5 5.94 18.52 0.38 2.98 1.90 17.38 Table 3. Particle size distribution of the sludges / wt% +100 -100+140 -140+200 -200+320 1.41 5.16 2.75 2.70

LOI 2.62 12.98

-0.4 52.98

-320 87.98

As shown in Tables 2 and 3, the granularity of the fine ores is coarse, with nearly 30 wt% larger than 1mm. The sludge granularity is smaller, and about 87.98 % sludges pass through 320 meshes. The composite binder used was developed by Central South University, which is composed of modified humic substances and inorganic binders. The FT-IR of modified humic substances is shown in Fig.l. The loss of ignition of composite binder is about 50 wt%, including organic matters and inorganic substances.

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Figure 1. FT-IR spectrum of composite binder The absorption at 3398 cm"1 is attributed to stretching of-OH groups (fig.l). Within the wide frequency range from about 1550 to 1790 cm-1, principally assigned to protonate carboxylic (-COOH), carboxylate anión (-COO-) and ester carbonyl groups (-COOR). It is noteworthy that there are the protonated -COOH groups or the ester carbonyl groups (-COOR) in the composite binder. Therefore, CB can improve the hydrophilicity of the sludges and improve the briquette strength. However, CB is difficult to disperse evenly in the mixtures, especially in the sludges containing high moisture, and wet-grinding process is used to pretreat the sludges. Method Experimental process included mixing, wet-grinding, briquetting, drying, reduction roasting and cooling. Wet-grinding of mixture was used with O500* 500 mm Mill Run. In this research, the grinding parameters are set as the follows: moisture was 9 % and grinding time was 3 min. Briquetting of mixture was in the universal testing machine. Roasting was carried out in a high temperature baking furnace, and roasted briquettes were cooled to room temperature naturally. Compression strength testing was in accordance with ISO 4700. Drop strength of briquette refers to the drop timesfromthe height of 0.5 meters before they are broken. Results and Discussion Effects of Raw Material Proportionine Schemes In this study, proportioning schemes affecting qualities of briquettes were researched first. The alkalinity and iron grade of briquettes from different proportioning schemes are shown in Table 4. Table 4. Alkalinity and iron grade of briquettes from different schemes

Ratio of sludge to fine ore CaO/Si0 2 (CaO+MgO)/(Si02+Al203) Iron Grade of Roasted Briquettes

7:3 2.91 3.60 56.05

5:5 1.81 2.18 57.88

3.5:6.5 1.24 1.47 59.21

2:8 0.81 0.87 60.49

As shown in Table 6, with the decline of ratio of sludge to fine ore from 7:3 to 2:8, CaO/Si02 is decreasedfrom2.91 to 0.81, while (CaO+MgO)/(Si02+Al203) decreasefrom3.60 to 0.87. As far as iron grade of roasted briquette is concerned, it is increased from 56.05 wt% to 60.49 wt%. Experimental conditions were set as follows: composite binder dosage of 3.5 wt%, briquette pressure of 3500 N/cm2, roasting time of 40min and roasting temperature of 1050 °C. The final results are given in Table 5. It can be seen from Table 5 that, with the increase of the sludge proportion, drop strength of wet and dry briquettes increase gradually. The sludge has very fine granularity as well as strong

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adhesion. The more sludges are added, the stronger adhesion in the mixture is [6], and the strength of the briquettes is enhanced. However, the increase of the sludge proportion is unfavorable for the compression strength of roasted briquettes. During the roasting, the organic matter is burnt out so there are many pores left in the roasted products, which reduce the strength of the products. In order to utilize the sludges, the ratio of sludges to fine ores of 7:3 was selected to study. Table 5. Effects of proportioning schemes on the briquette strength Compression Compression Ratio of Drop strength of Compression strength of wet strength o froated sludges to iron wet briquette strength of dry briquette bruquette ores (times /0.5m) briquette (N/cm2) (times/lm) (N/cm2) 1786 2:8 5.8 2.6 359 1426 3.5:6.5 20.4 3.6 558 1388 5:5 31.4 5.6 425 1157 7:3 43.0 7.8 370 Effects of Briquetting Parameters Briquetting Pressure. Experiments of briquetting pressure were carried out under the conditions of composite binder dosage 3.5 wt%, briquetting moisture 13 wt%, roasting temperature 1000 °C and roasting time 40 min. The findings are listed in Table 6. Table 6. Effects of briquetting pressure on the quality of briquettes Drop Drop Compressive Compressive Compressive strength of strength of Briquetting strength of strength of wet strength of dry pressure wet roasted dry briquette briquette (N/cm2) briquette briquette briquette (N/cm 2 ) (N/cm2) (times/0.5m) (times /lm) (N/cm 2 ) 1592 7.0 127 1.0 236 1430 2070 8.2 194 3.2 389 1701 2548 16.0 219 7.6 546 1719 3025 28.6 260 9.0 670 1776 3500 46.8 267 9.4 803 1736 As seen from Table 6, with increasing briquetting pressure, strength of wet and dry briquettes increased gradually, drop strength of wet briquettes and compressive strength of dry briquettes also increased. Considering comprehensively, the optimal briquetting pressure should be not less than 2070 N/cm2. When the briquetting pressure is 2070 N/cm2, the drop strength of wet briquettes, compressive strength of wet briquettes, drop strength of dry briquettes, compressive strength of dry briquettes, compressive strength of roasted briquettes were 8.2 times/0.5 m, 194 N/cm 2 ,3.2 times /lm, 389 N/cm2, 1701 N/cm2, respectively.

172

Briquetting Moisture. Effects of briquetting moisture was studied when composite binder dosage was 3.5 wt%, briquetting pressure was 3500 N/cm2, roasting temperature was 1050 °C and roasting time was 40min. The experimental results are shown in Fig. 2. 15« 153D

1380

Briquette Moisture (%)

Figure 2. Effects of briquetting moisture on wet briquette strength The strength of wet briquettes was determined by the interaction between the iron ore particles and binder molecules. With increasing briquetting moisture, the strength of wet briquettes increased atfirstsignificantly and then decreased sharply (fig.2). When briquetting moisture was 13%, the strength of wet briquettes was at highest. Higher moisture content makes the wet briquettes easily to crack during the drying, which would affect the strength of dry briquettes. Effects of Roasting Parameters. The main purpose of roasting is to oxidize the ferrous oxide and promote liquid-solid or solid-solid reactions, which is beneficial to the strength of thefinishedproducts [7], Fig. 3 shows the effects of roasting temperature on the strength of roasted products. Compression strength of the roasted products is increased with the roasting temperature rising. The main reason is that higher temperature accelerates the interactions and promotes the crystallization of iron oxides, so the products with high strength are obtained.

173

940

960 960 1000 1020 WO 1060

Roasting TemperEÈjre f °C)

Figure 3. Effect of roasting temperature on product strength

Ftoasting Time (min)

Figure 4. Effect of roasting time on product strength The effects of roasting time on the product strength are shown in Fig. 4. When the roasting time increases from 30 min to 40 min, the product strength increased sharply from 1460 N/cm2 to 1687 N/cm2. However, if the time prolonged, the strength gradually decreased. During the tests, we found there were much liquid phases appearing in the products when the roasting time was over 50 min. Comprehensively comparing, the optimal roasting parameters should be at roasting temperature of 1000 CC and the roasting time of 40 min. Effects of Composite Binder Dosage According to the experience, composite binder dosage has obvious influence on the quality of

174

the briquettes. Effects of composite binder dosage on the strength of briquettes were also studied. The experimental included, the moisture of mixture of 13 %, briquette pressure 2070 N/cm2, roasting temperature 1000 °C, and roasting time 40 min. The results are presented in Table 7.

Composite binder dosage (wt%) 2.0 2.5 3.0 3.5

Table 7. Effects of the composite binder dosage on the quality of briquettes Compressive Drop Drop Compressive Compressive strength of strength of strength of strength of wet strength of dry wet roasted dry briquette briquette briquette briquette briquette 2 2 (N/cm ) (N/cm ) (times/0.5m) (N/cm2) (times /lm) 1965 2.2 2.4 350 170 2050 3.0 410 5.5 180 1920 7.0 3.2 395 190 1701 3.2 389 8.2 194

From the data in Table 7, it can be seen that if the composite binder dosage was 2.5 wt%, the finished products had the compression strength of 2050 N/cm2, which can be high quality burdens for blast furnaces. The suitable composite binder dosage was much smaller than the inorganic bentonite reported in literatures (generally more than 5%) [3,4]. Moreover, after the composite binder burnt, about 50 wt% of CB is residual in the finished briquettes, which can obviously enhance the iron grade of products compared with the usual inorganic bentonite. Conclusions Agglomerates composed of metallurgical sludges with a new-type composite binder prepared for blast furnace in this research. First, the converter sludges and fine ores were subjected to wet-grinding pretreatment. Under the optimal conditions of briquetting pressure of 2070 N/cm2, briquetting moisture content of 13%, roasting temperature 1000°C and roasting time of 40 min, the qualities of briquettes met the requirements of the industry standards. The finished products had the compression strength of more than 2000 N/cm2, which can be used as high-quality burdens for blast furnaces. Composite binders were effective for the agglomeration of the converter sludges with addition of fine ores. Moreover, the composite binder enhances the iron grade of products compared with the usual inorganic bentonite. Acknowledgements The authors want to express their thanks to National Science Fund for Distinguished Young Scholars (No.50725416), National Natural Science Foundation of China (No.50804059) and National Key Program of Science and Technology (No.2008BAB32B06) for financial supporting of this investigation.

175

References [1] H.M.Zhu, "Technology on Utilization of OG Sludge at MaSteel", Research on Iron & Steel,35 (2006), 51-53. [2] H.Q.Liao, et al, "Technology and Its Clew for High Efficiency Cyclic Utilization of Dust Mud with Iron Content in Iron and Steel Processing," Metallurgical Environmental Protection, 6(2007), 286-289. [3] P.S.Yang, R.Huang, X.W.Qing, "Problem and Countermeasures on Reutilization of OG Sludge in Sintering Plant of LIUZHOU Iron & Steel Group Company" Henan Metallurgy,n(2009), 26-29. [4] Z.B.Shen, Y.Z.Sha, "Study on Cold Bonder Briquette Technology of Dust from Steel Industry", iron and steel, 38 (2003), 1-5. [5] GZ.Qiu, et al. Derect reduction of Cold-bonded pelelts. (Changsha: Central South University Press,2001),4. [6] Y.B.Zhang, et al, "Study on Agglomeration and Reduction of Metallurgical Dusts and Sludge," Journal of Iron and Steel Research International, 16(2009), 475-479. [7] J.Y.Fu, T.Jiang, D.Q.Zhu, Theory of Stintering and Pelletizing. (Changsha: Press of Central South University of Technology, 1996), 2.

176

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

Determination of FeO containing liquid slag surface tensions using the sessile drop method Clemens Schmetterer and Patrick J. Masset Freiberg University of Mining and Technology Center for Innovation Competence VIRTUHCON Fuchsmühlenweg 9, Reiche Zeche, D-09596 Freiberg, Germany Keywords: surface tension, slag, sessile drop technique Abstract The surface tension represents an important property for the modelling of high temperature processes. In this study the surface tension of two AI2O3 - CaO - SÍO2 model slags and the influence of the admixture of 5 up to 30 wt.% FeO were studied. The surface tension measurements were carried out using the sessile drop technique on BN substrates under Ar atmosphere (10 6 atm 0 2 ) up to 1500 °C. Introduction The aim of VIRTUHCON1 is the virtualization of high temperature conversion processes, a task that requires thorough understanding of these processes. In many of these conversion processes slag plays an important role to control the properties of the product, to protect the reactor lining or simply as waste. Slag surface tension is one of the materials properties that is needed for the description of such processes, since it influences the interactions between the components in the reactor, e.g. at the slag - Fe interface in the blast furnace. Surface and interfacial tensions have thus received considerable attention in research, but the number of investigations in this particular pseudo-quaternary system is surprisingly low. Furthermore, the literature data are subject to different measurement conditions and techniques which complicates the data analysis and comparison. The slag atlas [1] offers a compilation of the available literature until 1986 (shown in Table I), but literature search has not revealed more recent studies of this system until now. For example, Elliott and Mounier [2] determined the surface tension of ACSF slags using the double tube maximum bubble pressure technique at 1200 °C. They kept their slags in contact with Cu-S-F mattes, and the slag therefore contained a certain amount of sulphur which acts as a surfactant and lowers the surface tension. The compositions of their slags were scattered over a wide composition range and they did not use a common ACS base mixture for their tests, either. The surface tension values vary between 417 and 446 mN/m, but due to the reasons just mentioned no trend in the measurements can be observed and the influence of FeO admixture on the surface tension cannot be deduced. Further information including the temperature dependence is available from Pavlov et al. [3], Ishchanov and Shushkov [4] and Murav'eva and Kaplun [5]; the diagram on page 446 in Ref. [1] has been drawn from these sources. Both positive and negative corresponding author: Dr. Patrick J. Masset [email protected] for more information about this project see: www.virtuhcon.de

177

temperature coefficients are reported in these works. SÍO2 is the only pure constituent that has a positive temperature coefficient [6], but a correlation of the temperature coefficient with the SÍO2 content was not found to be conclusive. Table I: Surface Tension ofpseudo quaternary AhCh-CaO-SiOa-FeO slags A1203

Compos ition (wt.%) CaO Si0 2 FexO

Technique

Conditions

Refs.

double tube maximum bubble pressure method

1473 K in contact with Cu-S-F mattes

PI

unknown

1473 K

[5]

unknown

1473 K

[3]

unknown

1473 K

[4]

0

(mN/m)

8.58

2.88

33.15

56.66

426

9.96

3.92

32.3

50.28

416

13.81

7.93

34.9

42.57

442

9.88

7.99

32.94

46.04

417

35.7

42.05

441 419

10.63

11.47

10

3

34

53

12.09

12.26

34.87

41.02

421

4.6

18.3

31.7

35.7 | 9.8 *>

421

9.1

17.4

30.8

33.919.3

424

5.69

5.08

50.68

38.55

340

5.87

22.24

49.77

22.11

370

5.57

28.84

51.21

14.19

377

5.49

37.09

51.92

5.49

397

11.15

14.18

44.5

29.26 | 0.24

350

11.20

19.35

44.75

24.63 | 0.32

368

*) indicates amount of FeO and Fe2Ch

The measurement temperature of 1200 °C seems to be low, but according to the phase diagram [1] the liquidus temperatures of the quaternary mixtures can be even below. Furthermore, silicate based slags tend to form glasses which soften below the melting range of the crystalline phases. Softening of the samples between 900 and 1000 °C was observed in the present work, too, but the surface tensions from this range have been omitted from this work. From the various literature information it is clear that FeO - containing slags are complex substances. They impose additional experimental problems and the experimental conditions have to be controlled precisely. In the light of the available literature (see below), the need for a systematic investigation in the pseudo quaternary system Al203-CaO-Si02-FeO with respect to a number of parameters was recognized: • FeO content • Atmosphere / O2 partial pressure • Temperature Therefore, in this work the surface tension determination of model slags of AI2O3 - CaO - SÍO2 mixtures (further on called ACS) with varying FeO content (denoted ACSF) is described with the emphasis on the experimental procedure. In addition, the first results so far obtained will briefly be discussed.

178

Experimental Section Experimental Procedure Model slag mixtures were prepared from the high purity substances AI2O3 (99.995%, Alfa Aesar), CaO (99.99%, Sigma Aldrich), Si0 2 (99.995+%, Sigma Aldrich) and FeO (99.9%, Sigma Aldrich) dried at 900 °C before use. First, ternary AI2O3 - CaO - SÍO2 base mixtures were prepared, powdered and portioned for the addition of FeO. They were then remelted in a stream of Ar / 2% O2. All melting procedures were carried out in BN crucibles in a tubular furnace (Ar atmosphere, which corresponds to p(Oj) = 10"6 atm). All BN parts were cleaned by immersing them in ethanol for a couple of days and heated in vacuum at 900°C. During sample preparation it turned out that FeO may be reduced to metallic Fe in Ar atmosphere (see also Experimental Considerations). This was indeed noticed in the present work with a number of samples (compositions in wt.%) so that finally only the following samples were used for analysis: ACS1 ACS2 ACSF1 ACSF2

A13C48S39 A19.5C38.1S42.4

ACSl + 5%FeO ACS2 + 30 % FeO

Surface tension measurements were carried out according to the sessile drop technique (see below). Before use, the BN substrates were ground with SiC paper of 240, 1000 and 4000 mesh in order to obtain a smooth surface. The Sessile Drop Method The sessile drop method for surface tension determination can basically be performed easily and does not require a sophisticated experimental setup. In this technique, the sample rests on an inert substrate in a suitable furnace and is lit from the back. The shadow is recorded and for evaluation the images are analysed. In the present work, a TOM-AC optical dilatometer (Fraunhofer ISC, Würzburg, Germany) was used for measurements of the surface tension under Ar only, because the furnace is equipped with a graphite heater. The evaluation of the surface tension from a sessile drop is in the first place image analysis. First the baseline is determined and the recorded drop shape (the shadow of the drop) is analyzed. With the use of a size standard the drop area is calculated. The volume is obtained under the assumption of axial symmetry and together with the sample mass allows the calculation of the density. The density can therefore be determined individually for each temperature, but the assumption of axial symmetry may not always be valid; for the measurements presented in this study it was found to be well fulfilled. The shape of a drop is described by the Laplace equation (SI units are used throughout): &p=0\— + — Ap...pressure difference at the surface

equ.l

a...surface tension

179

r„...principal radii

This equation can be rewritten as: =<7 — + 1—

Apgz+Ap0

-

)

equ.2

A/9... density difference g... gravitation z... vertical coordinate in drop description Apo... pressure difference at a certain datum plane Apo can also be substituted resulting in the following expression: Apgz+—

2(7

=a

1 1 —+ —

equ.3

r 0 ... radius at drop vertex The principal radii of the drop contour are then expressed by geometrical relations, e.g.: X

A

equ.4

r2 =

sin0 x... coordinate of tangent axis to drop vertex 0...turning angle between tangent to the interface and the datum plane The coordinates are then themselves expressed as functions of the arc length s measured from the origin, e.g.: dx = COS0 equ.5 Y ds — = sin è equ.6 ds Substitution and rearranging leads to: dé 2 Ape sin0 -2- = — + -C—2-2+ ?ds r0 a x

equ.7

The set of differential equations 5, 6 and 7 needs to be solved in conjunction with a least squares fit to the experimental drop contour. A full description of this method was given by Rotenberg et al. [7], whereas in the present work this evaluation was done using a computer program called "Drop" developed at TU Chemnitz, Germany [8]. Experimental Considerations Though simple in its basic outline, there are a number of important points to consider for a sessile drop experiment. The measurement of liquid slag surface tensions imposes further experimental complications that shall be outlined in the following. Substrate selection is a crucial point because any reaction between sample and substrate has to be avoided. Furthermore, the sample must not wet the substrate, because wetting angles higher than 90° are favourable for a reliable evaluation. In the present work, boron nitride (BN) was chosen, because it is not wetted by the slag. A drawback is the restriction of use to vacuum or protective gas above 1000 °C.

180

A further issue is encountered during the measurement of FexO - containing slags. The thermal stability of iron oxides is determined by the partial pressure of oxygen (p(C>2)) in the surrounding atmosphere with respect to the temperature. Whenever p(C>2) is lower than the oxygen pressure caused by the decomposition of FeO, the sample loses oxygen and therefore changes its composition; even (metallic) iron may be formed under these conditions. This situation causes two undesired effects: first, the resulting change in composition does not allow for precise measurements and, second, the formation of pure Fe introduces another liquid liquid or liquid - solid interface into the system. This loss of oxygen was noticed in the present work, too, where the sample appeared to "boil"; the "boiling" temperature was found to decrease with increasing FeO content and above this temperature the surface tension could only be estimated. The third important point concerns image recording for data evaluation. Images used for profile analysis should not be blurred and the camera needs to be focussed on the sample. This setup must not be changed between recording of the size standard and the actual measurement. High image quality is furthermore of particular importance at the intersection of baseline and drop contour, because at this point the calculated drop contour needs to be extrapolated. In blurred images there is a transition zone between drop and background via various grey shades; at the intersection of drop and substrate the grey shades overlap preventing a reliable fit of the calculated drop shape in this critical area. Results and Discussion Indeed the experimental difficulties described above were encountered during the work but have been largely solved so that preliminary results are available. These are compiled in Table II and are shown in Fig. 1 in graphical form. In the text, when unspecified, the units for the surface tension and temperature are mN/m and °C, respectively. : Expenmental surace tens,ion of ACS andA(^SF slags as a function of temt ACS1 ACS2 ACSF1 ACSF2 T(°C) a (mN/m) T(°C) a (mN/m) T(°C) a (mN/m) T(°C) a (mN/m) 1302 1324 1352 1376 1400 1425 1453 1474 1500

441 442 438 416 422 420 410 375 385

1274 1300 1324 1352 1377 1401 1426 1450 1475 1500

592 560 551 580 577 430 375 404 380 390

1232 1249 1266 1283 1300 1318 1350 1374 1402 1427 1475 1500

181

555 550 540 525 505 500 480 452 447 435 420 410

1426 1474 1500

400 360 350

For all investigated slags a linear decrease with rising temperature, i.e. a negative temperature coefficient, was obtained, but for the individual mixtures the values differ significantly, as can be seen from the relations2: ACS1: o- = 863.21-0.3189-r ACS2: cr = 2014.5-l. 1 0 2 8 ? ACSF1: a = 1276.5-0.5875 T The values obtained for the ACS slags are between 441 and 385 mN/m and 592 and 390 mN/m, respectively. These values are in agreement with surface tension values reported for the ternary ACS system [1], but for mixture ACS2 there are large deviations from the fitted curve; e.g. the value at 1426 °C is quite below the fitted line. The effective low p(C>2) of 10"6 atm (nominal Ar atmosphere) did not influence the stability of these slags in the observed temperature interval. For ACSF slags, the oxygen level of 10"6 arm was low enough to result in the reduction of FeO to Fe as already described in Experimental Considerations; this reaction was found to start at rather low temperatures depending on the FeO content. While below this temperature the evaluation was straight forward, surface tension values obtained from higher temperatures were estimated from suitable images of the measurement; they were generally found to be consistent with the lower temperature values. Sample ACSF1 appeared to slightly boil (minimal loss of oxygen due to the reduction of FeO) already at 1300 °C. The surface tension of this sample at low temperatures is -50 mN/m lower than that of the corresponding sample ACS2 at comparable temperatures, but due to the smaller temperature coefficient value this is reversed at higher temperature. However, taking into account the many experimental uncertainties at T>1300°C, the surface tensions of these two samples can be regarded as similar at 1500°C.

Fig. 1: Temperature dependence of surface tension Sample ACSF2 was found to start to boil already below 1200 °C, i.e. around its liquidus temperature. Therefore the surface tension could only be estimated for three temperatures and no For ACSF2 no trend was evaluated due to the low number of datapoints

182

expression for the temperature dependence was established. The obtained values are at or below 400 mN and represent the lowest values found in this study. The surface tension of sample ACS2 appears to be decreased by the admixture of FeO (samples ACSF1 and 2), but the surface tension of sample ACS1 lies between the values of the ACSF samples (see Fig. 1). More measurements are planned in order to further clarify this behaviour. A comparison with literature values is difficult due to the many different compositions used in these studies. In general, the literature values range from 320 - 450 mN/m in the whole temperature range from 1100 to 1500 °C [1]. The values determined in the present study are about 150 mN/m higher for ACSF1 compared with the 5.49 wt.% FeO mixture from Pavlov et al. [3]; however, the alumina content is 10 wt.% higher in the present mixture and vice versa for silica. According to Pavlov et al. [3] the surface tension has a positive temperature coefficient for high FeO (-38 wt.%) content, and a negative one for low FeO contents (< 25 wt.%). On the other hand, Murav'eva and Kaplun [5] report decreasing surface tensions for even higher FexO concentrations; however, in their mixtures the silica content is lower. In the present study only negative temperature coefficients for FeO contents up to 30 wt.% were found. Conclusion and Outlook The surface tension of two ACS model slags and two FeO-containing slags was determined under inert gas (Ar, p(O2)~10"6atm) as a function of temperature in this work. For all investigated mixtures the surface tension decreases with rising temperature and the admixture of FeO generally lowers the surface tension. This result, however, is tentative and further work is required for a stronger experimental basis. Future work will also comprise the measurement of the surface tension under air and at various p(Oa). Experiences made with the application of the sessile drop technique to this type of sample showed that there is no universal substrate material, and for the measurements under oxidizing conditions BN will have to be replaced by Pt-plates. Acknowledgement The financial support of this work by the German Federal Ministry of Education and Research is gratefully acknowledged. The authors also would like to thank S. Starke, M. Kurkova and M. Schreiner, Freiberg University of Mining and Technology, for their help with experimental issues. References 1. 2. 3. 4. 5. 6. 7. 8.

M. Allibert, H. Gaye, and et. al., Slag Atlas. 2nd ed. (Dusseldorf: Verlag Stahleisen GesmbH, 1995/2008). J.F. Elliott and M. Mounier, Canadian Metallurgical Quarterly, 21(4) (1982), 415. A.V. Pavlov et al., ¡Completen. Ispol'z. Miner. Syr'ya, 2 (1980), 39. T.K. Ishchanov and K.V. Sushkov, Met. Metall., 3 (1974), 81. EL. Murav'eva and L.I. Kaplun, Adgez. Raps. Paika Mater., 12 (1984), 26. W.D. Kingery, Journal of the American Ceramic Society, 42(1) (1959), 6. Y.L. Rotenbergn, L. Boruvka and A.W. Neumann, Journal of Colloid and Interface Science, 93(1) (1983), 169. M. Köhler, Drop-Programm (Computer Software), Diploma Thesis, TU Chemnitz, Chemnitz, Germany.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiarm-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooks TMS (The Minerals, Metals & Materials Society), 2011

PREPARATION OF PARTIALLY STABILIZED ZIRCONIA AND INTERFACE STRUCTURE ANALYSIS Dongbo Li1,2,Shenghui Guo1'2, Lijun Liu1'2, Jinhui Peng u ,Libo Zhang12, Chengdong He1,2 'Faculty of Metallurgical & Energy Engineering, Kunming University of Science & Technology, No.253 XUEFU Road, Kunming 650093, PR China 2

Key Laboratory of Unconventional Metallurgy, Ministry of Education, Kunming University of Science and Technology, Kunming 650093, PR China Keywords: partially stabilized zirconia(PSZ); interface structure; cubic ZrC^; monoclinic Zr0 2 Abstract In present study, partially stabilized zirconia (PSZ) was prepared by CaO doped fused zirconia heated by microwave furnace at the temperature of 1450 °C and holding time of 120 min. The XRD results showed that untreated fused zirconia mainly consists of crystalline compounds of cubic Zr0 2 phase; while the roasted one mainly is composed of crystalline compounds of cubic Zr0 2 phase and monoclinic Zr0 2 phase. It is shown through the optimization of structures of two phases that structure of cubic Zr0 2 and monoclinic Zr0 2 reject each other, revealing its difficulty in forming coinciding interface and linkage of two phases being of Ca atom. Introduction Due to advantages of zirconia, such as high melting point, chemistry inertia, low heat transmission and so on. Zirconia is always taken as megathermal refractory material[l]. However, in the heating and cooling processes of pure zirconium oxide, martensitic transformation will take place associated with greater shear force and volume change, which severely limits the industrial application of pure zirconia in some areas. Such as: Refractory materials, grinding materials and high temperature insulation materials[2]. Through adding a certain amount of MgO, CaO, or Y 2 0 3 stabilizer, etc.[3,4], the zirconia can change into fully stabilized zirconia, which exists in cubic phase. But this high purity fully stabilized zirconia cracks easily because of thermal stress. Therefore, the thermal 1

Corresponding author: Jinhui Peng Tel.: +86 871 5192076; fax: +86 871 5191046. E-mail address: [email protected].

185

shock resistance variation is one of the most important quality indicators of zirconia-based materials. According to the heat treatment, the fully stabilized zirconia can be changed into the partially stabilized zirconia (PSZ)[5], which exists in cubic phase and monoclinic phase and has superior mechanical properties, such as anti-heat, low thermal conductivity, large coefficient of thermal expansion and stability, as well as good wear resistance[6]. In recent years, as a new method for green metallurgy and the preparation of materials, microwave heating has been developed as an outstanding field of new interdiscipline[7], which has a lot of incomparable advantages, such as selective heating, higher temperature rising rate, higher heating efficiency, lower reaction temperature, shorter reaction time and easier management, etc.. As the bridge of connecting the two kinds of materials with different chemical and physical properties, interface plays an important and decisive role in composite materials, which have transfer, blocking, absorption, scattering and other effects. But due to the thickness of interface is usually in the range of dozens of atoms[8,9], it is not easy to characterize its structure through common analytical techniques. So the method of computer relaxation optimization is employed to analyze the interface structure of PSZ in this paper. In the present work, the PSZ was prepared from CaO doped fused zirconia by microwave heating. Using the first principles, the interface of Ca-doping cubic partially stabilized zirconia(c-Zr02) and monoclinic partially stabilized zirconia (m-Zr02) have been studied, which provides the theoretical basis for the development of high quality PSZ. Experimental Materials, In the present study, fused zirconia was obtained from Yingkou City, Liaoning Province, China. Fused zirconia containing more than 92% Zr0 2 was prepared by electric melting method. The chemical composition of fused zirconia is listed in Table 1.

Material Mass M/%

Table 1 Chemical composition of starting materials CaO A1203 Si0 2 Ti0 2 Zr0 2 92.0

7.1

0.3

0.2

0.2

M/% Fe 2 0 3

others

0.1

0.1

Methods And Characterization Firstly, 500 grams of Ca-doping fully stabilized zirconia was broken into a certain size. And then, appropriate raw materials were placed inside a quartz holder. It has been proved that the best thermal shock resistance stability of materials occurs when the monoclinic zirconia (m-Zr02) content is about 30%. And the sample holder was placed inside the microwave furnace and heated to 1450 °C at a average heating rate of 400 °C

186

/min and held at this temperature for 120min. The temperature of the raw materials during roasting process was monitored using a thermocouple inserted into the raw materials. After finishing the roasting process, the roasted raw materials were naturally cooled to room temperature in the microwave furnace. The XRD for the fused zirconia before roasting is shown in Fig.l. It can be seen that the cubic Zr0 2 is the main crystalline phase in the fused zirconia. PSZ prepared after roasting at 1450 °C for 120 min was also checked by XRD, Fig.2.

Fig 1. XRD pattern of fused zirconia

Fig 2. XRD pattern of PSZ

It can be found that the diffraction peaks of cubic Zr0 2 phase are broadened and their intensities are decreased after roasting process. From Fig.2, the PSZ has peaks at 29=28.2°and 31.5°, which are the strongest and the second strongest peaks of monoclinic Zr0 2 , while the third strongest peak is peak at 29=24.0°. Comparing with XRD of fused zirconia and XRD of PSZ, the cubic fused zirconia transformed from single pure c-Zr02 phase into the mixture of the c-Zr02 phase and the m-Zr02 phase after roasting. It can be inferred that cubic Zr0 2 phase is partially converted to monoclinic Zr0 2 phase. Structure of Two Phase and Relaxation of Interface Structure In this section, the first-principles research of Ca-doping c-Zr02 and m-Zr0 2 assisted by using computer software has been carried out. Considering the premise of the computer capacity, the thickness of interface is 15nm in the model, which can fully meet the requirements of the actual cases. And we approximately consider there are about 15 layers of atoms disordered in the interface layer (this number will be larger in practical situations, and the distance effected by relaxation will be longer). In the model, it is reasonable Zr and Ca are selected as the linkage between the Zr0 2 and CaO, Due to the lower surface energy of O working as the linkage between the Zr0 2 and CaO. The crystal structure of Ca doping c-Zr02 and m-Zr0 2 before relaxation is shown in

187

Fig.3 (a). It can be seen that: in the initial structure construction, ZrC>2 and CaO in cubic phase and monoclinic phase are both well ordered. However, the two-phase lattice does not match well. The main reason may be the difference of two-phase lattice parameters is very obvious. According to the formula (1), we can calculate mismatch from the lattice constant. 8=(aa-aß)/aa (1) Where a„ is the interatomic distance of Zr0 2 , ap is the interatomic distance of CaO , 5 is the mismatch. when 8 is between 0.25 and 0.5, the interface belong to non-coherent interface. After relaxation, the phase diagram of Ca-doping c-ZrÛ2 and m-Zr02 is shown in Fig.3 (b). It can be seen that greatly changes have taken place in the two-phase interface. For example: Interface atoms are scramble, the original lattice is seriously distorted, the space lattice periodically arranged is destroyed and the interface structure is obviously widened. Besides, notable changes have taken place at Ca atoms near the two interfaces. As follows: The bond angle becomes smaller and the bond length between Ca and O atom is becomes shorter, making the Ca closer to the O atoms in monoclinic phase and the whole energy of the interface minimum. Finally the structure becomes more stable.

a b Fig 3. Phase diagram of Ca-doping c-ZrC>2 and m-Zr02 before (a) and after (b) relaxation The results above show that: due to large lattice mismatch, wide interface and large interfacial energy, the c-Zr02 and m-ZrÛ2 are incompatible with great rejection. When Ca is added in the system, CaO as solute and ZrÛ2 as solvent, CaO displaces Zr0 2 , leading to lattice distortion and higher internal energy. Therefore, after relaxation, Solute atoms

188

(CaO) move into the loose grain boundary area, decreasing the internal energy. For this reason, the added Ca links the c-ZK>2 and m-ZrCh closely. Conclusion 1. Partially stabilized zirconia can be prepared by microwave heating, under the conditions of 1450 °C at a heating rate of 400 °C /min and held at this temperature for 120min. 2. The characterization of XRD results show PSZ is composed of crystalline compounds of cubic Zr0 2 phase and monoclinic Zr0 2 phase. 3. The relaxation optimization showed that the C-Z1O2 and m-ZrÛ2 were incompatible, with great rejection against each other. Ca plays a positive role in two-phase combination. Acknowledgements The research is supported by National Basic Research Program of China (No: 2007CB613606) and National Natural Science Foundation of China (No: 50734007). References [1] Kurumada M, Hara H, Iguchi E. Acta Materialia, 2005, 53(18):4839-4846. [2] Pawlowski Andrzej, Bucko Miroslaw M, Pedzich Zbigniew. Mater Res Bull, 2002, 37(3):425-438. [3] L.Haoa, J.Lawrence, G.C. Limb, H.Y. Zheng, Opt. Lasers Eng, 2004, 42:355-374. [4] L.Hao, J.Lawrence, D.K.Y.Lowb, G.C.Limb, H.Y.Zheng. Thin Solid Films, 2004, 468: 12-16. [5] N. Shekhar, B. Sunny, B. Bikramjit, Int. J. Appl Ceram Technol, 2008, 5:49-62. [6] Q.Y. Li, J.H. Du, Z.P. Xi, Trans. Nonferrous Met. Soc. China, 2007,17:560-564. [7] Jinhui Peng, Xianwan Yang. The New Applications of Microwave Power. Yunnan science and technology press, 1997. [8] Finnis M.W, Kruse C, Schiinberger U. Ab intoio calculations of metal/ceramic interfaces: What have we learned, what can we learn. Nanostruct Mater, 1995, 6(1):145-155 [9] James M.H. Comparison of the atomic structure, kinetics, and mechanisms of interfacial motion in martensitic, bainitic, massive and precipitation face-centered cubic-hexagonal close-packed phase transformations. Mater Sei Eng A, 2006, 438-440(11):35-42

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann- Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

CHARACTERISTIC OF MINERALIZATION OF SPECULARITE IRON ORES DURING COMPOSITE AGGLOMERATION PROCESSING Helei Zhang, Heng Yu, Guanghui Li, Yuanbo Zhang, Qian Li, Tao Jiang (School of Minerals Processing & Bioengineering, Central South University, Changsha, Hunan 410083, China) Key words: Composite agglomeration process; spéculante; sintering; pelletizing; ironmaking Abstract Large reserves of specularite iron ores have been found all over the world, which is characterized by high total iron grade, low gangues content, as well as low commercial price. However, it has not been widely used in sinter or pellet production due to its inferior sintering and roasting behaviors. It has shown that the proportion of specularite ores was increased as high as 40%~50% and good results was achieveed by using Composite Agglomeration Process (CAP), an innovative agglomeration process invented by Central South University. But the mineralization of specularite is still unknown during composite agglomeration processing. In this paper, the induration characteristic of specularite ore is studied by simulating CAP in laboratory, as well as microstructure analysis of CAP product. The results show that specularite pellet is mainly indurated through recrystallization of Fe2C>3, calcium ferrite within high basicity sinters matrix interweaves with the recrystallized Fe2Û3 on the surface of the pellet, which forms an integral microstructure, and ensures high strength of CAP product for ironmaking. Introduction Large reserves of specularite iron ores have been found all over the world, especially in Minas Gerais of Brazil, Xuanhua city and Anshan city of China, Elba island of Italy, St.Gotthard of Switzerland, Cumberland of Britannic, and so on. The specularite ores have a smooth surface, as well as a good crystallization and dense texture. When specularite is used to produce pellets, the specularite ore grains are difficult to adhere onto green ball due to its poor ballability. Moreover, specularite pellet requires high roasting temperature over 1300°C [1]. When specularite ore is used in sintering, the permeability of sintering bed gets bad and the quality of sinter is diminished. The practice in a few sintering plants finds that the output and strength of sinter decrease with the increase of specularite proportion, and the consumption of coke increases [2-4], Therefore, the utilization of specularite has been limited. Composite Agglomeration Process (CAP) is an innovative technology invented by Central South University [5-7]. The process has the following merits: making reasonable utilization of different iron ore resources; improving the quality and output of sinters; manufacturing sinters with low

191

basicity; significantly reducing energy consumption, etc[8-9]. It has been shown that large proportion spéculante ores, as high as 40%~50% can be used well by CAP [10], but the mineralization of spéculante pellet is still unknown during composite agglomeration processing. In this investigation, the induration characteristics of spéculante ore pellets is studied by simulating CAP in laboratory, and the phase composition and microstructure of CAP product are also characterized by using microscope and Leica image analyzer (Germany, Leica DM RXp). Experimental Materials The spéculante sample was taken from Brazil. The main chemical compositions of spéculante are listed in Table 1. The total iron grade of spéculante is as high as 67.03% and gangue composition, such as SÍO2 and AI2O3, are relatively low. The size distribution of specularite is 7.88% undersize 0.075mm, and it is coarser than the requirement of balling. By using SEM, the morphology of specularite particle is characterized, and the results are shown in Fig.l. The SEM results indicate that the particles of specularite exhibit irregular and smooth surface. Wet-grinding followed high pressure roller grinding is carried out to improve the ballability of specularite, and undersize 0.075mm content reaches 95% before pelletizing it. Tab. 1 The main chemical compositions of specularite ( % ) TFe 67.03

FeO 0.39

Si0 2 2.02

CaO 0.20

AI2O3

1.02

MgO 0.074

P 0.017

S 0.013

LOI 0.46

Fig. 1 SEM results of specularite ore Methods Simulation tests were conducted in an 88mm diameter vertical tube furnace in laboratory. Mixed gas with N2 and O2 was injected into the vertical tube furnace from its bottom, and gaseous flow was fixed at 16.5L/min. To simulate the variable roasting atmosphere, the proportion of O2 within the mixed gas was changed. At the beginning of test, the container with dried pellets was put on the top of furnace where the temperature is about 480°C for 30 seconds, then dropped

192

down to the zone at 700°C within 1 minute, and went down further to the zone with the highest temperature (1250°C, 1300°C, 1350°C, 1400°C) within another 2 minutes. The pellets underwent calcination for a given period to simulate combustion zone during composite agglomeration processing. Afterthat, the samples were pulled away from the furnace evenly within 4 minutes for cooling. Finally, the compression strength was measured after the roasted pellets were cooled to the ambient temperature. Some of them were sampled for phase and microstructure study. Results and Discussion Induration of Pellet During Composite Agglomeration Processing The effects of roasting temperature and time on compression strength of finished pellet are studied when the oxygen partial pressure is fixed at 9% and 21%, respectively. It can be seen from Fig.2 that the effect of roasting temperature is remarkable, and the compression strength of pellets increases with the roasting temperature below 1400 °C, as to the roasting time, the compression strength of pellets also increases with time at relatively low temperatures, but decreases at high temperature, especially at 1400°C.

(a): 9% oxygen partial oxygen (b): 21% oxygen partial oxygen Fig.2 Effects of roasting on compression strength of finished pellet Compared with higher oxygen partial pressure (21%), the maximum compression strength of the pellet indurated under 9% is lower, and the compression strength begins to reduce in shorter time and lower temperature. The main reason is that the temperature for the decomposition of FeîCh is attributed to the oxygen partial pressure, with which its decomposition temperature decreases. The theoretical decomposition temperature of Fe23 is 1305°C and 1342°C under 9% and 21%

193

oxygen partial pressure, respectively. Generally speaking, the temperature at the combustion zone of sintering varies in the range of 1100oC~1500°C, and time keeps about 5~7min. The experimental results indicate that the compression strength of pellet reach about 1800 N/P when roasted at 1300°C ~1350°C for 2~3min. It can be concluded that good strength of pellet can be achieved during composite agglomeration processing. Microstructure of Fired Pellet The microstructure of the finished pellet roasted at 1350°C for 3 minutes under oxygen partial pressure of 9% is shown in Fig. 3. It has been shown that the main connection pattern of hematite grain is the growth and «crystallization of Fe2Ü3 at high temperature [11]. Fig. 3 shows that almost all of hematite grains have recrystallized with each other although there are a few grains which lie adjacent. Moreover, the microstructure of internal pellet is almost the same with the external. Therefore, good compression strength can be achieved for pellets after processed by CAP.

Fig. 3 The microstructure of the finished pellets (a: internal, b: external; l-Fe2C>3) Pellet and Its Connection With High Basicity Sinter Matrix in CAP Product It is shown that the product of composite agglomeration includes acid part and high basicity matrix and the transition structure in-between[12]. Fig. 4 presents that Fe2Cb in specularite pellet is recrystallized and interconnected with each other. This result is coincided with those obtained by simulating tests. However, a few magnetite (Fe 3 0 4 ) grains (Fig.4 b) are interwoven with the recrystallization of Fe2Û3. Magnetite is contributed to the decomposition and reduction of FeaCh at high temperature. Compared with Fig. 3, it can be found that the recrystallization of Fe 2 0 3 of the pellets in CAP product is much better than in the simulating tests. Therefore, it can be concluded that the

194

compression strength of pellets in CAP product is higher than in the simulate tests.

Fig. 4 The microstructure of pellet in composite agglomeration product (a: internal, b: external; 1-Fe203, 2-Fe3Û4) In high basicity matrix, the calcium ferrite is well-developed (Fig. 5), and most of them appear as acicular shape, which improves quality of composite agglomeration product. In the transition part, calcium ferrite interweaves with the recrystallized FeaCh ( Fig.6). Therefore, an integral microstructure is formed in composite agglomeration product, which ensures good quality.

Fig. 5 The rricrostructure of hgh basicity sinter matrix (1-calcium ferrite)

Fig.6 The microstructure of transition between sinter and pellet (1-calcium ferrite; 2 - Fe2Û3) Conclusion

The simulating tests show that good strength can be achieved when spéculante pellet is treated under the conditions of composite agglomeration process. The pellet in composite agglomeration product is mainly indurated through recrystallization of Fe2Û3. Calcium ferrite is the main phase composition of the high basicity matrix. And calcium ferrite in the high basicity matrix interweaved with the recrystallized Fe2C>3 on the surface of the pellet, forming an integral microstructure, and ensures high strength of composite agglomeration product.

195

References [I] Aoping He, "A study of the Process and Mechanisms of Pelletizing Brazilian Spéculante" ( M.E.thesis, Central South University, 2005), 7-8 [2] Deqing Zhu, Zhi-yuan Wang, Jian Pan, et al., "Improvement of Sintering Behaviors of Brazilian Spéculante Concentrate by Damp Milling," Iron and Steel, 1, (2007), 12-16. [3] Xuhui Yang, "On Granulation Efficiency of the Mix Composed of 100% Iron Ore Concentrates," Sintering and Pelletizing, 2, (1994), 8-12. [4] Zhiyuan Wang. "Mineralization of Sintering Brazilian Spéculante and its Provemenf"(M.E. thesis, Central South University, 2006), 13-15. [5] Tao Jiang, Guang-hui Li, You-ming Hu, et al., "Composite Agglomeration Process of Iron Ores," Patent Number(s): CN200510032095, 2005. [6] Guanghui Li, Tao Jiang, Ke-cheng Zhang, et al., "A Method of Production of Low Bisicity Sinter," Patent Number(s): CN200610031342, 2006. [7] Tao Jiang, "The Concept and Study of Composite Agglomeration Process of Iron Ores," Proceeding of 2006 ' Annual Conference on Sintering and Pelletizing Technology Of China, (2006), 1-6. [8] Guanghui Li, Jing-hua Zeng, Tao Jiang, et al., "Study and Application of Composite Agglomeration Process of Fluoric Iron Concentrate," Journal of Iron and Steel Research International, (2009), 149-153. [9] T. Jiang, G H. Li, H. T. Wang, et al., "Composite agglomeration process (CAP) for preparing blast furnace burden," Ironmaking & Steelmaking, 1, (37), (2010), 1-7. [10] Yuan-bo Zhang, Tao Jiang, Guang-hui Li, et al., "Study on Iron Ore Sintering with High Proportion of Specularite Concentrates," Proceeding of 10' Chinese Conference of The Production Technology and Academic of Ironmaking, (2010), 187-192. [II] Jun Fan, YanZhong Jia, Delan Liang, "The Research On The Baking Progress For Hematite Pellet," Proceeding of 05 'Forum of Metallurgical Engineering, (2005), 26-30. [12] Tao Jiang, Kecheng Zhang, Youming Hu, "Investigations On Mineralization Mechanisms of Composite Agglomeration Process of Iron Ores," Proceedings of 06' International Conference on Agglomeration of Iron Ores, (2006), 36-41.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

2nd International Symposium on

High-Temperature Metallurgical Processing

Ferrous and Nonferrous Metals Session Chairs: Mark Cooksey Tao Jiang

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooks TMS (The Minerals, Metals & Materials Society), 2011

ENHANCING THE PELLIZATION OF BRAZILIAN HEMATITE BY ADDING BORON BEARING ADDITIVES Wei YU1'2, Deqing ZHU1, Tiejun CHUN1, Jian PAN1 1-School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, Hunan,PR China. 2- WISCO Minerals, Wuhan 430080, Hubei, PR China. Keywords: hematite, boron bearing additives, kinetics of crystal grain growth, fired pellets. Abstract In this paper, improving the pelletization of Brazilian hematite by adding a boron bearing additive was carried out, and the pelletization parameters were also optimized. The mechanism of boron bearing additive was revealed in the kinetics of crystal grain growth by using LEICA microscopy and its software to measure hematite grain growth. Good qualities of green balls are manufactured with drop numbers above 5.7 times/0.5m, compressive strength higher than 13 N/pellet and thermal shock temperature over 620°C under the following conditions: 1.2wt% bentonite, 0.4wt% boron bearing additive, 7wt% moisture and pelletizing in disc for 15 min. When the green balls were preheated at 1000°C for 8 min and fired at 1280°C for 15 min, the compressive strength of the fired pellets containing boron bearing additive is about 3096 N/pellet, which is far beyond the standard requirement for good quality pellets. Compared to the fired pellets without adding any boron bearing additives, the compressive strength of fired pellets containing boron bearing additive is increased by 1225 N/pellet. It is revealed that adding boron additive into the hematite pellets help to form more liquid phases, lower the activation energy of the hematite grain growth by 135.2 kj/mol and increase grain growth rate by 33.6%, leading to higher strength of fired pellets due to forming more coarse and dense recrystallized hematite and energy saving by firing at lower temperatures for shorten time. Introduction With the rapid development of iron and steel industry in China, there has been a significant increase in the demand for iron ores resulting in a high dependency on

199

imported iron ores in the past decade. The imports of iron ores have increased from 54.8 million ton in 1998 to 628 million ton in 2009 [1]. The imported Brazil iron ores possess high iron grade, low impurity content and rich reserve, which is becoming more and more popular in China steel companies. However, some imported hematite ore are poor ballability, higher firing temperatures and lower compressive strength of fired pellets. There have been reseach on improving the refractory of hematite. Ball mill and high pressure roller were used to pretreat the hematite or blend with domestic magnetite concentrate before balling [2-4]. Ball grinding and high pressure roller milling can improve the fine size fraction and specific surface area of hematite but limit the roasting performance. Blending with domestic magnetite concentrate is also limited due to the shortage of high iron grade magnetite [5-6]. Boron bearing additives such as borax, boric sludge and paigeite have already been applied in magnetite pelletization and research show obvious improvement to the pellet roasting property [7-8]. It is of great importance to Chinese iron and steel industries to apply additive to imported hematite and improve the roasting performance. In this study, improving the pelletization of Brazilian hematite by adding a boron bearing additive was carried out, and the pelletization parameters were also optimized. Experimental Materials The chemical compositions of Brazil hematite and its size distribution are given in Tables 1 and 2. The concentrate is a superior material for pelletization due to the high iron grade, low silicon and low harmful impurities. The size of 92.4% passes 0.074mm and specific surface area is 1626 cm2/g which can meet the normal standard for the pellet materials [9-10]. Table TFe 65.72 8.36

Types Hematite Bentonite

1. Chemical composition of iron ore concentrate (wt%). P S FeO Si0 2 A1203 CaO MgO Na 2 0 1.51 0.04 0.06 0.008 0.036 0.011 0.53 1.25 0.85 47.51 16.15 1.59 3.33 1.05 0.011 0.25

LOI 2.29 14.38

Table2. Size distribution and specific area of iron concentrate (wt%). Size ( mm )

+0.106

0.106-0.074

0.074-0.045

0.045 - 0.0385

-0.0385

Content

1.72

5.87

12.92

8.44

71.05

Moisture absorption

Table 3. Physical characteristic of bentonite. Expanding Colloid Montmorillonite content /wt% /vol% Volume /ml-g~'

200

-0.074mm /wt%

/wt% 187.5

6.5

91.0

88.6

100.0

The chemical composition and physical characteristic of bentonite are shown in Tables 1 and 3, and all the properties of this bentonite meet the specification of iron pelletization [10-11]. Boron bearing additive used in this study was a chemical purity. Experimental Procedure The experimental flow sheet is shown in Figure 1. Green balls were produced using pelletizing disc 1000 mm diameter and the moisture of green balls was fixed at 7%. Pellets were made from the mix of pretreated manganese fines and boron bearing additive in a disc pelletizer of 0.8 m in diameter and 0.2 m rim depth, rotational speed at 38 rpm and inclined at 47° to the horizontal. The drop number and compressive strength of the finished green balls were measured to evaluate the ability of the green balls to remain intact and retain their shape during handling, respectively. Dry balls were preheated and roasted in a tube furnace of 50 mm in diameter and 600 mm width. The compressive strength was measured to evaluate the quality of the roasted balls. Hematite Water

Green .

Blending

Balling

Bentonite

Drying, preheating,

Additive

roasting

Compressive

balls

properties test

strength

1—* ^

Porosity

1—»

Microstructure

Fig .1 Test flow sheet of experiment. Study included examination of microstructure of pellet and kinetics of the growth of hematite grain. Leica DMRXP optical microscope was used in microstructure study. The size of hematite grain can be tested by Image Analysis Software Qwin, according to M. Hillert formula [12-13]. The grain growth rate equation was as follows:

D"=yexp(--^) = *i

(1

where, K is grain growth constant. Taking the logarithm of the both side of equation, and equation 2 was obtained. «lnZ) = lnA: + liU It is seen that InD is linear with lnt, where slope is 1/n and intercept is (lnk)/n:

201

(2)

where, InD is linear with 1/T; slop is -Q/nR; Q is grain growth activation energy. Results and Discussions Effect of Boron Bearing Additive Dosage on the Properties of Green Balls Figure 2 shows the effect of additive dosage on the properties of green balls. The drop numbers of green balls increases from 4.5 times/0.5m to 9.1 times/0.5m when additive dosage was increased from 0.2wt% to 0.6wt%, and the compressive strength remain unchanged. Therefore, the boron bearing additive can effectively improve the drop numbers of green balls. Effect of Boron Bearing Additive on Preheated and Fired Pellet Figure 3 shows the preheated pellet compressive strength trend under various temperatures with different additive dosages. The higher preheating temperature, the higher compressive strength of preheated pellets. It also shows that the more dosage additive, the higher compressive strength of preheated pellets under the same preheating conditions. The compressive strength of 500 Newton of preheated pellets without additive was obtained when preheating temperature was over 1000°C, Howe/a, the preheating temperature of 950°Cis needed to achieve the same compressive strength of preheated pellets with 0.2wt% additive. Therefore, the boron bearing additive plays a significant role in reducing preheating temperature and improving the compressive strength of preheated pellets.

202

ü

Dosage of additive/ wt%

Figure 2. Effects of boron bearing additive dosage on properties of green balls (1.2wt% bentonite dosage and balling at 7wt% moisture for 15 min).

940

960

980

1000

1020

1040

Preheating temperature/°C

Figure 3. Effect of preheating temperature on the compressive strength of preheated pellet (preheating for 8 min).

203

1500 1400 1300 «

1200

¡5.1100 a

1000

I 900 U

800

70

I ° I **> |

600

400

300 6

7

8

9

10

11

12

Preheating time/rrin Figure 4 Effect of preheating time on the compressive strength of preheated pellets (preheating at 1000°C). The effect of preheating time on the compressive strength of preheated pellets is demonstrated in Figure 4. The compressive strength of preheated pellets increases slightly with prolonging preheating time. However, the boron bearing additive can observably improve the compressive strength of preheated pellets. The effect of firing temperature on fired pellet with various boron bearing additive is shown in Figure 5. It can be found that the compressive strength of fired pellet without additive show little increase with the increase in firing temperature. The compressive strength was only 1865 N/pellet when temperature of 1300°C was used, which is far from the requirement of 2500 N/pellet and presented poor firing ability. However, the compressive strength of fired pellets was elevated significantly after adding the boron bearing additive. The compressive strength of 3096 N/pellet was achieved when adding 0.4wt% additive and firing at 1280°C for 15 min.

204

3000

g 2500

55 2000

E 1500

8

1000

1200

1220

1240

1260

Firing temperature/ °C

1280

Figure 5. Effect of firing temperature on the compressive strength of fired pellet (preheating at 1000°C for 8 min and firing for 15 min).

8

10

12

14

16

18

Firing time/rrin Figure 6. Effect of firing time on the compressive strength of fired pellet (preheating at 1000°C for 8 min and firing at 1280°C). The effect of firing time on the compressive strength of fired pellets is shown in Figure 6. With extending the firing time, the compressive strength of fired pellet without additive is elevated slightly, and the compressive strength is only 1968 N/pellet when firing at 1280°C for 18 min. However, the compressive strength of 2700 N/pellet was obtained when adding 0.4wt% additive and firing at 1280°C for 12min. The boron bearing additive

205

can efficiently improve the firing performance of fired pellets, and also reduce the temperature and short on the time. Action Mechanism of Boron Bearing Additive Microstructure. Figures 7 and 8 show the microstructures of fired pellets without additive and with 0.4wt% additive when preheated at 1000°C for 8 min and fired at 1280°C for 15 min. There was only a little interlinkage of recrystallised hematite. Most of hematite still acts as angular primary crystallite and with little liquid filling among the grain irregular shape holes were formed. The binding was not strong enough due to the little liquid phase and high porosity, resulting low compressive strength of fired pellet. However, large area of recrystallised hematite binding together and liquid phase filling into the corrosion and skeleton structures in pellets with 0.4wt% additive (Figure 8).

(a) without additive,

(b) wth 0.4v*% boDn hearing additive

A—Fe203, B—Hole

( x500 )

Figure 7 Microstructure of oxidized pellet Growth Kinetics of Hematite Grain. According to the mineralogy of fired sample, growth of hematite grain and recrystallised binding of hematite were the main reasons for the improvement of the roasting ability of pellet. Hillert grain growth kinetics model was used to calculate the hematite grain growth constant and grain growth activation energy which aimed to reveal effect of boron bearing additive on hematite growth. Effect of firing temperature on hematite grain size of fired pellet is shown in Figure 9. It is demonstrated that the increase of firing temperature increase the hematite grain size of fired pellet both without and with additive. The additive dramatically improved the hematite grain size under the same conditions. The pellet without additive possesses the hematite grain size of 14.5um at 1300CC compared to the hematite grain size of 17.7um

206

when adding 0.4wt% additive and firing at 1200°C.

18

I 14

1150

1175

1200

1225

1250

1275

13O0

Firing terrperature/°C

Figure 9. Effect of firing temperature on the hematite grain size (firing for 50 min).

S 12

Ï

20

30

40

50

Firing timafmin Figure 10. Effect of firing time on the hematite grain size (firing at 1250°C). The effect of firing time on hematite grain size in pellet is shown in Figure 10. The grain size increases markedly when extending the firing time. Boron bearing additive also can obviously increases the grain size under the same firing conditions. The hematite grain size of 13.9um was achieved when adding 0.4wt% additive and firing for 30 min, which is bigger than the hematite grain size of pellets without additive for firing 50 min.

207

Using equation (2) and results in Figure 10, the linear fitting between InD and lnt was obtained and is shown in Figure 11. The hematite grain growth constant k was calculated from the intercept (lnk/n) of the two lines. The hematite grain growth constant k of pellets with additive is 7.59><105(um)n,min"1,which increases 66.3% compared to the hematite grain of 5.04><105 (nm^'min"1 of pellets without additive. 2.72 2.70 2.68 2.66 2.64 2.62 2.60 Q Ü

2-68 2.56 2.54 2.52 2.50 2.48 2.46 2.44 2.42 3.4

lnt

Figure 11. Fitting lines of InD vs Int. 2.9

2.8

2.7

2.6 Q c - 2.5

24

2.3

2.2 0.00063 0.00064 0.00065 0.00086 0.00067 0.00068 0.00069 0.00070 0.00071

1/T

Figure 12. Fitting line of InD vs 1/T. Using equation (3) and dates from Figures 8 and 9, the linear fitting between InD and 1/T was calculated and is shown in Figure 12. The slope of this line is -Q/nR and the hematite grain growth activation energy of pellet with and without additive are 242.3

208

kj/mol and 377.5 kj/mol, Respectively. It can be seen that the dosage of additive can drop the hematite grain growth energy of 135.2 kj/mol and dramatically improve the growth of hematite. Conclusions 1. Good qualities of green balls are manufactured with drop numbers above 5.7 times/0.5m, compressive strength higher than 13 N/pellet and thermal shock temperature over 620°C under the following conditions: 1.2wt% bentonite, 0.4wt% boron bearing additive, 7% moisture and pelletizing in disc for 15 min. When the green balls were preheated at 1000°C for 8 min and fired at 1280°C for 15 min, the compressive strength of the fired pellets containing boron bearing additive is about 3096 N/pellet, which is far beyond the standard requirement for good quality pellets. 2. Compared to the fired pellets without adding any boron bearing additives, the compressive strength of fired pellets containing boron bearing additive is increased by 1225 N/pellet. It is revealed that adding boron additive into the hematite pellets helps to form more liquid phases, lower the activation energy of the hematite grain growth by nearly 135 kj/mol, leading to higher strength of fired pellets due to forming coarse and dense recrystallized hematite and energy saving by firing at lower temperatures for shorten time. References 1. Chen. Q "Imports of iron ore in China in 2009," China Customs, 4(2010), 30-32. 2. Huang. Z. C, et al., "Study on mainly made from hematite based oxidized pellets with addition of some magnetite," Iron & Steel, 39(4) (2004), 9-10. 3. Zhu. D. Q, et al., "Improve pelletitation of brazilian spéculante by using high pressure roller grinding," Journal of University of Science and Technology Beijing, l(2009),30-35. 4. Li. Z. H, "On increasing grade of pellets and comprehensive economic results," Sintering and Pelletizing, 23(6) (1998), 10-13. 5. Liu. G F, "Practice of increasing the pellet grade in Jigang," Sintering and Pelletizing, 26(6) (2001), 37-38. 6. Zhang. Y. X, et al., "Pelletizing test of adding various complex binders," Sintering and Pelletizing, 29(5) (2004), 9-11. 7. Zhang. K, Yang. Z X, "Present state and development trend of using boron-containing additives for agglomeration," Sintering and Pelletizing, 23(2) (1998), 20-28. 8. Zhang. Y. Z, "A study on the influence of boron-magnesium addition on agglomeration process of iron ore" (Master. Thesis, Hebei University of Technology, 2000), 1-6.

209

9. Zhao. Q. J, He. C. Q, Gao M H, "Reasonable Utiligation of Boronic Magnetite Ore," Jounal of Exist China University of Metallurgy, 14(3) (1997), 262-266. 10. Fu. J .Y, Jiang. T, Zhu. D. Q, Principles of sintering&pelletizing, (Changsha, Centre South University of Technology Press, 1996), 233 11 Fu. J. Y, Zhu. D. Q, Basic principles, techniques and equipments of iron oxidized pellets, (Changsha, Central South University Press,2004), 30-32. 12. Hillert M, "On the Theory of normal and abnormal grain growth," Acta Metallurgica, 13(1965), 227-238. 13. He X, "Synthsis of nanocrystalline Sn02, Ti02% BaTi03 and grain growth kinetics of nanocrystalline materials" (Master. Thesis, Central South University, 2007), 5-6.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cook TMS (The Minerals, Metals & Materials Society), 2011

STUDY ON IMPROVING T H E QUALITY OF PELLET MADE FROM VALE HEMATITE PELLET FEED Vinicius Mendes 12, Deqing ZHU1, Tiejun CHUN1, Jian PAN1, Marcus Emrich2 'School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, Hunan, P.R. China. 2 Vale, Rio de Janeiro 20030, Brazil. Keywords: Vale hematite; Magnetite concentrate; Pellets; Roasting; Metallurgical performance. Abstract Joint projects of studies of hematite acid pellet have been implemented to develop technological solutions for using Vale hematite pellet feed in Chinese palletizing plant. Beneficial improvements on the pellets quality were achieved by adjusting the blends composition and using new technology of pre-treating the blended pellet feed with HPGR (High Pressure Grinding Roller) on rotary kiln pellet technology. The drop strength of green balls increased from 3.4 to 6.0 times from 0.5m height, at the same time, the bentonite dosage decreased from 1.5wt% to 0.7wt%. The compressive strength of roasted pellets reaches 2500 N/pellet. The reduction swelling index of roasted pellets is below 15%. Introduction With the fast development of steel industry in China, the production of pellets has been further expanded. Pellets are used as a high quality charge for blast furnace and the annual output is around 80Mt pellets [1]. According to the developmental trend, sinter will be replaced by pellets due to the larger environmental impact of sintering plants when compared to pelletizing plants. For example, the specific CO2 emission in sintering process is around 2.5 times higher than in pelletizing process. However, the steel industry of China has been confronted with a serious shortage of domestic pellet feed due to the low iron grade of Run-of-Mine ore despite of huge reserves of iron ores [2]. Therefore, more pellet feed, mainly hematite, has been imported to make pellets. Brazil has abundance in iron ores reserves, especially hematite, and Vale is the major supplier of pellet feed of hematite type [3]. But the portion of Vale hematite has different behaviors in palletizing contrasting with domestic magnetite, the coarse grain size, poor

211

wettability of particles and small specific surface area of iron ore concentrate influence negatively the output and quality of pellet. So the problem how to use this kind of hematite needs to be solved. In this context, some important research on how to utilize Vale hematite make pellet has been run, and many remarkable technological solutions have been proposed. Vale Iron Ore Resource As shown in Figure 1, Vale owns three production systems in Brazil: the Northern, the Southeastern and the Southern System. All of them are constituted by mine, railroad and port [3]. The Northern System - It started industrial operation in 1985. It is constituted by the Carajás mines, the Carajás Railroad (892 km) and by the Ponta da Madeira Port, in Säo Luis.

1) Proven & probable reserves, as of December 2004. 2) Under US GAAP Figure 1. Vale Northern, Southeastern and Southern production systems. The total ore reserve in Carajás is about 1.8 billion tons, being Carajás the Brazilian largest mineral province. The production commercialized in 2005 was 73.4 million tons, being 54.9 million tons of sinter feed, 8.7 million tons of lump ores for blast furnace and 3.4 million tons of pellet feed. The run of mine is already a high grade ore (66wt% Fe) and so it is only classified to produce lump ore, sinter feed and pellet feed. The Southeastern System - It is constituted by 14 mines, in Minas Gérais State, by the Vitoria-Minas Railroad (905 km), seven pelletizing plants and by the Tubaräo Port, being these last two located in Vitoria.Total ore reserve in the Southeastern System is 1.1 billion

212

tons of hematite and 3.4 billion tons of itabirite. They are concentrated in beneficiation plants. Production commercialized in 2005 was 111 million tons, being 65.8 million tons of sinter feed, 14.5 million tons of pellet feed and 10.0 million tons of lumps. The Southern System - This last system includes operations of MBR (Mineraçoes Brasileiras Reunidas). It is constituted by 6 mines, in Minas Gérais state, by the MRS railroad (510 km) and by two ports: Itaguai and Guaiba, both located in Rio de Janeiro state.The total ore reserve in Southern System is 0.6 billion tons of hematite and 0.6 billion tons of itabirite. Production commercialized in 2005 was 47.6 million tons, being 27.7 million tons of sinter feed, 12.8 million tons of lumps, and 7.1 million tons of pellet feed. With the new investments, Vale iron ore production reached 300 million tons in 2009. Raw Materials and Experimental Raw Materials The chemical composition and physical characterisitc of Vale hematite pellet feed and domestic magnetite pellet feed are shown in Table 1 and 2, respectively. Table 1. Composition of raw materials (wt%). Ore type

TFe

FeO

Si0 2

A1203

CaO

MgO

Vale hematite

66.79

0.19

2.47

0.82

0.10

0.038 0.042

64.39

25.13

4.31

0.86

1.86

0.78

Domestic magnetite

P

S

LOI

0.017

0.64

0.007 0.066 0.77

Table 2. Physical characteristic of two types iron ores. Ore type

Vale hematite Domestic magnetite

-0.074mm (%)

-0.043mm (%)

Specific surface area (cm2/g)

Static pelletability index

91.14

61.59

520

0.190

80.40

40.70

1875

0.651

Vale hematite ores possess higher iron grade of 66.8wt% Fetotai and low silica content. The size of Vale hematite ores is 91.1% passing 0.074 mm, with a specific surface area of 520 cm2/g and static pelletability index of 0.190. It is necessary to reduce the particle size even more, and obtain adequate specific surface area of 1500-1700 cm2/g for improving the pelletizing process. The grade of domestic magnetite pellet feed is 64.4wt% Fetotai, It is lower than hematite

213

pellet feed content, and FeO content is 25.1%. Moreover, the specific surface area of domestic magnetite pellet feed is much higher than Vale hematite pellet feed, resulting in an excellent ballability which the static pelletability index of 0.651. Methods The pellet feed and bentonite were homogeneously mixed at a given ratio and prepared for balling. Green balls were made in pelletizing discs, and followed by drying and firing [4-6]. The chief equipment of Rotary-Kiln which roast pellet is shown in Figure 2.

Figure 2. The equipment of Traveling-Grate & Rotary-Kiln for drying,preheating and roasting pellets. Various characteristics of raw materials, green balls and fired pellets were determined, including size, particle morphologies, specific surface area compressive strength, drop number, thermal stability, metallurgical performance, mineralogy and chemistry of pellets. The specific surface area was measured by Blaine measurement device according to the standard of GB8074-87. The compressive strength of preheated and fired pellets was measured according to the standard of ISO 4700 (1996). The metallurgical performance, including reducibility, reduction swelling were determined according to the standard of ISO 7215 (1995), ISO 4698, respectively. Results and Discussions The Ballability and Roasting Performance of 100% Vale hematite and 100% Domestic magnetite Figure 3 illustrates the effect of bentonite dosage on the strength of green balls. With the bentonite dosage increasing from 0.0% to 2.5%, the drop numbers and compressive strength of green balls are heightened for Vale hematite and domestic magnetite. However, comparing with domestic magnetite, the strength of green balls made from Vale hematite is improved so less. The drop number of Vale hematite balls is only increased from 0.8 times/0.5m to 3.2 times/0.5m whereas the drop

214

number of domestic magnetite balls is risen from 2.83 times/0.5m to 9.5 times/0.5m. The compressive strength of Vale hematite balls is also below the compressive strength of domestic magnetite balls. From the big difference in strength of green balls between Vale hematite and domestic magnetite, it can be seen that the ballability of Vale hematite is unaccepted and not as good as domestic magnetite, with the small specific surface area and low static pelletability index.

o.o-

00

0.5

1.0 1.5 Bentonite Dosage(wl%)

2.0

2.5

1-the drop nurrber of 100% Vale hematite

2-the compressive strength of 100% Vale hematite

3-the drop number of 100%domestic magnetite

4-the compressive strength of 100%domesacmagnetite

Figure 3. Effects of bentonite dosage on the strength of green balls.

I O.

lfe

'S

I

S

42

Conditions:

3000 -

Preheating temperature 1150°C, Preheating timel5nin, Induration timel5nin

L

2000 -

1160 1180 1200 1220 12« 1260 1280 1300 1320 1340 Temperature (°Q 2-100% domestic magnetite

l-100%Valehematite

Figure 4. Effect of temperature on the compressive strength of fired pellets.

215

The effects of induration temperature on the compressive strength of fired pellets are shown in Figure 4. It can be found that the domestic magnetite possesses the more excellent firing performance than Vale hematite. The compressive strength of domestic magnetite fired pellets is about nearly 4000 N/pellet when firing at 124CFC, whereas the compressive strength of Vale hematite fired pellets is lower than 1000 N/pellet. Improving the Ballabilitv and Roasting Performance of Vale Hematite Test # 1 : Blending Vale Hematite with Domestic Magnetite Vale hematite pellet feed was blended with different amounts of magnetite pellet feed and then grinded on ball mill. The mixed and grinded pellet feed was blended with 1.0wt% to 2.5wt% of bentonite, pelletized for 15 min at 10% moisture, then dried and fired. Table 3 shows the chemical composition of the fired pellets, together with the analytical characteristics of pellets of Vale hematite blended with different ratios of domestic magnetite pellet feed. It can be seen that the drop number reached 3.4 times from 0.5m height. The compressive strength of fired pellet reached 2,906 N/pellet with magnetite weight ratio of 30%. The oxidation of magnetite is exothermic reaction, which contributes to good firing and lower fuel consumption. The newly formed hematite from magnetite oxidized has higher reactivity during solid state reactions. Exothermical reaction of magnetite during oxidation to hematite can raise the temperature inside pellets [7-8]. It can also be seen that the RI and RSI of fired pellets blended with magnetite was on the limit (Table 3). It was proposed to reproduce again test # 1, but using high pressure roller press instead of grinding ball mill aiming at increase balling property by HPGR (high press grinding roller) of improving iron ore surface characteristic. Test # 3 described in further part of the paper. Test # 2: Using 100% Hematite Pellet Feed and Adjusting Basicity The same methodology adopted on test # 1 was again used, but this time using 100% hematite and adjusting basicity by adding limestone. Table 3 shows the characteristics of 100% Vale hematite pellets at two different acid basicity levels. It can be seen that the compressive strength of fired pellets increased from 2,228 N/pellet to 2,624 N/pellet with the basicity of pellets rising from 0.2 to 0.3. The additive accelerates solid state reaction, boosts up the intensity of induration and reduces negative effect on métallurgie performance of pellets [9]. But there is suitable basicity according to different Si0 2 content. It is very important to choose suitable basicity of pellets which can guarantee not only good compressive strength of fired pellets but also good

216

métallurgie performance of pellets. In conclusion, the method of adjusting basicity by adding some additives can improve the characteristic of hematite pellets. Test # 3: Pretreating Blending Vale Hematite with Magnetite by HPGR For test # 3, 70wt% of hematite was blended with 30wt% of domestic magnetite and then pre-treated the mixed material by HPGR. Table 3 shows the characteristic of the pellets with the feed pretreated by the roller press. A roller press circuit with close recirculation circuit was tested. It can be seen that the characteristics of pellets was improved because the pellet feed was pretreated by roller press before balling. Static ballability index achieved the value of 0.798 for the recirculation circuit. The drop strength of green balls increased from 3.4 to 6.0 times from 0.5m height. At the same time, the bentonite dosage decreased from 1.5wt% to 0.7wt%. The finer the pellet feed pre-treated by HPGR, the lower of reduction swelling degree of fired pellets, which is below than 15vol%. Reducibility of pellets also increased to 67.18% when the pellet feed is pre-treated by HPGR. After pre-treating pellet feed blend, the specific surface area reached nearly 1500 cm2/g, with much rougher surface and more super fine particles occurring, as shown in Figure 4. The activation of crystal lattice and an increase in specific surface area by roller press can enhance balling dynamics and speed of solid state reaction [6,9]. Pre-treating hematite is a good technique which can wreck the disfigurement of crystalloid and activate the crystal. Note on Figure 4 that the hematite have good mineral composition and microcosmic structure, which can improve the quality of the pellets. In general, pre-treating hematite pellet feed on the roller press is an optimal technique to improve pellets quality.

Figure 4 Morphologies of Vale hematite after pre-treating by HPGR

217

Table 3. Chemical, physical and metallurgical characteristics offiredpellets. Metallurgical Physical Chemical composition characteristic characteristic Test types DNGP CSFP RI RSI K™ TFe FeO Si0 2 A1203 CaO 3 m (Times/0.5m) (N/pellet)

(%)

r""

(%) " (%)

VH80wt%

66.18 0.47 4.04

1.86

0.47

3.2

2,741

63.44 35.86

2.74

VH70wt%

65.82 0.42 4.14

1.92

0.64

3.4

2,906

58.82 20.36

2.07

VH100wt% Basicity

65.82 0.43 3.51

1.05

0.70

3.3

2,228

64.36 25.96

2.67

65.55 0.40 3.54

1.03

1.05

3.1

2,624

68.79 21.37

1.24

66.18 0.47 4.04

1.86

0.47

6.0

2,478

67.18 13.08

2.35

°-20 VH100wt%

Basicity 0.30 VH 70wt% #3 with Roller Press

*Footnote: VH-Vale hematite; DN-drop numbers of green pellets; CSFP-compressive strength of fried pellets; RI-reducibility index; RSI-reduction swelling index; RDI-reduction degradation index. Conclusions (1) According to studies undertaken at the Central South University and Vale on hematite acid pellet, technological solutions were achieved for the Chinese pellet industry. Beneficial improvements on productivity and quality were achieved by using Vale hematite as pellet feed on rotary kiln technology. The final result was optimized by using some technological solutions, such as: - blending hematite with different proportion of magnetite pellet feed (20 and 30wt%) and grinding it using ball mill. - using 100wt% hematite pellet feed grinded by ball mill and adjusting basicity by adding limestone. Blending hematite with 30wt% of magnetite and pre-treating the blended pellet feed by roller press is an optimal technique to improve pellets quality. Good quality was reached on the fired pellets, highlighting: chemical composition, ballability, compressive strength and metallurgical characteristic.

218

(2) Vale believes it is not only prepared to supply hematite pellet feed, based on superior quality, but to develop and offer customized technological solutions for the Chinese pellet industry. Moreover, Vale is quite confident that it will manage to consolidate our attendance in the Chinese market as a reliable supplier ready to establish long-term relationships. References 1. Li. M, Chen. R.M, "The Status and Prospect of pellets Production Inside and Outside," China Metallurgy, 11 (2004), 1-3. 2. Peng. Z.J, Luo. H, "Experimental Study on Oxidized Pellet Fabrication with Adding Brizil Hematite Ore Fines and Its Roasting Properties," Research on Iron & Steel , 6 (2005), 1-4. 3. Chen. H, "Scenario of World Iron Ore Resources and Production," Iron & Steel, 36(11) (2001), 69-73. 4. Fu. J.Y, Jiang. T, Zhu. D.Q, Sintering and Pelletization, (Changsha: Central South Universityof Technology Press, 1996), 38. 5. Fu. J.Y, Zhu. D.Q, Theory,Technics and Equipment of Iron Ore Pellets, (Changsha: Central South University Press, 2005), 234. 6. Zhu. D.Q et al., "Mechano-chemical Activation of Magnetite Concentrate for Improving Its Pelletability by High Pressure Roll Grinding," ISIJ International, 44(2) (2004), 310-312. 7. Zhang. H.Q, Qi. Y.G, "Appl ication of roll press in improvement of qual ity of iron ore pellet," Mining Engineering, 3(1) (2005), 37-40. 8. Zhang, Y.B, Yan. Y.Y, Zhu, Q, "Grinding Test of Iron Ore Concentrates by High Pressure Roller Mill," Sintering and Pelletizing, 30(3) (2005), 14-17. 9 Qiu. G.Z, et al., "Improving the Oxidizing Kinetics of Pelletization of Magnetite Concentrate by High Press Roll Grinding," ISIJ International, 44 (1) (2004), 69-70.

219

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

DECOMPOSITION AND OXIDATION OF BISMUTHINITE IN NITROGEN-OXYGEN ATMOSPHERES R. Padilla, R. Villa, M. C. Ruiz Department of Metallurgical Engineering, University of Concepción Edmundo Larenas 285, Concepción, Chile Keywords: Bismuthinite, Decomposition, Vaporization Abstract Bismuth is an impurity in copper minerals where it occurs mainly as the mineral bismuthinite (BÍ2S3). In smelting of copper minerals, most of the bismuth reports to the white metal, as a result, a substantial amount of bismuth accompanies the copper to the electrorefined cathodes. Thus understanding the behavior of BÍ2S3 at high temperatures is crucial to eliminate bismuth before electrorefining. In this paper, some experimental data on the decomposition/volatilization of bismuthinite in nitrogen- oxygen atmospheres is discussed. The results indicated that in nitrogen atmosphere, bismuth volatilization occurs through a fast decomposition of bismuthinite to metallic bismuth and sulfur with subsequent volatilization of bismuth in the range 850-1100 °C. Similarly, in the presence of oxygen, bismuthinite decomposes to bismuth too, followed by oxidation to the non-volatile bismuth trioxide at temperatures higher than 750 °C. Both, temperature and partial pressure of oxygen affect significantly the bismuth volatilization and the oxidation rates. Introduction Copper sulfide minerals are the major source for the production of metallic copper by pyrometallurgical processes, which are the dominant processes in the production of copper by selectively oxidizing the sulfide sulfur to SO2 and the accompanying iron in the concentrates to iron oxide, which can be conveniently removed as a stable slag. Generally, copper concentrates contain some deleterious minor elements such as arsenic, antimony, bismuth and lead, which are present as sulfides. These elements depending on their concentration may originate problems in the copper smelting processes and they can ultimately lead to a final copper product with unacceptable high levels of impurities. During the treatment of these copper concentrates by smelting, the impurities generally distribute themselves in the condensed and gas phases and they are eliminated mainly through the gas phase [1-3]. Thus, smelter off gases contains substantial amounts of volatilized impurity elements. For example, in the case of bismuth, the estimated distribution of bismuth in flash furnace smelting [2] was 15 % in the matte, 5 % in the slag and 80 % in the gas phase, and recently, for Outokumpu flash smelting Davenport [3] reported the distribution of bismuth as 3075% in matte, 5-30% in slag and 15 -65 % in the off gas.

221

The present paper is concerned with an experimental study on the behavior of bismuth at high temperatures in neutral and oxidizing atmospheres. Experimental Work The experimental work in this study was carried out in the temperature range of 550 to 1100°C under nitrogen and mixtures of nitrogen-oxygen atmospheres. Bismuth tnsulfide of 95.43 % BÍ2S3 was obtained from Aldrich Chemicals for the experiments. The nitrogen gas utilized was of high purity and the mixtures of nitrogen-oxygen contained 1- 21 % O2 by volume. The volatilization/oxidation reaction of the BÍ2S3 in any particular atmosphere was studied in a thermogravimetric apparatus. This apparatus consisted essentially of a vertical tube furnace with a temperature controller unit, an electronic balance, data acquisition and gas distribution systems. The experiments were isothermal in nature, and the experimental procedure started by heating the furnace to a pre-specified temperature in nitrogen atmosphere and allowed to reach thermal stabilization under a constant flow of gas, nitrogen or mixtures of nitrogen-oxygen. At the set temperature, 50 - 200 mg of the BÍ2S3 sample contained in a 5 ml crucible was suspended in the furnace tube and the instantaneous weight loss was recorded as a function of time. Results Preliminary experiments were carried out at 1050 °C to determine the effect of the gas flow rate on the decomposition/volatilization of the bismuth sulfide in our experimental setup. The studied flow rates were in the range 0.6 to 1.5 L/min. The results indicated that the gas flow rate had negligible effect on the weight loss that the sample experienced. Thus, 1.5 L/min was adopted as the standard flow rate to minimize mass transfer effects in our system. Decomposition/Volatilization of Sb2Sj in Nitrogen Atmospheres Isothermal experiments were carried out to study the behavior of BÍ2S3 in nitrogen atmosphere in the temperature range of 800 to 1100 °C. The results are shown in Figure 1, where the data are shown as the fraction of weight loss of the bismuth tnsulfide sample as a function of temperature. We observe in this figure that temperature has a significant influence on the weight loss of the sample. For all the temperatures studied, one can see a very rapid initial weight loss followed by a change in the slope of the weight loss curve. Thus, we can deduce from this data that the sample losses weight due to two distinct processes which is very evident at the lower temperature (850 °C) by the abrupt change in the slope of the curve. This change occurs at around 0.18 fractional weight loss, which would correspond to the theoretical weight loss for the decomposition of SD2S3 to metallic bismuth calculated as 0.18. In order to verify this decomposition process, several experiments were carried out at 1050 °C to obtain partially reacted samples to identify the compounds present in the residues for the times indicated by number 1 and 2 in Figure 2. These residues were analyzed by X-ray diffraction spectroscopy (XRD) and the results are shown in Figures 3 and 4. In Figure 3, we can observe diffraction lines for both Sb2S3 and Bi; this result indicates that the Sb2S3 sample decomposes

222

indeed initially very rapidly to metallic bismuth in nitrogen ambient. At longer times, only bismuth is present in the sample as confirmed by the results shown in Figure 4. Thus, the following reactions would represent a two stage decomposition/volatilization of bismuthinite at high temperatures in nitrogen atmospheres,

¡5" |

0.4

20

40

60

80

100

120

Time, min)

Figure 1. Bismuth volatilization from BÍ2S3 in nitrogen atmosphere as a function of temperature.

0

20

40

60

80

Time, rrin Figure 2. Decomposition of BÍ2S3 sample at 850 °C in nitrogen atmosphere. The numbers on the curve indicate sampling times for XRD analysis. and

Bi2S3 =Bi +S 2 (g)

(1)

Bi = Bi(g)

(2)

The monomeric Bi gas was considered in reaction (2) because the concentration of other polymers such as BÍ2 gas at this temperature has been reported to be negligible [4].

223

10

20

30

40

50

2Theta, degrees Figure 3. X-ray diffraction spectra for the residue corresponding to the time indicated with number 1 in Figure 2.

3000 .■§■

1

2000 1000 o|!Wiii^M,i l i J l ( . ( » w .fr > i,|, WHMiHÉtiUi 10 20 30

40

50

60

2Theta, degrees Figure 4. X-ray diffraction spectra for the residue corresponding to the time indicated by number 2 on the curve shown in Figure 2. Decomposition/Volatilization of Sb?S^ in Oxygen-Nitrogen Atmospheres Figure 5 shows the experimental results obtained in reacting BÍ2S3 with an atmosphere containing 5% O2. In this figure, the effect of temperature on the weight loss rate of the sample is shown for the range of 600 to 725 C, which is below the melting point of bismuthinite (750 °C). On the other hand, Figure 6 shows the results for the higher range of temperature of 900 to 1050 °C. We can observe in both figures that the rate of weight loss of the sample is significantly affected by changes in the temperature. We can also observe that at the lowest temperature of 600 °C, the sample gradually loses weight at a very low rate compared to higher temperatures (700 and 725 °C) where the sample losses weight very rapidly up to about a fraction of 0.18 and then begins to gain weight slowly in time.

224

0.30

0.25

^ 0.20

ï

I 0.15

0.05

0.00

0

10

20 30 Time, min

40

50

Figure 5. Decomposition/oxidation of bismuthinite in 5% oxygen in the gas phase in the temperature range of 600 to 725 °C. At the higher temperatures of 900 to 1050 °C, the weight loss of the samples is very rapid and all the curves exhibit a maximum at the weight loss fraction of 0.18; for example, at 950 °C and higher the maximum is reached in less than 4 min. The following weight gain of the sample proceeds down to a limiting value of 0.10. In order to determine the intermediate compounds formed in the process, partially reacted samples were produced for XRD analysis; the times considered are indicated by point 3 and 4 on the 1050 °C curve shown in Figure 7. The XRD result depicted in Figure 8, which corresponds to the time indicated by point 3, shows diffraction lines solely for bismuth. This result indicates that even in oxidizing atmosphere Bi2S3 firstly decomposes to Bi before any oxidation takes place. On the other hand, the XRD result corresponding to point 4 is shown in Figure 9, where we can see diffraction lines only for Bi 2 0 3 . Based on diese results, the following reactions can be written for the behavior of bismuthinite in oxygen containing atmosphere at high temperatures. and

Bi2S3 = Bi +S2(g)

(3)

Bi + 0 2 (g) = Bi 2 0 3

(4)

The decomposition oxidation of bismuthinite through reactions (3) and (4) is supported by the agreement between the experimental values of consecutive weight loss and weight gain of the samples with the theoretical values according to reactions (3) and (4).

225

0.30

025

0.20

0.15

m 0.10

0.05

0 00 6

10

20

A

Figure 6. Decomposition/oxidation of bismuthinite in 5 % oxygen concentration in the gas at 900 to 1050 °C. 0.3

|

02

! lo,

0.0

0

2

4

6

8

10

Time, frin

Figure 7. Typical weight loss of a bismuthinite sample reacting in 5 % oxygen at 1050 °C. Points 3 and 4 in the figure indicate the reacting time for XRD analysis. Effect of Oxygen Concentration on the Decomposition/Oxidation of Bismuthinite Experiments were carried out using gas mixtures with 1%, 10% and 2 1 % O2 at 750 °C. The results are shown in Figure 10, where we can observe that the decomposition rate and the oxidation rate are strongly dependent on the oxygen concentration at this relatively low temperature. The same behavior has been found at higher temperatures, too. These experimental results indicate that bismuth elimination through the gas phase from copper concentrates that contain bismuthinite at roasting or smelting temperatures in oxygen-nitrogen

226

atmospheres would be difficult since Bi2S3 decomposes at a fast rate to metallic bismuth, which subsequently oxidizes with also fast rate to stable bismuth trioxide.

12000 10000

§.

aoco

¿

6000

I M

_

4000 2000 0 10

20

30

40

50

60

2Theta, degrees

Figure 8. XRD pattern of a partially reacted bismuthinite sample (corresponding to point 3 in Figure 7) in 5% oxygen at 1050 °C. 5000 4000 §■ 3000 j5 2000 "~ 1000 0 10

20

30

40

50

60

2 Theta, degrees Figure 9. XRD pattern of a reacted bismuthinite sample (point 4 in Figure 7) in 5% 0 2 at 1050 °C. Conclusions - In nitrogen atmosphere in the range 800-1100 °C, bismuthinite decomposes very rapidly to bismuth and sulfur followed by a slow volatilization of bismuth. - Temperature is a variable that has a significant influence on the decomposition and volatilization processes. Bi could be eliminated effectively only at temperatures higher than 1100 °C. - In oxygen containing atmosphere in the range 600 to 1100 °C, bismuthinite also decomposes at very fast rate to bismuth followed by oxidation to stable bismuth trioxide. - Temperature and oxygen concentration in the gas phase affect greatly both the decomposition rate and the oxidation rate.

227

- Bismuth elimination through the gas phase from BÍ2S3 at roasting or smelting temperatures in the presence of oxygen is difficult since the Bi formed by the decomposition oxidizes rapidly to stable bismuth trioxide.

Figure 10. Effect of oxygen concentration in the gas phase on the decomposition / oxidation of hkmnthinite bismuthinite at 750 °C Acknowledgements The authors would like to express their gratitude to the National Council for Scientific Research of Chile, FONDECYT, for the financial support of this investigation through Project No. 1080296. References L.Winkel, J. Wochele, C. Ludwig, I. Alxneit, M. Sturzenegger: ., "Decomposition of Copper Concentrates at High Temperatures: An Efficient Method to Remove Volatile Impurities." Minerals Engineering, Vol. 21(10) (2008) pp. 731-742. J. Steinhäuser, A. Vartiainen, and W. Wuth, "Volatilization and distribution of impurities in modern pyrometallurgical copper processing from complex concentrates". JOM, 36(1) (1984), pp. 54-61. W. G. Davenport, M. King, M. Schlesinger, A. K. Kumar, Extractive metallurgy of copper (Kidlington, Oxford, UK: Elsevier Science Press, 2002), 73-89. P. C. Chaubal, M. Nagamori, (1982), "Volatilization of Bismuth in Copper Matte Converting-computer Simulation." Met. Trans. B, vo!.13B (1982), pp 339-348.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

PYROMETALLURGICAL CONTROLS OF SILVER-RESIDUE SMELTING IN A SHORT ROTARY FURNACE Atsuhiro Nabei and Ken-ichi Yamaguchi Mitsubishi Materials Corp. Central Research Institute 15-2, Onahama-aza-fukimatsu, Iwald,FukushimaPref., Japan, 971-8101 [email protected] Keywords: Slime Smelting, PbO-SiCç slag, Short Rotary Furnace Abstract Wet chlorination process of de-copperized anode slime is followed by de-chlorination stage in which lead-sulfate and metallic silver form silver residue. The silver residue is subsequently reduced to Ag2Se matte and PbO-SiCh slag to remove lead to the slag in a short rotary furnace. Although this smelting process is proven well in actual operations, few papers have investigated its metallurgical aspects. This paper will describe the effectiveness of controlling parameters adopted at the Naoshima Smelter & Refinery and discuss the decisive factors influencing silver losstoPbO-SiCh slag. Introduction The precious metals plant at Naoshima smelter & refinery was established in 1989, and has produced electrolytic silver, refined gold and platinum-group metals (PGM) from de-copperized anode slime as by-products of electrolytic copper. The process for gold production has been innovated from conventional smelting process to hydrometallurgical process since 2004 [1], while the process for silver production has still remained the same. The process for silver production at the Naoshima smelter & refinery is illustrated schematically in Figure 1. De-copperized anode slime is, at first, concentrated in the froth flotation process, then leached in the wet-chlorination process, and converted to silver residue in the de-chlorination process. The silver residue, which is a mixture of lead-sulfate and metallic silver, is charged to a short rotary furnace for the oxidation roasting to volatilize selenium dioxide and subsequent reduction smelting. In the smelting stage, silver residue and reverts containing selenium are smelted with flux and carbonaceous materials to form PbO-Si02 slag and Ag2Se matte. Furthermore, Ag2Se matte is transferred to a rotary cupell and oxidized by air with the addition of sodium carbonate and sodium nitrate to produce silver anodes while removing impurities into litharge. Silver anodes are electrolytically refined by the Moebius process. Electrolytic silver is washed, dried and,finally,cast into 30kg-bars.

229

Figure 1. Schematic processflowof silver production at Naoshima smelter & refinery. Although the smelting process above is proven well in actual operations, few papers have investigated on its metallurgical aspects. Barbante et al. [2] have described actual smelting of de-copperized anode slime and discussed the smelting reactions thermodynamically. Ranldn et al. [3] have investigated on the thermodynamics of Ag-Cu-Se matte. These papers have concerned mainly on the converting of selenide matte to Dore metal. In contrast, no paper has discussed on silver loss to slag in the smelting stage. The present paper aims to discuss on which operational parameters are effective for reducing the silver loss to slag in the smelting stage. From an analogy with copper losses in copper smelting process [4], silver loss to slag is classified into two categories; chemical dissolution of silver to molten slag, and mechanical suspension of matte droplets in the slag phase. The latter loss becomes quite severe when the settling of matte droplets is retarded due to high viscosity of homogeneous slag or high apparent viscosity of heterogeneous slag. Accordingly, the predominant mode of silver loss to actual slag will be identified by the examination offrozenslag, utilizing the electron-probe micro-analyzer (EPMA). Such silver loss will also be discussed utilizing metallurgical fundamentals tofindout better slag controls which can be followed in the actual operation.

230

Procedure of the Present Investigation While tapping of molten slag out of the short rotaryfiimace,slag sample was quenched on a steel plate at the tip of the slag launder. The slag samples were crushed to small pieces and molded with resin to polish a surrace. Their mineralogical texture was examined by EPMA. The slag samples were also analyzed chemically.

Results Microstructure of the Slag Typical microstructure of frozen slag is demonstrated in Figure 2. Figure 2 (a) shows backscattered-electron image of frozen slag, and Figure 2 (b) to (d) show specific X-ray images of Ag, Si, and Se, respectively, in the same area of Figure 2 (a).

(c) Si distribution

(d) Se distribution

Figure 2. Typical microstructure of frozen slag sampledfromthe short rotaryfiimaceshowing (1) silicate matrix, (2)Ag2Se matte, (3) lead-sulfate, and (4) magnetite.

231

It is seen from the Figure 2 that: - The frozen slag comprises glassy matrix of silicate, particles of Ag2Se matte, agglomerates of lead-sulfete crystals, primary crystals of magnetite, - Silver is concentrated in Ag2Se matte and, in contrast, distributed scarcely in silicate matrix, lead sulfate and magnetite. It was also observedfromphotographs not demonstrated in the present papers that: - AgzSe matte droplet is occasionally included in thefrozenparticles of sodium-sulfate, - Silver is not concentrated in sodium-sulfate. From theresultsabove, it is noted that silver loss to slag is due to the mechanical suspension ofAgîSe matte droplet and the effect of chemical dissolution of silver on silver loss is negligible. Discussion on the Mechanical Suspension of A&Se Matte Droplets As silver loss is due to mechanical suspension of Ag2Se matte in the slag phase, the mechanism of the matte suspension will be discussed below. Plant data will be analyzed on a metallurgical basis focusing on the viscosity and apparent viscosity of the smelting slag. Viscosity of Molten Slag on Silver Loss Numerous studies have been conducted to measure or estimate the viscosity of silicate melts and the comprehensive information is available for the systems of PbO-SiC>2fromprevious literatures [5 - 8]. The viscosity of homogeneous slag can be expressed by the statistical thermodynamic model proposed by Weymann [5]. The temperature dependence of the viscosity is given by the equation (1), where /* is viscosity of molten slag and T is the absolute temperature in K. A and B are dependant solely on the chemistry of the melt and have been quantified for PbO-SiCh slag by Urbain et al. [6].

(1)

^=ATexp\-^^j

Figure 3 shows the relation between the viscosity estimated by equation (1) and the composition of PbO-SiCh slag at 1000 °C. The dashed line expresses the boundary of SÍO2 saturation. As the SiQ> content increases, the viscosity of PbO-SiQi slag increases steeply in logarithmic scale, especially when slag composition is close to the boundary of silica saturation. These behaviors of PbOSiC>2 slag suggest that with increasing silica content of PbO-SiCh slag, the terminal velocity of matte droplets falling in the slag layer decreases and, thus, the settlingtimeof matte droplets increases greatly. It is therefore noticed that an excessive addition of siliceous materials may increase silver loss to slag due to the retardation of matte settling. Primary measure for reducing silver loss should be the reduction of silica content of the slag.

232

Figure 3. Relationship between the viscosity of PbO-SiCh slag and the SÍO2 content at 1000 °C. Since actual slag contains minor oxides, direct comparison between actual slag and PbO-Si02 slag is of no reason. Thus, the Pb/SiQj ratio is used in the plant practice so that the chemistry of the actual slag is projected on PbO-SiCh binary system. The influence of the Pb/SiQ2 ratio on sflver loss to the actual slag is shown in Figure 4. Y-axis shows standardized silver loss to slag, with the annual average value equal to 1, and the dashed line expresses the boundary of SÍO2 saturation. It is definite from the Figure 4 that silver loss deceases with increasing the PTVSÍO2 ratio as noticedfromthe behavior of the PbO-Si02 system. The approximation using the Pb/Si02 ratio may be effective when the impact of minor oxides on the slag properties is negligible.

Figure 4. The influence ofthe Pb/SiCh ratio on silver loss to slag.

233

Effect of Solid Particles on the Apparent Viscosity of Slag Roscoe el al. [9] has shown the effect ofsolid particles on the viscosity of liquid by equation (2), V = Mo(\-l35Vy2i

(2)

where p. is the apparent viscosity of heterogeneous system in which spherical particles are suspended in a liquid phase, and uo is the viscosity of the liquid Fis the volume fraction of solid phase in the system. It is seen from equation (2) that solid suspension increases the apparent viscosity even if the viscosity of the liquid is kept constant. Therefore, heterogeneous slag including solid particles may retard the settling of matte droplets leading to a high level of silver loss to the slag. The actual slag includes various crystals, such as magnetite and lead-sulfate. Among these crystals, magnetite is precipitated easily from the slag because the oxygen potential of the melts in the smelting stage is much higher than that of the FeO/Fe304 equilibrium. Accordingly, an excessive charging of iron-bearing materials may increase the apparent viscosity of the smelting slag Figure 5 shows the relation between the silver loss and the magnetite content in the actual slag estimated from the chemical analysis of iron in the slag. Y-axis is the same as in Figure 4. It is evident from Figure 5 that magnetite crystals included in the slag enhance the apparent viscosity, resulting in the gradual increase of silver loss to the slag.

Figure 5. Relation between silver loss and magnetite content in the actual slag.

234

Effect of Liquid Emulsion on the Apparent Viscosity of Slag The apparent viscosity of a emulsive liquid in which small droplets of one liquid are suspended in the other was modeled by Einstein [10,11]. The apparent viscosity (Uej) is expressed by the equation (3) where c is the volumetricfractionofthe droplets.

It is seen from equation (3) that the more a liquid is emulsified, the more the liquid becomes viscous. Accordingly, heterogeneous slag emulsified with insoluble liquids may also retards the settling of matte droplets leading to a high level of silver loss to the slag It is supposed that droplets of sodium sulfate may enhance the apparent viscosity of the actual slag, resulting in an increase of silver loss to the slag. However, the operational data have notrevealedpositive correlations between sodium and silver content of the actual slag clearly. It may be because silver loss is influenced by lead-sulfate and sodium-sulfate simultaneously and the content of both species could not be quantified by chemical analysis so far. Concluding Remarks The present paper has investigated the texture of frozen slag sampled from the short rotary furnace for the smelting of selenide matte at Naoshima smelter & refinery. The results have demonstrated that silva loss to the actual slag is caused by the suspension of selenide matte and primary measure for reducing the silver loss should be the control of the Pb/SiQ2 ratio ofthe slag. Although much information regarding the smelting of selenide matte has been collected in Naoshima smelter &refinery,definite correlations have not beenrevealedin many cases. In order to overcome the difficulties in analysing the plant data, the authors have been conducting fundamental experiments tofillthe gap between the theory and practice of the silver loss.

References 1. K, Komori, et al, "Hydrometallurgical Process of Precious Metals in Naoshima Smelter & Refinery," Copper2010,(Clausthal-Zellerfeld, Germany; GDMB, 2010), 1403-1411. 2. G G Barbante, D. R. Swinboume and W. J. Rankin, "Pyrometallurgical Treatment of Copper Tank House Slimes", Pyrometallurgy for Complex Materials and Wastes, M. Nilmani, T. Lehner and W. J. Rankin, Eds. (Warrendale, PA; TMS, 1994), 319-337. 3. W. J. Rankin, G G Barbante and D. R Swinboume, "Laboratory-scale Smelting of Copper-anode Slimes" Trans. InstnMin. Metall. 104(1995), C59-69

235

4. R. W. Ruddle, The Physical Chemistry of Copper Smelting, (London, UK; MM, 1953). 5. H. D. Weymann, "On the Hole Theory of Viscosity, Compressibility, and Expansivity of Liquids", Kollid. Z. Polymers, 181(1962), 131-137. 6. G Urbain and M. Boiret, "Viscosities of Liquid Silicates", Ironmaking and Steelmaking, 17(1990), 255-260. 7. Suresh K. Gupta, "Viscosity of PbO-SiOi Melts", Metallurgical and Materials Transitions B, 26B (1995), 281-287. 8. R. G Reddy and J. Y Yen, "Effect of Solid Particles on the Viscosity of Slags", Extractive Metallurgy of Copper, Nickel and Cobalt, Vol. /, Ed. R. G Reddy and R. N. Weizenbach (Warrendale, PA; TMS, 1993), 309-323. 9. Roscoe et al., "The Viscosity of Suspensions of Rigid Spheres", British Journal of Applied Physics, 3(1952), 267-269. 10. A. Einstein, "Eine neue Bestimmung der Moleküldimeiisionen"/l«/i Phys., 19(1906), 289-306 H.A. Einstein, "Berichtigung zu meiner Arbeit: Eine neue Bestimmung der Moleküldimensionen", Ann Phys., 34(1911), 591-592

236

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann- Yang Hwang, Jaroslav Drelich, Jerome Downey, Tao Jiang, and Mark Cooks TMS (The Minerals, Metals & Materials Society), 2011

A STUDY OF PEIXETIZATION OF MANGANESE ORE FINES Deqing ZHU1, Vinicius Mendes 12, Tiejun CHUN1, Jian PAN1 'School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, Hunan, P.R. China 2 Vale, Rio de Janeiro 20030, Brazil Keywords: manganese ore fines; high press grinding roller; preheated pellet; fired pellet Abstract Fine manganese fines are characterized by large amount and low price compared to lump manganese ores. However, fine manganese must be agglomerated as EAF (electric arc furnace) feed. In this paper, high quality fired pellets were produced by simulating grate-kiln process to supply the feed for EAF from some imported manganese ore fines containing high combined water. Manganese ore fines assaying 44.47wt%Mn, 5.89wt%Fe, 12.09%LOI and with 72.7wt% ranging between 0.5mm and 8mm, are pretreated up to the particle size of 84.7wt% passing 0.074mm by HPGR (high press grinding roller) twice in open circuit. The finely ground manganese fines were mixed with 1.2wt% bentonite, balled at 15% moisture for 7min in disc pelletizer. The drop numbers, compressive strength and thermal shock temperature of green balls reach 11 times/0.5m, 11 N/pellet and 360°C, respectively. After the green balls drying in the oven, the dry pellets were preheated at 1050°C for 12 min and then fired at 1300°C for 12 min in electric tube furnace. The compressive strength of fired pellets reaches 2621 N/pellet. Therefore, these pellets are high quality burden for blast furnace or EAF due to high mechanical strength and high Mn grade of 49.11 wt% and low detrimental elements. Introduction Manganese, mainly using as deoxidizing agent, desulfurization agent and alloying elements, is widely used in the steel company. With the rapid development of iron and steel industry in China, domestic manganese ores cannot meet the requirement, resulting large quantities manganese ores import to China [1-4]. Compared to lump manganese ores, fine manganese ores containing high combined water are characterized by large amount and low price. However, they must be agglomerated as EAF (electric arc furnace) feed. There are three ways to agglomerate the manganese fine ores, sintering, briquetting and pelletizing [5-6]. Compared to the other two methods, pelletizing technology possesses the following superiorities, such as uniform size distribution of product, stabilizing chemical composition of product and lower energy consumption in producing. Therefore, pelletizing is a superior way to treat the fine manganese ores [7]. The

237

systematic research of producing manganese oxide pellets by fine manganese ores containing high combined water was studied in this paper. Experimental Raw Materials The raw materials include the manganese ore fines, bentonite, all of which were imported from Brazil. Their chemical composition of materials is shown in Table 1. The manganese grade of ROM (run-of-mine) manganese ores is 44.47wt%Mn and the LOI is high to 12.09%. In the meantime, most of manganese minerals are manganese oxide as shown in Table 2. The size of ROM manganese ores and the pretreated manganese ores by HPGR are illustrated in Tables 3 and 4. The size of manganese is about 84.7wt% passing 0.074mm after ROM pretreated by HPGR. The pretreated manganese fines possess the static ballability index and specific surface areas of 0.60 and 2764cm2/g, respectively. The quality of bentonite is excellent due to higher content of 65.6wt% montomorillonite, higher swelling volume of 13mL/g and water adsorption of 405wt% and particle size of 95.2wt% passing 0.074 mm. Table 1. Chemical composition of materials (wt%). P Mn,oai Mn0 2 Fe,0,ai Si0 2 A1203 CaO

Types Manganese ores bentonite

Items Distribution Content

Size/mm Content

S

LOI

44.47

68.22

5.89

5.66

9.09

0.28

0.097

0.035

12.09

-

-

9.10

50.97

17.32

2.89

0.053

0.035

11.91

Table 2. Composition of Mn containing minerals (wt%). Mn Oxide Mn Silicate Mn Carbonate 96.94 2.34 0.72 1.04 43.11 0.32

Mn,otal 100 44.47

Table 3. Size distribution of manganese ore fines (wt%). 0.5-1 0.3-0.5 0.1-0.3 >8.0 5-8 3-5 1-3 8.50 8.33 5.14 13.56 11.98 25.9 21.23

<0.1 5.36

Table 4. Size distribution of manganese ore fines pretreated by HPGR (wt%). Size/mm >0.3 <0.038 0.3-0.1 0.074-0.1 0.044-0.074 0.038-0.044 Content 3.44 5.38 6.44 19.11 6.03 59.6

238

Experimental Procedure The flowsheet of pelletizing tests in this paper include the traditional section, such as manganese pretreated by HPGR, mixing pretreated manganese fines and bentonite at a given ratio, balling in a disc pelletiser and firing in a tube furnace to make fired pellets. Manganese ores were pretreated by the HPGR of 250 mm in diameter and 120mm width at 300rpm and passed twice in open circuit. The manganese fines at 7% moisture were pretreated by HPGR at a normal feed rate of 30 kg/min, and then mixed with bentonite in a kneader-mixer at 350 rpm with a 150 rpm orbital motion for 5 min. Green pellets were made from the mix of pretreated manganese fines and bentonite in a disc pelletizer of 0.8m in diameter and 0.2 m rim depth, rotational speed at 38 rpm and inclined at 47° to the horizontal. The drop number and compressive strength of green balls were measured to evaluate the abilities of the green balls to remain intact and retain their shape during handling, respectively. Dry pellets were preheated and roasted in a tube furnace of 50mm in diameter and 600mm width. The compressive strength was measured to evaluate the quality of the roasted balls.

Balling

Results and Discussions

Bentonite Dosage. The effects of bentonite dosage on the properties of green balls are shown in Figure 1. The drop numbers and compressive strength of green balls improve markedly with the bentonite dosage adding from 0.9wt% to 1.8wt%. All the properties of green balls can meet the requirement of over 4 times/0.5m when bentonite dosage is at 1.2wt%, which means this manganese ore fines have better balability, agreeing well with that of static ballbility index. Therefore, the drop numbers of 6.5 times/0.5m and compressive strength of 14 N/pellet of green balls were obtained at 1.2wt% bentonite dosage, balling at 13wt% moisture for 9 min. Balling Moisture. As shown in Figure 2, drop numbers of green balls climb significantly with increasing balling moisture. However, compressive strength of green balls appears to adverse trend and drops significantly when balling moisture is elevated. The optimum moisture takes place at 15wt% moisture, drop numbers of 16 times/0.5m and compressive strength of 13 N/pellet being achieved. The balling moisture of manganese is higher than that of iron ore concentrate (8-10wt%) due to the porosity on the surface of manganese.

239

Figure 1. Effects of bentonite dosage on the properties of green balls (balling at 13% moisture and for 9 min).

Figure 2. Effects of balling moisture on the properties of green balls (1.2wt% bentonite dosage and balling for 9 min). Balling Time.The effects of balling time on properties of green balls is presented in Figure 3. The drop numbers climbs dramatically and the compressive strength of green balls increases slightly when prolong balling time. The drop numbers of 11 times/0.5m, compressive strength of 11 N/pellet and thermal shock temperature of 360°C were achieved when balling for 7 min, which can meet the requirement in the pelletizing plants. Therefore, the suitable balling time is 7min, much shorter than that of 10-13 min in producing iron ore concentrates pellets.

240

Figure 3. Effects of balling time on the properties of green balls (1.2wt% bentonite dosage and at 15wt% moisture). Preheating and Roasting of Pellets Green balls are made under the following optimum conditions: 1.2wt% bentonite and balling at 15wt% moisture for 7 min. Then the green balls were dried in the oven at 105°C for 2h and preheated and fired in tube furnace. Preheating Temperature. Figure 4 shows the effect of preheating temperature on the compressive strength of preheated pellets. With an increase in preheating temperatures from 970°C to 1090°C, the compressive strength of preheated pellets climbs from 317 N/pellet to 741 N/pellet. The compressive strength of 600 N/pellet was achieved when preheating at 1050°C, which meets the requirement of 500 N/pellet in pelletizing plants. Preheating Time. Figure 5 demonstrates the effect of preheating time on the compressive strength of preheated pellets. With extending preheating time the compressive strength of preheated pellets increases obviously and reaches the peak value of 701 N/pellet when preheating for 12 min. However, the compressive strength declines slightly when preheating time is beyond 12 min. Enough preheating time is the requirement not only for many reactions to finish but also for the preliminarily bond to form. Therefore, the suitable preheating conditions are preheating at 1050°C for 12 min.

241

1000

1040

1080

Preheating temperatureTC

Figure 4. Effect of preheating temperature on the compressive strength of preheated pellets (preheating for 10 min).

ÉÍ A

'S eso 2" ïj 1 640 ■a

.S s

600

A

0 A

"

560

8

10

12

14

Preheating duration/min

Figure 5. Effect of preheating duration on the compressive strength of preheated pellets (preheating at 1050°C). Firing Temperature. The compressive strength of fired pellets increases significantly from 1490 to 3314 N/pellet when firing temperature is elevated from 1260 to 1320°C in Figure 6. The reason is that oxidation reaction is easy to occur in the pellet at high temperature, resulting Fe 2 0 3 and Mn0 2 recrystallization, leading to higher compressive strength of fired pellets. The target compressive strength of fired pellets at 2621 N/pellet was achieved at 1300°C, which meet the requirement of 2500 N/fired pellet in pelletizing plant.

242

1280

1300

Firing temperatureA;

Figure 6. Effect of firing temperature on the compressive strength of fired pellet (preheating at 1050°C for 12 min and firing for 15 min).

9

12

Firing timefrrin

Figure 7. Effect of firing duration on the compressive strength of fired pellet (preheating at 1050°C for 12 min and firing at 1300°C). Firing Time. Figure 7 illustrates the effect of firing time on the compressive strength of fired pellets. With an extension of firing time from 6 to 12 min, the compressive strength of fired pellets increases dramatically from 1984 to 2621 N/pellet, and remain unaffected when firing duration is above 12 min. Hence, the firing conditions are suggested at 1300°C for 12 min. It can be concluded that the manganese pellets possess good firing performance.

243

The chemical composition of fired pellets can be found in Table 5, and the results show that final product possess high Mn content of 49.11 wt% , and low detrimental elements, which can be used excellent feed for blast furnace or EAR

Element Content

Table 5. Chemical compositions of final product (wt%). Mn,oai Fetotal Si0 2 A1203 CaO MgO K 2 0 Na 2 0 49.11 6.48 7.75 10.98 0.62 0.21 1.36 0.097

P 0.11

S 0.010

Conclusions (1) The imported manganese fine ores must be pretreated before pelletizing due to coarse grain. The particle size of 84.7% passes 0.074mm and the specific surface areas of 2764 cm2/g were achieved after pretreated by HPGR. The drop numbers, compressive strength and thermal shock temperature of green balls are 11 times/0.5m, 11 N/pellet and 360°C, respectively, under the following conditions: mixing pretreated manganese fines with 1.2% bentonite, balling at 15% moisture for 7min in disc pelletizer, (2) The final manganese pellets, assaying 49.11wt%Mn and the compressive strength of 2621 N/pellet of fired pellets, were obtained when the green balls were preheated at 1050°C for 12 min and then fired at 1300°C for 12 min in electric tube furnace, which can be used the excellent burden for blast furnace or EAR References 1. Luo. H, Peng. Z. J, "Experimental research on cold bonding briquette of Mn Si alloy powde," Multipurpose Utilization of Mineral Resource, 10 (5) (2006), 9-10. 2. Zhu. D.Q, Guo. Y. P, Fu. S. C, "Cold bonding briquette of manganese oxide with adding high-efficiency binder," China's Manganese Industry, 13(3) (1995), 43. 3. Ding. Y. H, Chen. X.Y, Wang. X, "Sintering process experiment in Yunan jianshui manganese mine," Yunan Metallurgy, 33(1) (2004), 18-19. 4. Xing. F, Gao. D.Y, Zhang. Y L, "The research on cold bonding briquette of manganese fines," China's Manganese Industry, 25(3) (2007), 24-25. 5. Yuan. M. L, Mei. X. G, Zhuang. J. M, "Research on rhodochrosite pelletizing," Sintering and Pelletizing, 21(5) (1996), 9-10. 6. Zhu. D. Q, et al., "Mechanization activation of magnetic concentrate for improving its pelletability by high press grinding roller," 1SIJ International, 44(2) (2004), 310-312. 7. Zhu. D.Q, Xu D. L, Pan. J, "Study on high properties pellets produced by manganese ore fines containing high combined water," Iron and Steel, 44(5) (2009), 11-13.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

REDUCTION OF CARBON-BURDENED CHROMITE PELLETS IN THE PRESENCE OF ADDITIVES Guanghui LI, Jianchen LI, Mingjun RAO, Guohua BAI, Tao JIANG1 School of Minerals Processing & Bioengineering, Central South University, Changsha, Hunan 410083, China Key Words: chromite, solid-state reduction, additives, metallization Abstract By using thermogravimetry, optical microscope and chemical phase analysis, the solid-state reduction of carbon-burdened chromite pellets in the absence and presence of additives were investigated. The effects of reduction temperature, reduction time and additives on the metallization of chromite pellet were studied. The results indicate that the pre-reduction ratio of carbon-burdened chromite pellets is improved with the increase of reduction temperature and the addition of additives. The metallization of chromium is increased from 63.9% to 92.8% at 1300 °C for 4h with the addition of 2 wt.% of sodium borate. Investigation suggests that the reduction of carbon-burdened chromite pellets is diffusion controlled. The apparent activation energy without additive is 192.1 kJ/mol, while it decreases to 170.4 kJ/mol in the present sodium borate. Introduction Ferrochrome is used to produce high strength, corrosion resistant, wear-resisting, high temperature resistant and oxidation resistant special steels, and it is becoming one of the most important alloy materials for the production of stainless and high alloying ferritic steels [1-5]. Demand for ferrochrome is steadily on the rise due to the growth of stainless steel industry. Nowadays, more than 80% of the world's ferrochrome is utilized in the production of stainless steel [6-7]. Ferrochrome is mainly produced in submerged-arc furnaces (SAF) with chromite lump ore. However, lump ore merely accounts for 20% of the chromite ore output, while fine ore (< 8 mm) accounts for about 80% [8]. Therefore, agglomeration of chromite fine ore is necessary before charging into SAF. In order to save energy, carbon-burdened chromite pellets are usually reduced in a rotary kiln before charging into the SAF [9-12]. However, it is still of high energy consumption to obtain a high metallization ratio of reduced pellets due to the requirement of high reducing temperature and a long period. 1

Corresponding Author: Dr. Tao Jiang, Email: [email protected]. Tel: +86-731-88877656

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In this paper, additives were investigated to enhance the reduction of carbon-burdened chromite ore pellets. The effects of reducing temperature, reducing time and additives were investigated, and the reduction kinetics in the presence of sodium borate was also analyzed. Experimental Materials The chromite ore was ground to 81.44% passing 0.074 mm beforehand (Table 1), and its chemical composition is presented in Table 2, and the chromite sample contains 43.03% Cr203. Mineral composition (Fig.l) obtained by XRD indicates that the sample is composed of (Mg, Fe)0-(Cr, A1)203, (Mg, Fe)Cr204, Fe304 and Fe0( Cr, A1)203. Anthracite, 90% passing 0.074mm, is used as reducing agent, and the chemical composition and industrial analysis is presented in Table 3.

Figure 1. XRD pattern of the chromite sample Table 1. Particle size of chromitefinessample +0.074 18.56

Size fraction/mm Content/wt %

TFe 20.04 V205 0.37

0.074-0.045 20.49

-0.045 60.95

Table 2. Chemical composition of chromite sample (wt%)

FeO 17.46 K20 0.012

Cr 2 0 3 43.03 Na 2 0 0.045

MgO 9.15 Co 0.01

Si0 2 4.03 Ni 0.027

LOI: loss on ignition

246

A1203 13.91 P 0.002

CaO 1.06 S 0.01

Ti0 2 0.76 LOI 0.56

MnO 0.086

Table 3. Industrial analysis result of reduction coal (wt%) Main chemical composition of ash Chemical properties /% Ad Mad CaO MgO A1 2 0 3 Si0 2 FCad Fe 2 0 3 Vdaf Stotal 74.11 13.51 10.84 1.54 2.02 5.75 0.53 0.68 0.07 4.46 Mad: Moisture; Ad: Ash content; Vdaf: Volatile content; FCad: Fixed carbon content Methods Reducirte roastins The ground anthracite and ground chromite ore was blend at C/O molar ratio of 1.2 (C/O refers the ratio of the fixed carbon amount contained in coal to the oxygen amount in iron and chromium oxides), and the mixture was made into 10-12 mm balls in diameter in a disc pelletizer. The green balls were dried statically at 110°C for 2h in electrically-heated oven . The dried pellets were loaded into the corundum crucible and roasted at a desired temperature for a given period. After cooled to ambient temperature with 2 L/min argon gas, the reduced pellets were crushed and sampled for chemical analysis. Chromite Grinding Additive-

1 Balling L

Anthracite

Reduction roasting

Ï

Pre-reduced pellets fSAF burden ")

Figure 2. Experimental flowsheet of reduction of chromite pellet The experimental flowsheet is shown in Fig. 2, and the metallization ratio of iron, chromium and reduced pellet were calculated as follows: ^Fe = - ^ X l 0 0 % _MCr *7cr =

xl00%

247

(MFe+MCr)xl00% (TFe+TCr) where Mpe is the metallic iron content, %; Mcr is the metallic chromium content, %; 7>„ is the total iron content, %; 7b- is the total chromium content, %;

TJFS:

is the iron

metallization ratio, %; ^Cris the chromium metallization ratio, % and //Peis the pellet metallization ratio, %. Kinetics The isothermal thermogravimetry was applied to study the reduction kinetics of carbon-burdened chromite pellets. The weight change of the pellet during the experiment was measured every 10 minute. Results and Discussion Effects of Reducing Time and Temperature Fig. 3 shows the results of reduction of pellet as the function of reducing time and temperature. The metallization ratio of metals is improved with the increase of reduction temperature firstly; when temperature exceeds 1300°C metallization keeps approximately a constant because the formation of close-grained metallic layer covers over the unreacted core, which hinders the diffusion of reducing gas into the pellet. Moreover, it can be found from Fig. 3 that the reduction of iron oxide is easier than that of chromium oxide. The metallization ratio of iron is 84.9% when reduced at 1200°C for 4h, which is much bigger than that of chromium (58.2%). The microstructure of reduced pellets confirms the marked effect of reducing temperature (Fig. 4). When reducing temperature increases from 1150°C to 1400°C more ferrochrome grains occur, and the size of metallic grains grow obviously .

248

Figure 3. Effects of reducing time and temperature on the metallization of chromium ore pellets

249

Figure 4. Micrographs of pellets reduced at different temperatures for 4h (a) 1150°C, (b)1200°C, (c) 1250 °C, (d) 1300°C, (e) 1350 °C, (f) 1400°C. Effects of Additives on Metallization of Chromium Additives including sodium borate (Na2B4(V10H2O), sodium sulfate (T^SCv), sodium hydroxide (NaOH), sodium carbonate (Na2C03), sodium chloride (NaCl) and sodium nitrate (NaNOs) are mixed into the carbon-burdened chromite ore pellets in order to enhance the reduction of chromite ore. The Na2B4O7*10H2O dosage is fixed at 2.0 wt% (mass fraction to chromite ore), and based on the same level of sodium ions, dosage of other additives is calculated as follows: 0.74 wt% for Na2SC>4, 0.42 wt% for NaOH, 0.55 wt% for Na2C03, 0.6 wt% for NaCl and 0.89 wt% for NaN0 3 . The reduction results at varied temperature for 4h are shown in Fig. 5. All of the additives are able to improve the reduction of chromite, among them sodium borate performs best. The metallization of chromium is increased from 58.2% to 74.9% at 1200 °C, and from 63.9% to 92.8% at 1300 °C with the addition of 2.0 wt% of sodium borate.

250

i

Without NaOH Na£03 Nad

NaN03 N^SO, Naßfi^OHp

Figure 5. Effects of additives on metallization of chromium (reducing time: 4h) Kinetics The isothermal thermogravimetry was applied to study the reduction kinetics of carbon-burdened chromite pellets. The experimental data are plotted according to gas-solid reaction kinetic equations. It was found that Ginstling and Brundshtein equation for diffusion-controlled kinetics fits the reduction of carbon-burdened chromite pellets well (Fig. 6):

l-f/-(l-/)2/3=*/ where/is the reducing weight loss ratio, %; k is the rate constant and t is the time.

imin Figure 6. Plots of [\-2ßf-(l-ßm]

vs /

(1/7riO"5 Figure 7. Arrhenius plots, In (k) vs \IT

Fig. 6 suggests that the reduction of carbon-burdened chromite pellet is controlled by diffusion of reductant through the product layer. After \nk is plotted against l/T(Fig. 7), apparent activation energy has been obtained through Arrhenius equation. The apparent activation energy of the reaction is found to be 192.1 kJ/mol without sodium 251

borate and 170.4 kJ/mol in the presence of sodium borate. Conclusions 1) The reduction of carbon-burdened chromite pellets is improved in the presence of sodium salt additives, and sodium borate performs best. The metallization ratio of chromium is increased from 63.9% to 92.8% at 1300 °C for 4h with the addition of 2.0 wt% of sodium borate. 2) The apparent activation energy of the reduction of carbon-burdened chromite pellets is decreased from 192.1 kJ/mol without additives to 170.4 kJ/mol in the presence of sodium borate. Acknowledgement The authors want to express their appreciation to National Science Fund for Distinguished Young Scholars (50725416) for financial support of this research. References [I] P. Weber, R. H. Eric. "The reduction of chromite in the presence of silica flux", Minerals Engineering, 2006, 19(3): 318-324. [2] C. Takano, A. P. Zambrano, A. E. A. Nogueir, et al. "Chromites reduction reaction mechanisms in carbon-chromites composite agglomerates at 1 773 K", ISIJ International, 2007,47(11): 1585-1589. [3] Y. D. Yang, I. D. Sommerville, R. F. Johnston, A. Mclean. "Use of solid state carbothermic reduction in production of transition metals and their carbides", J. Iron & Steel Res., 2000, 7(2):15-22. [4] J. F. Yan, J. X. Chen, L. Hu. "Chromium metallurgy", Beijing; Metallurgical Industry Press, 2007: 119-120. [5] V. Singh, S. M. Rao. "Study the effect of chromite ore properties on pelletisation process", Int.J.Miner.Process, 2008, 88(1/2):13—17. [6] Courtesy Dow Jones Newswire. "Ferrochrome Facts 2007", http://www.estainless steel.com/ferrochrome.Shtml. [7] J. Chen, J. Zhao, M. Zhang. "Carburization of ferrochromium metals in chromium ore fines containing coal during voluminal reduction by microwave heating", Journal of Central South University of Technology, 2009, 16(l):43^t8. [8] D. Q. Zhu, J. Li, J. Pan, A. P. He. "Sintering behaviours of chromite fines and the consolidation mechanism", Int.J.Miner.Process, 2008, 86 (l/4):58-67. [9] O. Soykan, R. H. Eric, R. P. King. "Reduction mechanism of a natural chromite at 1416 °C", Metallurgical transactions. B, 1991, 22B(1): 53-63. [10] A. Lekatou, R. D. Walker. "Solid state reduction of chromite concentrate: melting of prereduced chromite", Ironmaking and Steelmaking, 1995, 22(5): 378-392. [II] A. Lekatou, R. D. Walker. "Mechanism of solid state reduction of chromite concentrate", Ironmaking and Steelmaking, 1995, 22(5): 393-404. [12] A. B. Hazar-yoruc. "Reduction mechanism of chromite spinel with carbon", 252

Minerals & Metallurgical Processing, 2007, 24(2): 115-120.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

Production of Srontium Metal from Strontium Oxide Using Vacuum Aluminothermic Reduction

Yeliz DEMIRAYa, Onuralp YUCELb Istanbul Technical University, Department of Metallurgical and Materials Eng. 34469 Maslak, Istanbul, Turkey a

[email protected],b [email protected]

Keywords: Strontium, celestite, metallothermic reduction Abstract Production of strontium metal from its oxide was studied under the pressure of 1-5 mbar by methalothermic process. SrO with 99 % purity was used in the experiments. Effects of Al powder addition (100, 200, 300 % of stoichiometric ratio) and time were investigated on recovering of metallic strontium. Effects of BaO addition (100 %, 200 %, 300 %) was also investigated. The final residues were examined for their chemical composition. XRD, AAS and Flame Photometer were used for chemical analysis. More than 96 % of strontium metal recovery can be obtained. Introduction Strontium is an element which is used in many advanced technology applications. Barium strontium titanat, strontium bismuth titanat, and strontium bismuth tantalat thin films are promising materials in ferroelectric and Schottky-based microelectronics technologies, especially for memory applications. Physical vapor deposition (PVD) techniques as magnetron sputtering, thermal evaporation and molecular beam epitaxy (MBE) are the most widely used methods for growth of these thin films [1-3]. In mese techniques high-purity of Sr metal is used as evaporation sources and sputtering targets. Furthermore, metallic strontium is used as a "getter" in electron tubes, and as an alloying elements in aluminum alloys [4]. The most used Sr compound is S1CO3 and widespread process used in production of SrCCh is the hydrometallurgical carbonation of celestite. Although there are considerable amount of Selestite reserves, SrC0 3 and strontinum metal are not being produced in Turkey, yet [5]. Concentrated celestite (SrSC>4) is conversed to SrCC>3 in carbonate media. Black ash and direct conversion processes are currently used methods. There are many satisfactory studies about hydrometallurgical carbonation of celestite and decomposition of S1CO3 [6, 7]. Since strontium metal is produced from it's oxide (SrO), S1CO3 needs to be decomposed to SrO. Strontium is produced only by the thermal method for industrial use. Timminco Metals, Ontario, Canada used a method similar to the Pidgeon Process in the production of strontium. However, complete details on the actual strontium production route are not determined clearly [8], In this study, production of strontium metal from its oxide was studied. The present work aims to investigate the parameters effecting the aluminothermic reduction of SrO using aluminum metal.

255

In the experiments, charge composition, amount of reducing metal and time were taken as variables in order to obtain high recovery efficiencies. Theoretical Investigations AG°-T diagram for oxide formation indicates that SrO reduction with aluminum is only possible beyond 2000 °C. In order to obtain strontium metal at lower temperatures for Eq. 1, partial pressure of the strontium vapour and/or activity of Al should be lowered. For this aim, vacuum technique and additions of BaO for formation of a basic slag (BaO.AhCh) are needed. 3 SrO + 2A1 + BaO = 3 Sr + BaO.A1203

(Eq. 1 )

Using the FactSage 5.2 thermodynamic program, Figure 1 was simulated for aluminothermic reduction of SrO by Al powder at different temperatures under 1 bar and 1 mbar of vacuum atmosphere [9]. As can be seen in the figure, reduction of strontium oxide starts at around 1050 °C and it totally reduced to metallic Sr under pressure of 1 mbar.

Fig. 1 : AG - T diagram of AI2O3 and SrO under pressure of 1 bar and 1 mbar. Experimental Raw Materials In the experiments, synthetic SrO is used as strontium source. Al powder was used as reducing agent and BaO was used for slag formation. Reduction Technique The schematic diagram of the experimental set up was illustrated in Fig. 2. The temperature was measured by 6RhPt-30RhPt (EL-18) thermocouple. The reduction process was carried out in a

256

retort made from Incoloy 800H/HT alloy [10]. In order to pump the retort for vacuum, a two stage compact rotary vane pump was used. The stoichiometric ratios of Al and BaO were together selected 100 %, 200 % and 300 %. The desired weight ratios of SrO, BaO and Al metal powder were mixed throughly. The mixture was weighted and bnquetted. Briquettes were put in an alumina boat. The charged boat was inserted into the retort at room temperature. After closing the retort cover, the inside pressure was decreased to 1-5 mbar. In order to condensate the strontium vapour formed during the reduction reactions, a cooling copper tube was mounted at the retort cover. After the furnace attained to the required temperature, the retort was inserted into the furnace. Since the furnace temperature decreases after inserting the retort, initial time was started when the furnace reached the desired temperature. At the end of the reduction experiments, the retort was left in the furnace at the same vacuum values and was cooled to room temperature. Then, the cover was opened and the condensed strontium metal on the cooling section and the residue left in the boat were weighted and analyzed chemically. The degree of Sr metal recovery was calculated as Sr Recovery, % = 100 - [(Wt x % Sr,)/(W0 x % Sr0)] x 100 where Wt the weight of the residue at time t, % Srt the weight percent of the strontium in the residue at time i, Wo is the weight of the briquette, and % Sro the weight percentage of the strontium of the briquette.

Fig. 3. Schematic diagram of experimental setup. 1- Furnace, 2- retort made from Incoloy 800H/HT alloy, 3-Briquetted charge, 4 Sr condensation section, 5-Cooling water in and out, 6Vacuum connection, 7-Vacuum pump, 8-Digital pressure gauge, 9- EL 18 thermocouple, 10-Ice water box, 11-Temperature measuring unit. Results and Discussion The effects of time and the stoichiometric ratio of reducing agent on recovery were investigated in this work. It's observed that with increasing stoichiometric ratio of Al and BaO the recovery

257

of strontium metal is increased. For example 40,1% and 96,89 % of recovery of strontium is observed while the stoichiometric ratios of Al powder and BaO were 100 % and 300 % respectively. The experiments results which were carried out at 1250 "C for 60, 120, 180 and 240 minutes were shown in Fig. 4 and Fig 5. It is observed that with increasing time the recovery of strontium metal also increased. For example, recoveries of strontium were determined to be 40,1 % after 60 min. and 96,89 % after 240 minutes reduction time. Sr concentration in the residue decreased with increasing time. For example strontium concentration in residue were determined to be 69 % after 60 min. and 2,25 % after 360 minutes reduction time.

I

Time (min)

Fig. 4: Effect of time on strontium recovery

Fig. 5: Effect of time on concentration of Sr in residue Conclusions On the basis of the results of the present study of aluminothermic reduction of SrO by Al powder under 1-5 mbar pressure for different stoichiometric amount of reducing agent and time, following conclusions can be drawn:

258

Highest Sr recoveries were determined as: 1) 96,89 % with addition of 300 % stoichiometric Al and BaO for 240 min., 2) 96,87 % with addition of 300 % stoichiometric Al and BaO for 180 min., 3) 86,45 % and 85,28 % with addition of 200 % stoichiometric Al and BaO for 240 min and 180 min respectively . References [1] J. Robertson, C. W. Chen, Schottky barrier heights of tantalum oxide barium strontium titanate lead titanate and strontium bismuth tantalite, Applied Physics Letters Vol. 74/8, 1168-1170, 1999 [2] R. Moazzami, Ferroelectric thin film technology for semiconductor memory, Semiconductor Sei. Technology, 10,39-85-390, 1995 [3] Matichny S., Fabrication and characterisation of ferroelectric lead zirconate titanate and strontium bismuth tantalate thin films, Otto-von-Guericke University Magdeburg, Doctrate thesis, 2006 [4] Davis, J. R., 1996: ASM Speciality Handbook, Aluminum and Aluminum Alloys, ASM International, Materials Park, Ohio, USA [5] DPT (State Planning Organization), Eight Five-Year Development Plan, 2001, Ankara. [6] M. Erdemoglu, S. Aydogan, M. Canbazoflu, A kinetic study on the conversion of celestite (SrSC>4) to SrCO¡ by mechanochemicalprocessing, Hydrometallurgy, 86 (2007), 1-5. [7] M. Erdemoglu, Carbothermic reduction of mechanically activated celestite, Journal of Mineral Processing 92 (2009), 144-152 [8] J. Langlais, Strontium Reduction by Aluminothermic Reduction, McGill University, master thesis, 1991 [9] C.W. Bale, A.D. Pelton, and W.T. Thompson: FactSage 5.1: Thermochemical Software for Windows. "'Montereal, Quebec: Thermfact Ltd., (2002) [10] ASTM B 409, B 408, B 407, B 564.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

2nd Interaaflonal Symposium on

High-Temperature Metallurgical Processing

Treatment of Metals and Pellets Session Chairs: Jeffrey Schoonover Stephen Kampe

2nd International Symposium on I ligh-Temperarure Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

HEATS OF REACTION IN THE FORMATION OF TIB2 REINFORCED TITANIUM ALUMINIDE COMPOSITES Andrew H. Baker1, S.L. Kampe1, Tony Zahrah2 'Michigan Technological University, Department of Materials Science and Engineering 1400 Townsend Drive, Houghton, MI 49931-1295 2 Matsys Inc., 504 Shaw Road Suite 215, Sterling,VA 20166 Keywords: reaction synthesis, bomb calorimetry, titanium diboride, intermetallics Abstract Experimental heats of reaction for the formation of several Al3Ti + TiB 2 intermetallic composite formulations have been determined utilizing bomb calorimetry. To overcome the kinetic constraints and achieve ignition in these formulations, a boron (B)/potassium nitrate (KNO3) initiation aid was incorporated within the blended reactant compact. The heat contribution from the initiation aid was regressed to provide estimates of the heat of reaction of the nominal formulations. Powder x-ray diffraction was utilized to identify the products of the reaction and discern any differences between the predicted equilibrium products of the formulation and the actual products. Introduction Reaction synthesis is a technique for the creation of ceramics, intermetallics, and in-situ composite materials [1], Reaction synthesis is defined by its self-propagating mechanism, where the energy liberated from a localized reaction within a green powder compact is enough to carry the bulk reaction to completion. Achieving a self-propagating state can be achieved in variations of two primary methods of initiation: uniform volumetric heating and directional propagation [2]. Uniform volumetric heating involves subjecting a green compact of the unreacted formulation to uniform heating in a furnace to encourage diffusion to initiate the reaction (Figure la) throughout the bulk material. Directional propagation utilizes a point heat source at one section of a green compact of unreacted formulation (Figure lb). The localized exothermic reaction at one point in the compact then initiates the reaction in the section of the compact adjacent to it and fully propagates throughout the entire compact until completion. Initiation of the localized reaction is generally associated with the melting of the constituent with the lowest melting point, with instantaneous initiation of the reaction usually occurring close to that melting point [1]. This melting point can be high in some systems of interest, such as the titanium (Ti) and boron (B) system in equation 1:

263

Ti + 2B -» TiB2 , initiation at ~Tmp of Ti = ~1670°C

(1)

Another constituent can be added to the green formulation with a much lower melting point to act as an intermediary solvent to encourage diffusion of the reactants and more readily initiate the reaction at a lower temperature [3]. For the Ti and B system, Al can be added to serve this purpose as shown in equation 2: xAl + Ti + 2B -> xAl + TiB 2 , initiation ~Tmp of Al = ~660°C

(2)

Introducing an excess amount of Ti can lead to the formation of an intermetallic of Al and Ti, as shown in equation 3: 3A1 + 2Ti + 2B -» Al3Ti + TiB 2 , initiation ~T mp of Al = ~660°C

(3)

The success of using reaction synthesis for creation of the expected equilibrium products from a formulation depends on several intrinsic and extrinsic characteristics of the pre-reacted powder blend. These include a variety of thermodynamic and kinetic relevant variables, including the magnitude of the heat of reaction, the temperature of the reaction, the time scale over which the reaction occurs, and diffusional characteristics. A plethora of effects relative to the homogeneity of the reactant blend such as the size and dispersion of the powder reagents also contribute. Because of the many potential kinetic barriers in play during a reaction, the actual heat of reaction of a formulation may or may not coincide with predicted values associated with the equilibrium products. In instances where equilibrium products do not form, actual heats need to be experimentally measured.

(a)

(b)

Figure 1. Modes of initiation in reaction synthesis: (a) directional propagation (b) uniform volumetric Bomb calorimetry offers a means to accurately measure heats of reaction in relatively large sample sizes. Bulk samples more realistically take into account uncertainty associated with the volume effects; for example, the effects of large surface to volume ratios and inhomogeneity of the dispersion of the powders. However, due to kinetic constraints, traditional means of ignition in bomb calorimetry can be difficult for certain formulations. In such instances, utilization of a newly developed bomb calorimetry technique involving the blending of a B/KNO3 initiation aid

264

with the green reagent compact to sensitize it may offer a viable approach for measuring heats of reactions for reaction synthesis [4]. Creation of the equilibrium products for the AI3TÍ + TÍB2 intermetallic composite system from raw reagent powder compacts of Al, Ti, and B using reaction synthesis have been previously successful in the literature [3], This research aims to use and validate the novel B/KNO3 initiation approach in bomb calorimetry to estimate the heat of reaction in the formation of various volume proportions of AI3TÍ and TÍB2 through regression of the initiation aid.

Experimental Procedures 100 g master batches of Ti, Al, and B formulated in stoichiometric proportions to form various composites of AI3TÍ + TÍB2 were created. The blends were sealed in a container with cylindrical zirconia media under argon in a glove box and allowed to mix overnight. KNO3 was crushed to a fine powder using a silica mortar and pestle. Preparation of 10 g formulated samples for the calorimetry experiments utilized a portion of the master batch as well as K.NO3 plus additional B as the internal initiation aid fraction. To create the charges, the powders were densified in a cold isostatic press at 240-270 MPa. Small pieces of the 10 g charge were removed and a corollary external initiation aid (separately densified under the same procedure) was added (shown in Figure 2) to it in a silica crucible. The sample was then evaluated using a standard operating procedure in a Parr 1341 Plain Jacket Bomb Calorimeter. The bomb in the calorimeter was charged with argon (99.998% purity) to seal it and provide an inert environment for the reaction. Following reaction synthesis, x-ray diffraction was conducted on selected samples to determine the resulting phase constituency of the reacted product.

Figure 2. Blended compact setup with B/KNO3 and Al, Ti, B [4].

265

Results and Discussion Process Variable Screening Before a regression model can be developed to estimate the heat of reaction of a sample, the processing variables must be understood so as to be able to produce a model that contains a high degree of accuracy. In previous work involving the creation of TÍB2, it was shown that the pressure of argon used to charge the bomb in the calorimeter did not have a significant effect on the energy change [4]. Since no new elements are being added that would produce a gas, it is assumed that the pressure will continue to be insignificant. Utilizing a 26"2 screening design of experiments (DOE) matrix, the following factors in Table I were analyzed for main effect, twoway and three way-interaction significance. Table I. Screening factors and levels for DOE. Factor Levels Initiation Aid (IA) Fraction 0.25, 0.50 Fraction of IA in Compact 0.667, 0.80 0.70, 0.85 KNO3 Fraction of IA TÍB2 vol. Fraction (Formulated) 0.20, 0.40 Titanium Particle Size 44um, 149 urn 44um, 149 urn Aluminum Particle Size

Figure 3. Normal plot of the effects for the screening factors displaying significance

266

The normal plot above (Figure 3) estimates the significance of the variables in Table I. The total initiation aid fraction (IA Frac), the fraction of the IA in the compact (Int Frac), the K.NO3 fraction of the initiation aid, the formulated volume fraction of TiB2 in the product (TiB2 Frac), and an interactive effect between the IA Frac and the KNO3 Frac are significant. These effects are all expected and pose no further problems for developing a regression. The initiation aid and TÍB2 volume fraction are the basis of the regression, and therefore would be expected to have a large effect. The fraction of the IA in the compact (Int Frac) being significant implies an interaction between the constituents in the compact and the initiation aid added to the compact. Dealing with this phenomenon will be addressed in the regression development section. The presence of the KNO3 effect and interaction between KNO3 and IA Frac is expected because in previous work it is shown that the levels chosen will have a different external initiation aid enthalpy value [4]. Keeping the KNO3 fraction constant will negate this effect as the enthalpy will then depend only on the total level of initiation aid and the internal fraction of the initiation aid. Regression Development In order to develop a refined regression, additional levels were added at certain TÍB2 levels. The experimental enthalpy data input into the regression are a result of the factors in Table II. K.NO3 fraction of die initiation aid was kept constant at 0.70. TiB 2 Levels IA Frac Levels Int Frac Levels

Table II. Factors and levels for the regression model. 0.2 0.4 0.25, 0.375, 0.50 0.25, 0.375,0.50 0.401,0.514,0.667,0.80 0.401,0.514,0.667,0.80

0.6 0.25, 0.375, 0.50 0.667, 0.80

For creation of a regression model, the internal and external contributions were separated to eliminate collinearity. This was accomplished by creating the following terms: IA Frac*Int Frac (the internal contribution to the enthalpy change from the initiation aid) and IAFrac*(l-Int Frac) (the external contribution to enthalpy change from the initiation aid). Using these newly created terms and the TiB2 factor, the reduced regression model is determined to be equation 4: AH (kJ/g) = -1.67 - 3.25 Ext Contrib - 8.95 Int Contrib - 4.01 TiB2 Frac + 8.98 Int*TiB2 Statistical information displaying the significance of each term in the model (p-value) and the standard error of each coefficient (SE Coef) are shown in Table III: Table III. Statistical information for the terms of the regression model Predictor SE Coef T P Coef 0.2532 -1.671 Constant -6.60 0.000 Ext Contrib -3.2546 0.4707 -0.69 0.000 -8.9514 0.8923 -10.03 0.000 Int Contrib -4.0088 0.6685 -6.00 0.000 TiB2 Frac Int*TiB2

2.416

8.984

267

3.72

0.001

(4)

Where 'Ext Contrib' is the external initiation aid component, 'hit Contrib' is the internal initiation component, TiB 2 Frac' is the formulated IÏB2 product volume fraction, and 'Int*TiB2' is the interactive effect between the internal initiation component and the formulated TiB 2 product volume fraction. The adequacy of the model was determined to be sufficient with an Rsq (adj) value of 93.3%. Enthalpy of reaction values were determined using the regression model for the experimental TÍB2 volumes of 0.2, 0.4, and 0.6 fraction and also extended to predict values at 0.0, 0.8, and 1.0 TÍB2 volume fraction. These values are obtained by entering the particular T1B2 fraction value and zeroing the contributions from the initiation aid. Figure 4 displays the experimental and predicted values from the regression model next to the predicted values from the literature. The enthalpy is converted to kj/mol from kj/g assuming the correction proportions of products were formed in order to evaluate the regression values versus the literature values. The Barin [5] data utilizes enthalpy data for both TiB2 and Al3Ti from [5], while the (Frankhouser + Barin) [5,6] data utilizes enthalpy data for TiB2 from [6] and for AI3TÍ from [5].

2 x

Formulated Volume % of TiB2 in AI3Ti + TiB2

Figure 4. Experimental and predicted enthalpy change versus vol. % of TiB2 formed. Note: Only 20%, 40%, and 60% are directly regressed from experimental data. 0%, 80%, and 100% are extrapolated using the model. The regression model follows closely for the experimental values of 0.2, 0.4, and 0.6 volume fraction of TiB2 and its predictive value is reasonable with the literature values for 0.8 and 1.0. However, the predictive value of the model fails when attempting to predict the value of 0.0 volume fraction of TiB2, single phase AI3TÍ. As apparent in the model, this divergence likely occurs because of the large enthalpy change estimated at the 0.2 TÍB2 volume fraction level. While close in value, it falls outside the range of the Barin and Barin+Frankhouser values and

268

alters the trend. It is possible that this 0.20 volume fraction value is skewed due to an outlier. An analysis of large residuals yields one of the experimental enthalpy values at 0.20 TÍB2 volume fraction occurring at 2.56c away from the predicted value of the model. Further investigation at the 0.20 TÍB2 volume fraction and smaller volume fractions must be conducted in order to determine if the technique will yield a regression model with predictive ability that yields enthalpy values similar to the literature enthalpy values. X-Rav Diffraction X-ray diffraction of selected samples at 0.40 (Figure 5) and 0.20 revealed formation of the intended equilibrium products TiB2 and Al3Ti. Additional byproducts of AI2O3, TiN, and Al were also identified. AI2O3 likely forms from the oxidation of Al with KNO3. While at this particular particle size, KNO3 will not readily oxidize aluminum, once Al is in the liquid form, it should be possible for this reaction to take place. TiN is likely formed from titanium absorbing nitrogen gas, which is a byproduct of the reaction of KNO3 and B [7]. If these are the mechanisms for the formation of AI2O3 and TiN, as the initiation aid is regressed to zero, these byproducts should disappear.

a. u

28(Deg.)

Figure 5. X-ray diffraction pattern of a product with formulation factor levels: 0.50 IA Frac, 0.80 Int Frac, 0.4 TiB2 Frac The presence of Al in the x-ray diffraction spectrum is likely due to the forced imbalance of the formulation when the initiation aid is added. If the prevalent reaction is the formation of TiB2, the TiB2 reaction would abscond Ti from the formation of Al3Ti, putting the effective formulation titanium deficient and therefore in the phase field of Al + Al3Ti [8]. Under this

269

reasoning, in the presence of the initiation aid, Al would be expected to be seen in the x-ray diffraction spectrum of the product. Conclusions The B/KNO3 initiation aid blending technique for bomb calorimetry appears to be a viable way for measuring the heat of reaction of a bulk sample of blended powders of Al, Ti, and B formulated to form various volume proportions of AI3TÍ + TiB2. The regression model of experimental values were near literature values and extrapolation of the model to values outside the experimental range were in good agreement except for the nominal value of single phase AI3TÍ (0 vol% TÍB2). This deviation may be due to an outlier in the experimental data which may be altering the trend. Further experimentation is needed in both the high volume percent TÍB2 region and the low volume percent TÍB2 region to verify these findings. The low volume percent region requires particular attention to determine if trend deviation is an artifact of the outlier or a lack in the technique's ability to produce the intended products. X-ray diffraction confirmed the presence of the intended equilibrium products and the identified byproducts are reasonably explainable. The identified byproducts should disappear as initiation aid is removed. Further x-ray diffraction must be conducted to establish quantitative values of the volume percent of the products to verify this predicted diminishing effect of the byproducts and to confirm that the relative volume proportion of the intended products nears the intended value. References 1. A S . Rogachev, A S . Mukas'yan and A.G. Merzhanov, "Structural Transitions in the Gasless Combustion of Titanium-Carbon and Titanium-Boron Systems," Dokl. Phys. Chem., 297 (1987), 1240-1243. 2. John J. Moore and H.J. Feng, "Combustion Synthesis of Advanced Materials: Part II. Classification, Applications and Modelling," Progress in Materials Science, 39 (1995), 275-316. 3. R. Martin et al., "Microstructure/Processing Relationships in Reaction-Synthesized Titanium Aluminide Intermetallic Matrix Composites," Metallurgical and Materials Transactions A, 33A (2002) 2747-2753. 4. A.H. Baker, "A Technique to Measure the Heat of Reaction of TÍB2 Reinforced Intermetallic Composites" (Paper presented at the 2010 TMS Annual Conference, Seattle, Washington, 15 February 2010), 457-463. 5. I. Barin, Thermomechanical Data of Pure Substances, 2nd ed.(Weinheim: VCH, 1973) 6. W.L. Frankhouser, K.W Brendley, and M.C. Kieszek, Gasless Combustion Synthesis of Refractory Compounds (Park Ridge, NJ: Noyes Publications, 1985), 52. 7. Yutaka Yano, "Condensed Phase Reaction of Boron with Potassium Nitrate," Propellants, Explosives, Pyrotechnics, 14 (5) (1989), 187-189. 8. F.H. Hayes, Ti-Al Phase Diagram, ASM Alloy Phase Diagrams Center, P. Villars, editorin-chief; H. Okamoto and K. Cenzual, section editors; ASM International, Materials Park, OH, 2006.

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2nd International Symposium on High-Tempera ture Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

HOT WORKABILITY OF 1.2690 LEDEBURITIC TOOL STEEL AND DEVELOPMENT OF MICROSTRUCTURE M. Tercelj'.G. Kugler1 'Department of Materials and Metallurgy, University of Ljubljana, Askerceva cesta 12, 1000 Ljubljana, Slovenia Keywords: Tool steels, Hot workability, Microstructure, Carbides. Abstract The 1.2690 high alloyed tool steel is usually applied for cold working and is alloyed with carbide-forming elements Cr, V, W and Mo. It exhibits very poor hot deformability. Since there is no enough available data in literature for elucidation of this problem, hot workability and development of microstructure during hot deformation was studied. Hot compression tests were carried out in temperature range 850-1150 °C and strain rates range 0.001-6 s"1. SEM and OM were used for observation of microstructure. Results revealed very complex precipitation of carbides that depends on deformation temperature as well as on strain rate. It was found that especially at lower temperatures of hot working range the precipitation of carbides strongly depends on strain rate. Three different hot deformation behaviors' were observed depending on temperature and strain rate range. Introduction The hot workability of highly alloyed high speed steels is usually described as relatively poor in comparison to other steels due to their tendency of cracking during hot working especially when inappropriate hot working conditions are applied. Alloying elements such as Cr, W, Mo and V in ledeburitic tool steels form carbides that improve hardenability, strength, hardness, wear resistance, etc., but they decrease hot workability and narrow hot working temperature range in comparison to conventional steels. The size, distribution, type and fraction of carbides, as well as the thermomechanical history of material and hot working parameters have major influence on the hot workability. The microstructure of ledeburitic tool steels during hot deformation can be characterized as a two-phase system consisting of an austenite matrix and a combination of ledeburitic carbides [1-4]. In the soft-annealed condition the total volume fraction of carbides is in the range 19-30% and in the quenched condition in the range 9-25%, depending on the chemical composition. During heating before deformation a large quantity of carbides dissolve, but non-dissolved carbides can have an extremely diverse influence on the hot workability due to variations in their composition, size, distribution, hardness and shape, and their different behavior during the deformation phase [5-7]. In order to prevent hot cracking and microstructural damage during hot-working it is necessary to understand the influence of carbides on strain hardening and softening mechanisms. In the present investigation the evolution of microstructure during hot deformation of 1.17C-11.3CrL48V-2.24W-l.35Mo ledeburitic tool steel was studied using hot compression testing and metallography.

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Experimental Procedure and Material The chemical composition of the tool steel used in this investigation is given in Table 1. The cylindrical specimens with dimensions <|>10mmxl5mm were machined from forged and annealed square 95-mm billet. The center lines of the specimens were perpendicular to the longitudinal direction of the billet. Due to some microstructural inhomogeneity across the section of the forged billet, all the compression specimens were taken at the same depth from the surface of the square billet. Table 1: Chemical composition of the tool steel in wt.\%. ~C 1.17

Si Mn 0.24 0.26

Cr W V 11.3 1.48 2.24

Mo 1.35

The hot-compression tests were carried out on a Gleeble 1500D thermo-mechanical simulator. For a determination of the flow curves and in order to study the influence of the deformation parameters on the evolution of the microstructure, the following testing conditions were selected: temperature range 850-1150°C, strain rates between 0.001 and 6s"1 and a true strain of 0.9. The specimens were heated to 1150°C with a heating rate of 3°C/s, which was followed by holding them for 10 min at this temperature, and cooling them to deformation temperature with 2°C/s, holding them at this temperature for 10 min, which was followed by hot compression and water quenching. In order to avoid inhomogeneous deformation and sticking between the specimen and the compression tool, tantalum foil with a thickness of 0.1 mm was inserted between the cylindrical specimen and the compression anvil, and a Ni-based lubricant was used. The temperature of the deformed samples was measured and controlled with a type-S thermocouple. The deformed samples were visually inspected and then longitudinally cut for the preparation of the metallographic samples. Optical microscopy (OM, Carl Zeiss AXIO Imager.Alm) and a field-emission scanning electron microscope (FE SEM) were applied for the observation of the microstructure. The deformed specimens for the optical microscopy were polished with a sequence of grinding papers from 180 to 1000 meshes of granulation, followed by polishing with diamond paste of 1 and 0.25um granulation, and then etched with Vilela's reagent. The deformed specimens for EBSD were metallographically prepared using a standard procedure, followed by polishing with silica oxide for 3 minutes and cleaned in an ultrasonic bath. Results and Discussion The initial microstructure of the applied samples is shown in Figure la. It is clear that the softannealed microstructure consists of small, spheroidised carbides and ledeburitic carbides in a ferrite matrix. Figure lb shows the quantity and distribution of the ledeburitic carbides in the quenched and tempered microstructure, consisting of tempered martensite and secondary carbides. The combined EDS and EBSD analyses of the soft-annealed microstructure revealed mat the ledeburitic carbides with a size of a few micrometers are of the M7C3 type and the small spheroidised carbides are of the M23C6 type. In order to study the deformation behavior and the influence of temperature and strain rate on the microstructure, a series of compression tests was carried out for various temperatures and strain rates. Typical stress-strain curves are shown in Figure 2a. The flow curves for all the tested temperatures and strain rates exhibit a distinct maximum, which is followed by a decreasing of the flow stress towards the steady state that indicates that the dynamic recrystallization (DRX) was the main softening mechanism. This will be further discussed later.

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Figure 1: The initial microstructure of a 95mm, forged, square billet: soft annealed (a) and quenched and tempered microstructure - quantity and distribution of carbides (b), OM. In order to present more clearly the relationship between the deformation temperature, the strain rate and the maximum stress, the peak stresses are plotted against the temperature for different strain rates in Figure 2b. Activation energy for deformation was determined by fitting the maximum flow stresses to the hyperbolic sine equation [8]: Z = eexp(2//?7') = /(sinh"(acr) (1) where Z is the Zener-Hollomon parameter, R is the gas constant, Q is the apparent activation energy for hot deformation, and A, n and a are constants. These parameters were determined using the procedure described in [9] and the following values have been obtained: Q=681kJ/mol, a=0.0045MPa"', n=5.85, A=8.121026s"'. Obtained activation energy lies in the range of activation energies reported in the literature for tool steels [5-7].

True strain,/

Temperature, *C

Figure 2: Stress-strain curves in the range of 850°C C to 1150°C and strain rate of Is"1 (a), dependence of the peak stresses on the temperature for different strain rates (b). Optical microscopy carried out on the deformed samples revealed fully recrystallized microstructures with equiaxed grains at all strain rates for temperature of 1150°C, and at strain rates of 6 and Is"1 for a temperature of 1100°C, as shown in Table 2. At lower strain rates the microstructures are only partially recrystallized, but for temperatures of 1000°C and below the DRX grains cannot be clearly distinguished due to the intensive precipitation of the secondary carbides and etching problems. On the other hand, the shapes of the flow curves suggest that DRX has taken place during deformation.

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Table 2: Fractions of the recrystallized microstructure (in %) for various deformation temperatures and strain rates, x denotes that the fraction of recrystallized microstructure is not clearly evident from our observation of the microstructure T[°C] 1150 1100 1050 1000

950 900 850

6.0 100 100 100 75 30 X X

Strain rate [s1] 0.1 0.01 0.001

1.0 100 100 50 40

100 75 50 30

100 50 30 30

100 40 25 25

X X X

X X X

X X X

X X X

Upper and Medium Temperature Ranges A fully recrystallized microstructure is composed of retained austenite, martensite and carbides (see, e.g., Figure 3a). In samples deformed at 1100°C some fine secondary carbides that precipitated at the primary non-dissolved carbides are visible as a serrated surface. However, on the grain boundaries these secondary carbides were not observed. For strain rates of 6s"1 and Is"' the grains are fully recrystallized, and with a decreasing strain rate the fraction of DRX grains decreases, but their mean size increases (see Figures 3a-c and Table 2). A similar dependence of the fraction and the mean size of the DRX grains on the strain rate was also observed at a deformation temperature of 1050°C. But at this temperature a larger amount of secondary carbides precipitated at the primary non-dissolved carbides in comparison to a deformation temperature of 1100CC, which consequently became more serrated (see Figures 3d-f). Furthermore, at this temperature the individual, fine, secondary carbides mat precipitated on the grain boundaries are also visible.

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Figure 3: Microstructures of deformed samples in the upper temperature range: 1100°C, 6s"1 (a), 1100°C, 0.1s"1 (b), 1100°C, 0.001s"1 (c), 1050°C, 6s"1 (d), 1050°C, 0.1s"1 (e), 1050°C, 1050°C, 0.001s"1 (f). Lower Temperature Range and Higher Strain Rates For deformation temperatures of 1000°C and below, and strain rates of Is"1 and 6s"1, the microstructure consists of deformed, non-recrystallized austenitic grains, carbides and DRX grains on the grain boundaries of non-recrystallized grains (see Figures 4a-c).

Figure 4: Microstrucrures of deformed samples in the lower temperature range strain rates of 6s"', OM: 1000°C (a), 950°C (b), 850°C (c). Note that the secondary carbides precipitated on the grain boundaries and on the non-dissolved ledeburitic carbides. It should be mentioned here that the grain boundaries were not clearly revealed by the etching. Figure 4a shows the microstructure of a sample that was deformed at a temperature of 1000CC and a strain rate of 6s'1. In this sample a fraction of the recrystallized microstrucrure, which was estimated on the basis of the coarseness and the orientation of the martensitic needles, amounts to approximately 75% (see Table 2). With decreasing temperature and strain rate the fraction of DRX decreases (see Figures 4a-c). Note that the same behavior of a decreasing DRX fraction with a decreasing strain rate, as was found for the medium temperature range, is also observed here. Lower Temperature Range and Lower Strain Rates For a deformation temperature of 1000°C at lower strain rates (0.001-O.ls"1) the microstrucrure consists of elongated, deformed, primary austenite grains with a necklace of DRX grains around their grain boundaries and carbides. The fraction of DRX grains is about 30% for the two higher strain rates and about 25% for the lower strain rate (Figure 5a). Note that by lowering the strain rate the DRX grains become coarser. At deformation temperatures of 950CC and below the deformed microstrucrure consists of elongated, deformed, primary austenite grains and occasionally also of twin lamellae. The base microstrucrure consists of different amounts of retained austenite and martensite as well as of primary, non-dissolved ledeburitic carbides. The recrystallized crystals along the grain boundaries of the primary grains were not clearly revealed by etching, but intensive precipitation of the secondary carbides is clearly visible (see Figures 5b-f). This intensive precipitation of secondary carbides along the primary grain boundaries influences the chemical composition in the vicinity of the precipitated carbides, i.e., the concentrations of the carbide-forming elements are decreased. As a consequence, the loweralloyed martensite along the primary grain boundaries is etched more strongly (coloured brown) in comparison to the base austenite-martensite grains (coloured white), as shown in Figure 5. With a decreasing deformation temperature and strain rate the bands of lower-alloyed martensite become wider, which means a larger amount of precipitated secondary carbides on the grain

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boundaries. Additionally, the microhardness measurements for the sample deformed at 850°C and a strain rate of 0.001s"1 gave values between 536 and 592 HV0.025, for the white-colored areas, and between 488 and 536 HV0.025, for the brown-colored areas. The amount of precipitated carbides increases with decreasing strain rates, as can be seen from Figures 5c-f. This can be explained by the longer time that is available for precipitation during the deformation at lower strain rates. Additionally, from the thermodynamic point of view, the tendency of the carbides to precipitate is more emphasized at the lower deformation temperatures.

Figure 5: Microstructures of deformed samples in the lower temperature range at lower strain rates, OM: 1000°C, 0.1s-1 (a), 950°C, 0.1s"1 (b), 950°C, 0.01s"1 (c), 900°C, 0.001s"1 (d), 850°C, 0.001s"1 (e), 850°C, 0,01s"' (f). An inspection of the microstructure revealed that the precipitated carbides, which precipitate mainly on the grain boundaries, have a large influence on the growth of the DRX grains. Since the fraction and the size of the DRX grains decreases with the increasing fraction of secondary carbides it can be concluded that they are pinning the DRX grain boundaries, thereby hindering their movements (the size of the DRX grains was assessed on the basis of the coarseness of the martensitic needles). Furthermore, they also weaken the grain boundaries, which leads to cracking at lower temperatures (850°C) and lower strain rates 0.01-O.OOls"1 (see Figure 5f). Therefore, the lower fractions of DRX grains and consequently the higher values of maximum flow stresses for the temperatures 850-950°C and the strain rates O.l-O.OOls"1 can be attributed to the role of the precipitated secondary carbides. Secondary precipitation also results in increasing flow stresses at the lower strain rates. Thus, the peak values of the flow stresses at 0.1s approach the peak values of the flow stresses at Is"1 (see Figure 2). In order to investigate the precipitation of the secondary carbides at lower temperatures, microstructures obtained with a FE SEM were compared for samples that were soaked at 1150°C, then cooled to 900°C, held there for 10 min and then quenched or deformed with strain rates of 6s"1 and 0.01s"1, respectively. It can be seen from Figure 6a that during cooling from the soaking temperature to the deformation temperature and holding there, secondary carbides precipitate along the primary grain boundaries and on non-dissolved primary carbides. The precipitated carbides already form an uninterrupted carbide film on some parts of the grain

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boundaries. These films of carbides hinder the grain-boundary movements, but with the progress of deformation at 900°C the carbide film is crushed into discrete parts (see Figures 6b-c). It can also be seen that during deformation the secondary carbides precipitate along the grain boundaries. For a higher strain rate of 6s"1 the carbide film is more crushed and the amount of precipitated secondary carbides is lower than for a strain rate of 0.01s"1, as shown in Figures 6bc. Note that for the lower strain rate the secondary carbides are partially stick-shaped.

Figure 6: Microstructure of samples soaked at 1150°C and cooled to 900°C and held there for 10 min, FE SEM: only quenched without deformation, BE (a); deformed at a strain rate of 6s"1, SE (b), deformed at a strain rate of 0.01s"1, BE (c). Conclusions The main findings of the present investigation are: 1. On deformation temperature the microstructure consists of austenite and carbides of type M7C3, M23C6 and M6C. Strain rate influences on fraction of precipitated carbides and consequently also on fraction of recrystallized grains. 2. The lower limit of the safe working range is determined by the precipitation of the secondary carbides, which is influenced by the applied strain rate. During the deformation at temperatures below 1000°C an intensive precipitation of secondary carbides at the grain boundaries, especially at lower strain rates, takes place, which results in increased work hardening and weakening of the grain boundaries. 3. It was observed that with decreasing of strain rate also the fraction of dynamically recrystallized microstructure is decreasing. Thus, at a strain rate of 6s"1 and at 1100°C the recrystallized fraction is 100%, while at a strain rate of 0.001s"1 mis fraction amounts to 40%. At a deformation temperature of 1000°C the recrystallized fractions were 75% and 25% for the highest and lowest strain rates, respectively. 4. From the results of the present investigation it can be suggested that during hot deformation the combination of low temperatures and lower strain rates should be avoided. 5. The apparent activation energy for 1.2690 ledeburitic tool steel in temperature interval 850-1150°C for strain rates of 0.001-6s"1 is 681 kJ/mol. References [1] M. Boccalini, H. Goldstein, Solidification of high speed steels, International Materials Rewiews 46/2 (2001 ) 92-115. [2] D. Peidao, S. Gonqi, and Z. Shouze, A Scanning Electron Microscopy Study of Carbides in High-Speed steels, Materials Characterization 29 (1992) 15-24.

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[3] D. W. Hetzner, Refining carbide size distributions in Ml high speed steel by processing and alloying, Materials Characterization 46 (2001) 175-182. [4] S. Kheirandish, Effect of Ti and Nb on the Formation of Carbides and the Mechanical Properties in As-cast AISI-M7 High-speed Steels, ISIJ International 41 (2001) 15021509. [5] C. A. C. Imbert, G. J. McQueen, Hot Ductility of tool Steels, Canadian Metallurgical Quarterly 40/2 (2001) 235 - 244. [6] C. Rodenburg, M. Krzyzanowski, J. H. Beynon, W. M. Rainforth, Hot workability of spray-formed AISI M3:2 high-speed steel, Materials Science and Engineering A 386 (2004) 420 - 427. [7] J. Liu, H. Chang, R. Wu, T.Y. Hsu, X. Ruan, Investigation on hot deformation behaviour of AISI Tl high-speed steel, Materials Characterization 45 (2000) 175-186. [8] J.J. Jonas, CM. Sellars, W.J. M. Tegart, Strength and structure under hot working conditions, Metall. Rev., 130 (1969) 1-24. [9] G. Kugler, M. Knap, H. Palkowski, R. Turk, Estimation of activation energy for calculating the hot workability properties of metals, Metalurgija 43/4 (2004) 267-272.

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2nd International Symposium on fugh-Temperahire Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

EFFECTS OF BINDERS ON OXIDIZED PELLETS PREPARATION FROM VANADIUM/TITANIUM-BEARING MAGNETITE Guihong Han, Yuanbo Zhang *, Yanfang Huang, Zengqing Sun, Guanghui Li, Tao Jiang School of Minerals Processing & Bioengineering, Central South University, Changsha, Hunan, 410083, China Keywords: Oxidized Pellets, Vanadium/Titanium-bearing Magnetite, Organic Binder Abstract Because of small specific surface area and poor ballability, vanadium/titanium-bearing magnetite is difficult for production of oxidized pellet. A novel organic copolymer binder, modified humic acid (MHA), has been authorized recently in China. Effects of the novel MHA binder on the quality of V/Ti-bearing magnetite pellets are studied in this paper. Experimental results show the MHA binder can obviously improve the strength of green pellet. Because of better heat endurance, the MHA binder pellet has a higher strength than Peridur pellet under the same preheating conditions. Moreover, TFe grade of pellet with the MHA binder is much higher than that of the bentonite pellet. The strength of preheated pellet containing MHA binder of 0.25%~1.0% meets the requirements of oxidized pellet production by grate-kiln process. Comparatively, MHA is a promising orgainc binder for iron ore pellets. Its price is only one tenth of Peridur price and about 3 times of bentonite price. Introduction The iron-making practices show that the use of pellet is beneficial to blast furnace life, productivity and fuel savings. Therefore, many iron-making plants have expanded the production of oxidized pellets in order to meet the needs of modern large-scale blast furnaces [1]. Presently, bentonite is widely used to improve the quality of iron ore pellets. The negative effects of using bentonite include decreasing total iron (TFe) grade and degrading the metallurgical properties of the final products [2, 3]. Many studies, therefore, have been conducted on the application of organic binders to overcome the weakness of the inorganic bentonite. In comparison with inorganic binders, one of the biggest strengths of organic binders is low residual silica content [4]. However, those studies showed that the economics of foreign organic binder addition into pellets was not attractive. For example, the compression strength of preheated pellets and roasted pellets are both too low, resulting in powder generation and ring formation in the industrial rotary * Corresponding author: Dr. Yuanbo Zhang. Tel: +86 731 88877214; Fax: + 86 731 88830542; E-mail: [email protected].

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kiln [3,4]. One novel organic copolymer binder with many functional groups, named MHA binder, has been patented in China recently [5], V/Ti-bearing magnetite, as one of typically complex iron ores, is partly utilized for production of oxidized pellet currently in China. Because of low surface areas and poor balling behavior however, V/Ti-bearing magnetite is difficult for pelltizing [6, 7]. Effects of different binders on qualities of V/Ti-bearing magnetite pellets are studied emphatically in this paper. Experimental Raw Materials The chemical compositions of the V/Ti-bearing magnetite concentrates are shown in Table I. The concentrate contains 55.36% TFe, 10.00% Ti0 2 and 0.72% V 2 0 5 (wt %). Then, size distribution of iron concentrates was tested. The content of the sample below 0.074 mm is only 39.60% and the specific surface area is just 993 cm2/g. Table I. Chemical Compositions of the Iron Concentrates (wt %) TFe FeO Si0 2 A1 2 0 3 CaO MgO Ti0 2 v2o5 P S 2.12 55.36 21.36 4.40 5.23 0.81 10.00 0.72 0.018 0.045 * LOI means loss of ignition.

LOI* 0.65

The MHA binder mainly consists of fulvic acid and humic acid. In the MHA binder, the proportion of ful vie acid to humic acid is about 1:10. Chemical compositions of the MHA binder were given in Table II. Thermoanalytical results showed that MHA binders were characterized by decomposition and combustion of volatile components and fixed carbon. Table II. Chemical Compositions of the MHA Binder Other composition, % Volatile matter, % Si0 2 Na 2 0 AI2O3 CaO MgO 1 Fe 28.99 20.62 3.82 11.98 0.50 0.33 | 6.08

Fixed carbon 20.52

The SEM photo and FTIR spectrum of the MHA binder were presented in Fig. 1 and Fig. 2, respectively. As presented in Fig. 1, the MHA binder mainly exists in the shape of aggregate structure. Fig. 2 shows the MHA binder contains a lot of carboxyl and hydroxyl groups, as well as aromatic ring. Relevant research shows that the adsorption of MHA binder onto iron ore particle surfaces is able to increase the hydrophilicity of the particles, and to improve the ballability of iron ore particles and quality of the pellets [8].

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Fig. 1 SEM photo of the MHA binder

Fig. 2 FTIR spectrum of the MHA binder

The chemical compositions of bentonite were shown in Table III. The granularity of bentonite below 0.074 mm is 95.2%. The organic Peridur has been used for the process of travelling grate induration machine abroad [3, 4], The main characteristic of Peridur had been presented in former study [9], Fe 2 0 3 3.13

Si0 2 58.37

Table III. Chemical Compositions of the Bentonite (wt %) AI2O3 15.15

CaO 2.70

MgO 4.28

K20 0.95

Na 2 0 2.55

P 0.14

S 0.025

LOI 10.75

Methods The experimental procedure included blending, wet-grinding, balling, drying, preheating, oxidation roasting, cooling, etc. For each batch balling test, 5 kg of the iron concentrate was mixed with given binder. Because of high viscosity of the MHA binder, the original mixture was ground in a cylindrical grinder with dimensions of 1000 mm diameter and 500 mm length. The grinding conditions were fixed as follows: the moisture content of raw materials 7.0%, grinding time 5min, rotary ratio 35 r-min"1. Green pellets of 10 mm to 12 mm in diameter were prepared using a laboratory balling disc with a diameter of 1000 mm, an edge height of 200 mm, and a tilt angle of 45°. Green balls were dried at the temperature of 105 °C. Preheating and roasting experiments were respectively conducted in an electric heated horizontal tube furnace. Then, the preheated and roasted pellets were naturally cooled to room temperature. Results and Discussions By comparison with bentonite and Peridur, effects of the MHA binder on the compression strength of green balls, preheated and roasted pellets were mainly investigated. Based on the exploratory studies, the suitable balling time and moisture content was given as 12 min and 8.5%,

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respectively. Effects of Binders on the Strength of Green Pellets Effect of Binder Variety The effects of different binders on the strength of green balls were shown in Table IV.

Binder Bentonite Peridur MHA

Table IV. Effects of Binder Variety on the Strength of Green Balls Dosage Drop strength Compression strength Cracking temperature /% /times/0.5m /N/P /°C 2.0 4.8 11.0 >600 0.15 10.1 10.5 >600 0.50 9.6 11.2 >600

As seen from Table IV, the difference in the strength of green balls is embodied mainly as the drop strength. Under the condition of 2.0% bentonite, the drop strength of green balls is 4.8 times/0.5m. When it comes to green balls with 0.15% Peridur, the drop strength can reach 10.1 times/0.5m. When the MHA dosage is 0.50%, drop strength of green pellets is 9.6 times/0.5m.. It can be concluded that organic binders can obviously improve the drop strength of green balls. It can be also seen from Table3 that the compression strength and cracking temperature of green pellets satisfy the demands of pellet production. Compared Peridur with MHA binder, though the dosage of Peridur is less than that of MHA, the price of Peridur is much greater than that of MHA. Effect of MHA Dosage The effect of MHA dosage on the quality of green balls was investigated. The results were given in Table V. Table V. Effect of MHA Dosage on the Strength of Green Balls Dosage of MHA Drop strength Compression strength Cracking temperature /times/0.5m /N/P /°C /% 0.25 5.2 11.2 >600 0.50 9.6 11.2 >600 0.75 13.7 10.6 506 1.00 18.4 10.2 425 It can be seen from Table V, the drop strength of green balls increases obviously with the increase of the MHA dosage. However, the thermal cracking temperature is remarkably declined if the MHA dosage is in excess of 0.75%. When the MHA dosage is 0.25%, the drop strength is

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5.2 times/0.5m and the compression strength is 11.2N/P, which are still higher than the green balls with 2.0% bentonite (shown in Table IV). Effects of Binders on the Strength of Preheated Pellets Effect of Binder Variety The effects of different binders on the strength of preheated pellets were tested and the results were listed in Fig. 3.

Bentonite2.0%

PeridurO.15%

MHAO.50%

Type and dosage of binder

Fig. 3 Effect of binder variety on the compression strength of preheated pellets (Preheating temperature 950 °C, preheating time 10 min) The results in Fig. 3 clearly show that the compression strength of preheated pellets with organic binders is lower than that of bentonite pellet. The compression strength of preheated pellets with benonite is 623 N/P. While the preheated pellets with Peridur is only 287 N/P. Because of better heat endurance than Peridur, the MHA pellet has a higher strength than the Peridur pellet under the same preheating conditions. And the compression strength of prehated pellets with 0.50% MHA is 558 N/P. The results indicate that the strength of preheated MHA pellets meet the demands of industrial grate-kiln process. Effect of MHA Dosage Under die condition of preheating temperature 950 °C and preheating time 10 min, the effect of MHA dosage on the compression strength of preheated pellets was measured and given in Fig. 4. Fig. 4 displays the compression strength of preheated pellets is clearly decreased with increase of MHA dosage. The compression strength comes to 563 N/P when the MHA dosage is only 0.25%. However, the compression strength of preheated pellets decreases to 470 N/P if the dosage

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increases up to 1.0%. The results indicate that the increasing of MHA dosage is unfavourable for the compression strength of preheated pellets. The reason is that decomposition and release of volatile matter contained in the MHA binder weakens the oxidizing atmosphere. With increase of MHA dosage, gaseous product increases the resistance to oxygen diffusion into pellets, which results in the slow magnetite oxidation and decrease of compression strength of preheated pellets [9, 10].

I

Ô

Dodage of the MHA /%

Fig. 4 Effect of dosage of MHA on the quality of preheated balls Effects of Binders on the Strength of Roasted Pellets In the roasting experiment, preheated condition was fixed at preheating temperature 950 °C, preheating time 10 min. Effect of Binder Variety

Bentenite20% Pedir 0,15% NMAO.50% Type and dosage of bi nder

Fig. 5 Effects of binder variety on the compression strength of roasted pellets

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At the roasting conditions of roasting temperature 1250 °C and roasting time 10 min, the effects of variety binders on the strength of roasted pellets were tested. The results were shown in Fig. 5. Seen from Fig. 5, the effect of binder variety on the compression strength of roasted pellets is not obvious. And the compression strength of roasted pellets with different binder is all higher than 2500 N/P, which meets the demands of industrial production. Effect of MHA Dosage The effect of dosage of the MHA binder on the compression strength of roasted pellets was studied under the conditions of roasting temperature 1250 °C and roasting time 10 min. The experimental results are shown in Fig. 6.

Dosage of the MHA/%

Fig. 6 Effects of MHA dosage on the strength of roasted pellets Seen from Fig. 6, it can be concluded that the compression strength of roasted pellets gradually decreases with increasing of MHA dosage. The compression strength is reduced to 2563 N/P when the MHA dosage is 1.0%. However, the compression strength of all roasted pellets in the range of 0.25%~1.0% MHA is higher than 2500 N/P, and meets the demands of large-scale blast furnace ironmaking. Effects of Binders on the Iron Grade of Roasted Pellets The effects of binders on the iron grade of finished pellets were also investigated in this paper. Qualified finished pellets with different binders are obtained under the condition of preheating temperature 950 °C, preheating time 10min, roasting temperature 1250 °C and roasting time 10 min. The iron grade of finished pellets with different binders was analyzed and presented in Table VI.

285

Table VI. Iron Grade of Finished Pellets with Different Binders Binders Dosage /% TFe /% Bentonite 2.00 54.25 Peridur 0.15 55.29 0.50 MHA 55.15 From Table VI, it can be seen that the TFe grade of pellet with the MHA is 0.9% higher than that of the bentonite pellets. Conclusions Compared with inorganic bentonite, the MHA binder can significantly improve the strength of green pellets. Adopting the wet-grinding process, the dosage of the MHA binder can be decreased to 0.25%. Because of better heat endurance, the MHA binder preheated pellet has a higher strength than the Peridur pellet under the same preheating conditions. However, increasing the MHA dosage is also unfavorable for the improvement of compression strength of roasted pellets. The TFe grade of 0.5% MHA roasted pellets is 0.9% higher than that of the 2.0% bentonite pellets. The qualities of pellets made with MHA binder meet the requirements of pellet production by grate-kiln process. Acknowledgements The authors want to express their thanks to National Science Fund for Distinguished Young Scholars (No.50725416), National Natural Science Foundation of China (No. 50804059), National Key Program of Science and Technology (No.2008BAB32B06), and the Graduate Degree Thesis Innovation Foundation of Hunan Province and Central South University (No.CX2010B063) for financial supporting of this investigation. References 1. C.Y. Li "Discussion on Increasing Grade of Pellets and Comprehensive Economic Results," Sintering and Pelletizing, 1998, 23(6): 10-13. 2. Y.F. Huang et al., "Investigation on Sodium-modification of Ca-based Bentonite via Semidry Process," Journal of Central South University ofTechnology (English Edition), (in press) 3. Y.B. Yang et al., "Study on Preparation of Oxidized Pellet by New-style Organic Binder," Journal of Central South University (Science and Technology), 2007, 38(5):850-856. 4. D.L. Murr, D.J. Englund, "Development and Production of Organic-limestone Pellets" (Paper presented at the 56th Iron-making conference proceeding, 1989), 715-732. 5. T. Jiang et al., "A Complex Type from Iron Ore Pellets Organic Binder and Its Usage" (Patent 200910309383 in China). 6. X.S. Wang, "Discussion on the Direct Reduction Technology of V/Ti-bearing Magnetite,"

286

Panzhihua Sci-Tech & Information, 2005, 30(l):3-8. 7. J. Deng, X. Xue, and G.G. Liu, "Current Situation and Development of Comprehensive Utilization of Vanadium-bearing Titanomagnetite at Pangang," Journal of Materials and Metallurgy, 2007, 6(2): 83-86. 8. GH. Han et al., "Adsorption Behaviors of Humic Substances onto Iron Ore Particle Surface, (Paper presented at the XXV Intemaltional Mineral Processing Congress, Brisbane, 6, September, 2010), 163-171. 9. T. Jiang et al., "Effects of Composite Binder (CB) on Oxidation Behavior of Iron Ore Pellets" (Paper presented at the 2010 TMS Congress), 373-381. 10. J.Y. Fu, T. Jiang, and D.Q. Zhu, Sintering and Pelletizing (Changsha: Central South University Press, 1996).

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Atetáis & Materials Society), 2011

CONSTITUENTS AND POROSITY OF LEAD CONCENTRATE PELLETS PRODUCED IN THE TREPCE PLANT Ahmet Haxhiaj1' and Jaroslaw Drelich2 'Faculty of Mining and Metallurgy, University of Pristina, Mitrovica, Republic of Kosovo department of Materials Science and Engineering, Michigan Technological University, Houghton, MI 49931, USA

Keywords:

agglomeration, lead sulfide, smelting

Abstract Both composition and porosity of pellets are the main parameters influencing the reductive melting process in water-jacket furnace during the roasting of sulfide lead concentrate. Roasting in a smelter located in Trepçê is done at about 900°C, for pellets which are composed of PbS, CaCÛ3, Fe2Û3, and SÍO2. In order to faciltiate the formation of slag in the reductive melting, the pellets must have the proper content of CaO, SÍO2, and FeO, and porosity. The plant's practice indicates that pellets with porosity of 18% do not melt adequately in the water-jacket furnace. The optimum porosity for the proper melting of pellets needs to be 32-40% and can be achieved when pellets are loaded with more than 10 wt.% CaO. Pellets with less than 10% CaO have a limited network of pores, such that reactive gases cannot reach the entire ore structure. Also, the composition of the pellets must satisfy the ratio of CaO : SÍO2 : FeO = 1 : 1.2 : 1.8 to facilitate the formation of slag in the reductive melting process.

Introduction Agglomeration roasting of sulfide lead concentrate is an important phase in the production of lead by reductive melting [1-6]. The reductive melting process is a complex process with contradictory effects of several technological indexes which many times are hard to measure and determine. In this process, loading the water-jacket furnace with pellets of optimized composition and structure can reduce the consumption of coke, and achieve high level of lead recovery. Pellets (agglomerates) should be porous, but at the same time they should have necessary mechanical strength. 'Self-melting' ability of pellets is controlled by the types and quantities of fluxes blended into the pellet composition. Three major additives and their role are: Calcite (CaCÛ3) - absorbs the heat released during the oxidation of sulfide and lowers the temperature of the load to keep it in a friable-porous state;

* Corresponding author: [email protected]

289

Iron oxide (Fe203) - plays the role of thermal regulator, serves as de-sulphurizing agent, and also lowers vaporization of lead oxides; and Quartz (SÍO2) - is a thermal regulator in the oxidation of sulfldes, and if contacted with lead oxide (PbO) reacts to form silicate: PbO + Si02 = PbSiO¡. In this communication, we briefly review the analysis of lead concentrate agglomerates production, through which the increase in quality and porosity of pellets was accomplished at the plant in Trepçë. A schematic of the agglomeration plant is shown in Figure 1 and more discussion on the technology was presented in earlier publications [7-11],Over the last decade, the industrial and laboratory research has confirmed positive effects related to the control of fluxes on the quality and porosity of pellets, as discussed in the following sections.

Figure 1. Schematic of one level agglomeration roasting: 1-concentrate's bunker, 2-smelters' bunker, 3- bunker of reverted agglomerate, 4-dusts' bunker, 5-mixer of the load, 6-load's bunker, 7-bunker of the ignition layer, 8-bunker, supplementary load layer, 9-agglomeration machine, 10-ignition room, 11-agglomeration fissure, 12-bolter, 13-cylindrical mill, 14-furnace's bunker, 15-gasses rich with SO2, 16-removal of dust from rich gasses, 17-gas for the production of sulphuric acid, 18-removal of dusts from the gasses, 19-ventilator of gasses of recirculation, 20the ventilator of fresh air, 21-ventilator for initial combustion, 22-dust, 23-reverted agglomerate, 24-room of rich gasses with SO2, 25-room of the gas of recirculation, 26-preparation of reverted agglomerate.

290

In the first part of the paper, both the material balance of lead concentrate pellets and daily analysis of pellet composition are presented. In the second part, we discuss the influence of zinc on the addition of fluxing agents into the pellets. In the last segment of the paper, examples of pellet porosity measured and recorded daily during the industrial operation at Trepçë are presented. The porosity values are usually much below the recommended threshold ones of 3240%.

Material Balance for Agglomeration In this section, material balance for the lead concentrate agglomeration process is analyzed with the aim of controlling the process and evaluating the environmental and economical effects. Thermal balances were analyzed in refs [7, 9, 10] and will not be repeated here. The load of agglomerate for melting in water-jacket furnaces is defined by the quantity necessary for the production at Trepçë, which is: Gagi=535t/24h [7, 8]. Constituents of pellets include compounds of lead such as oxides, silicates and ferrites, metallic lead, and lead sulfide, as well as other minerals typical to the ore used [5]. The ore comes from the Trepçë mine located in southern part of Kopaunikut near Mitrovica. A schematic of the agglomeration technology is shown in Figure 1. The typical load includes: Pb (48.2 wt%), S (5.2 wt%), Si0 2 (9.1 wt%), FeO (12.6 wt%), CaO (7.7 wt%), ZnO (5.4 wt%), A1203 (2.2 wt%), but the concentration of constituents in pellets vary depending on the composition of the ore and amount of added fluxes. For example, the content of PbO, FeO, and S in agglomerates produced in the plant operation varied during one of the production shifts as shown in Table 1 and Figure 2.

Table 1. Example of the daily analysis of agglomerate content. Sample Collection Time (h) 11 12 13 Mean

%PbO

%FeO

%S

42.21 46.34 42.14 43.56

14.12 13.67 14.79 14.12

1.10 0.98 0.93 1.00

Figure 2. The chart of agglomerate content.

291

The quantity of each component (Gcompo„„¡) in agglomerate was calculated with the formula: Gcomponeni=Gagi x (concentration in agglomerate). Therefore, the typical quantities of components are: lead oxide (PbO): Gno=233t/24h; iron oxide (FeO): GFeo=75.3t/24h; and sulfur (S): Gs=5.35t/24h. Three oxides, including zinc oxide (GZM), silica (Gsm), and calcium oxide (Gcao) combine for the remaining content of agglomerates (GRemai„i„goXides), which is calculated using the formula: G Remaining oxides = Gagl — GpbO — GpsO — G$ = 221.

lt/24t.

The results of calculations are presented in Table 2.

Table 2. The content of agglomerate: PbO, FeO, S and the remaining oxides. Component PbO FeO S Remaining oxides

(%) of component 43.60 14.12 1.00 41.28

Quantity (t) 233.00 75.50 5.35 221.1

Optimum Amount of Fluxes Porosity and mechanical strength of pellets for reductive melting depends on the ratio of fluxes used and should be either CaO:Si02:FeO = 1:1.2:1.8 or 1.3:1.6:1.8 [5]. Over the years of research done in the plant at Trepçë, it was found that the application of CaO:FeO:Si02 = 1:1.2:1.8 ensures the production of porous agglomerates of sufficient integrity which melt well in the water-jacket furnace. A later section provides experimental details on porosity measurements. Therefore, the ratio of CaO:Si02:FeO = 1:1.2:1.8 should be used to optimize the amounts of fluxing oxides in the pellets to maintain correct porosity and melting characteristics. Examples of calculated concentrations necessary for production, for three cases from Table 1, are shown in Table 3. Using the average concentration values, the optimal amount of CaO in agglomerate should be: GCa0 =Gagi x %CaO=60.3 t/24h; iron oxide (FeO): GFe0 =Gagi x %FeO=75.3 t/24h; and silica (Si0 2 ): Gs,o2 =Gagix %Si02 =121.2 t/24n

Table 3. The optimum concentration of fluxing agents in agglomerates. Time (h) 11 12 13 Mean

%FeO 14.2 13.67 14.79 14.12

%CaO 11.3 10.94 11.84 11.3

292

%Si02 22.72 21.87 23.66 22.75

CaO and Si0 2 in Agglomerate: Plant Practice In this research, amounts of the fluxes in the plant operation were examined in order to find the ratio between calcium oxide, silica, and iron oxide. The pellets produced in the Trepçë plant were weighted on a special weighting machine and chemical analysis was conducted in analytical laboratory. Examples of results for a set of three measurements performed during one shift (8h), regarding the contents of CaO and SÍO2 are shown in Table 4 and Figure 3. Chemical analysis also included the content of lead and sulfur with average concentration of 42% and 1%, respectively (note that these values are different from the load composition). Finally, compositional analysis of slag was also conducted and results are presented in Table 5. Neither strength nor porosity of agglomerates were measured in this set of experiments. As shown by the data in Table 4 and Figure 3, the ratio of Si0 2 to CaO is about 2, much above 1.2 required for optimum quality of the pellets. This clearly indicates that the plant operation requires optimization regarding amount of fluxes to be added to the ore.

Table 4. The optimal quantity of CaO and SÍO2 in agglomerate. Sample

The quantity of agglomerate (t) 534 534 534

The quantity of CaO (t¡_ 60.342 58.419 63.225

The quantity S1O2 (t) 121.218 116.421 126.024

Figure 3. The relation between CaO and SÍO2 components.

293

Influence of Zinc Zinc is present in the lead concentrate agglomerate as ZnO, ZnS, and ZnSC>4 and its content can vary in the ore used. In the melting process that takes place in a water-jacket furnace, zinc sulfate undergoes reduction according to the following reactions: ZnSO4->ZnO+SO2+0.5O2 ZnS04 +4CO=ZnS+4C02 Zinc sulfate also dissolves partially in ZnS and ZnO. It is a very damaging chemical to the quality of agglomerates, and also enables the formation of sinter, an inter-product of the melting process in the furnace [5]. Zinc oxide, on the other hand, dissolves easily in a slag made of CaO, FeO, and SÍO2. Dissolution of zinc oxide increases with increasing content of iron oxide present in the slag (Table 5). The quantities of Si0 2 , FeO and CaO in the slag can be determined according to the content of ZnO in the slag. The ratio between main components of the slag (CaO, FeO, SÍO2) in the presence of ZnO are shown in Table 5.

Table 5. The chemical composition of slag as a function of the zinc oxide content (data from the production plant). No ZnO(%) 0.0 1 5.0 2 10.0 3 4 15.0 20.0 5

Si02(%) 34.20 30.62 27.57 23.47 19.89

FeO<%)

37.80 38.26 38.72 39.18 39.64

CaO(%) No 18.00 6 16.12 7 14.23 8 12.35 9 10.47

ZnO(%) Si02<%) FeO(%) 25.0 16.21 40.10 30.0 12.73 40.75 41.20 35.0 8.16 40.0 5.58 41.48

-

-

-

CaO(%) 8.59 6.70 4.32 2.44

50

,_^ 40 S! % 30 +■»

cOl c0

0

20 10

ZnO(%)

Si02(%)

FeO(%)

CaO(%)

Figure 4. Chemical composition of slag affected by the content of zinc oxide.

294

-

The data in Table 5 and Figure 4 show that with the increasing amount of zinc oxide in the slag, and also with the aim of achieving good conditions for slag formation, the quantity of iron oxide increases, although this increase is small. At the same time, the quantities of Si0 2 and CaO decrease and the changes are at a much larger rate than for FeO.

Porosity of Pellets The important objective of this research was to solve the problem of producing high quality agglomerates for the reductive melting in the Port-Piri furnaces [7, 10]. This includes examination of the effect of fluxes on quality of pellets, as discussed earlier. The pellets produced should not compromise the conditions of oxidation process of the load. In the technological process in the plant at Trepçë, it was found that the optimum amount of iron oxide (FeO) in agglomerates is 75.3t/24h (see section on mass balance). Iron oxide decreases vaporization of PbO and decreases the losses of lead with gases, making the process economically and environmentally sustainable. The optimum amount of quartz SÍO2 in agglomerate is 121.2 t/24h, as presented earlier. This quantity enables the formation of lead silicates, contributes to the agglomeration process of roasted load and positively affects the agglomerate porosity. Fragile pellets are those poorly enriched with CaO, FeO, SÍO2, especially with SÍO2. In this case, the roasted material has no opportunity to agglomerate. Several years of industrial and laboratory research conducted in a lead smelter in Zveqan, Mitrovica, indicates that the optimum porosity of agglomerates should be 32-40%. In this study, several samples of agglomerates were picked from the production line in the Trepçë plant and their porosity was determined volumetrically to examine whether porosity of pellets is close to optimum value of 32-40%. The laboratory set-up is shown in Figure 5. Agglomerates before and after melting were hung on strings and then immersed in water in graduated cylinders. The porosity was determined based on the difference in volume of raw and molten agglomerates. Table 6 and Figure 6 present the results of the porosity of lead agglomerate monitored during the production in Trepçë. It was found that pellets produced in the Trepçë's plant had reduced porosity, especially the ones that had less than 10% CaO. CaO absorbs the heat released during the oxidation of sulfide and lowers the temperature of load to keep it in friable-porous state. The pellets were too dense, with pores of small dimensions. Such solid pellets do not melt appropriately. Only pellets with more than 10 wt% CaO melted correctly in the water-jacket furnace. This research clearly indicates that control over the composition of pellets needs to be improved in the plant at Trepçë. Current practice reveals significant fluctuation in quality of the pellets, with porosity usually remaining below threshold of optimum values, 32-40% (Figure 6).

295

Figure 5. The equipments used in measurements of the lead agglomerate porosity (1-graduated cylinder, 2-crucible for melting, 3-agglomerates attached to strings) Table 6. Results of porosity measurements. Sample

Porosity

Sample

Porosity

Sample

Porosity

17.66 11.20 18.73 19.16 17.10

6 7 8 9 10

37.01 16.88 27.00 22.14 25.35

11 12 13 14 15

26.30 12.01 14.50 24.49 12.30

%

1 2 3 4 5

x

%

%

On average, the pellets consisted of: PbO 44.2 wt%, ZnO 4.98 wt%, FeO 15.21 wt%, Cu 0.86 wt%, CaO 9.98 wt%, Si0 2 10.81 wt%, S 0.69 wt%, A1203 3.02 wt%, MgO 1.80 wt%, As 0.38 wt%, Sb 0.23 wt%, Bi 0.07 vrt%, plus remaining rare and precious metals. 40% 35% >

30%

O

25%

Q-

20%

15% 10% 1

2

3

4

5

6

7

8

9

10 11 12 13 14 15

Observations

Figure 6. The porosity of lead agglomerates. Broken line indicates the median of the experimental results.

296

Concluding Remarks Based on the analytical work of lead agglomerates produced in the Trepçë plant and calculations made, we suggest that in order to produce good quality agglomerates, detailed chemical analysis of agglomerate composition and the oxidation process of lead concentrate are necessary. In this paper, we discussed the effects of fluxes on the quality of agglomerates for efficient melting in water-jacket furnaces. In current industrial operation at Trepçë, the porosity of agglomerate remains at 18% (median), but this porosity does not enable efficient melting in the water-jacket furnace. The industrial and laboratory research conducted in the last decade suggests that the optimum porosity of the agglomerates should be kept at 32-40% and can be achieved with loading of more than 10 wt% CaO which also ensures efficient melting of pellets. The agglomerates with less than 10 wt% CaO are small, with not enough pores for the gasses to penetrate the load. The optimum quantities of fluxing agents should follow the ratio of CaO : SÍO2 : FeO = 1:1.2:1.8, or close to it, to ensure optimum porosity of agglomerates, amounts which also facilitate the formation of slag in the reductive melting process. The above mentioned factors could enable economical and environmental sustainability in waterjacket furnaces during the production of technical lead in Trepçë.

References [I] Hofrnan HO. Metallurgy of Lead. New York and London: The Scientific Publishing Company, 1899. [2] Ingalls WRE. Lead Smelting and Refining. New York and London: The Engineering and Mining Journal, 1906. [3] Vesely V, Hartman M, Svoboda K, Mracek J. Roasting of zinc sulfide, lead sulfide, copper sulfide and iron sulfide. Chem. Listy 1991;85:9. [4] Battle TP, Hager JP. Viscosities and activities in lead-smelting slags. Metallurgical Transactions B-Process Metallurgy 1990;21:501. [5] Agolli F. Metalurgjia e Metaleve me Ngjyrë. Prishtinë, 1985. [6] Beilstein D. The Herculaneum Lead Smelter Lead and Zinc. New York, 1970. [7] Haxhiaj A. Materialna i Toplotna Bilansa PORT-PIRI Peci. Zagrebu, 1989. [8] Haxhiaj A. Bilanci i Nxehtësisë ne Zonen e Materialit të Shkrirë ne Furrat Water-Jacket. Prishtinë, 2004. [9] Haxhiaj A, Rizaj M, Murati N, Ibrahimi S. Thermal balance in the melting zone of the water-jacket blast furnace. Kërkime 2004;12:145. [10] Haxhiaj A, Elezi D, Gashi Z. Balance of heat on reactions of the lead zone in Port-Piri furnace. Acta Chimica Kosovica 2005;14:85. [II] Haxhiaj A, Elezi D. Ambience and drill of the thermal gas in pre-heating zone of waterjacket blast furnace in Trepçë. Kërkime 2009;17:263.

297

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

OXIDIZED PELLET PREPARATION FROM REFRACTORY SPECULARITE CONCENTRATES USING MODIFIED HUMIC ACID (MHA) BINDERS Guohua Bai, Daoyuan Zhang, Yuanbo Zhang *, Guihong Han, Zijian Su (School of Minerals Processing and Bioengineering, Central South University, Changsha, Hunan, 410083, China) Keywords: specularite concentrate; organic binder; oxidized pellets; humic acid Abstract Specularite ores have poor hydrophilicity and ballability as well as inferior high-temperature reactivity, which restrict the large-scale application of them in the pelletizing and sintering production. Modified humic acid (MHA) organic binders are firstly used for pellet preparation from specularite in this paper. The results show that the drop strength of green pellets is 3.7 times/0.5m and compression strength is 12.5 N/P by using 1.0 wt% MHA as binder. Under the optimal experimental conditions of preheating temperature 980 °C, preheating time 12 min, roasting temperature 1280 °C and roasting time 10 min, the preheated pellets have the compression strength more than 400 N/P, and that of roasted pellets is 2747 N/P. The strength of finished pellets meets the requirements of blast furnace iron-making. MHA is a kind of more effective organic binder for specularite pellets compared with the inorganic bentonite. Introduction Acid Pellets have advantages of good strength, uniformity particle size, high iron grade and good reducibility, which can be used as high-quality burdens for blast furnaces [1]. With fast development of Chinese iron and steel industry .domestic magnetite concentrates with high grade of iron and good firing performances become shortage. Imported specularite ores have advantages of low price and high grade of iron (generally above 67 wt%), so the application of them in pellets is an effective way for sustainable development of pellet industry in China. However, the spéculantes have not been utilized in pelletizing and sintering at present because of the poor ballability and high-temperature reactivity [2, 3]. Most of the pellet plants use bentonite as binders, the amount of which is about 2~3 wt%. The pellet production practice shows that 90 wt% bentonite is still residual in finished pellets and reduces the iron grade [3]. Developing a new organic binder instead of bentonite is very important to the high-quality pellet production. On one hand, the consumption of bentonite is obviously reduced, resulting in decreasing the ironmaking coke ratio and slag discharge; On the other hand, the iron grade of finished pellets are improved, which brings the benefit for the Corresponding author: Dr. Yuanbo Zhang, E-mail: [email protected].

299

ironmaking production. One novel organic copolymer binder named MHA binder has been developed by Central South University. In this paper, this new binder was used to prepare oxidized pellets from specularite ores and the optimal parameters are obtained. Experimental Raw Materials The samples of specularite are taken from Brazil. Chemical compositions and particle size distribution of them are shown in Table 1 and Table 2, respectively. Specific surface area of specularite is tested as 421 cm2-g"'.

TFe 67.03

FeO 0.39

Table 1. Chemical compositions of specularite/wt% Si0 2 CaO P MgO AI2O3 2.02 0.20 1.02 0.074 0.017

S 0.013

Table 2. Particle size distribution of specularite Particle size/mm

> 0.196

0.196-0.104

0.104-0.075

0.075 - 0.045

< 0.045

Percentage /wt%

9.04

30.92

18.98

16.57

24.49

The data in Table 1 show that oxidation degree of specularite is very high. It can be seen from Table 2 that the proportion below 0.075 mm of the sample is only 41.06 wt%. Meantime, specific surface area of the specularite is 421 cm2-g"'. Therefore, the original specularite is very coarse and difficult for pelltizing. The MHA binder is a kind of novel organic copolymer, the main components of which are humic substance extracted from lignite. Chemical compositions and main functional groups of the MHA binder are analyzed and shown in Table 3 and Table 4, respectively. Table 3. Chemical compositions of the MHA binder /wt% Na 2 0 Fe K20 A1203 CaO MgO P S Si0 2 3.82 20.62 0.55 11.98 0.50 0.33 0.096 1.06 6.08 * LOI means loss of ignition, including organic matters.

Total acidic group/ mmol-g"' 5.37

LOI* 49.61

Table 4. Main functional groups of the MHA binder Carboxylic Phenolic Carboxylic acid group Phenolic hydroxyl hydroxyl group acid group to Total acidic group group to Total acidic /mmolg"1 /mmol-g"1 /wt% group /wt% 0.55 4.82 10.24 89.76

The FT-IR spectrum of MHA binder is shown in Fig.l. It can be seen from Fig.l that, notably, the absorption at 3398 cm"1 comes from stretching of -OH groups. Within the wide frequency

300

range from about 1550 to 1790 cm-1, principally assigned to protonate carboxylic (-COOH), and carboxylate anión (-COO-). It is noteworthy that there are the protonated -COOH groups or the ester carbonyl groups (-COOR) in the binder. Therefore, MHA binder can apparently improve the hydrophilicity of the spéculante so as to improve the strength of green pellet.

Figure 1. FTIR spectrum of the MHA binder Compared with bentonite, effects of MHA binder on quality of oxidized pellets were studied in this research. The proportion below 0.075 mm of the bentonite is 99.50 wt% and the content of montmorillonite is 81.72 wt%. Method The experimental procedure included ball-grinding combined with high pressure roller grinding, blending and damp-grinding of mixture, balling, drying, preheating, oxidized roasting, cooling, etc. In order to improve specific surface area and ballability of the specularite, ball grinding combined with high pressure roller grinding was carried out firstly. The specification of ball mill is 0460620 mm. The specification of high pressure roller is 0250x300 mm, roller speed 68 r/min and roll pressure 30 KN. For each batch balling test, 5 kg of the iron ore fine was mixed with the given amount of binder. Because of high viscosity of the MHA binder, the original mixture was ground in a cylindrical grinder with dimensions of 1000 mm diameter and 500 mm length. The grinding conditions were fixed as follows: the moisture content of raw materials 7.0 wt%, grinding time 5 min. Green pellets of 10 mm to 12 mm in diameter were prepared using a laboratory balling disc of 1000 mm in diameter, an edge height of 200 mm, and a tilt angle of 45°. Green balls were dried at the temperature of 105 °C. Preheating and roasting experiments were respectively conducted in an electric heated horizontal tube furnace. Then, the preheated and roasted pellets were naturally cooled to room temperature. Results and Discussion

301

Under the condition of ball grinding combined with high pressure roller grinding, specific surface area of the specularite with 1609 cm2g"' was obtained. The size distribution of the specularite particle is shown in Table 7. Table 7. Particle size distribution of specularite after pretreatment Particle size/mm

> 0.196

0.196-0.104

0.104-0.075

0.075-0.045

< 0.045

Percentage Avt%

0.80

1.25

3.10

35.45

59.40

As seen from Table 7, adopting the ball grinding combined with high pressure roller grinding; the content of the specularite below 0.075 mm can be increased to 94.85 wt%. Compared with the data in Table 2, the content of the specularite below 0.045mm is enhanced notably, which leads to the increase of specific surface area and balling index of the specularite. Effects of Dosage of The MHA on Green Pellets Adopting the MHA as binder for pelletizing, the effect of binder dosage on the quality of green balls was investigated. The results were given in Table 8. Table 8. Effect of dosage of MHA on the quality of green balls Thermal Shock Compression Drop Strength Dosage of MHA Strength/N-P"1 /(0.5m)-times"' Temperature/°C 0.5 wt% 0.75 wt% 1.0 wt% 1.25 wt% Bentonite 2.0 wt%

9.7 10.3 12.5 13.8 13.7

2.6 3.1 3.7 4.1 3.9

345 336 330 322 365

It can be seen from Table 8, the drop strength and the compression strength of green balls are increased with the increase of the MHA dosage. However, the shocking temperature is gradually declined slightly. When the dosage of the MHA is 1.0 wt%, the drop strength is 3.7times/0.5m and the compression strength is 12.5 N/P. Comparatively speaking, the drop strength is 3.9 times/0.5m and the compression strength is 13.7N/P for the pellets of 2.0% bentonite. The quality of the two kinds of pellets almost equal. Effects of The MHA Binder on the Preheated Pellets The effects of the MHA binder and bentonite on the strength of preheated pellets under different preheated temperature and time were studied and the results were plotted in Fig.2 and Fig.3. Under the preheating time 10 min, experimental results in Fig.2 shows that, with the increase of preheating temperature from 950 °C to 1020 °C, the preheated pellet compression strength is

302

gradually increased. The MHA pellets compression strength is only 272 N/P when the preheating temperature is 950 °C. When the preheating temperature increased 980 °C, the preheated pellet strength is up to 405 N/P.

Preheated TaqmatimTC

Figure 2. Effects of preheating temperature on preheated pellet strength

Preheated Tíme/nin

Figure 3. Effects of preheating time on the preheated pellet strength As given in Fig. 3, at the same preheating temperature of 980 °C, the preheated pellet compression strength also increase with the extending of preheating time. The MHA preheated pellet compression strength is 347 N/P when preheating time is 8 min. The compression strength increased to 474 N/P when preheating time is 12 min. Comparatively speaking, the strength of bentonite pellets is higher than that of MHA pellets under the same conditions. The quality of MHA pellets preheated at 980 °C for 12 min can meet the requirements of industrial production. Effects of the MHA Binder on the Roasted Pellets Under the same preheating temperature of 980 °C and time of 12 min, the effects of roasting

303

temperature and time on the strength of roasted pellets were researched and the results were shown in Fig.4 and Fig.5.

6, z

! s 3 Roasting Temperature/ °C Figure 4. Effects of roasting temperature on roasted pellet strength

i.

z a s e

35

I Roasting Time/min Figure 5. Effects of roasting time on roasted pellet strength As shown in Fig.4 and Fig.5, the compression strength is gradually increased with the increase of roasting temperature from 1250 °C to 1320 °C. The MHA pellet compression strength is 2107 N/P under roasting temperature of 1250 °C. When roasting temperature comes to 1280 °C, the MHA compression strength reaches 2747 N/P. Moreover, the compression strength of roasted pellet is also improved with the prolonging of roasting time. When roasting time is 8 min, the strength of roasted pellet is 2351 N/P, and it reaches to 2906 N/P when roasting time extends to 12 min. Comparatively speaking, the strength of roasted MHA pellets is a little lower than that of bentonite pellets under the same conditions. The main reason is that the content of low melting minerals produced in MHA pellets is less than those in bentonite pellets. So the bentonite pellets

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have closer connection with each other [5-7]. Conclusions The original spéculante ores are very coarse and difficult for pelletizing. Using wet ball grinding associated with HGRP, the specularite is pretreated. When the granularity of specularite below 0.075 mm is 94.85 wt%, specific surface area of specularite is 1609 cm2-g"', and 1.0 wt% MHA binder is used for pelletizing, the drop strength of green pellets is 3.7 times/0.5m and compression strength is 12.5 N/P. When preheating temperature 980 °C, preheating time 12 min, roasting temperature 1280 °C and roasting time 10min, the compression strength of roasted pellet is 2747 N/P. MHA is a kind of more effective organic binder for specularite pellets compared with the inorganic bentonite. The strength of finished pellets meets the requirement of iron-making production. Acknowledgements The authors want to express their thanks to National Science Fund for Distinguished Young Scholars (No.50725416), National Natural Science Foundation of China (No.50804059), National Key Program of Science and Technology (No.2008BAB32B06) and Specialised Research Fund for the Doctoral Program of Higher Education (No. 200805331080) for financial supporting of this investigation. References [1] X.H. Fan, Y. Wang,"Strength Enhancement of Oxide Pellet With Organic Binder", Journal of Iron and Steel Research, 20 (5) (2008):6-7 [2] D.Q. Zhu, Y.Y. Tang, "Pretreatment of Brazilian specularite before pelletization by high pressure roller grinding ", Journal of Metal M«e,382(4)(2008): 67-68 [3] G.X. Huang. "Applies new organic binder preparation oxidation pelletizing research", Changsha, Central South University, 2007. [4] Y.B. Yang, G.X. Huang, "Application of Organic binder as substitutes for bentonite in pellet preparation", Journal of Central South University, 38(5) (2007):851-857 [5] X.L. Chen, M. Gan, "Concretion properties of organic-binder oxidate pellets and strengthen measures", Journal of Central South University, 40(i)(2009):551-555 [6] Y.M. Chen, Y.B. Zhang, "Study on crystallization rule of oxidized pellet", Research on iron and¿fóe/,33(3)(2005):10-12 [7] Y.M. Chen, J. Li, "Crystal rule of Fe2Û3 in oxidized pellet", Journal of Central South University, 38(1) (2007): 71-73

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

2nd International Symposium on

High-Temperature Metallurgical Processing

Raw Materials Processing Session Chairs: Ismail Dunían Wei Li

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cooksey TMS (The Minerats, Metals & Materials Society), 2011

AN INNOVATIVE PROCESS ON BENEFICIATION OF SUPERFINE LOW GRADE HEMATITE ORE Deqing ZHU1, Yongzhong XIAO1'2, Tiejun CHUN1, Jian PAN1 1

School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, Hunan, PR China 2 Lianyuan Steel of Valin Group, Loudi 417000, Hunan, PR China

Keywords: superfine low grade hematite; reverse flotation; direct reduction; low intensity magnetic separation Abstract An innovative process of reverse floatation-direct reduction-low intensity magnetic separation was developed to effectively beneficiate the superfine run-of-mine (ROM) ore mined from Hunan Province, China. Mineralogy was measured that the ROM ore is of superfine low grade hematite ore type, assaying 27.2wt%Fe,otai and with main valuable minerals of hematite occurring at size between 3 to 5um. The upgrading results show that the final iron concentrate, assaying 88.3wt%Fe totai and 94.5% metallization degree was obtained at an overall iron recovery of 69.9% under the following conditions: rough concentration by grinding of ROM up to 88.7% passing 0.074 mm and reverse flotation at pH=9.4, 100 g/t starch and 200 g/t dodecylamine (DDA), and coal reducing the rough concentrate pellets containing 12% calcium containing complex additive at 1200 °C for 120 min, and finally magnetic separation of the reduced pellets by grinding up to 89.2% passing 0.043 mm at 0.08 T field intensity. Introduction With the rapid development of iron and steel industry in China, there has been a soaring demand for iron ores in the past decade. The imports of iron ores have increased from 148 million ton in 2003 to 628 million ton in 2009 [1]. Iron making industries increasingly demand for high grade raw materials to improve product quality and reduce operation cost. However, availability of high grade natural ores is limited and enrichment of lower grade ores is necessary to meet the demand. Therefore, large quantities of superfine low grade iron resources are necessary to utilize in China at the moment. The characteristics of superfine low grade iron ores are as follows [2]: low grade iron, complicated disseminated with gangue minerals and main valuable mineral hematite

309

occurring at superfine size. Hence, the superfine low grade iron ores cannot be concentrated successfully by conventional beneficiation methods such as gravity, flotation and magnetic separation, and also are not upgraded successfully by combination technology such as magnetic separation-flotation, flocculation-reverse flotation [3-4]. During the past years, the reports about the utilization of superfine low grade hematite are very few. However, extensive research has been carried out on the low grade hematite ores, using reverse floatation, gravity separation and high-gradient magnetic separation [5-6]. The reverse cationic flotation is one of the main processes used by the iron ore industry to produce hematite concentrate. In this process, quartz, the main gangue mineral, is floated with cationic collectors. In many iron ore beneficiation plants, dodecylamine (DDA) is generally used as collector for low grade hematite. Hematite is generally depressed with starch [7-8]. Direct reduction-magnetic separation is another way to treat low grade iron ores [3, 9], and the results show that this technology is possible to obtain high grade and low impurity product. In the direct reduction Fe 2 0 3 of hematite can be reduced to the magnetic product (Fe). In addition, low intensity magnetic separation is the most effective way for concentrating the magnetized ores [10-11]. Therefore, the new technology of reverse flotation-direct reduction-magnetic separation was studied to upgrade iron from superfine low grade hematite ores. Experimental Raw Material The ROM superfine low grade iron ore samples were provided by Lianyuan steel company, Hunan Province, China. The size distributions of samples are given in Table 1. Chemical compositions analysis (Table 2) indicates that the iron content of ROM ore is only 27.2wt%Fe. Therefore, upgrading of iron has to be conducted to manufacture good concentrate for iron and steel industry.

Size/mm Content

Table 1. Size distributions of ROM iron ore sample (wt%). O.074 0.074—0.18 0.18—0.5 34.44 38.63 17.69

0.5—1 9.24

Element Content

Table 2. Chemical composition of ROM iron ore sample (wt%). TFe FeO Si0 2 A1203 CaO MgO Na 2 0 P S 27.23 2.91 49.51 3.72 1.45 1.34 0.13 0.063 0.025

Items Content

Table 3. The main mineral content of ROM iron ore sam pie (wt%). Hematite Limonite Quartz Chorite Illite Ankerite 31.9 2.1 47.4 10.2 5.3 2.6

310

LOI 3.97

Others 0.5

Table 3 demonstrates the main minerals of ROM iron ore. Hematite is the main iron mineral and the main gangue mineral is quartz. Other minerals include limonite, chlorite, illite and ankerite. The mineralogical results (Figure 1) present that the iron ore sample is complicated disseminated with gangue minerals and most hematite mineral occurring at size of between 3 to 5um. Therefore, the superfine low grade iron ores cannot be concentrated successfully by conventional beneficiation methods. The complex additive is the off-white power containing some calcium. The Qitai coal is used as reductant coal and its size 100% passing 1 mm. Its industry analysis and chemical compositions are reported in Tables 4 and 5. Table 4. Industrial analysis of reductant coal. Items Content

Element Content

w (FCad)/%

w(Mad)/%

w (Ad)/%

w (Vdaf)/%

53.80

16.06

5.44

32.21

Index of coking 2

Table 5. Chemical composition of reductant coal (wt%). TFe P CaO MgO S A1203 Si0 2 0.83 0.43 0.98 0.66 0.60 0.25 0.016

LOI 96.02

Figure 1. Super grain hematite (white) disseminate into gangue minerals at size of between 3 to 5um. Experimental Procedure The experimental flow sheet is depicted in Figure 2, including main procedures, such as reserve flotation of ground ROM ore, pelletizing of rougher concentrate, direct reduction of pellets, and low intensity magnetic separation of reduced pellets to manufacture final concentrate. The reverse flotation was used as follows: Finely ROM iron ores of 100g was mixed with suitable level of water to make slurry, the slurry was then added into flotation cell (0.5L). The flotation machine was started to stir the slurry for mixing for 2 min. Then the

311

suitable pH of the slurry was modified by adding NaOH solution and stirring for 3 min. Adding proper dosage of depressant starch solution was followed and stirring for another 3 min. When the suitable collector dosage (dodecylamine solution) was added for 3 min, the flotation began to control froth baffle time at 7 min and the rougher concentrate was obtained. The flotation equipment is a single cell of XFD-63 mechanical flotation machine. ROM ore

V

Blending

T

Wet grinding '[ Reverse flotation Rougher concentrate Tailings Pelletizing Drying Coal

Direct reduction ,

*

■

Wet grinding

T

Magnetic separation

I

Iron concentrate

I

Tailings

Figure 2. Test flowsheet of upgrading iron from superfine low grade iron ores. Pellets were made from the mix of rougher concentrate and complex additive in a disc pelletiser of 0.8 m in diameter and 0.2 m rim depth, using rotational speed at 38 rpm and inclined at 47° to the horizontal. The green pellets were transferred into the oven to dry at 105°Cfor2h. The direct reduction was done as follows: Dry pellets were covered with proper mass

312

coal (mass ratio of coal-to-pellets was 2 in order to keep enough reducing atmosphere) in the stainless steel pot (1/3 overall coal in the pot bottom and 2/3 overall coal on the dried pellets). The steel pot was then put into the reduction furnace while the reducing temperature was elevated to the target value. When the reduction time was ended, the hot steel pot was taken out and covered by pulverized coal to cool down, and prevent pellets from being re-oxidized. High temperature muffle furnace, controlled automatically, was used in this study. The magnetic separation was performed in the XCGS-73 Davis Tube after 20g reduced pellets were finely ground in the cone ball mill of XMQ240x90. The final concentrate was obtained after magnetic separation. Results and Discussions Reverse Flotation The rougher concentrate, assaying 33.7wt% Fe and 39.4wt%SiOî were prepared at the iron recovery of 78.4%, yield of 63.3% under the conditions of at pH 9.4, 1000 g/t starch and 200 g/t dodecylamine (DDA) and grinding fineness of 88.7% passing 0.074 mm. Using the reverse floatation, iron content was uspgraded and tailings were discarded. It is important to reduce the feedstock of direct reduction by the pre-concentration technology of reverse flotation for cutting investment and operation cost in the further process. Direct Reduction Direct Reduction Temperature. The effect of direct reduction temperature on the metallization rate of reduced product is presented in Figure 3. As the temperature climbs from 1000°C to 1200°C, the metallization degree of reduced pellets increases from 85.1% to 95.5%. The reason is probably that increasing temperature can promote the direct reducing reaction. However, the metallization declines when the temperature was increased above 1200°C. This can be explained that if the reduction temperature was above 1200°C, resulting in lots of liquid phase on the surface of pellets and preventing the reducing reaction due to the more silica content of pellets [10]. Direct Reduction Duration. Figure 4 illustrates the effect of direct reduction time on the metallization degree of reduced product. As reduction time increases from 40 min to 120 min, the metallization degree of reduced pellets with complex additive climbs from 86.3% to 96.2%. The metallization degree drops to 93.5% when reduction time extends to 140 min. Therefore, the suitable time is recommended about 120 min.

313

g

i

Reduction temperature/ °C

Figure 3. Effect of direct reduction temperature on the metallization of reduced product (containing 12% complex additive and reduction for 100 min).

1 f .8

1 Reduction duration/min

Figure 4. Effect of direct reduction time on the metallization of reduced product (containing 12% complex additive and reduction at 1200 °C). Dosage of Complex Additive. The effect of dosage of complex additive is shown in Figure 5. As the dosage of complex additive adds from 0% to 12%, the metallization of reduced pellets is elevated significantly from 88.4% to 96.2% and then keeps steady when the dosage of complex additive is beyond 12%. The complex additive can effectively improve the reduction of pellets and the dosage of complex additive is suggested at 12%.

314

Dosage of complex addti ve/wt%

Figure 5. Effect of dosage of complex additive on the metallization of reduced pellets (reduction at 1200°C for 120 min). Magnetic Separation Grinding Fineness. The influences of grinding fines on the quality of final iron concentrate can be found in Figure 6. The iron content increases with increased amount of fines and reaches the peak value of 88.3wt%Fe at about 90% particle size passing 0.043 mm. However, the iron recovery keeps steady when the particle size passing 0.043 mm is less than 90%. The iron content and iron recovery decline markedly when the content of fines is above 90% passing 0.043 mm. It can probably be explained that at proper contend of fines, metal irons are liberated from gangue minerals, resulting in higher iron grade. However, too fine ore causes sliming, leading to lower selectivity during magnetic separation. For example, fine particles of gangue minerals containing metal iron were collected into concentrate, and some biggish slimed particles of metal iron and gangue minerals were collected into tailings, leading to lower iron content and recovery. Therefore, the proper grinding is suggested to produce about 90% particle size less than 0.043 mm. Magnetic Field Intensity. Figure 7 presents the effects of magnetic field intensity on the quality of final iron concentrate. The results of iron content appear to contradict the results of iron recovery with the increasing magnetic field intensity. When the magnetic field intensity rises from 0.06T to 0.12T, the iron content drops from 89.5wt% to 78.9wt% and the iron recovery climbs from 83.9% to 91.8%. The better results of 88.3wt%Fe and 89.2% iron recovery were achieved at magnetic field intensity of 0.08T.

315

Fraction of fines /passing 0.043mm wt%

Figure 6. Effects of grinding fineness on magnetic separation (magnetic field intensity at 0.08T).

Magnetic field intensityfi'

Figure 7. Effects of magnetic field intensity on magnetic separation (grinding fineness of 89.20% passing 0.043mm).

Element Content

Table 6. Chemical composition of final iron concentrate (wt% . P TFe f'Cmetal Si0 2 A1203 CaO MgO Na 2 0 S 1.03 0.34 0.10 0.061 0.065 88.31 83.41 6.72 0.96

LOI 0.18

The chemical composition of final iron concentrate is shown in Table 6. The iron concentrate, assaying 88.3wt%Fe, 94.5% metallization degree was achieved. The final product can be used for steel-making by electric furnace to replace sponge iron or scrap steel, which is a real short flowsheet to make iron with coal as reductant instead of coke. No sintering and coking and blast furnace are required.

316

Conclusions (1) A rougher concentrate, assaying 33.7wt%Fe, yield of 63.3% and iron recovery of 78.4% was obtained after the ROM ores with 27.2wt% Fe and main valuable mineral hematite occurring at size of between 3 to 5 um was pre-upgraded by reverse flotation. (2) The iron concentrate, assaying 88.3wt%Fe, 94.5% metallization degree and with a recovery 89.2% (the overall iron recovery is 69.9%) is achieved under the conditions of grinding of reduced pellets to 89.2% passing 0.043 mm, and magnetic separation at 0.08T magnetic field intensity after the dried pellets with 12% containing calcium complex additive were reduced for 120 min at 1200 °C. (3) The innovative technology of reverse flotation-direct reduction-magnetic separation proposed in this paper, can not only beneficíate the superfine low grade hematite ores efficiently upgraded, but also directly manufacture iron burden from ROM ore for steelmaking by electric furnace to replace sponge iron or scrap steel. References 1. China iron & steel association, http://www.chinaisa.org.cn/, 2010. 2. Zhu, D Q and Zhai, Y, "Beneficiation of super micro-fine low-grade hematite ore by coal-based direct reduction-magnetic concentration process," Journal of Central South University (Science and Technology), 3(12) (2008), 1132-1133. 3. Xu, B, Zhuang, J M and Bai, G H, "Study on coal-based direct reduction process in utilizing the low-grade and hard-to-separate iron ores," Multipurpose Utilization and Mineral Resources, 7(6) (2001), 20-23. 4. Sun, B Q, "Progress in China's beneficiation technology for complex refractory iron ores," Metal Mine, 2(3) (2006), 11-13. 5. Li, Y C and Tao, Q Q, "Experiment research and production practice of ultra-lean magnetite ore beneficiation," Metal Mine, 10(11) (2001), 37-39. 6. Yuan, Z T, Han, Y X and Li, Y J, "Advancement and developing trend of iron ore concentrating," Non-ferrous Mining and Metallurgy, 22(5) (2006), 10-13. 7. Mattedi, V A and Oliveira, J F, "Adsorption of starch onto apatite and magnetic and their select flotation", VI Southern hemisphere meeting on mineral technology, vol.1 (2001), 271-274. 8. Paclovic, S and Brandao, P R G, "Adsorption of starch, amylase, amylopectin and glucose monomer and their effect on the flotation of haematite and quartz," Minerals Engineering, 16(2003), 1117-1122. 9. Zhu, Z Z and Zhang, B H, "An experimental research on coal-base reducing poor siderite to sponge iron," Journal of Chongqing University( Science and Technology), 21(2) (1998), 101-105. 10. Zhu, D Q, Guo, Y F and Qiu, G Z, "Catalyzing the direct reduction of cold-bound

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pellets from titano-magnetite concentrate," Journal of Central South University (Science and Technology), 31(3) (2000), 208-211. 11. Xiao, Y Z, Zhu, D Q and Chun, T J, "Upgrading of superfine low grade hematite Ores by reverse flotation direct reduction and magnetic separation process," XXV International Mineral Processing Congress, (2010) 1653-1657.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

CALCINATION BEHAVIOR OF SIVRIHISAR LATERITE ORES OF TURKEY Ender Keskinkilic1, Saeid Pournaderi2, Ahmet Geveci2, Yavuz A. Topkaya2 2

'Atilim University; Department of Materials Engineering; Incek, Ankara 06836, Turkey Middle East Technical University; Department of Metallurgical and Materials Engineering; Ankara 06531, Turkey Keywords: Calcination, Latérite, Ferronickel Abstract

This study investigated calcination behavior of one of the Turkish latérite deposits, which was recently found in Sivnhisar region. Representative Hmonitic latérite samples (1.26% Ni) taken from Yunusemre Karasivritepe and Kucuksivritepe location were first subjected to drying. Removal of chemically bound water and other volatiles were then studied, in detail. In the calcination experiments, temperature and time were the main experimental variables. Thermal treatment was conducted at the specific temperatures in 250 °C - 800 °C range. The weight losses due to elimination of chemically bound water and other volatiles were reported to be approximately 10 per cent of the weight of the ore. For the particle size used in the current work, 700 °C and 40 minutes were determined to be the optimum calcination temperature and time, respectively. Introduction The nickel ores that are economically valuable can be grouped into two categories: sulphide type nickel ores and oxide type nickel ores, namely latérites. Over all reserves in the world, 40% are sulphide nickel ores and the remaining 60% are latérites [1]. Pyrometallurgical nickel extraction from sulphide type nickel ores is conducted via matte smelting. More recently, latérite ore bodies have been subjected to nickel extraction. Depending mainly on the chemical composition of the latérite and the process economics, various extraction techniques have been implemented: pyrometallurgical route involves ferronickel smelting. High pressure acid leaching (HPAL) is the commercial method for hydrometallurgical nickel extraction, while ammonia leaching (CARON Process) is a combination of pyrometallurgical and hydrometallurgical processes [2]. With the improvements in these processes, it is predicted that laterite's share in the world nickel production will rise to 50% in 2012, while it was reported to be 42% in 1,200,000 tons of annual production in 2003 [3], There are three main latérite deposits in Turkey. Two of them are in the region of Manisa, namely the Gordes and Caldag ores. The third was recently found in the region of Sivrihisar. Various hydrometallurgical & pyrometallurgical studies have been conducted for nickel extraction from Gordes and Caldag deposits [4-10]. In the first decade of the millennium, considerable amount of oxide type nickel ores were found in the Yunusemre and Mihaliccik locations near Sivrihisar, a town of Eskisehir. Since then, mining facilities have been continued in Yunusemre Karasivritepe and Kucuksivritepe peaks. Sivrihisar latérite ore (1.26% Ni) is a hmonitic one with its high iron content and low MgO composition. Low arsenic content of the ore makes ferronickel smelting a suitable method for nickel extraction.

319

In this study, calcination behavior of Sivrihisar latérite ores was investigated. Representative limonitic latérite samples taken from Yunusemre Karasivritepe and Kucuksivritepe location were first subjected to drying. Removal of chemically bound water and other volatiles were then studied, in detail. The results of these thermal treatment experiments provide useful information regarding the calcination characteristics of Sivrihisar latérites, which will hopefully be treated in the ferronickel plant in the Yunusemre region within the present decade. Experimental Four barrels of representative ore sample were taken from the Karasivritepe and Kucuksivritepe site. The barrels were weighed and the total amount of latérite was recorded. Approximately 400 kilograms of sample was thoroughly mixed. One representative quarter was split by the coning & quartering method. The other portions were stored in the barrels. The run of mine ore (ROM) sample prepared was then screened at 50 mm. The oversize lumpy materials containing chiefly gangue minerals were rejected and remaining undersize particles were used for experiments. The undersize particles used in the experiments were indicated as "-50 mm ore". Chemical analysis of -50 mm ore is illustrated in Table I. In the present work, the chemical analysis of the ore samples were performed with ICP method. Ni 1.405 AI2O3

3.23 K 0.2

Table I. Chemical Composition of-50 mm Ore (%) Fe Co Cr Si0 2 33.7 25.8 0.093 1.26 MgO MnO CaO Fe 2 0 3 1.29 0.74 48.2 1.65 Zn Pb S Ti0 2 0.03 0.01 0.08 0.03

As 0.04 P2O5

0.04 Cu 0.006

X-Ray diffraction analysis revealed that Sivrihisar latérites mainly contain quartz (Si0 2 ), goethite (FeO(OH)). Hematite (Fe 2 0 3 ), hisingerite (Fe 2 Si 2 0 5 (0H) 4 .2H 2 0), and some clay minerals are also present in considerable amounts. XRD analysis of the -50 mm ore sample is shown in Figure 1. In the calcination experiments, it was decided to use -50 mm ore which was crushed to -1 mm size. Screen analysis was conducted for dried sample. The particle size distribution of the ore used in calcination experiments is shown in Figure 2. An externally controlled muffle furnace was used in the calcination experiments. In each run, 100 grams of latérite ore (-1 mm size) were placed in a chamotte tray and charged to the furnace (kept at the experiment temperature). Experiments were carried out with temperature and time as experimental variables. Two different approaches were used. The first was done continuously until the charge weight became constant, called as "continuous to constant weight". The second was carried out with a specified time to evaluate the difference in the weight in a noncontinuous way. In the experiments conducted according to the first approach, the tray was taken out of the furnace periodically to record the weight and it was recharged immediately after weighing. In the experiments performed with the second approach, the tray was left in the furnace for a predetermined period. In both approaches, the furnace wall was opened systematically at specified times to perform mixing of the charge.

320

Figure 1. XRD Results for -50 mm Ore

Figure 2. Particle Size Distribution of the Ore Sample Used in Calcination Experiments With the first approach, experiments were performed at 250 °C, 300°C, 350°C, 400 °C, 500 °C, 600 °C, 650 °C, 675 °C, 700 °C and 800 °C until the change in weight became constant. In the experiments conducted with the second approach, which can be regarded as the reproduction of the first approach, the weight loss at the end of 40 minutes were determined and included in the scope of this work. Results and Discussion Drying experiments were carried out as the first stage of this study to determine the amount of physically bound water in the ore. The sample taken from the closed barrel of+50 mm fraction was left in the oven at 105 °C for one day. The weight loss, which represents the physically bound water, was 6.9%. The same procedure was applied for -50 mm ore and the amount of physically bound water was found as 13.6%.

321

DTA-TGA analysis was conducted by means of the differential thermal analysis device in Central Laboratory of Middle East Technical University. DTA-TGA curves were obtained by heating -20 mg of ground sample (<74 um) from 25 °C to 1000 °C with a heating rate of 10 °C/min. Differential thermal analysis and thermogravimetric behavior of the -50 mm ore is given in Figure 3.

Figure 3. DTA/TGA Curves of the - 50 mm Ore Sample Three main reactions were observed according to the DTA/TGA curves in Figure 3. The curves show three intense peaks at about 100 °C, 300 °C and 700 °C with corresponding mass losses of about 1.5%, 6.8% and 9%, respectively. The first endothermic peak at about 100 °C corresponds to the elimination of physically bound water. The second endothermic peak at about 300 °C corresponds to the dehydroxylation of goethite, transformation of goethite to hematite, which takes place according to Equation 1: 2 FeO(OH) -» Fe 2 0 3 + H 2 0

(1)

The third endothermic peak at about 700 °C most likely corresponds to dehydroxylation of clay minerals. Considering the results of DTA-TGA analysis, the first calcination experiment was conducted at 250 °C. In the experiment, the weight loss was found to continue for approximately 120 minutes. After the sample weight became constant, the weight loss was calculated as 5.88%. The variation of weight loss (%) with time is depicted in Figure 4.

322

I Time (min) Figure 4. Weight Loss (%) as a Function of Time at Various Temperatures The 300 °C experiment revealed that weight loss (%) increased with increasing temperature. The sample reached constant weight after 65 minutes, which corresponds to a 6.6% weight loss. Therefore, it was inferred that 250 °C was not sufficient for complete dehydroxylation of goethite for the particle size used in the current work. This result was expected and agreed with the DTA-TGA findings. Variation of weight loss (%) data at 300 °C and other temperatures are included in Figure 4. The 400 °C experiment produced a 7.15% weight loss; the 500 °C experiment produced 7.97% weight loss. Upon increasing temperature, continuous increase in weight loss was attributed to the removal of chemically bound water and other volatiles after dehydroxylation of goethite. In both experiments, the charge weight stabilized after 40-45 minutes. The sample weight was reached steady state after 40 minutes in the experiment conducted at 600 °C. The corresponding weight loss was 8.63%. The weight loss was found to increase further in the experiment carried out at 700 °C. This run yielded a weight loss value of 10.02%. The experiment performed at 800 °C produced a weight loss value of 10.04%, which was nearly equal to the one obtained at. From these results, it was concluded that chemically bound water and other volatiles in the ore were completely eliminated at 700 °C.

323

Although the weight loss values were nearly the same for both experiments, the time to reach to constant weight was considerably different. The sample weight reached a constant value after 60 minutes at 700 °C. On the other hand, it reached to steady state at the end of 45 minutes at 800 °C. An increase in temperature caused an increase in the rate of such reactions, so shorter time was sufficient at 800 °C, as expected. As a summary of these experiments, it was concluded that an increased temperature up to 700 °C increased weight losses and the elimination of chemically bound water and other volatiles were complete at 700 °C or at a temperature between 600 °C and 700 °C. Corresponding weight loss at the end of the thermal treatment was reported to be approximately 10% of the ore weight. Again considering the results of DTA-TGA analysis, an experiment was performed within the 300 °C and 400 °C range to precisely determine the temperature at which the goethite to hematite transformation goes to completion. Unless other reaction(s) overlapped goethite dehydroxylation, there should certainly be a distinct temperature at which weight loss data shows the behavior similar to the one observed at 700 °C and 800 °C. The 350 °C experiment at the midpoint of this interval gave a weight loss value of 6.67%, which was nearly the same as the one obtained at 300 °C. Therefore, it was inferred that goethite to hematite transformation goes to completion at about 300 °C. For the particle size studied in the current work, 300 °C was found to be nearly sufficient for dehydroxylation of Sivrihisar latérites. Also, the time needed to reach constant weight did not show considerable variation: the sample weight reached to steady state at the end of 65 minutes at 300 °C. This period was recorded as 60 minutes at 350 °C. In order to find the minimum temperature for complete elimination of chemically bound water and other volatiles, an experiment was conducted at the midpoint of 600 °C and 700 °C range. The weight loss value at 650 °C turned out to be 9.61%, which was greater than the weight loss at 600 °C but less than the one at 700 °C. Also, the time required to reach constant weight was reported as 95 minutes, which was considerably longer than the stabilization times at 600 °C and 700 °C. The replicate experiment at 650 °C revealed the same result. Therefore, it was found that 650 °C was not sufficient for complete removal of chemical bound water and other volatiles. A further experiment was performed at 675 °C. The weight loss was reported as 10%, which was nearly equal to the one obtained at 700 °C. The time to reach steady state was found to be 75 minutes. Therefore, it was concluded that a minimum temperature of 675 °C is necessary to completely eliminate the chemically bound water and other volatiles, assuming that sufficient time is allocated for this purpose. Experiments performed with the second approach revealed the similar results. The weight losses obtained from the second approach were found to be nearly the same as those reported from the first approach. However, slightly higher weight loss values were found at the end of the runs conducted with the second approach. This tendency was more pronounced in the lower temperature experiments. The weight loss values converged as the temperature increased. The variation of weight loss with temperature for both approaches are compared in Figure 5. The weight loss values were the ones obtained at the end of 40 minutes. The weight loss difference was the most prominent at the lowest experiment temperature, 250 °C. At the end of 40 minutes, the continuous to constant weight experiment led to a weight loss of 4.77%, whereas the experiment conducted with the second approach yielded 5.32%. The difference between the results is ~ 11.5%. The difference was 2.2% for 300 °C and 4.3% for 350 °C. At higher temperatures, the difference was found to stay in ± 2% band. The considerable

324

difference between the weight loss values at 250 °C was attributed to experimental procedure differences as well as the reversible behavior of goethite-hematite transformation reaction. In the first approach, as indicated previously, the sample was taken out of the furnace periodically for weighing. The sample taken out of the furnace does not show a significant temperature drop because the sample was not outside the furnace for a long time. On the other hand, the sample's interaction with the atmosphere, and therefore humidity, is longer in the first approach. Under these circumstances, it was inferred that goethite-hematite transformation reaction goes to reverse direction and the chemically bound water eliminated integrates back to the ore partially causing a lower weight loss value in the experiment conducted to constant weight. Although it was not given in the scope of the present work, the experiments conducted to evaluate the extent of reversibility showed that the weight of the calcined sample increased by a few grams when it was left outside for couple of hours. The change in the physical appearance of the ore also confirmed this behavior. At all temperatures, fawn-colored original ore converted to dark brown appearance after calcination. In all runs, it was observed that the surface of the calcined sample taken out of the furnace converted back to lighter color in a short time, which can be observed easily with naked eye. The final appearance is not the same as the original one but the color change is a probable indication of certain backward reaction(s).

j

f Temperature (C) Figure 5. Variation of Weight Loss (%) with Temperature for the Experiments Conducted with "Continuous to Constant Weight" and "Specified Time" Approaches (Calcination Time: 40 min.) In summary, it was found that complete removal of chemically bound water and other volatiles necessitates 75 minutes at 675 °C, 60 minutes at 700 °C and 45 minutes at 800 °C. Under these circumstances, 700 °C is recommended for effective calcination. The results of the experiments and the curves revealed that 98-99% of chemically bound water and other volatiles were removed in 40 minutes at 700 °C. This elimination was reported to require 25 minutes at 800 °C. These periods are ~ 33% and -45% shorter than the ones required for complete elimination at 700 °C and 800 °C, respectively. Considering the furnace heat requirements and the power costs, 700 °C and 40 minutes were determined as the optimum calcination temperature and time,

325

respectively. If the calcination process was to be performed at 800 °C, the recommended period would be 25 minutes. As indicated previously, die findings of this laboratory scale study reflect the results obtained with using "- 50 mm ore" and -1 mm particle size ground from this ore. The effect of particle size on calcination was not investigated in the scope of the present work. Conclusions Calcination behavior of Sivrihisar latente ores of Turkey was investigated. Drying experiments revealed that the amount of physically bound water was 6.9% for the +50 mm fraction, while it was reported to be 13.6% for -50 mm ore. The weight loss due to elimination of chemically bound water and other volatiles was approximately 10 per cent of the weight of the ore. For the particle size used in the current work, 700 °C and 40 minutes were identified as the optimum calcination temperature and time, respectively. Acknowledgements The authors would like to thank The Scientific and Technological Research Council of Turkey (TUBITAK) for the financial support given under the Project No: 109M068 and META Nikel Kobalt A.S. for supplying the lateritic ore samples of Sivrihisar. References 1. M.G. King, P. Zuluani, R. Schonewille, and P. Mason, "Technology Development for Processing Koniambo Latérites" (Paper presented at the ALTA 2005 Nickel/Cobalt Conference, Latérite Projects & Processes, Perth Western Australia, February 2005). 2. D.D. Ashok, W.G. Bacon, and R.C. Osborne, "The Past and the Future of Nickel Latérites" (Paper presented at PDAC 2004 International Convention, Trade Show & Investors Exchange, Toronto, Canada, March 2004). 3. A.E.M. Warner, CM. Diaz, A.D. Dalvi, P.J. Mackey, and A.V. Tarasov, "JOM World Nonferrous Smelter Survey, Part III: Nickel: Latérite", JOM, 58 (4) (2006), 11-20. 4. Yuksel, Melih, "Recovery of Nickel from Lateritic Caldag Deposit" (MS Thesis, Middle East Technical University, 1985). 5. Y.A. Topkaya, "Nickel Extraction from Lateritic Nickel Ores" (TUBITAK Project, Project No: 106M079,2009). 6. E. Buyukakinci, "Extraction of Nickel from Lateritic Ores" (MS Thesis, Middle East Technical University, 2008). 7. E. Buyukakinci and Y.A. Topkaya, "Extraction of Nickel from Gordes Lateritic Ore with Atmospheric Leaching" (Paper presented at the ALTA 2008 Nickel/Cobalt Conference, Perth Western Australia, June 2008). 8. E. Buyukakinci and Y.A. Topkaya, "Extraction of Nickel from Lateritic Ores at Atmospheric Pressure with Agitation Leaching", Hydrometallurgy, 97 (2009), 33-38. 9. V. Özdemir, "Ferronickel Production from Manisa Caldag Lateritic Ore" (Report, General Directorate of Mineral Research and Exploration, 2008). 10. C. Colakoglu, "Production of Ferronickel from Domestic Lateritic Ores" (MS Thesis, Istanbul Technical University, 2008).

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

FUNCTION OF HIGH PRESSURE ROLL GRINDING IN PRODUCTION of MAGNETITE OXIDIZED PELLETS Yufeng Guo, Haizheng Hao, Tao Jiang, Jianjun Fan School of Minerals Processing & Bioengineering, Central South University, Changsha 410083, China Keywords: High Pressure Roll Grinding, Oxidized Pellets Abstract Function of High Pressure Roll Grinding (HPRG) in production of oxidized pellets from magnetite concentrate as raw material was systematically studied. It was found that bentonite dosage decreased from 1.3% to 0.8% if specific surface area of magnetite concentrate pretreated by HPRG was up to 2553.41cm2.g"'. The green ball quality improved obviously due to the increase of specific surface area and fine particles content, and the improvements of size distribution and the ballability of pretreated materials. However, the decrease of pellet porosity accelerated the primary solid phase reaction and oxidation reaction, which made the pellets surface form compacted layer that handicapped the diffusion and migration of oxygen, promoting the formation of double layer structure of finished pellet. Consequently, the compression strength of preheated and roasted pellets was not markedly improved by HPRG.

Introduction With the increase of proportion of pellets in the BF recently, the technical-economic index of BF has obviously been improved. Consequently, the production and the quality of BF pellet are given more attention [1 l However, with the decrease of the quality of iron ore concentrate, the problems such as poor ballability, high bentonite dosage, high preheating and roasting temperature are becoming the great challenges to not only the quality of finished pellets, but also the energy consumption[2'3]. In order to overcome these problems, mechanical pretreating technologies such as the High Pressure Roll Grinding (HPRG) and Damp-grinding are widely applied. HPRG has a number of advantages such as low energy consumption, high productivity and easy-maintaining. It is more economic and environmentally friendly than the Damp-grinding technology [4' 5\ A great number of experiments have revealed that HPRG can improve the quality of hematite pellets. Although there are some studies of function of HPRG in the magnetite oxidized pellet production, they are not comprehensive and systematic, especially to the preheating and roasting system [1,2 ' 6' 7 l Accordingly, the influence of HPRG on the pelletizing system and the preheating and roasting system are systematically studied in

327

production of magnetite oxidized pellets in this article. Raw Materials and Research Methods Raw Materials The chemical composition, size distribution, specific surface area of the magnetite concentrate used in this work are given in Table 1 and Table 2, respectively. The physical properties of bentonite are given in Table 3. Table 1. Chemical compositions of magnetite concentrate /wt %

TFe

FeO

66.15

27.13

Si0 2 5.74

A1203 0.37

CaO

MgO

K20

0.63

0.42

0.034

Na 2 0 0.021

S

P

0.14

0.017

Ig

1.45

Table 2. Size distribution and specific surface area of magnetite concentrate Size distribution/% >0.074mm 2.27

0.074mm - 0.045mm 9.08

<0.045mm

Specific surface area /cm 2 .g ' 2161.70

88.65

Table 3. Physical properties of bentonite Colloid /%/(3g) 56.0

Swelling /ml/g 17.0

Moisture adsorption /% 333.90

Blue adsorption /g/100g 35.05

Montmorillonite

/% 79.30

Research Methods Ballability can comprehensively reflect the size distribution, specific surface area and hydrophilicity of iron ore concentrate. It is usually defined by the following formula: Kp=Wm/(Wc-Wm) Kp — Ballability index, W m —saturated molecular moisture/%,

W c —Capillary

moisture/%. The saturated molecular and capillary moisture are determined by using a press filtering and capillary water ascending methods, respectively [1,7] . The raw material was pretreated by HPRG with double rolls of 250mm in diameter and 120mm in width to improve its ballability, while the pressure of double rolls was IMP and the moisture of the raw materials was 8%. Effect of pretreating times, an important parameter of HPRG, on the quality of green, preheating and roasting pellet was investigated as well. The drop number and compressive strength were used to evaluate the quality of green ball, while the compression strength was the main index of the

328

quality of preheating and roasting pellet. The porosity used to evaluate the compactness of pellet was measured by the following method: 20 dry pellets were weighted (m,) and then covered with paraffin wax. The 20 pellets covered with paraffin wax were also weighted (iri2) and thereafter got the weight (1TI3) when they were suspended in the water. Finally, the pellet porosity was calculated by the following formula. e =

j__

mip,pwpz (p»-P,)m2+p!mi-pw

xlOO

£ — d r y pellet porosity (%), p,—paraffin wax density (g/cm ), pw—water density (g/cm3 ), pz—intrinsic density of dry pellet (g/cm 3 ) Results and Discussion Effect of HPRG on Quality of Green Pellet The effect of HPRG on the quality of green pellet was illustrated in Fig.l. The drop number of green pellet sharply increased with the increase of pretreating times of HPRG and the compressive strength was slightly improved as well. When the specific surface area of the pretreated materials was about 2553.41 cm2.g"' by HPRG for 3 times and the bentonite dosage was 1.3%, the drop number and the compressive strength of green pellet increased from 6.1 times/(pellet/0.5m), 10.5N/pellet to 14.7 times/(pellet/0.5m) 12.6N/pellets, respectively, compared with those of pellet of materials without pretreating.

Pretreating (times)

Figure 1. Effect of HPRG on the drop number and compressive strength of green pellet The improvement of the drop times and compressive strength of green pellet resulted from the substantial improvement of raw material characteristics after being pretreated by HPRG According to the influence factors of pelletizing, the most important one is the

329

iron ore concentrate characteristics on which the pelletizing directly depends, especially the ballability m . The experiment results were shown in Fig.2. The ballability index of raw materials was enormously improved by HPRG, and was up to 0.96 after being pretreated for 2 times. With the increase of pretreating time and ballability index, the hydrophilibicity of iron ore particles grew, which enhanced the particle-to-particle compaction interaction under the mechanical force of disc pelletizer. Therefore, the quality of green pellet was dramatically improved.

PianaongOims)

Pretnaing (liras)

Figure 2. Effect of HPRG on ballability index, content of-0.045mm fine particles and specific surface area of magnetite concentrate The second important factors on which the quality of green pellet depends are the specific surface area and size distribution of iron concentrate [8]. According to the mechanism of pelletization, when the material particle is finer and size distribution is better, the average diameter of capillary is smaller, resulting in better capillary action and particle-to-particle compaction. Therefore, the compactness of green pellet improved and the drop number and compressive strength enhanced. In addition, the content of -0.045mm is especially significant, which is crucial to the pelletizing process [7]. The results illustrated in Fig.2 and Fig.3 showed that after the raw materials was pretreated for 2 times by HPRG the content of -0.045mm and the specific surface area grew up from 88.65% and 1808.4cm2.g_1 to 94.11% and 2189.2cm2.g"' respectively, compared with those of raw materials without pretreating. Besides, the size range of iron ore concentrate became wider as well, resulting in the increase of pellet compactness.

330

Figure 3. Size distribution of pretreated concentrates (a)—pretreated materials by HPRG (b)—raw materials The particle morphology of magnetite concentrate is critical to the particle-to-particle contact area which influences the quality of green pellet, while the plate and flaws particle shapes are better than the cubic and spherical particle shapes, because the former shapes enhance the increase of particle-to-particle contact area during the pelletizing process [8 l The particle morphology of magnetite concentrate pretreated by HPRG was shown in Fig.4. It was seen that the particles of magnetite concentrate occurred as cubic shapes before pretreating, however, the particles of magnetite concentrate pretreated occurred as plate and flake shapes and had visible cracks and flaws, because HPRG was different from conventional crushers in breaking mechanism. The particles from conventional crushers were broken down singly by breaking medium, the particles from high pressure roller breaker were broken down by particle-to-particle interaction in a material layer under high pressure of rollers, which promoted formation of cracks and flaws on particles surface, and increased the specific surface area of raw materials [9' lo l Furthermore, the particle sharp edges and corners had seen a dramatic rise, which resulted in the increase of friction between particles and absorption of particle surface possessing higher energy [ n ] . Thereby, the improvements of the particle morphology and characteristics of particle surface of magnetite concentrate pretreated by HPRG all contributed to the improvement of green pellet quality.

Figure 4. Morphology of magnetite concentrates particles by SEM (a), (b)—materials without pretreating. Magnification of (a) and (b) is 1000_ and 5000_ respectively, (c), (d)—materials pretreated by HPRG. Magnification of (c) and (d) is 1000_ and 5000_ respectively. Comprehensively, HPRG improved dramatically the quality of green pellet and reduced the bentonite dosage. The results of experiment showed that when the specific surface area was about 2553.41 cm2.g"' after being pretreated by HPRG and the bentonite dosage decreased to 0.8%, the drop number and compressive strength of green pellet were up to 4.4 times/(pellet/0.5m) and 12.7 N/pellet, respectively, which were basically same to the quality of green pellet with 1.3% bentonite without pretreating. Therefore, the HPRG reduced the bentonite dosage substantially from 1.3% to 0.8%, which had

331

profound significance to production of high iron grade pellets. Effect of HPRG on Quality of Preheated and Roasted Pellet The results from Fig.5 showed that under the same preheating and roasting temperature and time, the compression of preheating and roasting pellets of the materials pretreated by HPRG did not increase markedly respectively, compared with that of materials without pretreating. When the specific surface area of magnetite concentrate was about 2553.41cm2.g"' by HPRG the compression of preheated and roasted pellets increased slightly from 439N/pellet and 2953 N/pellet to 442 N/pellet and 3066 N/pellet, respectively. Thereby, the HPRG did not improve markedly the performance of preheating and roasting pellets.

Pieh2atirgtençeraoire/*c

Roasting tençeratui^c

Figure 5. Effect of HPRG on compression strength of preheated and roasted pellet During preheating and roasting process, the magnetite pellet compression strength mainly depends on the oxidation of magnetite particles, formation and growth of solid bridges which can be dramatically accelerated by the improvement of pellet porosity and the particle-to-particle contact area of magnetite pellets [8]. The research showed that dry pellet porosity was reduced from 20.50% to 19.54% after the magnetite concentrate was pretreated by HPRG to increase the specific surface area to about 2553.41cm2.g"'. The increase of pellet compactness made particle-to-particle contact area see an obvious rise, which promoted the iron ions diffusion and migration, accelerating the primary oxidation reaction, but the accelerated oxidation reaction makes compaction layer of pellet surface more compact as illustrated in Fig.6, which handicapped the further diffusion and migration of oxygen [8' 12l

332

Figure 6. Macrostructure of finished magnetite pellets a - pellet without pretreating by HPRG b - pellet with pretreating by HPRG. Conclusions (1) HPRG can dramatically improve the quality of green pellet and reduce the bentonite dosage massively. When the specific surface area of magnetite concentrate pretreated by HPRG was about 2553.41cm2.g"', bentonite dosage was decreased from 1.3% to 0.8%, due to dramatic improvements in specific surface area, size distribution and fine particle content, particle morphology. (2) After the magnetite concentrate being pretreated by HPRG, the pellet porosity decrease and the pellet compactness increase, which accelerated the primary solid phase reaction and oxidation reaction, promoting the formation of double layer structure of finished pellet. Consequently, the compression of preheated and roasted pellets was not markedly improved. References [1] Huimin Li, "A Study of the Process and Mechanism of Improving the Pelletization of Brazilian Hematite by Adding Boron Bearing Compound Additives" (Master, thesis, Central South University, 2008),4-18. [2] Yanyun Tang, "A Study of the Process and Mechanism of Pelletizing PFC Specularite by Using High Pressure Roller Grinding"(Master. thesis, Central South University, 2008),26-50. [3] Deqing Zhu, Yanyun Tang, Jian Pan, Yong Zhai, "Pretreatment of Brazilians Specularite before Pelletization by High Pressure Roller Grinding," Metal Mine, 4 (2008), 67-92. [4] Changan Wang, Xinji Wang, Yimin Zhang, Tiejun Cheng, "Effect of High Pressure Roller Grinding on Activation Energy of Magnetite Materials of Pellet," Research on Iron and Steel, 1 (2005), 8-10.

333

[5] Lingzhi Li, "High Pressure Roller Grinding-a Effective Approach of Reducing Energy," Metal and Mining, 10 (2004), 251-284. [6] Changan Wang, "Study of Machinery Chemical Mechanism and Process of Improving the Production of Pellet" (Master, thesis, Central South University, 2005), 27-54. [7] Changan Wang, Shouan Xiong, Deqing Zhu, "Effect of High Pressure Roller Mill Pretreating Concentrates on Green ball Properties," Sintering and Pelletizing, 6(2002), 12-15. [8] Juying Fu, Tao Jiang, Deqing Zhu, Sintering and Pelletizing (Changsha, Central South University of Technology Press, 1996), 233-308. [9] Deshu Ren, "Principle of Particle Group Crushing and Application of Roll Press," Metal Mine, 12 (2002), 10-13. [10] Yunlong Li, Dianzuo Wang, Shensheng Huang, Shiying Zhang, "Application of High-pressure Comminution for Material-beds," China Ceramic Industry, 2 (2004), 29-31. [11] Danzuo Wang, Guanzhuo Qiu, Yuehua Hu, Minerals Processing (Beijing, Science Press, 2005),45-55. [12] Changan Wang, Xinji Wang, Chengfan Hu, Jincheng Huo, "Study on the Activation Mechanisms about Pretreating Magnetite Concentrate by High Pressure Roller Grinding," Sintering and Pelletizing, 2 (2005), 16-18.

334

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslav/ Drelich, Jerome Downey, Tao Jiang, and Mark Cook TMS (The Minerals, Metals & Materials Society), 2011

IMPROVING THE PELLETIZATION OF FLUXED HEMATITE PELLETS BY HYDRATED LIME Deqing ZHU1, Wei YU1'2, Tiejun CHUN1, Jian PAN1 School of Minerals Processing and Bioengineering, Central South University, Changsha 410083, Hunan, PR China 2 WISCO Minerals, Wuhan 430080, Hubei, PR China Keywords: hydrated lime; fluxed pellets; firing behavior; metallurgical performance Abstract In this paper, the technology of producing fluxed pellets by using hydrated lime as binder instead of bentonite was carried out. Green balls with 0.8% bentonite and others with 2% hydrated lime were obtained under the same conditions of 1.5% coke, 1.45 basicity and balling at 8.5% moisture for 13 min in disc pelletizer. The properties of two types green balls were as follows: drop numbers of 4.8 to 5.0 times/0.5m, compressive strength of 19.7 to 24.2 N/pellet and thermal shock temperature of 280 to 385°C, respectively. Comparing with bentonite as binder, the compressive strength of fired pellets with hydrated lime is elevated by 13.8% and climbs up to 3113 N/pellet under the following conditions: preheating at 1050°C for 10 min and firing at 1300°C for 15 min. The hydrated lime also improves the metallurgical performance of fired pellets. The reduction index was enhanced from 68.6% to 70.6%, RDI+3.i5 increased from 90.1% to 98.8% and reduction swelling index dropped from 22.4% to 7.4%, respectively. Therefore, hydrated lime can not only replace bentonite as binder to enhance the firing behaviors of green balls, but also effectively improve the metallurgical performance of fluxed pellets which can be used as high quality feed for blast furnace. Introduction Over the last 30 years, fluxed pellets have been slowly taking over the acid pellets as the primary burden of blast furnace around the world [1]. In fluxed pellets, limestone and sometimes dolomite, are added to the iron concentrate, which already contains silica gangue. By adding calcium (and magnesium) through the pellet burden, the blast furnace removes the need to charge limestone directly to the furnace and can more preciously control the distribution of slag forming components in the burden. Adding calcium flux to

335

iron ore pellets changes the chemistry of these liquid bridges from iron silicates to iron calcium silicates [2-3]. In producing fluxed pellets, MgO content significantly affects the firing performance of fluxed pellets and more MgO content leads to the lower compressive strength of preheated pellets [4-5]. With increasing the basicity of fluxed pellets, the metallurgical performance can improve noticeably [6]. Hydrated lime has been used as a binding agent for fluxed pellets in several plants, including Algorabo in Chile [7-8]. It is technically feasible to produce a fully fluxed pellet that has acceptable physical and metallurgical qualities using hydrated lime. Substitution of hydrated lime significantly decreases the total energy requirement for the process which provide direct cost savings [9-10]. When fluxed pellets are produced, bentonite is only used as binder to improve green balls properties, mainly in drop numbers, compressive strength and thermal shock temperature. However, hydrated lime acts as binder and flux, and fewer fluxes are required, which is helpful in improving the strength of preheated and fired pellets. Therefore, in this paper the use of bentonite and hydrated lime as binder to produce fluxed pellets was studied. Experimental Raw Materials The raw materials include the Brazilian iron ore fines, coke, bentonite, hydrated lime and limestone, all of which were supplied by Vale, Brazil. Their chemical compositionis shown in Table 1. It can be found that the iron ore fines are good quality for making pellets due to high grade assaying 65.5wt%Fe,otai, and low silica, alumina and other impurities, such as S, P, and nonferrous metals. The iron ore fines possess the static ballability index of 0.94 and the size of iron ore fine 75.8% passes 0.043 mm. Table 1 . Chem cal com position of materials (wt%). Elements Fetotal FeO Si0 2 A1203 CaO MgO Na 2 0 P Hematite 65.51 0.172 1.39 1.851 0.005 0.011 0.042 0.0315 — 51.82 21.77 0.30 2.60 1.42 0.0058 Bentonite 5.89 Hydrated 0.14 1.42 0.23 67.10 0.32 0.046 0.065 — lime limestone 0.42 — 1.49 0.45 52.22 2.10 0.048 0.0011 Coke 1.03 0.23 0.099 0.032 4.99 6.61 4.41 — breeze

S 0.012 0.023

LOI 1.635 13.20

0.084 29.33 0.060 42.62

—

79.41

The bentonite contains 90.5% montmorillonite, higher swelling volume of 57.5% and the water adsorption of 2.99 and fine size of 78.4wt% passing 0.043 mm. The hydrated lime

336

and limestone possess the CaO content of 67.1% to 52.2%, size of 76.8% to 80% passing 0.043 mm, respectively. Limestone was chosen as the flux to make sure 1.45 basicity of pellets. The coke breeze containing 75.2%FCad and the size of 85% passing 0.043 mm was used to reduce heat supplement from the combustion of gas fuel during preheating. The coke dosage of 1.5% was fixed in the experiment. Experimental Procedure The flowsheet of pelletizing tests in this paper include the traditional section, mixing iron ores with binder, limestone and coke breeze at a given ratio and basicity of 1.45, balling in a disc pelletizer and firing in a tube furnace to make fired pellets. The iron fine ores were pretreated by the HPRG of 250 mm in diameter and 120mm width at 300 rpm and in open circuit to improve its ballability by changing the passes through rollers. Result of the HPRG was evaluated by specific areas measurement in Blaine method. The moist iron fine ores was pretreated by HPRG at a normal feed rate of 30 kg/min. Fluxed Pellets were made from the mix of pretreated iron fine ores, binder and limestone in a disc pelletizer of 0.8 m in diameter and 0.2 m rim depth, rotational speed at 38 rpm and inclined at 47° to the horizontal. The drop numbers and compressive strength of the finished green balls were measured to evaluate the ability of the green balls to remain intact and retain their shape during handling, respectively. Dry balls were preheated and roasted in a tube furnace of 50 mm in diameter and 600 mm width in the experiment. The compressive strength and metallurgical performance of the roasted balls were measured to evaluate the quality of the roasted balls. Results and Discussion Balling Binder Dosage. The effects of binder dosage on green balls properties are shown in Figure 1. With an increase in binder dosage, the drop numbers and thermal shock temperatures sharply increase. However, the compressive strength of green balls reaches peak value at 0.8% bentonite to 2.0% hydrated lime and then remain unaffected The optimum binder dosage is suggested at 0.8% bentonite or 2.0% hydrated lime due to the drop numbers reach the requirement of 4.0 times/0.5m. However, the qualities of green balls containing 2.0% hydrated lime are superior to those with 0.8% bentonite under the same conditions of balling at 8.5% moisture for 13 min.

337

áentonite 8

1 2

1

'

l%dratáfl¡me

2 4

Binder dosage/%

Figure 1 Effects of binder dosage on green balls properties (balling at 8.5% moisture for 13 min). |

400

S

35

30

°

1 ° g 250

!

24

S

20

F^ g. 18

O

16

S 5

8.5 9.0 Balling moisture/%

338

Figure 2. Effects of baling moisture on green balls properties (balling for 13 min). Balling Moisture. As shown in Figure 2, drop numbers, compressive strength and thermal shock temperature of green balls increase with an increase in balling moisture from 8.0% to 8.5%, then the properties decline when the moisture increase above 8.5%. The optimum moisture takes place at 8.5% moisture, where drop numbers of 4.8 to 5.0 times, compressive strength of 19.7 to 24.2 N/pellet and thermal shock temperature of 280 to 385°C were achieved for green balls with 0.8% bentonite to 2.0% hydrated lime, respectively. Ball Duration. Figure 3 presents the effects of balling duration on the qualities of green ball. It indicates that the drop numbers of green ball with 0.8% bentonite and 2.0% hydrated lime increase when balling time increases from 10 to 13 min. It reaches peak values of 4.8 to 5.0 times/0.5m when balling for 13 min, then decline at 13 min to 17.5 min. However, the compressive strength of green ball keeps growing when balling was continued for 10 to 17.5 min. The thermal shock temperature of green ball drops while balling from 10 min to 13 min then remains unaffected when balling time increases from 13 to 17.5 min. Therefore, the optimum balling time is recommended for 13 min. Preheating and Roasting of Pellets Preheating Temperature. Figure 4 shows the effect of preheating temperature on the compressive strength of preheated pellets. With an increase in preheating temperature the compressive strength of preheated pellets climbs. The compressive strength of preheated pellets with 0.8% bentonite are higher than those preheated pellets containing 2.0% hydrated lime when the preheating temperature was less than 1050°C. However, the compressive strength for two types of pellets is close to each other when they were preheated to 1100°C.The preheating temperature for attaining the compressive strength of 749 to 557 N/pellet is 1050°C for preheated pellets with 0.8% bentonite and preheated pellets containing 2% hydrated lime, respectively. Preheating Time. Figure.5 demonstrates the effect of preheating time on the compressive strength of preheated pellets. It is shown that the compressive strength increases significantly as preheating time increased from 5 to 15 min. Enough preheating time is the requirement not only for many reactions to complete but also for the preliminarily bond to form. The compressive strength remains constant when preheating time is exceeded above 15 min. The compressive strength values of preheated pellets with 0.8% bentonite are superior to those for pellets with 2.0% hydrated lime when preheating at

339

1050°C.The suitable preheating time is recommended for 10 min due to the compressive strength can meet the requirement of 500 N/pellet for preheated pellets.

I

Balling time/min

Figure 3. Effects of baling duration on green balls properties (balling at 8.5% moisture).

Preheating temperatureTC Figure 4. Preheating temperature vs the compressive strength of preheated pellets (preheating for 10 min).

340

Preheating urne/min

Figure 5. Preheating time against the compressive strength of preheated pellets (preheating at 1050°C).

Firing temperaturas Figure 6. Effect of firing temperature on the compressive strength of fired pellets (preheating at 1050°C for 10 min and firing for 15 min). Firing Temperature. It can be seen from Figure 6 that with an increase in firing temperatures the compressive strength of fired pellets increases significantly. The compressive strength of fired pellets with 0.8% bentonite increases slowly when firing

341

temperature was above 1250°C. However, the compressive strength of fired pellets containing 2.0% hydrated lime climbs sharply when firing temperature is elevated from 1150°C to 1300°C. Comparing with the fired pellet containing 0.8% bentonite, the compressive strength of fired pellets with 2.0% hydrated lime is elevated by 13.8% and climbs up to 3113 N/pellet when firing temperature is 1300°C. In summary, the hydrated lime can improve the firing performance of fluxed pellets significantly at higher temperature. Firing Time. Figure 7 illustrates the effect of firing time on the compressive strength of fired pellets. The compressive strength of fired pellets increases when firing time extends from 5 to 15 rain but remains constant for 15-20 firing time. The compressive strength of fired pellets containing 2.0% hydrated lime is higher than the compressive strength of fired pellets with 0.8% bentonite when fired at 1300°C for 5 to 20 min. It can be concluded that the pellets comprising 2.0% hydrated lime possess the better firing performance.

I Firing ïmefirin Figure 7. Effect of firing time on the compressive strength of fired pellets (preheating at 1050°C for 10 min and firing at 1300°C). Table 2. Chemical composition of fired pellets (wt %). Element 0.8%betonite

Fe tot ai

FeO SiC-2 A1203 CaO MgO

61.80 3.60 3.23

1.70

4.68

342

1.47

Pb

Zn

S

P

0.001 0.003 0.002 0.006

2.0%hydrated lime

61.91 4.47 3.12

1.46

4.52

1.28

0.001 0.003 0.002 0.003

The chemistry of fired pellets is given in Table 2. Iron grade of fired pellets is higher than 61wt%Fetotai, and other impurities, such as sulfur, phosphorus and nonferrous metals are low, which is helpful to improve blast furnace performance. Metallurgical Performance of Fired Pellets It is seen in Table 3 that fired pellets containing 2.0% hydrated lime possess excellent metallurgical performance, such as reducibility index higher than 70%, reduction swelling index lower than 8% and RDI+315 higher than 98%, which are much better than those of fired pellets with 0.8% bentonite. Therefore, hydrated lime is superior binder as compared to bentonite in production of fluxed pellets.

Items 0.8%betonite

Table 3. Metallurgical performance of fired pellets. RSI RI

RDI+3.15

/%

/%

/%

68.6

22.4

90.1

98.8 70.6 7.4 2.0%hydrated lime ♦Footnote: RI- reducibility index ; RSI- reduction swelling index ; RDI-reduction degradation index Conclusions (1) Hydrated lime used to replace bentonite as binder not only increased the green balls properties and compressive strength of fired pellets, but also played an important role in improving the metallurgical performance of fired pellets. (2) Good quality green balls, the drop numbers of 4.8 to 5.0 times/0.5m, compressive strength of 19.7 to 24.2 N/pellet and thermal shock temperature of 280 to 385°C were obtained under the following conditions: binder of 0.8% bentonite to 2.0% hydrated lime, 1.5% coke and 1.45 basicity balling at 8.5% moisture for 13 min in disc pelletizer. Compared with the fired pellets containing 0.8% bentonite, the compressive strength of fired pellets with 2.0% hydrated lime is elevated by 13.8% and climbs up to 3113 N/pellet under the following conditions of preheating at 1050°C for 10 min and firing at

343

1300°C for 15 min. (3) The hydrated lime also improves the metallurgical performance of fired pellets, the reduction index being enhanced from 68.6% to 70.6%, RDI+3 )5 increased from 90.1% to 98.8% and reduction swelling index dropped from 22.4% to 7.4%, respectively.

References 1. M. G. Ranade and P. C. Chaubal, An Intensive Course—Blast Furnace Ironmaking, (Hamilton, McMaster University, 2004), 9-10. 2. J. J. Friel, and E. S. Erickson: Metall. Trans. B, 1 IB (1980), 233. 3. Andrew Robert, John Frederick, and Jeff Donald, "Phase Equilibria and Slag Formation in the Magnetite Core of Fluxed Iron Ore Pellets," ISIJ International, 11(48) (2008), 1485-1492. 4. Zhang H.Q, "Production and development of fluxed pellet," China Mining, 4(18) (2009), 89-92. 5. Wang X.F, Liu W.Y and Chen X.K, "Laboratory study of fluxed pellet production for Shougang," Sintering andpelletizing, 29(2) (2004), 5-8. 6. Zhang M.S, Zhou W.S, and Zhai L.W, "Development of Complex Fluxed Pellet," China Metallurgy, 16(3) (2006), 22-25. 7. Bleifuss, Rodney L, Goetzman, Harold E, "Replacement of limestone and dolomite with lime/dolomite hydrate for the production of fluxed pellet," Proceedings, Annual Meeting - Minnesota Section, AIME (1991), 195-206. 8. De Souza and Roberto Pimentel, "Production of pellets in CVRD using hydrated lime as binder is growing up fast," Ironmaking Proc Metall Soc AIME, 35th Annu Ironmaking Conf, (1976), 182-196. 9. Panigrahy, S.C, Jena, B.C, andRigaud, M, "Characterization of bonding and crystalline phases in fluxed pellets using peat moss and bentonite as binders," Metallurgical transactions. B, Process metallurgy, 21(3) (1990), 463-474. 10. Abouzeid, A.Z.M, Negm, A.A. and Kotb, I.M. "Iron ore fluxed pellets and their physical properties," Powder Technology, 42(3) (1985), 225-230.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

RESEARCH ON THE BALL MILLING AND FOLLOWED BY MICROWAVE REDUCTION OF PANZHIHUA LOW GRADE ILMENITE CONCENTRATE Ying Lei 1 ' 2 , Yu Li 1 ' 2 , Jinhui Peng1'2, Libo Zhang1-2, Shenghui Guo 1 ' 2 , Wei Li 1 ' 2 'Faculty of Metallurgical and Energy Engineering, Kunming University of Science and Technology, Kunming, 650093, China 2 Key Laboratory of Unconventional metallurgy, Ministry of Education, Kunming University of Science and Technology, Kunming, 650093, China 'Corresponding author. Tel: +86 871 5192076; E-mail address: [email protected] Keywords: Panzhihua low grade ilmenite; Ball milling; Microwave reduction; Optimization Abstract In this work, the Panzhihua low grade ilmenite and graphite were milled for 1, 2,4 and 8 h firstly, then the temperature rising behavior of milled sample in microwave field were studied. The average heating rate of samples milled for 1, 2,4, 8 h were 1.86, 4.08, 2.82, and 2.32 K/s; Before optimize the microwave reduction process, several experiments were conducted in order to confirm the parameters range. After milled for 4 or 8 h and reduced above 930 °C, the Fe metallization are higher than 90%; Then the optimization was investigated by using response surface methodology (RSM), the significance of predict model and each terms were analyzed, and the optimization parameters were given as: milling time is 4 h; reduction temperature is 1001 or 1070°C; the holding time is 37.5 or 26.9 min. Under these conditions, the Fe metallization given by predict model is 91 or 92 %. Introduction There is about 9.66 billion tons of Vanadic Titan-magnetite deposits in Panzhihua, China. In this reserve, there is about 8.73 x 10 8 t of TÍO2, accounting for 90.54% of the Chinese reserves, and 35.17% of the international reserves and ranking first in the world. [1] Titanium mineral deposits are classified into Rock type and Placer type. Rock type is the igneous mineral, which have high content of FeO (Relative to Fe2Ü3) and gangue. The Panzhihua ilmenite is a typical rock type ilmenite ore. The higher contents of MgO, CaO, SÍO2 and the complex mineralogy are making its direct reduction progressively more difficult. Mechanical activation, usually carried out by energetic milling or ball milling, was reported to significantly decreased the reduction temperature and accelerate the reaction rate of reduction of ilmenite with carbon, silicon and magnesium.[2"51 It was reported that the process temperatures were decreased to 600~1000°C and the annealing time was reduced to 30~120min after milling for 50-400 h. Microwaves are electromagnetic waves of frequencies between 3 x 10 8 Hz and 3xl0"Hz.

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Microwave processing of materials is a relatively new technology that provides alternative approaches for enhancing material properties. It provides several advantages through savings in energy, space and time; a reduction in the environmental impact of material processing and an opportunity to produce new materials and microstructures that cannot be achieved by other methods. In this work, the ilmenite and graphite were milled for 1, 2, 4 and 8 h firstly, then the temperature rising behavior of milled samples in microwave field were studied. Before optimizing the microwave reduction process, several experiments were conducted in order to confirm the rang of parameters, and then the optimization was investigated by using response surface methodology (RSM) |61 . Experimental Material and Milling Procedure In this experiment, ilmenite was supplied by a concentration plant in Panzhihua, Sichuan province, China. The chemical components are presented in Table 1. The TÍO2 grade is 38.36%. Major non-metal components are CaO, MgO and SÍO2. The total of CaO and MgO is 10.35%. The iron protoxide FeO content much more than the ferric oxide Fe 2 0 3 . As describe above, the raw material is a typical Rock-type low grade ilmenite. Figure 1 is the XRD pattern of ilmenite. It shows that the major phases of ilmenite raw materials are ilmenite (FeTiCh) and augite (Ca(Mg,Fe,Al)(Si,Al))Si206. Table 1. Chemical components of Panzhihua low grade ilmenite FeO Components Fe 2 0 3 AI2O3 TFe CaO MgO Ti0 2 Si0 2 Wt (%) 30.27 10.02 1.76 38.36 27.36 5.48 4.32 6.03

w a. u

2-Ttuta

Figure 1. XRD pattern of ilmenite (raw material) The high pure flakey graphite is used as a reducer in this work. The percent of added carbon by weight is greater than 99.6%.The ilmenite and graphite were processed in a ball mill (uni-ball mill). The weight ratio of graphite to ilmenite was 1: 4. With the steel ball amount and proportion, mixture quantity and feed density fixed, the milling time was systematically investigated. With milling times of 1, 2, 4 and 8 h, the mixtures were vacuum filtrated to eliminate most of the water and dried to a constant weight.

346

Microwave Reduction Apparatus A schematic illustration of the experimental apparatus is shown in Figure 2. Two microwave generators with 1.5 kW maximum power at 2.45GHz were employed. The milled mixtures were placed in a quartz tube in the middle of the cavity. Two thermocouples were used to measure the temperature of milled samples and cavity reactor. The control module analyzed the measured and settled temperature through a numerical controller to adjust the power output. The temperature was held at the settled temperature. In each experiment the location of the sample was exactly the same inside of the cavity. N2 with the flow rate of lL/min was employed as a protective gas during the reducing and cooling procedures. The experimental process can be observed from the viewport. The emissions were disposed during the experiment.

Figure 2. Schematic diagram of the microwave reduction apparatus Analysis and Characterization The iron metallization of the reduced product is calculated by Equation (1) ^ = (MFe/TFe)xlOO%

(1)

Where MFe is the metallic iron of reduced ilmenite, %; TFe is the total iron, %. The XRD patterns of ilmenite ore, both milled mixture and reduced ilmenite, were detected with a Rigaku diffractmeter using Cu K.a radiation; scanning angles were in the range from 10 to 90°(28) at a speed of 1.2°/min. The morphology of sample was observed with a scanning electron microscope (FEI Holland, Philip). Results and Discussion Temperature Rising Behavior in the Microwave Field The temperatures rising curves of ilmenite milled for different times are shown in Figure 3. Achieving a temperature of 1000°C needs 200~550s under different conditions. The rate of temperature increase in a microwave field can describe as Equation (2) and (3) [7! divagradT = W2T = CdT

-

P + W

X dl 1

P = 2n-f-E

.f 0 .£-V

(2)

X (3)

where Tis the temperature, Cis the specific heat, -lis the thermal conductivity, P i s the

347

microwave power absorbed by materials, W is the chemical reaction heat, / is the microwave frequency, E is the electric field strength, «-„is the permittivity of vacuum, eejr is the effective loss factor. Because of the same mixture compositions under different milling times, it is considered that s0 and ceff are fixed. The chemical reaction heat W is another factor affecting the temperature. Figure 4 shows the XRD pattern of reduced ilmenite under the conditions of activated 4h, temperature rise to 1000 °C and hold for zero min. It shows that reduction had occurred during the increase in temperature. The sample milled for 2h had the fastest heating rate. The average heating rate of samples milled for 1, 2,4, and 8h were 1.86,4.08, 2.82, and 2.32 K/s respectively.

Heating time,S

Figure 4. XRD pattern of reduced ilmenite (activated for 4h, temperature raised from room t o i 000 °C) Effect of Milling Time. Reduction Temperature and Holding Time on the Fe Metallization Table 2 Results of microwave low temperature reduction Temperature Holding time Sample Milling time Fe (min) number Metallization (h) (°C) Figure 3. Temperatures rising curves of different sample

No.l No.2 No.3 No.4 No. 5 No.6 No.7 No.8 No.9 No. 10 No. 11 No. 12

8 8 8 8 8 8 4 4 4 4 2 2

800 850 850 900 950 1000 830 880 880 930 920 970

348

30 30 15 30 30 30 30 30 15 30 30 30

72% 80% 47% 90% 94% 97% 44% 77% 46% 91% 52% 61%

No. 13

1

40

1000

85%

The microwave reduction experiments were conducted in temperature range of 800 to 1000 °C. The iron metallization of reduced ilmenite is presented in Table 2. The milling time, temperature and holding time have significant effect on the reduction results. The iron metallization increases with increasing milling time, decreases with reducing holding time, and increases with increasing process temperature. The highest metallization achieved was 97% at the highest level of these factors. Reduction Parameters Optimization The effects of three independent variables, %, (milling time), %2 (microwave reduction temperature),

Xj

(holding time) were investigated using central composite design (See Table 3).

The iron metallization of reduced ilmenite was taken as the response of the designed experiments. A total of 20 experiments consisting of 8 factorial points, 6 axial points and 6 replicates at the central points were performed. Experimental data obtained from the central composite design model experiments can be described in the form of the following Equation (4) '61

(4)

Y = ß0 +flßlZ, +tlß.xf +Y,ßl,XlXj i=i

i=i

,
where ß0 is a constant; and/? f , /?, are the linear, quadratic and cross product coefficients, respectively. The analyses of variance (ANOVA) and response surfaces were performed using the Design Expert Software (Version 7.15) from Stat-Ease Inc., USA. Optimization of milling activation-microwave reduction conditions was obtained using the software's numerical and graphical optimization functions. Table 3. Independent variables and their levels used for central composite rotatable design Independent variables Sym Coded variable levels +1 -1.682 -1 0 +1.682 bol Milling time (h) 4 6 8 9.4 2.6 Xi 1000 1100 831.8 900 1168.2 Xi Reduction temperature (°C) Holding time (min)

Run 1 2 3 4 5 6

13.2 20 30 z> Table 4. Experimental design matrix and results Activation -microwave reduction variables Holding time, Milling time, Temperature, X3(min) x,(h) x2(°C ) 20 900 4 20 8 900 20 4 1100 1100 20 8 4 900 40 900 40 8

349

40

46.8

Metallization, Y

(%) 54.7 71.2 90.7 94.4 80.6 93.5

4 94.1 7 1100 40 g 8 1100 40 97.6 9 2.6 (-1.68) 1000 30 78.6 10 9.4 (+1.68) 1000 30 98.1 11 6 30 61.8 831.8 (-1.68) 12 96.7 6 1168.2 (+1.68) 30 13 6 1000 13.2 (-1.68) 70.0 14 6 96.5 1000 46.8 (+1.68) 15 6 30 93.9 1000 16 6 30 93.6 1000 17 6 30 93.3 1000 18 6 93.2 1000 30 19 6 1000 30 93.0 20 6 1000 30 93.5 Table 4 shows the results of the experiments conducted. The Fe-metallization of reduced product was found to range from 54.7% to 98.1%. The lowest metallization of reduction (run 1, respectively), was obtained when the three variables were at their -1 levels. The highest metallization was obtained when the milling time was at its +1.682 level. Runs 15-20 at the central point were used to determine the experimental error. According to the sequential model sum of squares can be obtained. The models were selected based on the highest order polynomials where the additional terms were significant and the models were not aliased. The quadratic model was selected for metallization as suggested by the software. The final empirical models in terms of coded factors after excluding the insignificant terms for metallization is shown in Equation (5) r = 93.37+ 5.08x, + 9.92x2+7.2Sx,-2.nx,x1-0A7x,x,

-i.lOx^,

-1.46*,2 -4.68x22 -3.26x 2

(5)

The quality of the model developed was evaluated based on the correlation coefficient value. Values of "Prob > F" less than 0.0500 indicate model terms are significant. In this case, all terms and interaction terms are significant except x,x, which value of "Prob > F" is 0.3686. The closer the R2 value to unity and the smaller the standard deviation, the more accurate the response could be predicted by the model. The R2 value for Equation (5) was found to be 0.99. The R2 value obtained was close to unity, indicating that there was a good agreement between the experimental and the predicted metallization from models. Table 5. Analysis of varlance(ANOVA)for response surface quadratic model Sum of Source Prob>F Mean square F value squares Model 3136.655 348.5172 171.2172 < 0.0001 352.6186 352.6186 173.2321 < 0.0001 Xx 1344.29 660.414 1344.29 < 0.0001 Xi 722.9997 355.1905 < 0.0001 X, 722.9997 ZiZi 61.605 61.605 30.26489 0.0003 X\Xi 1.805 0.886748 0.3686 1.805

350

1 216.32 106.2722 216.32 < 0.0001 1 30.65109 30.65109 15.05806 0.0031 1 315.0646 154.7828 < 0.0001 315.0646 xl 1 153.2989 75.31165 153.2989 < 0.0001 xl 10 2.035527 Residual 20.35527 — — R2= 0.99; R2adj =0.99; adequate precision =39.04 (>4) The results obtained from the analysis of variance (ANOVA) proved the validity of the model. The ANOVA for the quadratic model for metallization is listed in Table 5. The model F value of 57.95 implied that the model was significant. There is only a 0.01% chance that a "Model F-Value" this large could occur due to noise. The "Pred R-Squared" of 0.99 is in reasonable agreement with the "Adj R-Squared" of 0.99, and it was shown that the above models were adequate to predict the metallization within the reasonable range of the variables studied. The checking of model adequacy is an important part of the data analysis procedure. "Adeq Precision" measures the signal to noise ratio. For a fixed model, adequate precision measure of the signal to noise ratio and a ratio greater than 4 is desirable. In the quadratic model, the ratio of 39.04 indicates adequate signal for the model to be used to navigate the design space.

xa> x]

Figure 5. Predicted vs. experimental Fe metallization

Figure 6. Response surface and contour plot of temperature VS milling time

Figure 7. Response surface and contour plot of holding time VS milling time

Figure 8. Cube graphs for different factors affecting the reduction process in term of metallization

351

Figure 5 shows the predicted metallization versus the experimental values. Actual values were the measured response data for a particular run, and the predicted values were evaluated from the model and generated by using the approximating functions. As can be seen, the predicted value obtained was close to the experimental values, indicating that the model was successful in capturing the correlation between the variables and the metallization. The three-dimensional response surfaces which were constructed to show the effects of the three variables on Fe metallization are show in Figure 6 and Figure 7. Figure 6 shows the effect of milling time and reduction temperature on metallization (holding time was fixed at zero level) while Figure 7 shows the effect of holding time and milling time on metallization (reduction temperature was fixed at zero level). The Fe metallization is clearly affected by milling time, reduction temperature and holding time. The milling time direct impact the reduction temperature, |2! and the interaction between temperature and holding time (process time) are also easy to understand. Figure 8 shows different factors affecting the reduction process in term of metallization. The cube graphs shows that the reaction is almost complete when the three variables all approach at the high levels. Table 6. Model valid Fe metallization (%) Variables Run *,(•») z3(min) y xXO 1 2

4 4

1001 1070

37.5 26.9

91 92

During titanium production, Fe metallization needs to exceed 90% to realize well separation of Fe and Ti. On this condition minimize the milling time can reducing energy consumption. After optimization, 100 solutions were given from the software, with all desirability equal 1.0. Two solutions among of them are given in Table 6. The results of Fe metallization were in agreement with the predicted value, and it means that the process optimizations in this work are successful. Conclusion In this work, the ilmenite and graphite were milled for 1, 2, 4 and 8 h firstly, then the temperature rising behavior of milled samples in a microwave field were studied. The average heating rate of samples milled for 1, 2, 4, 8 h were 1.86, 4.08, 2.82, and 2.32 K/s respectively; before optimizing the microwave reduction process, several experiments were conducted in order to confirm the range of parameters. After milling for 4 or 8 h and reduced above 930 °C, the Fe metallization is higher than 90%; Then the optimization was investigated by using response surface methodology (RSM), the significance of predict model and each terms were analyzed, and the optimization parameters were given in Table 6. Acknowledgements The authors would like to express their gratitude for the financial support of the Key National Basic Research Program of China (973 Program) (Grant No. 2007CB613606, No.

352

2007CB613605).

1. 2. 3. 4. 5. 6. 7.

References S.K. Guta, "Kinetics of Reduction of Ilmenite with Carbon at 1000 to 1100 °C," Metallurgical Transactions, 8B(6) (1987), 713-718. Y. Chen et al, "Ball Milling Induced Low-Temperature Carbothermic Reduction of Ilmenite," Materials Letters, 28(1996), 55-58. N.J. Welham, "A Parametric Study of the Mechanically Activated Carbothermic Reduction of Ilmenite," Minerals Engineering, 9(12) (1996), 1189-1200. N.J. Welham, "Mechanochemical Reduction of FeTiÛ3 by Si," Journal of Alloys and Compounds, 274(1998), 303-307. N.J. Welham, "Mechanically Induced Reduction of Ilmenite (FeTiC>3) and Rutile (TÍO2) by Magnesium," Journal of Alloys and Compounds, 274(1998), 260-265. R.H. Myers et al, Response Surface Methodology: Process and Product in Optimization Using Designed Experiments, 1st edition (New York: John Wiley & Sons, Inc, 1995), 1-728 A. Oktay et al, "An Analysis of the Fdtd Method for Modeling the Microwave heating of Dielectric Materials Within 3d Cavity System," Journal of Microwave Power, 31(1996), 199-214

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Too Jiang, and Mark Cook TMS (The Minerals, Metals & Materials Society), 2011

STUDY OF STRENGTHEN PELLETIZATION OF NICKEL LATERITE Jian PAN, Xianlin ZHOU, Deqing ZHU, Guolin ZHENG School of Minerals Processing & Bioengineering; Central South University; Changsha, Hunan 410083, P.R.China Keywords: Nickel latente, Agglomeration, Pelletization, Blast furnace, Fenonickel Abstract In ferronickel production, nickel oxide and part of the iron oxide are reduced to metal by agglomeration and smelting in blast furnace. Due to the physical and chemical properties of the latérite ores, a large fraction of fines is generated during the pre-treatment stages and the strength of agglomeration is weakened, which results in low productivity and bad quality. In this paper, the process of pelletization of nickel latérite was developed, and some parameters to strengthen the pelletization of nickel latérite have been optimized. The results show that the compressive strength of fired nickel latérite pellet can reach over 2000 Newton per pellet while firing at the temperature of 1220°C~1250 °C in 12 min. The product pellet can be used as high-quality burden for blast furnace to manufacture ferronickel. Introduction In recent years, the global stainless steel market was developed exuberantly, which greatly promoted the needs of the basic raw material of stainless steel smelting - metal nickel. Currently nickel sulfide ores are only 40% of global nickel reserves, but offer 60% of all the nickel production. However, with the deficit of resources, the nickel sulfide ore (high grade ore) is depleting, so there is an increasing focus on the utilization of low-grade nickel latérite. Recent research has focused on the wet treatment process to utilize nickel latérite, and less on the use of other technologies [1-8]. The economical and efficient technology relies on the reduction of nickel oxide and part of the iron oxide metal by agglomeration and smelting in blast furnace in China. Some domestic steel companies have adopted this process [9, 10]. Unfortunately a large fraction of fines is generated during the pre-treatment stages and the strength of agglomeration is weakened, which results in low productivity and bad quality of fired pellets. In this study, the process of pelletization of nickel latérite was developed firstly, and some parameters to strengthen the pelletization of nickel latérite have been optimized, and the product pellet can be used to produce stainless steel and ferronickel alloy by blast furnace.

355

Experimental Raw Materials Four types of materials were used as pellet feed, include one kind of magnetite concentrate, one kind of chromite ore and two kinds of nickel latérite ores. The chemical compositions (Table 1) indicate that the magnetite concentrate has high iron grade and low silica and alumina content, which is good for improving the iron grade of fired pellets. In contrast, chromite ore and two types of nickel latérite ores have lower iron grade and higher silica and alumina contents. Also the nickel latérite ores have low nickel grade which are lower than 2wt%, however, they possess high LOI, which affects the preheating and firing performance during pelletization. All the materials have low impurities like S and P. Table 1. Chemical compositions of raw materials (wt%). TFe CaO Samples FeO MgO Si0 2 Magnetite 65.70 27.40 4.36 0.31 0.44 0.62 Chromite 18.32 7.59 9.83 13.15 0.17 8.04 35.09 9.48 15.73 Nil# 21.41 0.27 Ni2# 2.42 32.85 13.47 Table 1. continued P Na 2 0 Ni K20 S Cr 2 0 3 0.097 0.022 0.26 0.011 0.085 42.09 0.35 0.091 0.013 0.072 0.150 0.013 0.072 2.28 0.25 1.420 0.130 1.59 0.24 1.850 0.010 0.050 0.050

A1203 1.68 11.88 8.45 4.74 LOI 0.58 2.37 13.45 12.87

Experimental Procedure Firstly, nickel latérite ores were dried to 12% of moisture and crushed to -3mm. Then all the materials and water were mixed at a given ratio, and the mixture was pretreated by high-pressure roller grinding (HPRG). Thirdly, green balls were made in disc pelletizer, and followed by drying and firing to manufacture product pellet. Various characteristics of green balls and fired pellets were tested, including size, strength, thermal stability, mineralogy and chemistry analysis of product pellets. Results and Discussion Effect of Nickel Latérite Dosage The effects of nickel latérite dosage on the performance of green balls, preheated and fired pellets are shown in Table 2 and Figure 1. When the dosage of nickel latérite was increased from 0% to 80%, the wet knock of green balls were improved significantly, and the wet crush strength was changed little, but the thermal shock temperature was decreased dramatically.

356

Nickel latérite contains clay minerals which play a role of binder. The bonding between mineral grains becomes more closely to improve the wet knock of green balls. Also the water diffusion resistance will be increased during green balls drying process, and that will result in the thermal shock temperature decreased dramatically. Table 2. Effects of nickel latérite dosage on green balls properties. Wet Crush Wet Knock Thermal Shock Dosage/% /timesO.Sm"1 /N-pellet"' Temperature/ °C 0 14.05 6.9 450 20 15.72 13.8 299 40 9.9 9.73 160 60 9.4 12.12 150 8.94 80 17.8 152 *2.4% of bentonite was added into the mixture when the dosage of nickel latérite was 0%. All the mixtures in the table above did not pretreated by HPRG Balling condition: balling time is 12 min and moisture of green ball is about 16%. 2800 2400 2000 § 1600 . | 1200

I*» 8

400

0

20

40

60

80

nickel latente dosage/% Figure 1. Effects of nickel latente dosage on the compressive strength of preheated pellets and fired pellets (Preheating condition: preheating temperature is 1000 °C and time is 12 min. Firing condition: firing temperature is 1250 °C and time is 12 min.). Fig. 1 shows that with the nickel latérite ore dosage increasing, the compressive strength of preheated and fired pellets decreased. The compressive strength of fired pellets with 80% nickel latérite was 486 N/pellet, too fragile to be used as the blast furnace burden. Mineralogical analysis (Fig.2) revealed that Fe203 develops slowly when nickel latérite is blended into pellet feed. Particles of Fe203 were small, without interlinkage, which indicates there is no enough combinational reaction at ~1250°C. This could be due to a

357

shortage of iron mineral and high melting point of nickel mineral (1420 °C), so that the crystals of Fe 2 0 3 and nickel oxide are smaller. Also, the compressive strength of fired pellets was weakened. There is a lack of liquid phase bonding in fired pellet, and the pores and cracks of pellets were increased with the nickel laterite dosage increasing (Fig.3), which will weaken the whole structure strength of pellets.

Figure 2. Mineral composition of fired pellets with 60% nickel laterite (><200 reflecting). Magnesium ferrite-color, granule; Fe3C>4-brown gray, granule; Fe203-simple white, interlinkage; nickel particles-light blue gray, granule; Pores-black, anomaly.

Figure 3. Pores and cracks in fired pellets with 60% nickel laterite under SEM.

358

Effect of Chromite Ore Dosage Chromite ore is an indispensable material of stainless steel smelting. In order to improve the compressive strength of fired pellets, from 10 wt% to 20 wt% of chromite ore was added into the mixture. In addition, because some of nickel laterite was aggregated together, the mixture was pretreated by HPRG in order to reduce the ratio of large particle size. At the fixed ratio of 60 wt% nickel laterite, the effects of chromite ore dosage on the performance of green balls, preheated and fired pellets are shown in Table 3 and Table 4. HPRG can change the size distribution of mixtures and increase the specific surface area. As the result, the ratio of micro-fine particles were increased, and at the same time, the specific surface area and surface morphological performance of materials were also improved [11]. Therefore, the wet knock of green balls increased significantly after the mixtures were pretreated by HPRG (seen in Table 3). However, the wet crush strength decreased because of plasticity improved. After adding chromite ore in the pellet feed, the compressive strength of preheated and fired pellets improved. With 15 wt% of chromite ore, the compressive strength of preheated pellets reached 615 N/pellet under the same preheating condition as described in a previous section. Compressive strength of fired pellets reached 1082 N/pellet under the best firing condition (firing temperature of 1250 °C, and fired of 12 min.) although it is still too low. Because of the large particles of chromite and better interlinkage, adding chromite ore into nickel laterite can improve the strength of fired pellets. The compressive strength still needs to be improved by decreasing the ratio of pores and cracks inside the fired pellets. Table 3. Effects of chromite ore dosage on green balls properties. Wet Knock Wet Crush Thermal Shock Dosage/% /times0.5m"' /N-pellet"1 Temperature/°C 10 49.7 5.99 <200 9.62 15 49.6 <200 20 7.25 58.3 <200 ♦The mixtures of all the blends above were pretreated by HPRG Balling condition as for Table 2. Table 4. Effects of chromite ore dosage on the compressive strength of preheated pellets and fired pellets (N/pellet). 10 15 20 Dosage/% Preheating pellets 558 615 531 1082 Fired pellets 1017 885 ♦Preheating and firing conditions were the same as specified for Figure 1.

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Effect of Additive Dosage Table 5 presents the effects of additive dosage on green balls properties. Just like chromite ore, additive did not improve the quality of green balls. The similar rule was occurred in preheated pellets. However, as the additive ratio increased from 5.7 wt% to 11.7wt%, the compressive strength of fired pellets was increased significantly, reached 2326 N/pellet with 7.7 wt% of additive under the firing temperature at 1220 °C for 12 min. The results are shown in Fig.4. The quality of pellets can meet the demands for strength for blast furnace to produce nickel-ferrous metal. Table 5. Effects of additive dosage on green balls properties. Wet Knock Wet Crush Thermal Shock Dosage/% /times-0.5m"' /N-pellef1 Temperature/ °C 13.3 5.7 6.41 220 7.7 >20 12.70 226 9.7 >17 16.34 230 11.7 >20 9.40 228 *The mixtures of all the blends above were pretreated by HPRG Balling conditions were as for Table 2. 2500 2250

t

2000 1750 g 1500 « 1250 1000 750 500 250 0.

7

8

9

10

11

12

additive dosage/% Figure 4. Effects of additive dosage on the compressive strength of preheated pellets and fired pellets (Preheating condition: preheating temperature is 1000 °C, time is 12 min. Firing condition: firing temperature is 1220 °C, time is 12 min.). The strength of fired pellets increased because calcium ferrous fayalite formed, which have large crystals and good consolidation with «crystallization (Fig.5). The melting point of silicate minerals in pellets by reacting with CaO, so the viscosity of silicate liquid decreased

360

in pellets was Fe 2 0 3 during was decreased enhancing the

cementation among the iron minerals. During high temperature, Ca2+ makes solid phase reaction easier and reduces the deformation temperature of Fe2C>3 crystal, thus promoting the interconnection of the crystals of Fe2Û3 in low roasting temperature. In addition, FeO reaction with SÍO2 are easy to form fayalite which can promote the crystallization of Fe2Û3. Therefore, it can increase the compressive strength of fired pellets. However, with additive dosage increased, the liquid phase inside the pellets increased, and a downward trend of compressive strength of fired pellets was presented.

Figure 5. Mineral composition of pellets with 7.7% of additive (x200 reflecting). Calcium ferrous fayalite-dark gray, granule; Fe304-brown gray, granule; Fe203-simple white, interlinkage; nickel particles-light blue gray, granule; Pores-black, anomaly. Table 6 presents the chemical composition of fired pellets with 7.7wt% of additive dosage which were fired at 1220 °C for 12 min. It can be seen that the product has low impurities of P and S. It can be used as the blast furnace burden to produce hot ferronickel metal.

TFe FeO 40.69 4.88

Table 6. Chemical compositions of fired pellets (wt%). P Si0 2 CaO MgO AI2O3 Ni Cr 2 0 3 K 2 0 Na 2 0 13.65 5.42 7.32 6.33 0.75 7.67 0.049 0.048 0.011

S 0.018

Conclusions (1) According to the balling test, nickel latérite plays a role of binder during balling, which improves the wet knock and crush of green balls significantly, but it decreases the thermal shock temperature. (2) With the increase of nickel latérite dosage, crystals of Fe2Ü3 develop slowly and without interlinkage, the pores and cracks in fired pellets increase, which weakens the

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whole structure strength of pellets. (3) HPRG and adding chromite ore can improve the wet knock of pellets to a certain extent. However, the compressive strength of fired pellets is still low because of the presence of pores and cracks inside pellets. Adding additive can improve the compressive strength of fired pellets because of the formation of plentiful calcium ferrous fayalite and other liquid phase, and good consolidation with Fe2Û3 recrystallization. (4) The nickel latérite pellet can be produced by common pelletization process and can be used as the charge of blast furnace to produce ferronickel. References 1. Pickles C A, "Microwave heating behaviour of nickeliferous limonitic latérite ores," Minerals Engineering, 17(6) (2004), 775-784. 2. Lee H Y, Kim S G and Oh J K, "Electrochemical leaching of nickel from low-grade latérites," Hydrometallurgy, 77(3/4) (2005), 263-268. 3. Mcdonald R G and Whittington B I, "Atmospheric Acid Leaching of Nickel Latérites Review (Part I). Sulphuric Acid Technologies," Hydrometallurgy, 91(1/4) (2008), 35-55. 4. Pickles C A, "Drying Kinetics of Nickeliferous Limonitic Latérite Ores," Minerals Engineering, 16(12) (2003), 1327-1338. 5. Liu Y., Cong Z.-F., and Wang D., "Primary Probe into Normal Atmospheric Leaching of Low-nickel Latérites," Non-ferrous Mining and Metallurgy, 20(5) (2007), 28-30. 6. Yin F. et al, "Experimental Study on Roasted Ore of Poor Nickeliferous Latérite Ore with Ammonia Leaching Technology," Mining & Metallurgy, 16(3) (2007), 4, 29-32. 7. Fu F.-M. et al, "Study about Hydrochloric Acid Leaching of Nickel Latérites," Hunan Nonferrous Metals, 24(6) (2008), 9-12. 8. Zhai Y.-C. et al, "A Green Process for Recovering Nickel from Nickeliferous Latérite Ores," Transactions of Nonferrous Metals Society of China, 2010, no.20: 65-70. 9. Wang Y.-D. and Xu F.-H., "Development of New Processes for Smelting of Stainless Steel With Hot Metal as Major Charging," Special Steel, 27(3) (2006), 35-38. 10. Fu Z.-Z., Shi G-M. and Zhu K.-L., "Stainless Steelmaking with BF Hot Metal," Shanghai Metals, 28(5) (2006), 14-20. 11. Zhu D.-Q. et al, "Improvement in pelletization of Brazilian spéculante by high-pressure roller grinding," Journal of University of Science and Technology Beijing, 31(1) (2009), 30-35.

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslav Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

WASTE TO WEALTH: PRODUCTION OF Fe-Ni FROM LATERITIC ORE/CHROMITE OVERBURDEN OF SUKINDA DEPOSIT IN ORISSA, INDIA Bhagyadhar Bhoi1, Chitta Ranjan Mishra2, and Hara Prasanna Mishra3 ''Scientist, Institute of Minerals and Materials Technology, Bhubaneswar- 751013,Orissa, India 2 Former Deputy General Maneger (R&D), National Aluminium Company Ltd, Bhubaneswar- 751 013, Orissa, India 'Former Chairman, Industrial Promotion and Investment Corporation of Orissa Ltd (IPICOL), Bhubaneswar-751 007, Orissa, India Keywords: Lateritic Orel, Chromite Over Burden(COB)2, Reduction Roasting3, Magnetic Separation^ Smelting Reduction5, Fe-Ni6, Abstract In the Sukinda Valley of Orissa, India, the total quantity of nickel ore reserves comprising of both lateritic and chromite over burden(COB) have been estimated at around 231 million tonnes containing 0.3 to 0.9% Ni. While mining of one tonne of Chromite ore, around 6-7 tonnes of chromite over burden are removed. These overburden materials are disposed as waste materials, which causes lot of environmental hazards and ecological imbalance. India imports a large amount of nickel and ferro-nickel to meet its growing demands in various sectors. For import substitutions, efforts have been made to produced ferro-nickel from the lean deposits of nickel bearing lateritic ores/ chromite over burden materials by reduction roasting followed by magnetic separation and then the nickel rich magnetic fraction is smelted to produce ferro-nickel. It has been possible to produce ferro-nickel of grade varying from 4-25% nickel. Introduction During mining of chromite ore, large quantities of overburden materials are being generated. The materials amount to approximately 10 million tones per year contain around 0.3-0.9% Ni, 0.020.06% Co, 20-50% Fe and 4-9% Cr203 depending upon the location and the depth of the mine. These materials presently do not find any application for the extraction of nickel due to lack of a suitable technology. The materials are also unsuitable for ferro-nickel making because of their very high iron to nickel ratio. The country is seriously looking for a suitable technology to utilize these resources as its entire domestic requirements of nickel is met though imports. In this connection, a systematic approach has been made at the Institute of Minerals and Materials Technology, Bhubaneswar to develop a process for ferro-nickel making from this indigenous resource. In India, the only potential nickel-lateritic and cobalt-bearing deposits that contain extractable chromite exist in four parallel bands of Sukinda region Orissa, India. It is estimated that this deposit contains around 130 million tonnes of chromite at a cut-off grade of 35% &2O3. While mining chromite ore, it becomes mandatory to remove the overburden materials from the

363

hanging and foot walls on either side of the chromite band. The chromite mining activities in the Sukinda region have been in progress for the last four decades. The overburden materials generated until now are considered waste and cause of environmental hazards and ecological imbalances. The total quantity of nickel ore reserves comprising of both lateritic and over burden has been estimated at around 231 million tonnes containing 0.3 to 0.9% Ni. Moreover, M/S Jindal Steel Limited is setting up a 2.5 million tonnes of stainless steel plant in Kalinganagar, near Sukinda Valley; a large quantity of ferro-nickel will be required for the facility. The Institute of Minerals and Materials Technology (IMMT), Bhubaneswar, has developed a state-of-the-art technology for complete utilization of nickel bearing lateritic ore / chromite overburden (COB) containing nickel by producing ferro-nickel, which can be used for manufacture of stainless steel. As a leading technology supplier in the field of minerals and materials technology in India, IMMT, Bhubaneswar is in a position to provide a novel and innovative process for commercial production of Fe-Ni, which will undoubtedly fulfill the demand of nickel consumption in India by way of import substitution. The complex nickel bearing lateritic ore/ COB can not be beneficiated strictly by physical means. The physical beneficiation process does not upgrade nickel or eliminate iron due to its complex and finely disseminated nature. To overcome this inherent problem, an innovative process has been developed by IMMT to convert the goethite into magnetite so that iron and nickel present therein can be easily separated from other gangue materials by employing techniques of reduction roasting followed by magnetic separation. This magnetic concentrate, which is rich in nickel, can be smelted to produce Fe-Ni very well in a commercial scale of the grade 4-25% Ni. This process is comparable to processes available elsewhere in the world for commercial Fe-Ni production. The process is a zero-waste process because the slag produced following the extraction of Fe-Ni can be efficiently utilized for cement manufacture. World Technology Scenario Nippon Yakin Kogyo Co. Ltd. developed a process for producing ferro-nickel by direct reduction of garnierite ore at the Oheyam works [8]. In this process, crude ferro-nickel is produced at low energy cost in a rotary kiln after pretreatment of the raw materials; the crude ferro-nickel is then used as the raw materials in the AOD stainless steel making process. A simplified mathematical model has been developed for the production of ferro-nickel from lateritic ores by the electric reduction furnace process (ERF Process) and the same has been applied satisfactorily at the plant of Larco at Larymna in Greece [9]. Nayak J.C [10] reported the results of an investigation for production of ferro-nickel from Sukinda latentes in rotary kiln-electric furnace process at Elkem's Laboratory in Norway. Nayak states that 20% nickel grade ferro-nickel was produced with high nickel recovery by selective reduction technique. The lateritic deposits in other parts of the world used in the extraction of nickel are given in Table-1.

364

Table 1: Composition of Lateritic ores used in Fe-Ni production in different countries Composit ion of Lateritic Ore

Plant and Source of Ore 1. 2. 3. 4. 5. 6. 7. 8.

Pacific Metal Co.Ltd., Japan (Philippines) Marro do Niquel S.A., Brazil S.A. Le Nickel, Doniamboo, New Caledonia Sumitomo Metal and Mining Co., Japan ( New Calendinia Ore) P.T. International Nickel, Indonesia Kovadarei, Yugoslavia Harm Nickel Smelting Co., Riddle, Oregon Larco, Greece

1.37

Ni%

Fe% 28.30

|

Fe/Ni 20.66

Grade of Fe-Ni Ni% 24-30

2.00 2.80

6.50 13.00

3.25 4.64

30.00 22-28

1.46

1.63

1.12

27

1.62

14.10

8.70

-

1.57 1.65

10.60 12

6.75 7.27

25.00 47.00

1.70

38.00

22.35

22-33

Indian Technology Scenario Nickel in India occurs in sulfide form along with copper mineralization in Singhbhum district of Bihar. It is being recovered as nickel sulphate crystals, as by-product of copper production. The other occurrence is nickeliferrous limonite in the overburden of chromite in Sukinda valley, Jajpur district Orissa [11]. This is an oxide ore and, hence, efforts are underway to develop a process for its utilization. Various organizations are trying to extract the metal values from chromite overburden materials. The process alternatives are: (i) production of Fe-Ni from imported and indigenous ore blend by Industrial Development Corporation (IDC), Orissa, (ii) extraction of pure nickel and cobalt from chromite overburden, ferro-nickel making and alloyed pig iron making from chromite overburden by Institute of Minerals and Materials, Bhubaneswar [12-16], and (iii) efforts are also being made to extract nickel from chromite overburden through biotechnology means. A typical composition of lateritic ore and chromite overburden in Sukinda region is given in Table-2. There is unfavorable Fe/Ni ratio in these materials, which require preferential treatment to remove large part of iron. The physical beneficiation process does not provide any significant nickel concentration or iron elimination due to the complex and finely disseminated nature of material. It is necessary to bring the ratio to an acceptable limit by some economic method prior to smelting. Reduction roasting is an effective unit operation in the treatment of iron oxide at elevated temperature, and it is carried out in the presence of both carbon, carbonaceous material and reducing gases like CO and H2. Conversion of CO by CO2 is the main prerequisite for reduction. The reduction process is represented by the equations [1], [2], [3], etc.

365

Table 2. Typical Composition of Chromite Overburden and Latentic ore of Sukinda. Constituents

Chromite Overburden

Latentic Ore ( Kansa)

0.50-0.90 0.02-0.06 28.00-42.00 4.00-9.00 0.50-2.50 5.00-12.00 9.00-38.00 9.00-1400

1.20-1.30 0.04 - 0.06 44.00 - 54.00 1.50-3.00 0.10-1.50 1.50-3.00 9.00-11.00 10.00-13.00

%

Ni Co Fe Cr 2 0 3 MgO AI2O3 S1O2

LOI

%

8FeO.OH + Si0 2 + 2C

Fe 2 Si0 4 + 2 Fe 3 0 4 + 4H 2 0 + 2CO

[1]

4Fe 2 0 3 + C + Si0 2

—

Fe 2 Si0 4 + 2 F e 3 0 4 + C 0 2

[2]

4FeO.OH + C

—

4FeO+ C 0 2 + C 0 2

[3]

2Fe 2 0 3 + C

—

4FeO + C 0 2

[4]

2FeO.Cr203+C

—

2Fe + 2 C r 2 0 3 + C 0 2

[5]

2Fe0.Cr 2 0 3 +S

—

2Fe + 2 C r 2 0 3 + S 0 2

[6]

4FeO.OH + Si0 2 +2C

—

Fe 2 Si0 4 + 2 Fe + 2 H 2 0 +2 C 0 2

[7]

2Fe 2 0 3 + Si0 2 +2 C

—

Fe 2 SiQ 4 + 2 Fe + 2 C 0 2

[8]

During preheating of the hydra :ed oxide mineral ore, goethite undergoes dehydration to form hematite. This reaction takes place in the range of 300-400°C. The Innovative Process Know- how Developed at IMMT, Bhubaneswar: A state-of-the-art technology has been developed by Institute of Minerals and Materials Technology (IMMT), Bhubaneswar for complete utilization of nickel bearing lateritic ore/ Chromite Over Burden (COB) containing nickel for commercial production of ferro-nickel, which can be used for manufacture of stainless steel. The nickel bearing lateritic ore/ COB is in a complex state that can not be beneficiated only by physical means. The physical beneficiation process does not provide any up-gradation of nickel or elimination of iron present therein due to its complex and finely disseminated nature. To overcome this inherent problem, an eco-friendly and zero waste innovative process know-how has been developed by IMMT to convert goethite phase present in ore bodies into magnetite phase so that iron and nickel present therein can be easily separated from other gangue materials viz. silica, alumina etc by employing the reduction roasting technique followed by magnetic separation. The magnetic concentrate, which is rich in nickel, can be smelted to produce Fe-Ni

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in a commercial scale of the grade 4-25% Ni. This process is very well comparable with any of the processes available elsewhere in the world for commercial Fe-Ni production. The main purpose of carrying out reduction roasting is selective conversion of the iron phase into magnetite, so that it can be removed magnetically. In mineral processing operation, the reduction is controlled to avoid formation of wustite, which is relatively nonmagnetic in nature. Since the nickel in Sukinda region is associated with iron phases, the adoption of reduction roasting technique is appropriate to separate nickel rich iron concentrate from the associate gangue. The nickel concentrate thus obtained by reduction roasting followed by magnetic separation route may be advantageous to enrich nickel content in reduced material. This reduced material then will be suitable for ferronickel making. A selective pre-reduction of the iron phase can be carried out in a rotary kiln with controlled parameters or a pan sintering unit designed and fabricated by IMMT, Bhubaneswar. In this process, efforts have been made to produce Ferro nickel of the grade 4-25% nickel from the lateritic ore or chromite over burden by a pyrometallurgical process which includes sintering followed by magnetic separation and smelting of the magnetic fraction rich in nickel using induction furnace facility at IMMT, Bhubaneswar. The slag produced by this method can be used to manufacture cement, thereby making it an environmental friendly zero waste process. Experimental Chromite overburden fines (-150u) containing 0.89% NiO and 45% Fe2C>3 were mixed with -150u coke breeze powder containing 65% fixed carbon in the required proportions. The mixture was granulated with 12% water in a disc pelletiser. The granules were then mixed with -2mm coke breeze and charged into the Down Draft Pan Sintering unit for reduction roasting (Figure 1).

Figure 1. Schematic diagram of down draft sintering unit

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A 10 X 10 X 310 mm laboratory-scale pan sintering was used in the experiments. About 50g of charcoal pieces were put on the top of the pan sintering unit for ignition. The bed height was maintained at 300-310 mm. Slow down-draft suction was initially applied for a minute to promote uniform ignition before the draft was increased to the required level. The sintered product was ground to -75u in a (wet) ball mill and then subjected to low intensity wet magnetic separation. The wet magnetic separation was carried out on a low-intensity (2000gauss) magnetic drum separator. The slurry was pumped into the drum at a rate of 300400cc/minute. The drum speed was maintained at 45rpm. Magnetic and nonmagnetic fractions were collected separately, and then dried, weighed and analyzed for nickel and iron. The results are summarized in Table 3. Table 3. Results of low intensity magnetic separation of Kansa Ore. Feed ground to 85% passing 75|xm, intensity of magnetic field- 2000 gauss. Details

Wt%

Roasted 100 Feed Mag. 73.30 Fraction Non.Mag. 26.70 Fraction 1st cleaning of magnetic Mag 67.10 Non Mag 6.20 2nd cleaning of magnetic Mag 64.00 Non-Mag 3.10

Fe%

Ni%

Fe/Ni 35.25

Ni Recovery 100

Fe Recovery 100

54.64

1.55

58.89

1.94

30.40

91.74

79.00

43.00

0.47

91.50

8.26

21.00

88.30 3.44

73.10 5.90

86.30 2.00

70.00 3.10

fraction 2.04 59.33 29.10 51.94 0.89 58.40 fraction after 1st cleaning 59.76 2.09 28.60 54.73 1.05 52.10

The nickel-rich magnetic fraction was smelted in an induction furnace (alumina crucible) to produce ferro- nickel (Figure 2). The chemical analysis of metal and slag are given in Table 4.

Figure 2. 35 kVA Induction furnace having three modules 5 kg, 10kg and 20 kg capacity

368

Table 4. Smelting Analysis of Magneticfractionof sinter product. Feed: Crucible material:

Magneticfractionof sinter product, containing 2.11%Niand60%Fe Alumina

Temperature:

1550°C

Exp.No

Charcoal

1 2 3 4

4.2 7.5 10.0 10.0

(%)

Wt% 3.70 30.00 36.00 42.00

Alloy %Ni 25.00 6.42 5.58 4.28

%Fe 76.80 91.60 93.30 86.57

Wt% 96.30 70.00 65.00 58.00

Slag %Ni 0.075 0.250 0.490 0.150

%Fe 22.480 40.210 16.20 6.140

The data in Table 4 indicate that ferro-nickel containing 4-25% nickel can be obtained using 410% of reductant required for reduction of iron oxide. Increase the amount of reductant increases the yield and decreases the grade of ferro-nickel. This trend is quite similar to data reported elsewhere [10,17]. Results and Discussions The chromite overburden/ lateritic ore contains iron in the form of goethite and nickel is present in the lattice of goethite phase of iron. The main objective of the reduction process is to convert this goethite phase of iron into magnetite phase and separate it from gangue materials such as alumina, silica, etc. If the entire goethite phase of iron is converted into magnetite phase, then the Fe/ Ni ratio will increase, thereby diluting the nickel grade during smelting. To overcome this problem, pan sintering technique was employed to selectively reduce iron oxide and nickel oxide such that the magnetic fraction becomes rich in nickel content. The fundamental study reveals that nickel oxide reduction occurs more rapidly than iron oxide reduction. During pan sintering process, iron oxide reduces into magnetite, wustite and metallic iron forms. When this reduced product is subjected to magnetic separation, magnetite and metallic iron along with nickel are separated. A nickel-rich magnetic fraction is obtained and it becomes the feed material for the smelting reduction process. The results of magnetic separation studies are shown at Table 3. The underlying philosophy of the study reveals that the subsequent stages of magnetic separation of the magnetic fractions shows higher content of nickel. The smelting reduction studies were carried out in an induction furnace using coke breeze fines as the reductant for production of ferro-nickel. The amount of reductant plays a vital role in smelting reduction process for selective reduction of iron for obtaining higher grade of ferro nickel. Ferro-nickel containing various percentages of nickel therein can be produced by selectively reducing iron oxide by controlling the use of different percentage of reductant as shown in Table 4.

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Conclusions A state-of-the-art reduction roasting process followed by magnetic separation and smelting techniques has been shown to produce ferro-nickel containing 4-25% nickel from chromite overburden / lateritic ore containing nickel. The technology is environmentally friendly and zerowaste process accounting for complete conversion of waste material chromite overburden. IMMT, Bhubaneswar, operating under the aegis of Council of Scientific and Industrial Research (CS1R), Government of India, is in a position to provide this technology to the industrial houses/ entrepreneurs for production of ferro-nickel in the pilot/ commercial scale any where in the world. A pre-feasibility report with plant capacity of 3000 tones per annum for production of ferro-nickel is available with IMMT Bhubaneswar for commercial exploitation of the process know- how. Patent application has been filed in India based on the above research findings. References 1. Watanable, T etal. Intl, J .Mineral processing May (1987), 173 2. Jena P.K. International Conference on recent developments in ferro-alloy production 2123 rd Jan. 1993, PP 1-23. 3. Nayak J.C.Trans .HM, Vol 38, No-3, June 1985, PP241-247 4. Bhoi B .etal "Monograph on mineral based industries in Orissa: a developmental effort" Vol-1: Nickel, RRL, Bhubaneswar, 2003. 5. Dasgupta R. etal Report on sintering of chromite overburden concentrate, RRL, Bhubaneswar T/PM/143/Feb./1996. 6. Bhoi B etal Report on " Pilot Scale Experimental studies on production of alloyed pig iron from beneficiated chromite overburden (COB) in Submerged Arc Furnace, Internal report of RRL,Bhubaneswar T/PM/375/Nov./2001. 7. Bhoi ,B. etal Report on Studies on Ferronickel making from Sukinda chromite overburden, RRL,Bhubaneswar. T/PM/334/June/2000. 8. Bhoi B etal Report on "An exploratory R & D study on Ferro nickel making using Nickel bearing Lateritic ores and chromite overburden, RRL, Bhubaneswar, 1996. 9. Lagendijk H and Jones, Production of Ferro nickel from nickel Latérites in DC-Arc Furnace. Nickel -Cobalt 97,36th Annual Conference of Metallurgists, Sudbury, Canada August 1997. 10. Bhoi.B, Nayak. B.D, Das .B, Tripathy.A.K, Murthy. B.V.R, Reddy B.R. and Dey D.N, Concentration of nickel bearing latente ore for ferronickel making. Proceeding of International Symposium on Beneficiation and Agglomeration and Environmental, Bhubaneswar, January 20-22, 1999, pp. 222-23

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

MINERALIZATION BEHAVIOR OF FLUXES DURING IRON ORE SINTERING Min Gan, Xiaohui Fan, Tao Jiang, Yi Wang, Lin Hu, Wenqi Li, Qiang Wang, Luben Xie School of Minerals Processing & Bioengineering, Central South University, Changsha, Hunan , 410083,P.R.China Key words: Sintering, Flux, Mineralization, Iron ore Abstract The mineralization behaviors of typical fluxes are studied by investigating the microstructures with optical microscope. The results show that the reactions begin with between iron ore and lime, limestone with small particle to form primary liquid. Then coarse limestone, dolomite and serpentine take part in reaction under the action of primary liquid. Limestone is the easiest to mineralization, next dolomite, and the last is serpentine. The higher basicity and smaller fluxes size favor the minralization process of dolomite and serpentine. Introduction The purposes of adding fluxes into sintering are that:(l) to increase the basicity of sinter, so as to reduce even eliminate addition fluxes into BF(blast furnace) and decrease the coke ratio of BF[1,2]. (2) to decrease the melting point of sintering mixture to ensure sintering can be carried out at a lower temperature and reduce the energy consumption. (3) to improve the qualities of sinter by forming binding phase with iron ore fines [3]. Flux is another significant component for sinter besides iron ore, especially for high basicity sinter[4]. The proportion of flux is usually higher than 10% in sinter mixture(excluding return mine). There were many reports about the influence of the flux types on sintering, such as limestone, dolomite, lime and so on[5-7]. It is very important to choose suitable fluxes for iron ore sintering. Although the effect of flux size on sintering also has been researched, the mineralization behaviors of the fluxes in the sintering process has not been illuminated. The paper will study the mineralization behavior of fluxes in sintering.

371

Materials and Methods Materials Properties The chemical compositions of materials are shown in Table 1. The iron ore is a hematite, which has high content of Fe, with few impurities except for 3.78% Si0 2 and 0.77% A1203. Four fluxes lime, limestone, dolomite and serpentine were used in the experiments. Lime is a pure chemical agent, and the content of CaO was 99.5%. The other three fluxes were from industrial production. Limestone and dolomite belong to carbonate. Limestone contains 52.31% CaO, and dolomite contains 33.08% CaO and 17.35% MgO. The theoretic molecular formula of serpentine is 3MgO-2Si0 2 -2H 2 0. Serpentine used contains 36.40% MgO and 37.12% Si0 2 . Table 1 Chemical compositions of materials Material types Iron ore

TFe 64.94

FeO 0.86

CaO 0.03

wt/%

MgO 0.01

Si02 3.78

A1,0,

LOI

0.77

1.67

0.52

42.95

99.50

lime Limestone

0.20

52.31

1.45

1.66

Dolomite

0.27

33.08

17.35

1.88

0.66

45.27

Serpentine

3.55

2.22

36.40

37.12

0.10

13.45

Experimental Methods The first step was agglomeration. Iron ore fines with the size of under 0.5mm and lime were blended with proper moisture. Then, 30g mixture was weighed and put into columniform mould with diameter of 30mm. The mixture in mould was pressed to a flat surface by a small pressure. Afterward, 2g flux with certain size(0.5-lmm,l-2mm or 2-3mm) was placed in mould evenly, and another 30g mixture was added into mould. Finally, the column material was again pressed to make the agglomerate under the pressure of 300Kg/cm2. The sketch map of fluxes distribution in agglomerate is shown in (a) of Fig. 1.

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Fig.l Sketch map of fluxes distribution in agglomerate And then agglomerate was sintered in a furnace whose temperature can be controlled by computer. The temperature was set at 1300 °C, and the time was 10 minutes. After cooling, agglomerate was fixed with resin to prepare polished section. Agglomerate was cut above centre line slightly and then cut section was polished till the fluxes were exposed. The polished section is shown in (b) of Fig.l. The mircostructures were studied by optical microscope. Results and Discussion Reactive Ability between Iron Ore and Fluxes The results are shown in Fig.2. From Fig.2(a), it could be seen that lime reacts with iron ore to form calciumferrite. The mineralization ability of lime is excellent. Fig.2(b),(c) and (d) show the mineralization of limestone, dolomite and serpentine with the size of 0.5~1.0mm, respectively. Limestone could be mineralized completely, but dolomite and serpentine remain unmineralized cores after sintering. It is difficult to find reation product surrounding cores, which indicates that dolomite and serpentine can hardly react with iron ore. It can be concluded that the reactions began with between lime, limestone of small particle and iron ore in preliminary stage of sintering, which can form primary liquid for sinter.

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( a ) Lime

( b ) Limestone with 0.5-1.Omm

( c ) Dolomite with 0.5-1.0mm

( d ) Serpentine with 0.5~1.0mm

CF—Calciumferrite, Do—Dolomite, Se—Serpentine Fig.2 Microstructures of fluxes reacting with iron ore Effect of Primary Liquid on Flux Mineralization How does the primary liquid affects the mineralization of fluxes? To simulate the forming of primary liquid, 9% lime was added into iron ore firstly. After mixed, fluxes such as limestone, dolomite and serpentine with size of 2.0~3.0mm were distributed in mixture, and then the prepared agglomerates were sintered. The results are shown in Fig.3. Limestone is mineralized thoroughly and the reaction product was assimilated to homogeneous microstructure. Dolomite and serpentine can react with iron ore but not mineralized completely. There are layers of reaction product around unmineralized cores, which indicates that primary liquid accelerates the reaction of these fluxes. Compared the size of remaining cores, it can be known that limestone has the strongest reaction ability under the assistant of primary liquid, next dolomite, and the last is serpentine.

a ) Limetone with 2.0-3.0mm

( b ) Dolomite with 2.0-3.0mm ( c ) Serpentine with 2.0-3.0mm

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Do—Dolomite, Dr—Reaction product of dolomite Se—Serpentine , Sr—Reaction product of Serpentine Fig.3 Microstructures of fluxes mineralization when 9% lime mixed with iron ore Conditions of Complete Mineralization of Flux Liquid formed in preliminary stage of sintering benefits the mineralization of fluxes, but how much liquid can make the fluxes mineralize completely is another interesting issue. The experiments changed the proportion of lime to form different content of primary liquid. First, certain lime was mixed with iron ore, then fluxes with size of 2.0-3.Omm were distributed in mixture. The influence of liquid content on fluxes mineralization is shown in Fig.4~ Fig.6. From Fig.4, the results show the mineralization behavior of limestone. With the increasing lime proportion, the degree of limestone mineralization increase. When lime is increased to 5%, small cores enclosed by the reaction product in sinter. Till lime is increased to 7%, the unmineralized cores are disappeared, and the transitional layer of reaction product also vanishes due to assimilating.

( a ) Lime 0%

( b ) Lime 5%

( c ) Lime 7%

Li—Limetone, Lr—Reaction product of limestone Fig.4 Effect of lime proportion in iron ore on limestone mineralization The effect of lime proportion on dolomite mineralization is shown in Fig.5. When the lime proportion is increased from 0% to 15%, the unmineralized cores became smaller and smaller till disappear. The investigation shows that the dolomite could be mineralized completely when the lime proportion is over 15%.

375

( a ) Lime W,

( b ) Lime 12.5%

( c ) Lime 15%

Do—Dolomite, Dr—Reaction product of dolomite Fig. 5 Effect of lime proportion in iron ore on dolomite mineralization

( a ) Lime

ff/o

( b ) Lime 15%

( c ) Lime 17.5%

Se—Serpentine Sr—Reaction product of serpentine Fig.6 Effect of lime proportion in iron ore on serpentine mineralization As to serpentine, the influence of lime proportion is shown in Fig.6. Serpentine is not mineralized completely even lime proportion is up to 15%. But when lime is increased to 17.5%, the cores could not be seen in sinter, which indicates that lime should be at least 17.5% for completed mineralization of serpentine. It can be concluded that the more the primary liquid produced, more complete the fluxes mineralization. The mineralization of serpentine and dolomite requires large amount of primary liquid, especially for serpentine. This suggests that these two fluxes can be mineralized completely under the condition of high basicity. When the liquid is not enough to mineralize dolomite and serpentine, the size should be reduced to ensure the fluxes to be mineralized completely. Conclusions

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(1) In preliminary stage of sintering, iron ore reacts with lime and small particle limestone to form primary sinter liquid. Dolomite and serpentine hardly react with iron ore directly. (2) Coarse limestone, dolomite and serpentine could be mineralized under the action of primary liquid. Limestone is easy to mineralize completely , while dolomite and serpentine are difficult to mineralize. (3) The liquid formed in primary sintering stage will prompt the mineralization offluxes.It is shown serpentine needs the largest amount of primary sinter liquid, next dolomite, and the last limestone for complete mineralization. Acknowledgement The authors want to express their thanks to Program for New Century Excellent Talents in University (NCET-05-0630), Program for Excellent Doctor's Degree Paper in Central South University, and Project(1343/74333001114) supported by the Postgraduate's Paper Innovation Fund of Hunan Province, China for financial support of this research. References 1. C.L. QI, J.L. ZHANG.Y.X. CHEN.and H. GUO, "Experimental study on high temperature characteristics of sintering fluxes.'The Chinese Journal of Process Engineering, 9(l)(2009):266-269 (in Chinese). 2. X.F. GUAN, "The way of hightening MgO content of sintering and the effect of BF," Sintering & Pelletizing, 26(2)(2001):l-5 (in Chinese). 3. J.L. FU, T. JIANG and D.Q. ZHU, Singtering and pelletizing(Changsha:Central South University of technology Press,1995),99 (in Chinese). 4. X.P. FENG, "The effect of bacisity on sintering strength in the condition of low-Si,"Sintering & Pelletizing, 29(2)(2004):9-12 (in Chinese). 5. S.H. ZHANG, H.Y. FENG, "Production practice on use of the roasted flux in No. 2 Sintering Plant of WISCO", WISCO technology, 46(l)(2008):4-8 (in Chinese). 6. X. JIANG, G.S. WU, "Effect of MgO on sintering process and metallurgical properties of sinter.'Tron & Steel, 41(3)(2006):8-12 (in Chinese). 7. J.T. SUN, "Investgation on rational bacisity of sintering,"Iron & Steel, 36(8)(2001):l-4 (in Chinese).

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslav/ Drelich, Jerome Downey, Tao Jiang, and Mark Cook TMS (The Minerais, Metals & Materials Society), 2011

MICROWAVE ASSISTED BREAKAGE OF METALLIC SULFIDE BEARING ORE Matthew D. Andriese1, Jiann-Yang Hwang1, Wayne Bell1, Zhiwei Peng1, Anish Upadhyaya2, Shivanand A. Borkar3 'Michigan Technological University, Houghton, Department of Materials Science and Engineering, Michigan, U.S.A 2 Indian Institute of Technology, Department of Materials and Metallurgical Engineering, Kanpur, India industrial Microwave Research Center, Pradeep Metals Ltd., R205 MIDC, Rabale, NaviMumbai, India Keywords: sulfide ore, microwave energy, Curie temperature, Bond work index Abstract A refractory ore body located in Michigan's Upper Peninsula contains high concentrations of nickel and copper chiefly occurring in the minerals pyrrhotite, chalcopyrite, and pentlandite. Refractory ore bodies are difficult to treat by conventional mineral processing methods so microwave pre-treatment of the ore is employed to increase metallic-particle/host rock liberation by making use the differential thermal expansion properties of the mineral phases that absorb microwave energy. The Curie temperature measurement of metallic-bearing particles is in agreement with the known value for pyrrhotite occurring at 325°C. A nickel-rich iron sulfide mineral is found to be in occurrence and also appears to be magnetic under BSE imaging. It is shown the ore particles heat rapidly when exposed to microwave radiation for short durations of time mainly due to the high concentration of ferromagnetic mineral phases. Rapid heating causes thermal expansion of constituent mineral phases that produce cracks within ore particles. SEM imaging shows fracture occurring along grain boundaries and throughout host rock matrix. Ball milling experiments show an increased grindability of the ore resulting in a decrease in work index values. Introduction Crushing and grinding operations consume 50-70% amount of energy used in mineral processing operations [1]. All these processes involve high energy impact of surfaces to develop compressive stresses that produce cracks to initiate breakage of particles. Microwave pre-treatment of certain types of ore might provide an energy savings to these type grinding processes by taking advantage of the tensional forces produced by selectively heating microwave absorbing type mineral phases. The heating of ore particles produces cracking by thermal expansion of the constituent mineral phases. Conventional heating has been investigated and shown to positively aid in grinding processes but determined energy inefficient to be used commercially.

379

Minerals that absorb microwave energy and rapidly heat are termed high loss. Not all minerals present within an ore particle will absorb microwave energy but will heat as a result of thermal conduction from neighboring high loss minerals. All minerals are subjected to heating but selectively heating high loss minerals causes different rates of thermal expansion between mineral grains. Thus, it is the difference that minerals expand and contract during heating and cooling processes that cause cracking. By selectively heating high loss minerals contained within the ore, crack formation can be produced in short durations of time with high power density. If the power density of ore particles under microwave irradiation is better understood, pre-treatment of ore can potentially result in a large energy savings in the crushing and grinding process. Though microwave heating of refractory ore particles is not well understood, the breakage characteristics after exposure are shown to increase with microwave energy exposure time. It is observed though that increasing exposure time can result in localized melting of mineral phases. The amount of thermal energy generated by an electromagnetic field within a particle exposed microwave radiation is known as the power absorption density [2]. It is dependent on the internal electric field strength within the minerals, the frequency of the microwave radiation, and on the dielectric and magnetic properties of mineral phases. If the electromagnetic field strength is known, the power absorption density per unit volume of a mineral can be approximated by the equation: P=

2xf0{eae"E2+ßoß"Hl)

Where Pj is power density (W/m3), / is the frequency of the electromagnetic radiation (Hz) eo is the permittivity of free space (8.845 * 10"12 F/m), er" is the relative permittivity of the mineral phase, E0 is the magnitude of the electric field, ft0 is the permeability of free space, and fir" is the relative permeability of the mineral phase, all of which are proportional to the microwave energy (V/m) propagating thru the mineral. Future work will investigate the effect of different microwave power and frequency but current work is performed with a conventional multimode 1000W, 2.45GHz microwave applicator. Understanding operating frequency and power levels insures optimum processing controls. Figure 1 displays the interrelated processing parameters that affect the power absorption density of ore particles when exposed to microwave radiation.

Fig.l. Relation of heating time, particle size and batch size during microwave treatment of ore particles

380

The minerals contained in the ore body exhibit different magnetic and electrical properties. A mineral that absorbs microwaves due to its electronic structure heats by loss caused by its relative permittivity. Conversely, a mineral heating by absorption and loss due to its magnetic structure heats by its relative permeability. As many minerals display a wide variety of ionic character and all minerals are classified according to their magnetic properties it is not always clear as to the mechanism of microwave absorption. Also, grain size, impurities and defects all contribute to the microwave absorption properties of a mineral. Experimental Magnetic Measurements Magnetic susceptibility investigations were carried out at IIT Kanpur, India. The equipment made available is used mainly for metals and ceramics to generate hysteresis loops and Curie temperature measurements, so it is only possible to test small sample sizes (>2g). Large ore particles were crushed into smaller pieces and particles with metallic appearance were hand-picked for testing. The Curie temperature measurement was made up to 400°C using conventional heating, which was high enough temperature to cause a rapid decrease in the magnetic susceptibility of samples. Microwave Treatment The microwave used to obtain materiel for grinding tests was a conventional multi-mode 1000W oven with a 1.1 ft3 cavity and glass turntable. Core samples were jaw crushed then later gyratory and roll crushed to obtain particles of various sizes for classification. After classifying a specific size range for a bulk sample of crushed ore product, the material is sent through a particle splitter to achieve homogenous bulk samples for grinding experiments. Samples were microwave treated as follows: For the gyratory 6+8 sample, 100 gram samples were treated in an alumina crucible placed in the center of the cavity. For jaw crushed -6+10 and -8+12, 500 gram samples were distributed evenly, single layered on the rotating glass tray of the microwave applicator. All samples were heated for the indicated times, removed from the applicator and cooled on a ceramic plate. A portion of microwave treated ore was then mounted in epoxy, polished to 0.05 Urn, and examined under the optical and SEM microscope to better understand morphology of the ore after being heated with microwaves. At various times throughout testing, water in a 100 ml beaker was heated for 30 sec to test if the magnetron was still functioning properly. Grinding Experiments Laboratory work index measurements are often time consuming with significant amounts of material needed for completion when using F. Bonds grinding procedures [3]. About 3-4kg is needed for completion and only one sieve size is used for testing for oversize and undersize material. Ore particles passing the 6 mesh sieve (3327um) were run through a particle splitter a dozen times to ensure homogeneity of the particles. A portion

381

of the sample was then treated with microwave energy. The ball mill was fitted with the 285 balls following Fred Bond's procedure for measurement of work index. Treated and untreated particles are ball milled following Fred Bonds procedure for grinding and crushing calculations. For determination of the work index (kWh per ton ore), the following equation was used: 44.5

W,= p023 G i p 0.82

JO V^O

10_ \

80

Where P¡ is the test sieve size in microns, Gbp is the grindability of the ore, P is the size in microns of 80 percent of the last cycle sieve undersize product passes, and F is the size in microns of 80 percent of the new feed passes. Results and Discussion Curie Temperature The strong magnetic property of the ore particles indicates that rapid heating will occur by microwave irradiation. Strong absorption of microwave energy within ore particles can be strongly attributed to the main metallic host mineral pyrrhotite (Fei.xS). When occurring with the 4C superstructure and stoichiometric composition of Fe7Ss the structure is ferromagnetic [4]. As other strongly ferromagnetic minerals are an occurrence within the ore, ferromagnetic spinel minerals, Curie temperature measurements were made on samples with metallic appearance. The samples shown in Figure 2 have magnetic moment measurements measured in electromagnetic units, emu, (0.001 Am 2 ) plotted as a function of temperature (°C). Though the plots have very different magnetic moment measurements, the shapes of the plots are in agreement and at roughly 300°C, randomly arrangement of magnetic dipoles rapidly decreases the measured magnetic moment. If ferromagnetic pyrrhotite is thermodynamically stable past 325°C, it is likely to have the effect of thermal runaway and high amounts of localized heating will occur in ore particles. 1

ff . O . 8 §

0.6

1"

O O

100 200 300 400 Temperature (Celolus)

500

Fig. 2. Curie temperature measurements of metallic ore particles. The rapid decline of magnetization in metallic particles occurs at roughly 300° C. This is

382

within 25°C of the known pyrrhotite Curie temperature measurement of 325°C [5]. The Curie temperature measurement is within a close approximation of the measured value for a pure pyrrhotite mineral. The slight difference in Curie temperature measurements could possibly be due to the high amount of nickel-rich pyrrhotite (Ni-Po) formed as a derivative superstructure of the pyrrhotite (Po). Rather than characteristic exsolution of the mineral pentlandite (FeNiS2) during cooling [6], nickel was able to be incorporated into the vacant hexagonal structure of the Po. An example of these minerals disseminated within an ore particle can be seen in the backscattered electron image in Figure 3. The Ni-Po is magnetic, as it appears grey in the BSE image due to an incomplete signal from the BSE detector with random scatter of electrons emitted from the beam.

Fig. 3. BSE image of dissemenated sulfide minerals pyrrhotite (Po), pentlandite (Pn) and nickel-rich pyrrhotite (Ni-Po) occuring with quartz (Qtz) Microwave Induced Cracking of Sulfide Ore Particles The images in Figure 4 are before and after optical images of the same polished surface of a jaw crushed +1/2" piece of ore before and after microwave radiation. The surface was re-polished to remove the partial oxidation that had occurred after microwave exposure for 30s. The mineral grains are seen in Figure 4b but with a cracking propagating thru the gangue between them with another crack propagating out of the field of view.

Fig. 4. Optical image of a.) Metallic sulfide grains b.) Same grains as in image a but microwave exposure for 30 seconds. (Scale bar not shown)

383

Fig. 5 BSE image of cracking between pyrrhotite (Po), Ni-rich pyrrhotite and the pyroxene host matrix.

Fig. 6. BSE image of a spinel mineral grain exposed to microwave radiation. Similar type cracking is seen in Figure 5 showing cracking between sulfide grains and Px gangue. Smaller grains of Ni-Po cause intergranular cracking of the iron sulfide surrondings. These particles hold value and should be magnetically washed during processing for the small percentage of nickel value they hold. In Figure 6, a Cr-bearing spinel mineral in liberated from its sulfide mineral neighbors also forming cracks within itself. Spinels are considered a refractory ore in mineral processing and difficult treat conventionally. Here it is seen completely liberated from its surrondings. Thus, physically seperating minerals at various sieve sizes might prove sucessful for recovery of metallic bearing minerals from gangue waste rock. Crushing and Grinding Experiments Grinding and sieve testing was performed on ore particles to better understand size classification of particles before and after treatment. These experiments are performed to predict the test sieve size that shows a decreased grindability during ball milling. Particles microwave pre-treated for shorter durations of time would consume less energy but may not heat to temperatures high enough to introduce significant amounts of stress within particles to induce cracks and increase the grindability of the ore. It is of interest to microwave samples for short time durations because of less energy consumption of

384

microwave pre-treating material. It can be seen in Figure 8 that particles microwave treated for 30s and 60s time duration show an increase in the cumulative percent passing for all particle size tested. The sample MW treated for 60s has a marked effect on the cumulative percent passing for all particle sizes seen in Figure 8.

Fig. 8. Cumulative percent passing for particles ball milled 100 revolutions The results of the grinding experiments are displayed in Table 1. The grindability is increased for all microwave treated ore particles compared to untreated ore particles resulting in a reduction of the ball mill work index. The largest particles, gyro -6+10, had the largest work index calculated for any of the samples. Conversly, the smallest particle size, jaw-8+12, with the microwave exposure time of 30s had the lowest calculated work index. As the feed particle size is increased, the reduction of size largly occurs by abrasion to the mineral surface rather then breakage at mineral grain boundaries. This is seen by lower grindability, or amount of undersize produced per mill revolution. The jaw particles have close values of work index specifically because of the way they were treated. Larger samples treated consequently absorb more microwave energy reducing the average power input for each particle. Heat tranfer between particles is also greatly reduced when placed on the glass tray rather than in the crucible. Table 1. Effect of Microwave Heating on Grindability and Work Index of Sulfide Ore.

Feed Material Gyro -6+10 As-received MW30s MW60s Jaw -6+10 As-received MW60s

Grindability, grams per Work Index, kW-h per ton Test Sieve mill Size (urn) revolution ore 147 147 147

0.79 0.82 0.86

20.81 20.52 19.87

417 417

1.83 1.97

17.16 17

385

Jaw -8+12 As -received MW30s

208 208

1.57 1.58

16.10 15.78

Conclusion The measured Curie temperture, 300°C, is in agreement with the known measured value for pyrrhotite, 325°C. The ore is heated by microwave radiation largely due to the high amount of ferromagnetic pyrrhotite present within the ore. BSE imaging of the nickelrich pyrrhotite shows it as another dissemenated magnetic mineral phase within the ore. The spinel minerals occuring within the ore are strong absorbers of microwave energy. Mineral grains heated under microwave energy produce cracks along mineral grain boudaries and across the host rock matrix in ore particles. An improvement in grinding characteristics is seen with longer heating intervals resulting in a reduction of work index values. When the liberated particles are of fine sizes compared to the feed material, it can be assumed that size reduction is taking place largely by abrasion. Abrasive wear to accomplish size reduction is ineffective because a high amount of energy is wasted mainly in the form of heat created by friction from collisions of balls and particles. Thus, it may prove better to "stamp" the ore particles after microwave treatment for liberation of metallic-bearing mineral phases. Reference [1] J.W. Walkiewicz, A.E. Raddatz, S. L. McGill, Microwave-Assisted Grinding, Reno Research Center, U.S. Bureau of Mines, 1989 [2]Understanding Microwave Assisted Breakage, D.A. Jones, S.W. Kingman, D.N. Whittles, I.S. Lowndes, University of Nottingham, Minerals Engineering 18, pp.659669, 2005 [3] Fred C. Bond, Crushing and Grinding Calculations Part 1, Canadian Mining and Metallurgical Bulletin, 1954, Vol 47, No. 507, 466-472 [4]E.N. Selivanov, R.I. Gulyaeva, and A.D. Vershinin, Thermal Expansion and Phase Transformations of Natural Pyrrhotite, Inorganic Materials, 2008, Vol. 44, No. 4, pp 438-442 [5]Electrical and Magnetic Properties of Sulfides, C. Pearce, R. Pattrick, D. Vaughan, Reviews in Mineralogy and Geochemistry, Vol. 61, pp. 127-180, 2006 [6]V. Rajamani and C.T. Prewitt, Thermal Expansion of the Pentlandite Structure, American Mineralogist, 1975, Vol. 60, pp. 39-48

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2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

2nd International Symposium on

High-Temperature Metallurgical Processing

Sintering and Synthesis Session Chairs: Xiaohui Fan Xuewei Lv

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

CRYSTALLIZATION BEHAVIOR OF CALCIUM FERRITE DURING IRON ORE SINTERING Xiaohui FAN, Lin HU, Min GAN, Tao JIANG, Wenqi LI, Qiang WANG Luben XIE, Zhiyuan YU School of Minerals Processing & Bioengineering, Central South University, Changsha, Hunan, 410083, P.R. China Key words: Iron ore sintering, Calcium ferrite, Crystallization behavior, Cooling rate, Si0 2 , A1203 content Abstract Crystal precipitation behavior of calcium ferrite (CF) during iron ore sintering process along with the effect of cooling rate, Si0 2 and AI2O3 on crystalline condensation were studied. The results show that CF, has strong ability of precipitation. The precipitation is barely affected by extenal factors. It begins to precipitate when the liquid phase cools down to about 1200°C. CF could crystallize rapidly at the cooling speed of 150 °C /min. Furthermore, decreasing of the cooling rate favors the development of crystal. Binder phase transformed from CF system to silicate system when Si0 2 content was increased to 5-6%, with a constant basicity (R) 2.0. Increasing A12C>3 content could promote the precipitation of CF and the development of crystal. Introduction Crystalline condensation is the last stage of sinter mineralization. It includes the processes that crystalline substance and amorphous substance precipitate from high temperature solution phase and the consolidation of material when melt cools down. During the crystallization process, minerals precipitate following the order from high to low melting point coupled with the decreases of free energy. Most of binder phase minerals and part of iron oxides crystallize in this process, which have an important influence on mineral composition and microstructure of sinter. They finally determine the sinter strength and metallurgical properties. Calcium ferrite contains A1203 and Si0 2 (SFCA). The needle-like CF, especially, has special features such as good strength, excellent reducibility and low formation temperature. Low formation temperature is suitable for developing low-temperature sintering, improving the quality of sinter, and reducing energy consumption'1"41. Crystalline condensation is a key stage for the forming and developing of CF, which directly affects its precipitation behavior and crystalline morphology. In this paper, we report on crystalline condensation mechanism of CF binder phase system based on minerals' precipitation behavior and crystallization condition,

389

which provide theoretical support on improving microstructure of sinter and optimizing its quality. Materials and Methods Properties of Materials The chemical compositions of raw material are summarized in Table 1. This mineral, which belongs to oxidized ore, has high iron grade, low gangue content, and the ratio of Total Iron (TFe) to FeO is much higher than 3.5. The chemical, flux (calcium oxide), and the additive (A1203, S1O2) are all from analytic grade reagents. Table 1. The chemical compositions of raw material (wt/%) Raw material Iron ore

TFe 64.94

FeO 0.86

CaO 0.03

MgO 0.01

Si0 2 3.78

AI2O3

0.77

LOI 1.67

Methods of Experiments We adopted mini-sintering test to simulate sintering process. The iron ore, which had been ground to powders of-0.074mm size, was well-mixed with a certain amount of flux, additive and water, and then the mixture was pressed into briquettes of 10 mm diameter at 300kg/cm2. After dried fully, briquettes were sintered at 1280 °C in horizontal tube furnace for 10min. Briquettes were cooled down from high temperature under various conditions, boiled in resin, lapped and polished. The microstructure of the obtained samples was carried out on a optical microscope, Leica DMRXE, at a magnification of 50-1000. Results and Discussion Crystal Precipitation Behavior on Cooling Process Calcium oxide was added (mass fraction 8%) to ensure the formation of CF melt. Briquettes were cooled down to 1280°C, 1250°C, 1200°C, 1150°C, and 1050°C at a cooling rate of 50°C /min respectively, and then quenched by water. Fig. 1 shows the micrograph of products at different quenching temperature. Dominated mineral composition were hematite and CF in each products, and microstructure was corrading, however the crystallization had obvious differences (Table 2). According to the results, we preliminary deduce that the precipitation temperature of crystal in CF system is close to 1200 °C. In order to verify this inference, briquettes were cooled with 50 °C /min to 1200 °C, and holding 10 min at this temperature, then quenched by water finally. It was found experimentally that as the holding time prolonged, crystalline morphology became significant needle-like. Although dominated mineral compositions were hematite and CF,

390

precipitation quantity of CF increased.

d) 1150°C Fig. 1

e) 1050°C

f) Holding 10 min at 1200°C

White—Hematite,Gray—CF,Black—Void Micrograph of different quenching temperature(x200)

Table 2 Mineral composition and microstructure of different quenching temperature Quenching Micrograph Mineral composition and microstructure temperature 1280°C

Fig. 1(a)

1250°C

Fig. 1(b)

1200°C

Fig. 1(c)

1150°C

Fig. 1(d)

1050°C

Fig. 1(e)

No crystalline was formed. The binder phase, which had no time to crystallize under high temperature quenching, kept the original morphology. The rudiment of the crystalline state CF was obtained, and its characteristic morphology was unidirectional extension. Melt were crystallized gradually, developed more fully, in addition, needle-like CF were precipitated.

Undercooling is an essential driving force for phase transition. When the temperature is lower than the melting point, undercooling AT (¿T=Tm-T, Tm—melting temperature, T— actual temperature) is obtained. It was found by spectrum, the component of CF was close to the

391

eutectic point of Fe 2 0 3 -CaO system. Phase diagram l5] shows the temperature of melt precipitation nearby eutectic point was just higher than 1200 °C, and the crystal had undercooling condition at this temperature. It also proved the conclusion that CF, which precipitated at about 1200 °C, had good crystallization capacity. The crystal precipitation process was closed to the equilibrium state. Effect of the Cooling Rate on Condensing Crystallization Process Cooling rate determines the time of crystal precipitation, which is an important factor of crystallization. Using controllable cooling design, cooling rate was controlled at 150 °C /min, 100°C /min, 50 °C /min in high temperature stage (>1000°C), and 50 °C /min in low temperature stage (< 1000 °C). Natural cooling method was adopted when briquettes was cooled down to 600 °C. 8% of calcium oxide (R=2.3) was added into raw material as flux. Fig. 2 and Table 3 provide a set of typical microstructure corresponding to the samples after sintering and cooling at different cooling rate.

a) 150°C/min

b)100°C/min

c)50°C/min

White—Hematite.Gray—CF,Black—Void Fig.2 Micrograph of different cooling rate(x500) Table 3

Mineral composition and microstructure of different cooling rate

Cooling rate

Micrograph

Mineral composition and microstructure

150°C/min

Fig. 3(a)

Crystal rudiment of CF was found,which further confirmed that CF had strong crystalize abiliity.

100°C/min

Fig. 3(b)

The crystal precipitation became obvious with decreasing cooling rate, and crystal tended to intensive.

50°C/min

Fig. 3(c)

Along with melt crystallizing fully, CF developed to needle-like morphology, which formed interlaced and corrasive structrue with hematite.

392

Based on the experimental results, it was found that CF crystal can precipitate in rapid cooling rate. On the other hand, slow cooling rate significantly favors further development of the crystal. In the actual production process, cooling rate of the upper and lower material layers are 120~130°C /min and 40~50°C /min respectively'61. During the production of high-basicity sinter, we can deduce that the cooling process of lower material layer is relatively slow, and this can be in favor of the crystallization of needle-like CF and development of melting structure. Effect of the Component of Gangue on Condensing Crystallization Process The gangue in iron ore can probably change the binder phase system and crystalline morphology. It also has an important influence on crystallization result. Therefore, we studied the function of gangue, especially SÍO2 and AI2O3, in the condensing crystallization process. After dried fully, briquettes were sintered at 1280 °C for 10min, and the cooling rate was controlled at 100 °C /min, 50 °C /min respectively in high and low temperature stage. Natural cooling method was adopted when briquettes was cooled down to 600 °C. Effect of the SiQ2 Content on Condensing Crystallization Process SÍO2 was added into raw material in order to control its content to 4%, 5%, 6% and 7%, meanwhile, inputting relevant amount calcium oxide to keep R=2.0. The microstructure is shown in Fig. 3 and Table 4.

d)Si0 2 =6% e) Si0 2 =7% White—Hematite,Gray—CF,Darkgray—Silicate,Black—Void

393

Fig.3

Table 4 m(Si0 2 )

<4.0%

Micrograph of different SÍO2 concentration (x200)

Mineral composition and microstructure at different SÍO2 concentrations

Micrograph Fig. 4(a) Fig. 4(b)

Mineral composition and microstructure Briquettes exhibited their original morphology without obvious shrinkage. Main phases were corrading structure of CF and hematite. A part of CF had small needle-like morphology, and the joint of CF and hematite was ablated obvious.

Fig. 1(c)

Briquettes had a certain extent shrink, and the ratio of binder phase was increased. At the same time, a certain amount of silicate mineral began to appear, which consolidated hematite with CF jointly, and the hematite was dominated by primary ore.

6.0%

Fig. 1(d)

All briquettes exhibited molten state after sintering completely. The ratio of binder phase was increased, and majority was silicate mineral with a certain crystal morphology. Secondary hematite developed rapidly among silicate mineral, and skeletal rhombohedral hematite was found at local area nearby the cavern that having rapid cooling rate.

7.0%

Fig. 1(e)

Despite the microstructure referred above,porphyritic structure consisted of glass and hematite was appeared accompanying with the ratio of secondary hematite increasing.

5.0%

On the one hand, in order to keep R constant, the ratio of calcium oxide in mixture would increase as the addition quantity of SÍO2 increasing. The per-contact probability of CaO, SÍO2, Fe203 was increasing, which provided a comfortable condition for solid state reaction and generation of original melt, and increased the ratio of binder phase in briquette. On the other hand, binder phase transformed from ferrite to silicate when m(Si02)=5%. Fe203, which enter into melt by the form of CaOFe203 initially, would precipitate. Only if CaO is excessive after reacting with SÍO2 and FeO, more CF crystal would appear. As Si0 2 increasing, more silicate is formed accompanying with CF+SÍO2—>CS+Fe203. Effect of the AbO? Content on Condensing Crystallization Process By adding AI2O3, we controlled its content to 1.5%, 2%, 3% and 5%, meanwhile, inputting relevant amount calcium oxide to keep its ratio at 8%. The microstructure is shown in Fig. 4 and Table 5.

394

d)Al 2 0 3 =3% e)Al 2 0 3 =5% White—Hematite,Gray—CF,Tawny—Al203,Black—Void Fig.4 Micrograph of different AI2O3 concentration (x200) Table 5

Mineral composition and microstructure of different AI2O3 concentration

m(Al 2 0 3 )

Micrograph

Mineral composition and microstructure

0.77%

Fig. 5(a)

The crystallization CF with imperfect morphology was formed, and the main minerals was CF and hematite.

1.5%

Fig. 5(b)

The amount of CF began to increase slightly accompanying with apperence of slender needle-like particles.

2.0%

Fig. 5(c)

The ratio of hematite was obvious decreased. CF binder phase further improve, and its appearance of crystallization transformed to rough needle and short rod forms.

3.0%

Fig. 5(d)

5.0%

Fig. 5(e)

Some A12C>3 mineralizing imperfectly (tawny area) was surround by CF. Simultaneously flake and plate-like morphology dominated the CF crystal, which occupied more than 80% of all minerals.

For CF, whether integrated morphology, or development of quantity, or centralized appearance surround residual A12Û3 particles, it all proved that A12Ü3 could promote precipitation of CF significantly. The primary cause is, alumínate (CaO*Al 2 03,Ca02Al 2 03) are easily formed in the high temperature stage, and this intermediate compounds had good stability

395

and reactivity. The solid solution reaction of CF and alumínate is beneficial to the steadily existing of CF in melt. Conclusions Crystalline CF began to precipitate at about 1200 °C. The lower the cooling temperature, the better the crystal growth, and the more the content of the precipitated needle-like CF. CF crystal has rapid growing speed, strong crystal ability, and little affection by dynamics and external factors. However,decreasing the cooling rate is obviously favorable to the development of crystal. Binder phase transformed from CF system to silicate system when SÍO2 content is increased to 5-6%, with a constant R=2.0. Along with the increase of AI2O3 content, the morphology of CF transform from acicular crystal to thick needle, short column, and finally to flake and plate. Increasing the ratio of AI2O3 could promote precipitation of CF and development of crystal. References [1] T. Miyashita, N. Sakamoto, and Hiroshi Fukuyo, "Influence of Mineral Structure on Sinter Agglomerate's Propeties, "Trans ISIJ , 22 (7)(1982):B-199. [2] Y. Ishikawa, "Improvement of Sinter Quality Based on the Mineralogical Propeties of Ores,"Ironmaking Proc, 42(1983): 17-29. [3] X.R. Liu, G.Z. Qiu, R.Z. Cai,and S.F. Lin, "The Mineralogy Research on the Ferrite of Low-temperature Sintering,"Sintering and Pelletizing, 25(2) (2000):7-10. (in Chinese) [4] X.H. Fan, J. Meng, XL. Chen, J.M. Zhuang, Y Li,and L.S. Yuan, "Influence Factors of Calcium Ferrite Formation in Iron Ore Sintering, " Journal of Central South University, 36(6) (2008) :1125-1131. (in Chinese) [5] X.M. Quo, The Formation of Calcium Ferrite and Mineralogy during Sintering Process (Beijin: Metallurgical Industry Press,1999) (in Chinese) [6] J.Y. FU, T. JIANG, and D.Q. ZHU, Sintering and Pelletizing (Changsha:Central South University of Technology Press,1995) (in Chinese)

396

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Jiann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

Enrichment Behavior of Phosphorus in CaO-Si02-FeO,-P 2 05 Based Slag Nan Wang, Yingying Shen, Zhen Tian, Min Chen School of Materials and Metallurgy, Northeastern University Shenyang, China Keywords: Phosphorus, Enrichment behavior, 2CaOSi0 2 particle, Phosphorus containing phase Abstract The enrichment behavior of phosphorus in CaO-Si02-FeOx-P20j slag was studied by making an investigation on the distribution of phosphorous content in slag and 2CaOSi02 phase. The research results showed that the existing 2CaOSi0 2 solid particles in slag were the sites for the phosphorous enrichment. The enrichment process of phosphorus in CaO-SiCh-FeOx-PiOs slag can be recognized as: mass transfer of phosphorus from bulk slag toward the surface of 2CaOSiC>2 particle, superficial dissolution reaction of phosphorus to 2CaOSiC>2 and diffusion of phosphorus through the product layer of 2CaOSi02-3CaOP205 solid solution to the inner of 2CaOSiC>2 particle. Compared to the 2CaOSiC>2 particles with larger size (500um to lOOOum), the completion time of enrichment was shorter for the particles with smaller size (less than 50um). Introduction During ironmaking and steelmaking, significant amounts of slag are produced. Contrast to the high utilization of ironmaking slag, the treatment of steelmaking slag is more difficult and challengeable, and therefore, an alternative approach is to recycle the steelmaking slag inside the ironmaking and steelmaking process itself. The main limitation to the reuse of steelmaking slag inside the process is that the refining process may be aggravated caused by the phosphorous accumulation if it is used in the blast furnace burden [1]. At the same time, the phosphorous content of steelmaking slag (approximately 1%~5% P2O5) is too low for it to be used as raw material for fertilizer [2, 3]. Moreover, owing to the diminishing supply of global iron ores with low phosphorous content, the control of phosphorus in iron and steelmaking processes has become an increasing problem, and as a result, more amounts of high phosphorus containing steelmaking slag are produced. Therefore, a valuable approach is quite appreciated in which the steelmaking slag is recycled within the ironmaking and steelmaking process and at the same time, the high phosphorus-containing phases in the steelmaking slag, suitable for use as fertilizer can be separated and recycled effectively [4]. It has been reported that the predominate phosphorus in steelmaking slag was found concentrated in the solid solution of dicalcium silicate and tricalcium phosphate, while very less phosphorus was found in glass phase and other mineral [5]. The maximum solid solubility of P2O5 in 2CaO-Si02-3CaO-P205 solid solution was reported to be up to 20% [6], unfortunately, the P 2 0 5 contents in the phosphorus containing phases are very low due to the fact that the basicity of converter slag is usually high and the precipitated amounts of 2CaOSiC>2 particles during cooling are very large, resulting in the P2O5 contents in 2CaOSi02-3CaOP 2 05 solid solution phase are much lower than its solid solubility limit. Although the precipitated 2CaO-Si0 2 particles can be considered as the concentrated sites of phosphorus in slag [7-10], the phosphorous contents in 2CaO-Si02-3CaO-P205 solid solution would be diluted with the increasing amounts of precipitated 2CaOSiC>2 particles. Therefore, in order to increase the phosphorous contents in 2CaO-Si02-3CaO-P20s phase, it is significant to control the precipitated

397

amounts of 2CaOSi0 2 particles strictly by adjusting the basicity and chemical composition of steelmaking slag to the proper range, and to improve the complete enrichment of phosphorus from slag to a small quantity of precipitated 2CaOSi02 particles. By this way, the higher phosphorus containing phases can be formed and separated from the steelmaking slag effectively. As we know, the distance of phosphorus transferring from bulk slag to 2CaOSi02 particles would be increased with the decreasing amounts of the precipitated 2CaOSiC>2 particles during cooling, and at the same time, larger solid P2O5 condensed grains are favorable to be separated from the steelmaking slag, but the enrichment of phosphorus to larger 2CaOSi02 particles may be incomplete during cooling which is mainly affected by the transferring rate of phosphorus to 2CaOSi0 2 particles. Therefore, in this present study, the enrichment process of phosphorus in CaO-Si02-FetO-P205 slag was studied and the mass transfer of phosphorus from bulk slag to 2CaOSiC>2 particle as well as diffusion inside the 2CaOSiC>2 particle was discussed by making an investigation on the P2O5 content in bulk slag and in 2CaOSi02 solid phase for different temperatures, reaction times and 2Ca0Si02particle sizes, and the theoretical basis to improve the phosphorous enrichment to 2CaOSi02 particle was established. Experimental Sample Preparation (l)2CaOSi0 2 Particle The CaO used in the present experiment was prepared through the calcination of reagent grade CaCC>3 at 1223K. Later the prepared CaO and reagent grade SÍO2 at the molar ratio of 2:1 was well mixed and pressed, followed by heating at 1073K. for 8h and then at 1773K for 8h. In order to prevent the dusting of 2CaOSi0 2 , lmass% B2O3 was added to the mixture. In the following, the sintered sample was ground and sieved, and the 2CaOSi02 particles with the size range of 44 to 74um as well as 500 to lOOOum were prepared. Figure 1 shows the composition confirmation of the prepared 2CaOSiC>2 by X-ray diffraction analysis.

Î 20

30

40 29/°

50

60

Fig. 1 XRD pattern of prepared 2CaOSi02 particle (2) CaO-Si02-FeOx-P205 Slag Sample The mixture of the prepared CaO, reagent grade SÍO2, reagent grade 3CaO-P205 and reagent grade C2Fe04-2H20 was heated at 873K for 2h to decarbonization and then the mixture was melted at 1723K. in an aluminum crucible under Ar atmosphere for 24h and then quenched quickly. In the following, the quenched sample was ground to below -5mm by a stainless steel ball mill. The chemical compositions of the synthetic slag are shown in Table 1. Procedure

398

The prepared 2CaOSi02 particles were uniformly mixed with the synthetic CaO-Si02-FeOxP2O5 slag powder in 5 mass% proportion, and then under high purity Ar atmosphere, a tablet obtained by pressing was heated in an aluminum crucible at 1623K and 1673K respectively. After different reaction times, the slag sample was taken and quenched by Argon quenching. The phosphorous contents in bulk slag and 2CaOSiC>2 particle were analyzed by SEM/ EDS. Table 1 Chemical composition of the synthetic CaO-Si02-FeOx-P2C>5 slag CaO(%) Si02(%) FeO„(%) P205(%) , Ci»0/sio2 (molar ration) 40.83 29.17 20.00 10.00 1.40 Results and Discussion Enrichment Process Figure 2 shows the SEM images of the samples heated at 1623K for 300s and 600s, respectively. During the reaction time, the enrichment of phosphorus to the smaller 2CaOSiC>2 particles (less than 50um) was complete and the smaller grains confirmed as 2CaO-Si02-3CaO-P205 (C2S-C3P) solid solution were observed (point A showed in Fig.2, C2S-C3P solid solution particle). On the other hand, for the same reaction time, only a thin layer of 2CaO-Si02-3CaO-P20s solid solution were found around the rim of 2CaOSi02 (C2S) particles for those larger 2CaOSi02 particles with the size larger than 200um (point B showed in Fig.2, larger C2S particle with a thin product layer of C2S-C3P solid solution). Moreover, the thickness of 2CaO-Si02-3CaO-P20s layer was found to be larger for 600s than for 300s.

Fig. 2 SEM image of 2CaOSi02 particle in CaO-Si02-FeOx-P205 slag at 1623K ((a)reactiontime 300s ; (b) reaction time 600s)

Fig. 3 SEM images of 2CaOSi02 particle in CaO-Si02-FeOx-P205 slag at 1673K ((a) reaction time 300s ; (b) reaction time 600s) Figure 3 shows the SEM images of the samples heated at 1400°C for 300s and 600s, respectively. Similarly, smaller 2CaOSi02-3CaOP2C>5 solid solution grains formed by dissolution of

399

phosphorus to the smaller 2CaOSi02 particles were observed, while only a thin product layer composed of 2CaOSi02-3CaOP2C>5 solid solution was confirmed around the larger 2CaOSiC>2 grains. Thus, the enrichment process of phosphorus from bulk slag to solid 2CaOSiC>2 particles can be considered as: mass transfer of phosphorus from bulk slag toward the surface of 2CaO-Si0 2 particle, superficial dissolution reaction of phosphorus to 2CaOSi0 2 and diffusion of phosphorus through the product layer of 2CaO-Si02-3CaO-P205 solid solution to the inner of 2CaOSi0 2 particle. Distribution of Phosphorus Content inside 2CaO SiO? Particle Figure 4 shows the distribution of phosphorus content inside the product layer as well as the inner of 2CaOSi02 particles for different reaction times. The phosphorus content is higher inside the product layer, while almost no phosphor was found inside the inner of 2CaOSiC>2 particle. The thickness of 2CaOSi02-3CaO-P20s product layer increases with the increasing reaction time.

Fig. 4 Line analysis of P-Ka of 2CaOSi02 grain at 1623K ((a) reaction time 300s ; (b) reaction time 1200s) Figure 5 shows the chosen positions from the bulk slag to the inner of 2CaOSi02 grain for phosphorus content analysis and the EDS results are shown in Table 2. The phosphorus content shows an increasing tendency from the bulk slag to the C2S-C3P product layer and almost no phosphorus was found inside the inner of 2CaOSi02 grain. Therefore, the presumption can be made that the phosphorus was concentrated from the bulk slag to the surface of 2CaOSiC>2 particle and then to the inner of 2CaOSiC>2 particle by diffusion gradually.

Fig. 5 Chosen positions of slag sample for EDS Table 2 P2Os contents at different positions from bulk slag to 2CaOSi02 particle Position P205( mass% )

1

2

3

4

5

6

2.46

15.04

15.33

14.42

0

0

400

Bulk slag

2CaOSi02-3CaOP205 solid solution layer

Inner of 2CaOSi02 particle

Impact of Temperature and 2CaOSiO;> Particle Size Table 3 shows the phosphorous contents inside C2S-C3P product layer at 1623K. as well as 1673K after reaction for 300s, 600s, 1200s and 3600s, respectively. It can be found that higher temperature is proper to the phosphorous enrichment to 2CaOSi02 particle due to the lower viscosity of slag as temperature gets higher and the improved dynamic condition of phosphorus transferring from bulk slag to 2CaOSiC>2. Table3 Average phosphorous content inside C2S-C;P layer Temperature

Reaction time(s) 300

600

1200

3600

1350°C

13.50

14.35

14.17

15.23

1400°C

14.95

14.70

15.10

16.48

As Figure 5 showed, the phosphorous enrichment for the 2CaOSiC>2 particles with 44~74um was basically completed during 300s at 1623K and 2CaOSiC>2 phase were transformed to C2SC3P5 solid solution completely. As shown in Figure 6, for the larger 2CaOSi02 particles with the size of 500~1000um, the phosphorous enrichment was incomplete during the same reaction temperature and time. The completion time of enrichment was shorter for the smaller 2CaOSi02 particles, but the larger ones were favorable for the separation of phosphorous-containing from bulk slag. Thus, it is necessary to improve the enrichment conditions of phosphorus enrichment to promote the enrichment of phosphorus to larger 2CaOSiC>2 particles.

Fig. 5 SEM image of 2CaOSi02 particle (a) in slag sample (b) P content inside 2CaOSi02 paticale (particle size 44~74um, 1623K, reaction time 300s)

Fig. 6 SEM image of 2CaOSi02 particle (a) in slag sample; (b) P content inside 2CaOSi02 particle (particle size 500~1000um, 1623K, reaction size 300s) Conclusions The enrichment behavior of phosphorous in CaO-Si02-FeOx-P2C>5 slag was studied and the controlling step of the transfer process of phosphorous from slag to 2CaOSi02 solid particles

401

was discovered by making an investigation on distribution of phosphorous content in slag, inside and outside the 2CaOSi02 particle at different temperature. The results are as follows: (1) Phosphorus CaO-SiCh-FeOx^Os slag was concentrated to 2CaOSiC>2 particle and the 2CaO-SiÔ2 particles were the enrichment sites. (2) Phosphorus was concentrated from the bulk slag to the surface of 2CaOSiC>2 particle and then to the inner of 2CaOSiC>2 particle by diffusion gradually. (3) Higher temperature can improve the phosphorous enrichment from slag to 2CaOSi02 particle, the phosphorous content in CaO-Si02-3CaOP2Os solid solution obtained at 1673K. was higher than obtained at 1623K. (4) 2CaOSi0 2 particles with size less than 50um were changed to 2CaOSi02-3CaOP 2 0 5 solid solution completely in 300s, but in the case of 2CaOSi0 2 particles with larger size (500um to lOOOum), only a thin layer of 2CaOSiC>2-3CaOP205 solid solution at the rim of particles was found for the same reaction time. Compared to the larger 2CaOSiQ2 particles. Acknowledgments This research work was supported by the National Natural Science Foundation of China (Grant No. 50874130) and the the National Natural Science Foundation of China (Grant No. 50974034). References [1] R. Dippenaar, "Industrial Uses of Slag (the Use and Re-Use of Iron and Steelmaking Slags)," lronmaking and steelmaking, 32 (1)(2005), 35-46. [2] K Morita et al., "Resurrection of the Iron and Phosphorus Resource in Steel-Making Slag," J Mater Cycles Waste Manag, 4 (2002), 93-101. [3] R. Boom, S. Riaz, and K. C. Mills, "Slags and Fluxes Entering the New Millennium, an Analysis of Recent Trends in Research and Development," lronmaking and steelmaking, 32(1) (2005), 21-25. [4] H. J. Li, H. Suito, and M. Tokuda, "Thermodynamic Analysis of Slag Recycling Using a Slag Regenerator," ISIJInternational, 35(9)(1995), 1079-1088. [5] L. S. Li et al., "Distribution of Phosphorus in Converter Steel Slags Stabilized in Different Ways," Metal Mine, 364(10) (2006), 78-89. [6] D. K. Agrawal, A. R. Maslowski, and J. H. Adair, "Evolution of the Formation of Inorganic Polymers in the CaO—SÍO2—P2O5 System Using Metal Alkoxides," J. Am. Ceram. Soc, 73(2)(1990), 430-434. [7] R. Inoue, H, Suito, "Phosphorous Partition between 2CaOSi0 2 Particles and 2CaO-Si0 2 Fe,0 Slags," ISIJ International, 46(2) (2006), 174-179. [8] R. Inoue, H, Suito, "Mechanism of Dephosphorization with CaO-SiCh-FetO Slags Containing Mesoscopic Scale 2CaOSi0 2 Particles," ISIJ International, 46(2) (2006), 188-194. [9] K. I. Shimauchi, S. Y. Kitamura, and H. Shibata, "Distribution of P2O5 between Solid Dicalcium Silicate and Liquid Phases in CaO-Si02-Fe203 System," ISIJ International, 49(4) (2009), 505-511. [10] X. Yang, H Matsuura, and F Tsukihashi, "Condensation of P2O5 at the Interface between 2CaOSi0 2 and CaO-Si0 2 -FeO x -P 2 0 5 Slag," ISIJ International, 49(9) (2009), 1298-1307.

402

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

NUMERIC SIMULATION OF THE COOLING PROCESS OF THE IRON ORE SINTER Jiaqing Yin, Xuewei Lv, Chenguang Bai, Guibao Qiu College of Materials Science & Engineering, Chongqing University, Chongqing 400044, China Emails: [email protected]. [email protected]. Keywords: Sintering, Cooling, 2-dimentional model, Flowfield,Temperature field

Abstract The iron ore sinter is one of the main raw materials for the blast furnace (BF) process. In the production of iron ore sinter, the cooling speed of the sinter determines its quality. In addition, the heat recovery devices are installed in the cooling machine for energy saving. The cooling parameters like gas flow rate and the size of sinter should be optimized for getting the gas with a higher temperature and a good quality of sinter simultaneously. In this study, the commercial simulation software COMSOL was used to establish a model for simulating the cooling process of the iron ore sinter. Moreover, three different gas flow rate (1.5, 3, 5m/s) and three different diameter of sintered particles (50, 70, 100mm) were investigated with the model. As a result, in certain porosity, a higher cooling speed was achieved when smaller size or larger gas flow rate was used. And lower gas flow rate was beneficial for the heat recovery. The model could intuitively present the exchange of heat between the sinters and air, as well as the flow and temperaturefieldsduring the cooling process. Introduction Iron ore sinter plays a great role in iron and steel industry for many Asian countries such as China, Japan and India. As a part of the sintering procedure, the impact of cooling process on the quality of iron ore sinter couldn't be neglected. On the other hand, the massive iron ore sinters which were at high temperature contain extremely large amount of heat. However, it was reported that there's very few of them recovered in China with an average recovery rate of nearly 10.5% at 2004 [1]. Thus, more reasonable cooling procedure should be developed to get high quality product and heat recovery rate simultaneously. Recently, some researchers had pointed out that the cooling schemes could influence the sinter's physical and metallurgical properties [2-4]. In addition, Dong et 403

al. investigated the factors that influence the cooling of iron ore sinter through experiments [5]. Leong et al. established a model with FLUENT to simulate the cooling process of iron ore sinter [6]. Eiki KASAI et al. had experimentally investigated the relationship between permeation characteristics and structural parameters of sinter [7-9]. A lot of work had been done which concentrate on simulation and analysis about schemes of heat recovery from packed bed. Caputo et al. represented a work on the optimization of heat recovery in gas-solid moving beds using simulation approach [10-11]. Laguerre et. al developed a transient model of heat transfer in a packed bed and compared it with simulation and experimental study [12]. Jang et al. successfully established a 3-D transient turbulent model for cooling process of sintered bed which was set as a packed 4-row bed of spheres [13]. The objective of this work is to study on how the gas flow rate and the size of sinter influence the cooling process of iron ore sinter. The flow fields and temperature fields of five cases were analyzed in this study. Methods Computational Model The main equations are shown as following [14]: p C p ^ + V ( - ( K + KT)VTf) = Q + q s T f -pC p uVT f K

T

=CP^

(1) (2)

p C p ^ + V ( - K + KTVTs) = Q + q s T s -pC p uVT s

(3)

The k-£ turbulent model was used in this model. Its governing equations are : p-^- + /9(ûV)û = V[-pî+(7+^ T )(VG + (VÛ) T --(VÛ)ï--pkï] + F dt 3 3

r)k p—+puVk at

^ + V(pu) = 0,p = p(p,T) at n 2 = V-[(77+-^)Vk] + 77 T P(u)-(-pk)Vu-/?i? 3 <7k

P^+pûV£ = V [ ( ^ + % V £ ] + Î ^ W ( û ) - ( ^ ) V Û ] P(ü) = VÜ:(Vu + (VG) T )--(V-ü) 2

404

(4) (5) (6)

Ce2p£2 k

(7) (8)

v=

d

d

dx dy

(9)

Where, CM=0.09, Cel=1.44, 0,2=1.92, ak=1.0, aE=1.3. Assumptions and Conditions

Fig. 1 The heat transfer modes in the model. The heat transfer modes in this study were shown as Fig.l. The geometries and initial conditions were fixed as following. The shape of iron ore sinter was supposed to be sphere. A few rows of separated iron ore sinters filled into a 1.38*0.3m2 area, so the height of sinter layer was 1380mm and cooling air was blew in from the bottom. The original temperature of sinter was 820°Q whereas cooling air blew in at an ambient temperature of 20°C The properties of the sinters were assumed to be similar to those of iron, whose density was 7870 kg/m3, heat capacity was 440 J/(kgK), and thermal conductivity was 76.2 W/(m • K). The properties of air were as a function of temperature and its values were obtained from the material library of COMSOL. The sphere-sphere radiation heat transfer was neglected and the wall of sinter cooler was supposed to be heat insulated. Five cases shown as Table I were designed in this study, in which three different particle sphere diameter (D=50,70,100mm) and three different gas flow rate (V¡n=1.5, 3, 5m/s) were investigated.

Case 1

Table I. Five cases studied in this paper D(mm) Diameter Porosity 0.583 100 405

V,„(m/s) Gasflowrate 5

Case 2 Case 3 Case 4 Case 5

70 50 70 70

0.583 0.583 0.583 0.583

5 5 3 1.5

Results and Discussions Different Size of Particles

Fig.2 Temperature distribution of sinters at different cooling time (Sec) for case 3.

Fig.3 The mean effective thermal conductivity andflowfield after 0.9s cooling for case 3.

406

Fig.2 illuminates the temperature field of iron ore sinter at various cooling time for case 3. It shows that the sinters was quickly cooled down to around 150 °Cwithin 30 minutes. However, the temperature of particles has a big range which indicates the cooling process is unhomogeneous. Besides, it could be seen that the sinter was cooled from outside to inside and those close to walls as well as those at a lower location were firstly cooled down. Fig.3 was plotted to explain this phenomena. The effective thermal conductivity (Keff) is defined as the sum of thermal conductivity (K) and turbulent thermal conductivity (KT). As K is influenced only by the temperature gradient in this case, the KT seems to play a great role for the value of Kejf. In Fig.3, the velocity field on the right hand shows that higher velocity was achieved at locations close to walls where the mean thermal conductivity is higher. For another reason, the air was blew in from bottom and slowly heated up as it flowed up, so the temperature gradient of the two phases was bigger at lower location. The average temperature of sintered particles in various sizes varying with the time is presented in Fig.4. Fig.4 indicates that the cooling speed of iron ore sinter is larger when the sintered particles were smaller. The two main reasons are as follows. Firstly, as the diameter of particle ranges from 100mm to 50mm, the contact area of two phases increases. Another main factor is the effective thermal conductivity. As mentioned before, the effective thermal conductivity consists two parts, K and KT. As the calculation results shows that K is at a lower order of magnitude compared to the KT, the KT was the main factor contribute to cooling process. On the other hand, KT was influenced by Cp, p, k and s, and these factors were greatly influenced by the flow field. The average velocity of air for particles in different size was presented as Fig.5. Compare the Fig.4 with Fig.5, curves for particles in larger diameters were smoother than those of smaller ones. This may indicate the velocity has a big influence on the turbulent thermal conductivity. Even so, the curves of velocity could not explain the cooling speed of iron ore sinter shown in Fig.4 completely. This is by reason of the change of temperature of air also took part in the changing of turbulent thermal conductivity.

Fig.4 The average temperature of sintered particles in various size vs. time.

Fig.5 The average velocity of air for sintered particles in various size vs. time

Fig.6 shows the temperature of air at the outlet for various size particles and 407

cooling time. It was found that the distribution of gas temperature along the outlet was determined by the size of sintered particles, as well as its distribution in the cooling machine. Since the smaller size particles were cooled more quickly than the bigger ones as displayed in Fig.4, the cooling air in this case was heated up to higher temperature. Besides, the velocity of air is another key factor influencing the gas temperature at the outlet. Therefore, the temperature distribution is different with the former ones when the cooling time was 701s. Comparing the third figure of Fig.6 with Fig.5, the same conclusion can be got.

Fig.6 The temperature distribution of air at the outlet for sintered particles in various size, while the cooling time was 11s, 101s and 701s, respectively. Different Gas Flow Rate The average temperature of iron ore sinter depending on cooling time is presented in Fig.7. It demonstrates that the higher gas flow rate could accelerate the cooling process. Nevertheless, this impact seems to be weaker when the flow rate was changed from 3m/s to 5m/s. It is meaning there's a extreme gas flow rate that will effectively cool the iron ore sinter down. Fig.8 shows the average velocity of air varying with the cooling time under various gas flow rate. It is obvious that the higher gas flow rate produces a flow field with higher average velocity in the whole cooling time. Thus, the turbulent energy and the turbulent heat conductivity were larger, achieving a better heat exchange condition.

Cooling time(Sec) Fig.7 The average temperature of sintered

408

Cooling time(Sec) Fig.8 The average velocity of air in various

particles in various gasflowrate vs. time.

gasflowrate vs. time

Fig.9 The temperature distribution of air at the outlet for various gasflowrate, while the cooling time was 101s, 701s and 1001s, respectively. Fig .9 illustrates the gas temperature distribution at the outlet in three stages during the cooling process. The higher gas flow rate resulted in the lower gas temperature at the outlet. It could be seen that the temperature of gas at the outlet was around 250°C after 701s under the condition of 5m/s, and 550°C under the condition of 1.5m/s. Meanwhile, the lower flow rate could get relatively higher temperature for a longer time which is beneficial for the heat recovery. However, when low gas flow rate was used, the cooling speed of sinter would be slow. The optimization of gas flow rate should be designed considering both cooling speed and the heat recovery. Conclusions The flow field and heat transfer during the cooling process of iron ore sinter were studied with the COMSOL. A two-dimensional transient model was established by considering the iron ore sinters to be spheres. The effects of different size of sinter and gas flow rate on the cooling process were examined in detail. As a result, the one with smaller size got a higher cooling speed and the gas temperature at the outlet was higher in the early stage of cooling process. Moreover, the lower gas flow rate produced higher gas temperature at the outlet but the cooling speed of sinter was slower simultaneously. The flow fields and temperature fields obtained in this work could be used to do further research about the cooling process of iron ore sinter, so the quality of product and heat recovery rate could be improved. Nomenclature p density, kg/m3 Cp heat capacity at constant pressure, J/(kg-K) qs represents production/absorption coefficient, W/(m 3 -K)

K thermal conductivity, W/(m • K) Kx turbulent heat conductivity, W/(m-K) 3 Q heat source, W/m Tf fluid Emperature, K

Ts

u

solid emperature, K

409

velocity vector, m/s

x] k PrT D Keff

dynamic viscosity, Pa's r\j turbulent khetic energy s turbulent Prandtl number cp diameter of particle, mm V¡n effective thermal conductivity ,W/(m • K)

turbulent dynamic viscosity dissipation rate of turbulence energy porosity gas flow rate, m/s

References [I] C.X. Zhang, Y.H. Qi et al., "Analysis on the Secondary Energy and pollutant for sintering and iron making" (Paper presented at the National Energy and Thermal Engineering Annual Meeting, Kun Ming, China, 9 November 2004), 578-582. [2] P. He, Y.L. JIN, and Q. H, "HAN. Study of Influence of Cooling Conditions on Properties of Sinters," Iron and Steel (in Chinese) , 42(11)(2007), 5-7. [3] W.S. Wang, Q. Lv et al., "Influence of Cooling Mode on Sintering Process and Strength of Vanadium-titanium Sinter," Journal of He be i institute of technology (in Chinese), 29(1)(2007), 12-14. [4] Y.X. Zhang, Y. Zheng et al., "Effect of Cooling Mode on Microstructure and Mechanical Properties of Ti (C,N)-based Cermets," Cemented Carbide (in Chinese, 26(1)(2009), 20-23. [5] H. Dong, J. Li et al., "Experimental Study on Cooling Process of Sinter," Journal ofNortheastern University (Natural Science) (in Chinese), 31(5)(2010), 689-692. [6] Jik-chang Leong, Kai-wun Jin et al., "Effect of sinter layer porosity distribution on flow and temperature fields in a sinter coole," International Journal ofMinerals, Metallurgy and Materials, 16(3)(2009), 265-272. [7] Eiki KASAl WJR, Roy R.LOVEL and Yasuo OMOR, "An Analysis of the Structure of Iron Ore Sinter Caker," ISIJ 29(8)(1989), 635-641. [8] Eiki KASAl BB, Yasuo OMORI, Noboru SAKAMOTO and Akria KUMASAKA, "Permeation Characteristics and Void Structure of Iron Ore Sinter Cake," ISIJ, 31(11)(1991), 1266-1291. [9] Eiki KASAl WJR, Roy R.LOVEL and F.GANNON, "The Effect of Raw Mixture Properties on Bed Permeability during Sintering," ISIJ, 29(1)(1989), 33-42. [10] A. C. Caputo, G Cardarelli, and P. M. Pelagagge, "Analysis of heat recovery in gas-solid moving beds using a simulation approach," Appl. Therm. Eng., 16(1) (1996), 89-99. [II] P. M. Pelagagge, A. C. Caputo, and G Cardarelli, "Comparing heat recovery schemes in solid bed cooling," Appl. Therm. Eng., 17(11)(1997), 1045-1054. [12] O. Laguerre, S. Ben Amara et al., "Transient heat transfer by free convection in a packed bed of spheres: Comparison between two modeling approaches and experimental results," Appl. Therm. Eng., 28(2008), 14-24 [13] Jiin-Yuh Jang, Yu-Wei Chiu, "3-D Transient conjugated heat transfer and fluid flow analysis for the cooling process of sintered bed," Appl. Therm. Eng., 29(2009), 2895-2903 [14] COMSOL user guide. The General Heat Transfer Application Mode. 2009.

410

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

AUTHOR INDEX 2nd International Symposium on HighTemperature Metallurgical Processing A

Andriese, M

B

Bai,C Bai, G Baker, A Bastrygin, V Bell, W Bhoi, B Bing, H Borkar, S Bystrov, V

c

Chen, D Chen, G Chen, M Choi, M Chou, K Chun, T Cui, Y

D

Dai, L Demiray, Y Dreiich, J

E

Emrich, M

F

Fan,J Fan, X Fan,Z

G

Gan, M Geveci,A Guang-bin, X

Guo, S Guo, Y

119, 185, 345 127, 327

51, 77, 379

H

Han, G Hao, H Haxhiaj, A He,C Hu, B Hu, L Hu, P Hua, W Huang, C Huang, D Huang, R Huang, X Huang, Y Huang, Z Hwang, J

25,403 245,299 263 11 51, 77, 379 363 15 379 59

151 119 397 3 161 69,151, 199,211,237,309,335 151

J

Jiang, T Jie,C

K

101 255 289

279, 299 327 289 185 111 371, 389 15 15 137 41 25 51 279 I l l , 169 51,77,379

33, 127, 169, 191,245,279,327, 371,389 161

Kamfjord, N Kamkin, R Kampe, S Keskinkilic, E Kugler,G Kuznetsov, A

211

L

Lei, L Lei, Y Li, D Li, G Li, J Li,Q Li,W Li, Y Liao, Z Liu, B Liu,J

327 33, 371, 389 33

371, 389 319 15

41 1

85 59 263 319 271 59

25 345 185 33, 169, 191, 245, 279 69, 245 191 345,371,389 119, 345 111 41 119

Liu, L Liu, S Luo, H Lv,X

M

Mamaev, A Masset, P Mendes, V Mishra, C Mishra, H Morita, K

N

Nabei,A Nœss, M

O

Okabe, T.

119, 185 127 119 25,403

Sun, Z

T

Tahautdinov, R Tercelj, M Tian, Z Topkaya,Y Tranell, G

59 95, 177 69,211,237 363 363 143

u

Upadhyaya, A..

229 85

Vernigora, A Villa, R

w

.143

Wang, H Wang,N Wang, Q Wang, Y Wiraseranee, C

P

Padilla, R Pan.J

221 69, 151, 199,211,237,309,335, 355 Panishev, N 11 Paretsky, V 59 Peng,J 41, 101, 119, 185,345 Peng,Z 51,77,379 Posazhennikov, A 11 Pournaderi, S 319

Q

Qiu, G

R

Rao, M Ruan, F Ruiz, M

S

Schmetterer, C Shen,Y Sohn, H Su,Z

X

Xiao, Y Xie, L Xie,Lu Xie, S Xin,X Xu, J

Y

25, 127,403

Yamaguchi, K Yang, Y Yi, L Yin,J Yu, H Yu, W Yu,Z Yücel, 0

245 161 221

177 397 3 299

Z

Zahrah, T Zhang, D

412

279

11 271 397 319 85

.379

59 221

3, 119 137,397 371, 389 371 143

137, 309 389 371 77 137 161

229 137 111 403 191 199, 335 389 255

263 299

Zhang, H Zhang, J Zhang, K Zhang, L Zhang, Y Zhang, Z Zheng, G Zhou, X Zhu, D

191 161 169 119, 185, 345 33,95, 111, 169, 191,279,299 137 355 355 69, 151, 199, 211, 237, 309, 335, 355 Zhu, H 101 Zhu-cheng, H 15 Zou, Z 137

413

2nd International Symposium on High-Temperature Metallurgical Processing Edited by: Mann-Yang Hwang, Jaroslaw Drelich, Jerome Downey, Tao Jiang, and Mark Cooksey TMS (The Minerals, Metals & Materials Society), 2011

SUBJECT INDEX 2nd International Symposium on HighTemperature Metallurgical Processing 2

2CaOSi0 2 Particle

D

397

Decomposition Dielectric Parameters Dielectric Polarization Direct Reduced Iron (DRI) Direct Reduction

A

Acid Leaching 127 Activated Carbon 77 Activation Roasting 127 Additives 245 Agglomeration 289, 355 Air Heat Exchanger 119 Aluminosilicate and Alumínate Systems.... 161 Ammonium Durante 41

B

Ball Milling Bentonite Dosage Bismuthinite Blast Furnace Blast Furnace Burden Bomb Calorimetry Bond Work Index Boron Bearing Additives Bottom Ash Boundary Layer

c

Calcination Calcium Ferrite CALPHAD CaO-Si02-FeOx-MgO System Carbides Carbothermal Reduction Celestite Chlorination Chromite Chromite Over Burden (COB) Coal-Based Direct Reduction Coal-Based Grate-Rotary Composite Agglomeration Process Composite Binder Comprehensive Utilization Condensed Silica Fume Cooling Cooling Rate Crystallization Behavior Cubic Zr0 2 Curie Temperature

E

End Point Temperature Enrichment Behavior Equilibrium Phase Relation

F

Fe-Ni Ferronickel FexO Reduction Field Attenuation Fired Oxide Pellets Fired Pellet Fired Pellets Firing Behavior Flow Field Flue Gas Circulation Sintering Flue Gas Desulphurization Flux Fluxed Pellets

345 327 221 355 169 263 379 199 137 85

41,319 389 137 137 271 25 255 151 245 363 15 69 191 169 169 85 403 389 389 185 379

H

Half-Power Depth Hematite High Press Grinding Roller High Pressure Roll Grinding High-Temperature Hot Workability Humic Acid Hydrated Lime Hydrogen Reduction

I

Ilmenite Injection Interface Structure Intermetallics Intrinsic Kinetics Iron Nugget Iron Ore

415

221 51 77 101 69, 111, 309

15 397 137

363 319,355 95 51 69 237 199 335 403 33 33 371 335

51 199 237 327 119 271 299 335 3

25 59 185 263 3 11 371

Iron Ore Sintering Ironmaking ITmk3

Kinetics Kinetics of Crystal Grain Growth

Ladle Refining Laterite Lateritic Ore Lead Sulfide Liquid Silicon Low Intensity Magnetic Separation

N

85 319 363 289 85 309

363 3,69,211 237 245 255 169 211, 335 151 271 51, 119 119 379 15, 101, 111 345 371 143 185

O

R

345 185 229 327 355 191 211 77 397 397 143 15 237 69 151 151 59 151 59

Reaction Synthesis Recycling Reduction Roasting Relaxation Time Response Surface Methodology Response Surface Methodology (RSM) Reverse Flotation Rhodium Roasting Rotary Hearth Furnace

S

Nickel Laterite Non-Ferrous Metals Nucleation and Growth Kinetics Nucleation and Growth of Iron Crystal Grain

Optimization Organic Binder Oxidation Oxidized Pellet

Panzhihua Low Grade Ilmenite Partially Stabilized Zirconia(PSZ) PbO-Si02 Slag Pellet Porosity Pelletization Pelletizing Pellets Permittivity Phosphorus Phosphorus Containing Phase Platinum Group Metals Porosity Preheated Pellet Preheated Pellets Prerduction Prereduced Pellet Processing Pyrite Cinder Pyrometallurgy

95 199

L

Magnetic Separation Magnetite Concentrate Manganese Ore Fines Metallization Metallothermic Reduction Metallurgical Dusts and Sludges Metallurgical Performance Metallurgical Slag Microstructure Microwave Microwave Absorbing Material Microwave Energy Microwave Heating Microwave Reduction Mineralization Molten Slag Monoclinic ZrO?

279,299 161

P

K

M

Oxidized Pellets Oxygen To Alumina Ratio

33, 389 3, 191 11

Sessile Drop Technique Short Rotary Furnace Siderite Sintering Si0 2 Al 2 0 3 Content Size Distribution Slag Slime Smelting Smelting Smelting Reduction Solid-State Reduction Specific Surface Area Spéculante Spéculante Concentrate

355 59 3 111

345 279,299 85 111

416

263 143 363 77 41 101 309 143 211 11

177 229 11 191,371,403 389 327 95,177 229 289 363 245 327 191 299

Strand Grate - Rotary Kiln Strontium Sulfide Ore Superfine Low Grade Hematite Surface Tension Synthetic Rutile

T

Temperature Field Temperature Rising Characteristics Titanium Diboride Titanium Dioxide Titanium Slag Tool Steels Total Uranium Two-Dimensional Model

u

U4+ U308

V

Vale Hematite Vanadium/Titanium-Bearing Magnetite Vanukov Furnace Vaporization Viscosity Vitrification

151 255 379 309 177 127

403 15 263 25 127 271 41 403

41 41

211 279 59 221 161 137

417

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