Impinging Streams Fundamentals- Properties- Applications
Impinging Streams Fundamentals- Properties- Applications
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Impinging Streams Fundamentals- Properties- Applications
Yuan Wu College of Chemical Engineering and Pharmacy Wuhan hlstitute of Technology Wuhan 430073 PR China
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Elsevier B.V. Radarweg 29, PO Box 21 l, 1000 AE Amsterdam, The Netherlands Chemical Industry Press No. 13, Qingnianhu South Street, Dongcheng District, Beijing 10001 l, P.R. China First edition in English 2007 Copyright © 2007 Elsevier B.V and Chemical Industry Press This edition is jointly published by Elsevier and Chemical Industry Press, P.R. China 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 or otherwise without the prior written permission of the publisher Permissions may be sought directly from Elsevier' s Science & Technology Rights Department in Oxford, UK: phone (+44) (0) 1865 843830; fax (+44) (0) 1865 853333; email:
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ISBN-13:978-0-444-53037-0 ISBN-10:0-444-53037-1 Impinging Streams: Fundamentals, Properties, Applications By Yuan Wu
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Foreword Impinging Streams (IS) is a novel technical method. As a scientific concept, it was first presented by Elperin in 1961, while its earliest emergence can be traced back to the development and application of the Koppers-Totzek gasifier in 1953. The original idea of IS is to send two solid-gas streams to impinge against each other at high velocity with the aim of enhancing transfer between phases. It is interesting to note that the core of the Koppers-Totzek gasifier first applied industrially effectively enhances transfer between phases. Without doubt, IS is very effective for this purpose. The results of a large number of investigations have shown that transfer coefficients between gas and solid in IS can be ten times, or even several tens of times, greater than those in traditional equipment. Because of the universality of transfer phenomena, IS has received widespread attention and has been the subject of extensive investigations. In the past 40 years and more, investigations into IS have been in three stages. The first, from 1961 to the early 1970s. was a newly established stage and the work was concentrated mainly in the former Soviet Union. Naturally, the target systems were mainly those with gas as the continuous phase, because the concept of IS was originally aimed at transfer enhancement in such systems, while the dispersed phase in IS was gradually extended to include liquid. In the second stage, from 1974 to the mid-1990s, the core of the research moved to Israel where work was mainly carried out by A. Tamir and his group. Most (about 80%) of the target systems were still those with gas as the continuous phase. Although several systems with liquid as the continuous phase, such as dissolution of salts, emulsion, extraction, etc'., were also studied, the ideas were simply analogous and the goal was still transfer enhancement. The depth and scope of investigations into the liquid-continuous phase were never comparable with those of the gas-continuous phase. The last ten years and more may be considered as the third stage. Researchers from more than twenty countries, including China, the United States, Canada, Germany, etc., have been engaged in investigations into IS, the number of investigations now overtaking those carried out in Russia and Israel. On the other hand, the emphasis at this stage changed to investigating IS with liquid as the continuous phase. As a technical method, IS can never be a universal tool. The method of gascontinuous impinging streams (GIS) is indeed very efficient in enhancing transfer. However, it has the fatal disadvantage of very short residence times (about 1 s only) in the active region and has much higher requirements for flow configuration arrangement than traditional devices. Some very fast processes, such as burning of sprayed liquid
fuel or powdery coal etc., can be carried out in GIS with greatly increased efficiency. In practice, however, many processes cannot finish instantaneously but have to last for a very long time. On the other hand, any arrangement of a multistage IS must lead to an excessively complicated system and these disadvantages, therefore, limit the application of pure IS to a considerable extent. For a long time, these problems were not fully addressed and this is the main reason for the slow progress made in the application of IS, although there might also be other reasons. One of the essential conditions for carrying out impinging streams is that there must be, at least, one continuous phase, which can be either a gas or a liquid. If a liquid is chosen, the dispersed phase should either be solid or another liquid, soluble or not. Otherwise, the employment of IS has limited use. Because of the difference in aggregation statuses, the physical properties of liquid and gas are quite different from each other. Normally, the density of liquid is of the same order as solid; it is greater than that of gas by 103 times, while liquid viscosity is about 102 times that of gas. As a result, the factors effectively enhancing transfer in GIS, e.g., very large relative velocity between phases, penetration of particles to and fro between the opposing streams, etc., no longer exist or become very weak in LIS, so that the transfer coefficient in LIS becomes essentially no different from that created by traditional methods. In fact, there already existed experimental data, including those by Tamir, showing such a situation. Unfortunately, LIS was still considered to be superior to traditional methods even in enhancing transfer, and thus few investigations were made into the property and performance differences between LIS and GIS before the 1990s. On the other hand, the high density of liquid brings new features to LIS. The discovery in the 1990s that LIS promotes micromixing very efficiently is the most important advance in the field of IS. It was also found recently that very strong pressure fluctuation exists in LIS. Such phenomena must relate to the fact that the impingement of two opposing streams of high density against each other leads to their much stronger interaction. Certainly, enhanced micromixing and pressure fluctuation are of significance for processes occurring at the molecular scale, especially those involving chemical reaction(s). The above discovery therefore enables the fields of IS application to be greatly expanded. In fact, the results of a number of investigations have shown the perfect performance and great application potential of LIS in the preparation of ultrafine products by reaction-precipitation, etc. Two books devoted to IS have previously been published. The first is "Transport Processes in Opposing Jets" by I. T. Elperin (Nauka I Tekhnica, Minsk, 1972; in Russian). It summarized the investigations in the field before 1970 and its depth and scope are consistent with the achievements made in the newly established stage of IS investigation. The second book is "Impinging-Stream Reactors: Fundamentals and Applications" by A. Tamir (Elsevier, 1994), which was translated into Chinese for my Chinese colleagues (The Chemical Industry Press of China, Beijing, 1996). It is a systematic summary of the works in the second stage of IS investigation. Based on the understanding that "almost any process in chemical engineering can be carried out by applying impinging streams", Tamir's investigations extended over almost all chemical
vi
unit operations. The topics discussed in the contents are numerous and the data are full and accurate. Without exaggeration, the book was almost an inspiration for me, even though it could not cover the achievements made in the last ten years, both in understanding and technologies, e.g., the perfect and valuable properties of LIS, etc. My investigation into IS began in 1992 and also started with GIS. Afterwards, the emphasis was diverted to LIS because of the intrinsic disadvantages of GIS and the perfect nature of LIS; consequently the latter will be one of the focal points of the discussions in this book, distinguishing it from the two books mentioned above. Science and technology continually grow and progress: and our understanding is constantly improved. This book cannot and should not be the last one related to IS. It is both believed and expected that, when the next book emerges, the industrial application of IS will have greatly evolved and become universal. This book was originally written in Chinese, and its publication was supported by the National Natural Science Foundation of China. This English edition is not simply a translation: some corrections, revisions and supplementary information have been made in order to improve its contents. I am very pleased that the publication of this English edition has enabled the book to be available to more of my colleagues, especially those from English-speaking countries. Yuan Wu Wuhan, China
January 31, 2006
vii
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Acknowledgments My investigations on Impinging Streams (IS) were twice supported by the National Natural Science Foundation of China (No29276260, 20176043); and also by the Key Laboratory of Open Investigation on Multiphase Reactions, Institute of Process Engineering, Academy of China; the Natural Science Foundation of Zhejiang Province, China: the Key Laboratory on Polymerization Reaction Engineering, Zhejiang University; and the Education Department of Hubei Province e t c . Without this support the author would not have been able to carry out the investigations involved in many projects and thus could not have written this book. I would like to express my sincere appreciation to all these supporters. During my investigation on IS. a number of my colleagues collaborated with me, undertaking sub-projects or specific research jobs to effectively push the investigations forward. They include Associate Professors Yuxin Zhou, Gao'an Wu, Jianmin Xu, Deshu Li, Chuanping Bao, Jun Yuan. Anqig Shu, Senior Engineers Xiaoping He and Lecturer Tielin Wang from Wuhan Institute of Technology; Associate Professors A'san Yang, Qin Sun, Rong Cheng, Yungen Chen and Huayan Liu from Zhejiang University of Technology; Senior Engineer Jingnian Xu from Beijing University of Chemical Technology, My PhD students for Huaiyu Sun, Qin Li, Jianwei Zhang; ME students Kai Huang, Yu Chen, Yang Xiao. Guochao Li, Zhen Chen and Fang Li, took various phases of the IS investigation as topics for their theses, making great efforts and substantially promoting the fundamental studies and technical developments. Many undergraduate students joined the investigations on IS and obtained a lot of useful experimental data. In addition, during many discussions with me, Dr. Xiaoxi Wu offered a number of helpful comments and suggestions All of the persons mentioned above have contributed significantly, and I would like express my thanks to them. I was also deeply moved that Academician Professor Yong Jin has warmly supported the publication of the Chinese edition of this book. Finally, I must mention my wife, Senior Engineer Yuqiong Huang. In order to support my work, she essentially forsook her own research and took over all the household jobs as an understanding wife and loving mother so that I could concentrate on my investigations. Her indirect contribution to this book has been invaluable. Once again, I would like to sincerely thank all those mentioned above. Yuan Wu
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Contents Foreword
........................................................................................................................
Acknowledgments
..........................................................................................................
v
ix
I n t r o d u c t i o n ....................................................................................................................
1 2 3 4 5
Part I
E n h a n c e m e n t of transfer between phases and origin of i m p i n g i n g streams ........................................................................................................ 1 Basic principles of impinging streams ........................................................ 4 Experimental evidence for enhancing transfer ........................................... 6 Other performances of impinging streams .................................................. 7 Extension of impinging stream technology ................................................ 9 5.1 Extension in flow configuration .................................................... 9 5.2 Extension in phase ...................................................................... I I Gas-Continuous
Flow of Continuous
1.1 1.2
1.3 1.4
Impinging
S t r e a m s ............................................................
P h a s e ..................................................................................
Flow characteristics .................................................................................. Velocity field in laminar impinging streams ............................................ 1.2.1 General equations ........................................................................ 1.2.2 Planar two-dimensional impinging streams ................................ 1.2.3 A x i a l - s y m m e t r i c impinging streams ........................................... 1.2.4 General three-dimensional impinging streams ............................ 1.2.5 Viscous impinging streams ......................................................... Experimental results for the flow field in impinging streams ................... Turbulent impinging streams ....................................................................
P a r t i c l e B e h a v i o r .................................................................................................
2.1
2.2 2.3
i
17 19
19 25 25 26 28 30 31 32 36 41
M o t i o n of a single particle in co-axial horizontal i m p i n g i n g streams ...... 41 2. I. 1 Qualitative description ................................................................ 41 2.1.2 Basic relationship for the particle motion ................................... 2.1.3 Solutions of the motion equation for various stages ................... 2.1.4 Residence time of the particle in the i m p i n g e m e n t zone ............. E x p e r i m e n t a l results on the behavior of a single particle in co-axial horizontal t w o - i m p i n g i n g streams ............................................................
43 45 51
Behavior of a single particle in co-axial vertical impinging streams ........ 2.3. l Description of motion p h e n o m e n a .............................................. 2.3.2 Motion equation and its solution ................................................. 2.3.3 Terminal velocity ........................................................................
56 56 57 58
xi
52
2.4
Behavior of particle crowds in impinging streams ................................... 2.4.1 Distribution of particle concentration in impinging streams ....... 2.4.2 Influence of particle concentration in feed streams ..................... 2.4.3 The influence of collision between particles ...............................
59 60 63 65
Residence Time of Particles and its Distribution ............................................. 67 3.1
3.2
3.3 3.4
3.5
Theoretical consideration ......................................................................... 3.1.1 Impinging stream device ............................................................. 3.1.2 Constituents of R T D of particles in the ISC ................................ 3.1.3 M o d e l for the overall residence time distribution ....................... M e t h o d for experimental m e a s u r e m e n t of particles' residence time distribution ................................................................................................ 3.2.1 Input signal ...................................................................................... 3.2.2 Data interpretation ........................................................................... Relationships for fitting data .................................................................... Major experimental results for R T D of particles ...................................... 3.4.1 M e a s u r e m e n t of tracer concentration .......................................... 3.4.2 C o m p a r i s o n b e t w e e n the results measured and simulated .......... 3.4.3 M e a n residence times of particles ............................................... Remarks ....................................................................................................
67 68 69 75 77 77 81 84 86 86 87 88 89
Hydraulic Resistance of Impinging Stream Devices ........................................ 91 4.1
4.2
4.3
4.4
Theoretical consideration ......................................................................... 92 4.1.1 F l o w through the accelerating tubes ............................................ 92 4.1.2 I m p i n g e m e n t b e t w e e n opposing streams .................................... 94 4.1.3 Resistance due to the structure of the IS device .......................... 95 4.1.4 Overall resistance of the IS contactor ......................................... 96 Experimental equipment and procedure ................................................... 96 4.2.1 Experimental equipment ............................................................. 96 4.2.2 Experimental procedure .............................................................. 97 M a j o r results from the experimental study ............................................... 98 4.3.1 Basic characteristics of pressure drop distribution ...................... 98 4.3.2 Resistance of accelerating tubes to pure air flow ........................ 99 4.3.3 Pressure drop due to acceleration and collisions of particles .... 100 4.3.4 Resistance due to structure of the device .................................. 102 4.3.5 M o d e l for the overall pressure drop .......................................... 103 Evaluation of p o w e r consumption and discussions related to application .............................................................................................. 105
Influence of Impinging Streams on Dispersity of Liquids ............................. 107 5.1 Statement of the p r o b l e m ........................................................................ 107 5.2 Experimental equipment and procedure ................................................. 109 5.2.1 Impinging stream device ........................................................... 109 5.2.2 M e t h o d for m e a s u r e m e n t of droplet size distribution ............... 110 5.2.3 A r r a n g e m e n t of sampling .......................................................... 110 5.3 M a j o r results of the investigation ........................................................... 111 xii
5.4
5.3.1 Size distribution of droplets ...................................................... 111 5.3.2 M e a n diameter of droplets ........................................................ 115 C o n c l u d i n g remarks ................................................................................ 117
Impinging Stream Drying ................................................................................ 119 6.1 Introduction ............................................................................................ 119 6.2
6.3
6.4
Earlier research and d e v e l o p m e n t ........................................................... 6.2.1 I m p i n g i n g stream spray drying ................................................. 6.2.2 Impinging stream drying of granular materials ......................... 6.2.3 Impinging stream drying combinations ..................................... Circulative impinging stream drying ...................................................... 6.3.1 Basic ideas for e q u i p m e n t design .............................................. 6.3.2 Structure and working principles of the dryer ........................... 6.3.3 Experimental model e q u i p m e n t scheme and procedure ............ 6.3.4 Major results of the model experiments .................................... 6.3.5 Influences of structural and operating parameters .................... Concluding remarks ................................................................................
121 121 123 128 134 134 135 137 139 142 151
Impinging Stream Absorption ......................................................................... 153 7.1 7.2
7.3 7.4
7.5
7.6
Adaptability of impinging streams for g a s - l i q u i d reaction systems ....... Earlier investigations .............................................................................. 7.2.1 M o d e l s tk)r absorption e n h a n c e m e n t ......................................... 7.2.2 Absorption equipments ............................................................. 7.2.3 Major results of the investigations ............................................ W e t desulfurization of flue gas (I) General considerations .................... W e t desulfurization of flue gas (II) Investigations in Israel ................... 7.4.1 Experimental equipment and procedure .................................... 7.4.2 Major results ............................................................................. W e t desulfurization of flue gas (III) Investigations in China ................. 7.5.1 Experimental e q u i p m e n t ........................................................... 7.5.2 Experimental scheme and procedure ......................................... 7.5.3 Data interpretation ..................................................................... 7.5.4 Results and discussion .............................................................. 7.5.5 Conclusions ............................................................................... Design of a device for large gas flow rates .............................................
153 155 155 156 160 162 164 164 166 169 169 172 174 176 186 186
Impinging Streams Combustion and Grinding .............................................. 191 8.1
M o d e l s for particles and droplets c o m b u s t i o n ........................................ 191
8.2 8.3
8.1.1 E v a p o r a t i o n - b u r n i n g equations tot a single droplet ................. 8.1.2 Burning equations for a single particle ..................................... Intensification of c o m b u s t i o n processes due to impinging streams ........ Impinging stream combustors .................................................................
8.4
8.3.1 Furnaces for gas and liquid fuels .............................................. 198 8.3.2 K o p p e r s - T o t z e k gasifier for p o w d e r y coals ............................. 199 Impinging stream grinding ..................................................................... 201 xiii
191 194 196 198
Part II Liquid-Continuous Impinging Streams .................................................... 205 Differences Between Properties of Continuous Phases and Classification of Impinging Streams ....................................................................................... 207 9.1 9.2
9.3
10
207 208 208 208 211
Micromixing In Liquid-Continuous Impinging Streams ............................... 213 10.1 10.2
10.3
10.4
10.5 10.6
11
Progress of investigation on liquid-continuous impinging streams ........ Differences between properties of continuous phases and their influences on the performance of impinging streams ............................. 9.2.1 Differences between properties of liquid and gas ..................... 9.2.2 Influences of property differences on the performance of impinging streams ..................................................................... Supplementary classification of impinging streams ...............................
Macromixing and micromixing .............................................................. Methods for investigation of mixing problems ....................................... 10.2.1 Macromixing ............................................................................. 10.2.2 Micromixing ............................................................................. Flow and macromixing in SCISR ........................................................... 10.3.1 Design ideas and basic structure of SCISR ............................... 10.3.2 Macromixing time ..................................................................... 10.3.3 Flow configuration and residence time distribution .................. Micromixing in SCISR ........................................................................... 10.4.1 Experimental equipment and procedure .................................... 10.4.2 Governing variable and its experimental measurement ............ 10.4.3 Experimental procedure ............................................................ 10.4.4 Major results for micromixing .................................................. 10.4.5 Comparison between micromixing performances of SCISR and STR .................................................................................... 10.4.6 Comparison between measured and theoretically predicted results for micromixing time ..................................................... 10.4.7 Relationship between macro- and micro-mixing ...................... Micromixing in impinging stream reactor without circulation ............... Comparison between the investigations on micromixing in LIS as concluding remarks ................................................................................
213 214 214 215 216 216 218 219 222 222 224 226 226 229 230 232 233 235
Pressure Fluctuation in the Submerged Circulative Impinging Stream Reactor .................................................................................................. 237 11.1
11.2
11.3
Investigation method of pressure fluctuation .......................................... 11.1.1 Meaning of pressure fluctuation ................................................ 11.1.2 Investigation method of pressure fluctuation ............................ Experimental equipment and procedure ................................................. 11.2.1 Experimental equipment ........................................................... 11.2.2 Measurement and control of the impinging velocity ................. 11.2.3 Arrangement of measuring points and sampling frequency ..... 11.2.4 Pre-treatment of the experimental data ..................................... Experimental results and discussion .......................................................
xiv
237 237 238 240 240 241 241 242 242
11.4 12
12.3
12.4
Qualitative analysis lbr the influences of pressure fluctuation and m i c r o m i x i n g ..................................................................................... 12.2 Crystal-growth kinetics of di-sodium phosphate ....................... 12.2.1 Basic principles ......................................................................... 12.2.2 E x p e r i m e n t a l investigation ........................................................ Kinetics of ethyl acetate saponification .................................................. 12.3.1 C h e m i c a l reaction and experimental m e t h o d ............................ 12.3.2 Major results ............................................................................. C o n c l u d i n g remarks ................................................................................
253 254 254 257 265 265 265 266
Preparation of Ultrafine Powders by Reaction-Precipitation in Impinging Streams I: "Ultrafine" White Carbon Black ............................... 269 13.1 13.2 13.3
13.4
13.5
14
242 245 246 247 249 250
Influence of Liquid-Continuous Impinging Streams on Process Kinetics .............................................................................................................. 253 12.1
13
11.3.1 Intensive fluctuation region ....................................................... 11.3.2 V o l u m e t r i c distribution of fluctuation intensity ........................ l 1.3.3 Definition of the i m p i n g e m e n t zone .......................................... 11.3.4 Influence of the impinging velocity on fluctuation intensity .... 11.3.5 Power spectrum analysis for pressure fluctuation ..................... Conclusions and discussion ....................................................................
Adaptability of liquid-continuous i m p i n g i n g streams for preparation of ultrafine powders ................................................................................ Properties of white carbon black and chemical reactions in its preparation by precipitation processes .................................................... Experimental e q u i p m e n t and procedure ................................................. 13.3.1 Experimental e q u i p m e n t ........................................................... 13.2.2 Experimental procedure ............................................................ Results and discussions .......................................................................... 13.4.1 Semi-batch operation ................................................................ 13.4.2 Continuous operation of the S C I S R .......................................... 13.4.3 C o m p a r a t i v e e x p e r i m e n t s in semi-batch operation ................... 13.4.4 Study of the final treatment of the reaction product .................. Conclusions ............................................................................................
270 271 273 273 274 275 275 278 279 280 281
Preparation of Ultrafine Powders by Reaction-Precipitation in Impinging Streams II: Nano Copper and its Surface Improvement ............ 283 14. I 14.2 14.3
14.4
Introduction ............................................................................................ Properties and main uses of nano copper ................................................ Principles and experimental m e t h o d ....................................................... 14.3.1 Chemical reactions in preparation of nano copper by reduction-precipitation ..............................................................
283 284 286
14.3.2 Results 14.4.1 14.4.2
287 288 288 290
Experimental e q u i p m e n t and procedure .................................... and discussions on the preparation of nano copper p o w d e r ....... Major results obtained in the first stage .................................... Results on the influences of various factors .............................. XV
286
14.4.3 Preparation experiments under optimal conditions ................... 296 Surface i m p r o v e m e n t of nano copper: preparation of Cu-Ag double metal powder .......................................................................................... 297 Conclusions ............................................................................................ 299
14.5 14.6
15
Preparation of Ultrafine Powders by Reaction-Precipitation in Impinging Streams III: Nano Titania ................................................................................ 301 15.1 15.2 15.3
Properties of nano titania and chemical reactions in its preparation ....... Experimental equipment and procedure ................................................. Results and discussions .......................................................................... 15.3.1 M a j o r results obtained in the first stage .................................... 15.3.2 Experiments and major results in the second stage ................... 15.3.3 Experiments of mass preparation and the results ...................... 15.3.4 Experiments of neutralization with aqua a m m o n i a ................... 15.3.5 Experiments for final optimization of conditions and the results .................................................................................. 15.3.6 Comparative experiments between SCISR and STR and the results .................................................................................. Conclusions ............................................................................................
15.4
16
313 313 314
Preparation of Ultrafine Powders by Reaction-Precipitation in Impinging Streams IV" Nano Hydroxyapatite ............................................... 317 16.1 Introduction ............................................................................................ 317 16.2
Experimental equipment and procedure ................................................. 16.2.1 E q u i p m e n t ................................................................................. 16.2.2 Procedure of the experimental operation .................................. Results and discussions .......................................................................... 16.3.1 Influences of some factors ........................................................ 16.3.2 Optimal conditions for synthesis of nano H A P ......................... 16.3.3 Characterization of nano H A P product ..................................... Concluding remarks ................................................................................
16.3
16.4 17
301 303 304 304 306 309 312
318 318 319 320 320 324 324 326
Research and Development of Liquid-Continuous Impinging Stream Devices and Application Forecasting .............................................................. 329 17.1 17.2 17.3
The vertical circulative impinging stream reactor .................................. 329 Impinging stream crystallizer ................................................................. 333 Prospects for the application of liquid-continuous impinging streams... 337
Postscript
....................................................................................................................
339
References
...................................................................................................................
341
..............................................................................................................
355
Index ...............................................................................................................
361
Nomenclature Subject
xvi
INTRODUCTION
1 ENHANCEMENT OF TRANSFER BETWEEN PHASES AND ORIGIN OF IMPINGING STREAMS Heat and mass transfer, especially mass transfer, in multiphase systems are problems commonly encountered in processing units in the chemical, petrochemical, and many other process industries. Because transfer rates significantly affect the efficiencies and technical-economic indexes of the processes, the enhancement of transfer has been a continuing topic of interest in chemical engineering since the late 1930s. A vast number of theoretical and experimental investigations have been carried out in the search for new methods of enhancing transfer between phases. According to the theory of transfer rates, the amount of heat or mass transferred per unit time can be expressed by Amount transferred per unit time =
driving force xinterface area specific resistance
(1)
Therefore one, or a combination, of the following measures can be used to increase the amount to be transferred per unit time: (1) Enhancing driving force; (2) Increasing interface area; and (3) Reducing specific resistance. All three measures are, of course, effective in principle. However, their potential to enhance transfer and the degree of difficulty in carrying them out are quite different in practice. The driving forces of heat and mass transfer are temperature and concentration gradients, respectively. To a considerable extent, they are limited by the characteristics of the specific processes involved, such as stocks, heat sources, and equipment materials, etc. In most cases only a limited increasing magnitude is permitted. Relatively, increasing the interface area, i.e., enhancing the dispersal of a liquid or a solid, is a measure that can be employed widely, and, in fact, has been applied successfully in a number processes, such as spray drying and cooling etc. However, it is also limited to an extent. For example, spray drying can only be applied in the production of powdery products, and excessive dispersion may give rise to difficulties in powder collection etc: while spray cooling is only applicable to the cases where moisture increase is permitted. On the other hand, in common equipment systems, the maximum relative velocity between phases is mostly just equal to the terminal velocity, u~, which
2
IMPINGING STREAMS
decreases sharply as the particle/droplet size reduces. This may partially offset the effect of the increase in interface area for enhancing overall transfer rate. In comparison, the reduction of specific resistance is an effective way of enhancing transfer between phases and has great potential. It is generally considered that there exist three resistances in series in transfer processes of gas-solid, gas-liquid, liquid-liquid, and liquid-solid systems: gas or liquid side resistance, the so-called external resistance, interface resistance, and internal resistance of particle/droplet. The interface resistance possibly results from the accumulation of impurities on the interface. Reduction of any one of these three types of resistance can enhance transfer processes. The overall specific resistance for heat or mass transfer is the reciprocal of the heat or mass transfer coefficient, U or K, where U and K are the common parameters characterizing heat and mass transfer rates, respectively, defined as U=~
Q
AAT
K -
nA
(2) (3)
AAC A
They represent heat and mass fluxes with unit gradients of temperature and concentration, respectively, and so can be considered as specific heat and mass transfer rates. In the last few decades, the results of a large number of investigations into drying, absorption, cooling, combustion etc, have shown that, apart from the natures of the systems involved, including dispersion degree, the major factor influencing heat and mass transfer coefficients is the relative velocity between phases, Ur. An increase in the relative velocity results in enhanced turbulence and reduced thickness of the boundary layer, and also favors surface renewing of the liquid side. As a result, the transfer resistance of the gas or liquid side is reduced. Synthesizing existing experimental results of various unit operations, it can be concluded that transfer coefficients are exponential functions of the relative velocity: t
U -///r " and K - ///r"
Depending on the substance systems involved, the types of equipment and the operating conditions, the exponent, n" or n, varies approximately in the range from 1/3 to 4/5. For example, for particles of spherical form, Ranz-Marshall [1] obtained the following relationship for prediction of the film heat transfer coefficient: Sh = 2 + 0 . 6 R e l / 2 S c 1/3
(4)
For bubbles or suspended particles smaller than 2.5 mm in size, the Levich empirical equation below was recommended by Calderbank and Moo-Yong [2] for prediction of the heat transfer coefficient:
INTRODUCTION
3
Sh_l.()l(dpurl I/3 IDaB )
(5)
All the results mentioned above lead to a simple and clear conclusion: increasing the relative velocity between phases is one of the most effective approaches to enhance transfer processes. In traditional processing devices, increase in relative velocity is limited by various factors. For example, in column equipment the operating velocity must be smaller than that of liquid-flooding; the limitation of relative velocity in common gas-solid or liquid-solid suspensions is the terminal velocity, etc. It seems that other approaches must be found in order to raise the relative velocity between phases to higher levels. The efforts to search for approaches to raising relative velocity between phases has led to the development and/or application of impinging stream contactors, and also some other devices. The original conception of impinging streams (IS) is to bring two solid-in-gas suspension streams to flow in opposite directions at a considerably high velocity and impinge against each other, yielding extremely high relative velocity at the instant of impingement, and thus greatly enhance transfer between phases. As a scientific concept, IS was first proposed by Elperin [31 in 1961; while its application can be traced back to the development and application of the Koppers-Totzek gasifier in the early 1950s [4], although the term "'impinging streams" was not used at that time. In the period from the 1960s to the early 1970s, a large number of theoretical and experimental investigations on impinging streams were carried out, mainly by Elperin and his group. On the death of Elperin, the research core was moved to Israel. Tamir [5] carried on for over 20 years from 1974 until the 1990s, and his researches extended over almost all the unit operations in chemical engineering. All the results of investigations involving transfer processes show that impinging streams can increase transfer coefficients by large amplitudes. For instance, the heat transfer coefficient obtained by Elperin [6] from the experiments of wet particles drying is as high as 5800 W.m--~.K- I while that calculated with the assumption of relative velocity being of the order of fluidizing velocity is only of the value of 470 W.m-2-K-~ Another kind of device that efficiently enhances transfer processes in gas (vapor)liquid systems is the rotating packed bed (RPB), also called "HIGEE", presented in the 1960s [7, 8]. The basic idea tbr RPB design is that extremely high relative velocity can be employed with the action of centrifugal force produced by rotating the packed bed at high speed to enhance strongly the transfer between phases. In comparison, in traditional column equipments, such as packed tower and sieve plate column etc', the permitted operating relative velocities are bounded to low levels due to the limitation of liquid-flooding. There is yet another method which also enhances transfer very efficiently, in which a stream is induced to impact a fixed wall surface, i.e., the impinging jet (IJ). Obviously, the flow configuration and the action of stream impingement of IJ are totally different from the impinging streams, although it uses also the term "impinging"
4
IMPINGING STREAMS
[9]. The impinging jet has important applications in rapid heating and cooling, drying of coating layer, reaction, and surface cleaning e t c , and investigations in that field are also very active; but it is beyond the scope of the present book. All the researches, developments, and applications of IS, RPB, and IJ show the extreme importance of increasing relative velocity for enhancing transfer between phases.
2
BASIC
PRINCIPLES
OF
IMPINGING
STREAMS
As mentioned above, the original concept of impinging streams presented by Elperin [3] is to bring two equal solid-in-gas suspension streams formed after fully accelerating solid particles by gas to flow in opposite directions at a considerably high velocity and impinge against each other at the middle point between the two accelerating tubes, as shown in Fig. 1. The gas velocity at the outlet of the accelerating tubes can be as high as 20 m.s -~ or even higher, and the particles can theoretically be accelerated to a velocity near that of gas. The impingement between the two-phase streams causes an impingement zone of high turbulence with the highest concentration of particles [5], which provides excellent conditions for heat and mass transfer. In the case where the difference in densities of the two phases, e . g . , in a solid-in-gas suspension, particles would penetrate from one stream into the opposing one, and, just at the instant of penetration into the opposite stream, the relative velocity between particles and gas flow achieves a maximum value. After that, particles are decelerated due to the friction force of the opposing gas flow until particles achieve zero velocity. Thereafter, particles are accelerated by that gas flow in the opposite direction towards the impinging plane, and then penetrate into the stream that the particles originally existed in. After several repetitions of penetration to and fro between the opposing streams, particles gradually lose their axial velocity due to dynamic energy consumption, and are finally carried by the radial gas flow to leave the impingement zone.
ug
.'i'i, t OD +~ ~
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•
•
•
•
•
•
"el
Air+Particles
eio
~
•.-
,/o!o"
• • •• •o •
.~ ~
Accelerating tube
"p
"
ug ~
',
..
//~f..-'...'/-_..'...'_..'..-'//...'...'..-'..-'..-;.;-'/...'...'..-n
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~..........+.._..._........._..._............. _...............
•
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-
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. . . . . . .
Alr+varncJes
~.,.,......,...,...., . . . . ....,.,.. ....... ._+.,..,
/
",°~o , ~ Accelerating tube ',o.o, ".[., Impinging plane Impingement zone
Ur- Up--(-- Ug)= Up+ Ug
Figure 1 Basic configuration and principles of impinging streams.
INTRODUCTI()N
5
The phenomena of penetration to and fro between the opposing streams can occur even in a homogeneous gas system. Bley et al [10] observed experimentally that, when a gaseous He stream impinges against another co-axial stream of a mixture of He and SF(, SF(, molecules penetrate deeply into the pure He stream. In the impinging streams of solid-in-liquid suspensions, penetration phenomena may occur theoretically. However, because of low operating impinging velocity and large friction resistance of the opposed stream, perceiving the penetration of particles is difficult. Elperin and Tamir considered that, in impinging streams with gas as the continuous phase, transfer between phases is enhanced by the factors below: (1) Relative velocity between particles and the opposite gas flow is greatly increased. The relative velocity round the impinging plane, u,., may be calculated roughly by u, - .p - i - u ~ ) - up + u,~
(6)
where the velocity of gas, u~, can be considered approximately as constant, while that of particles, u~,, varies from time to time during penetration. At the exact instant the particles penetrate the opposite stream, the relative velocity achieves a maximum value, and, in the idealized case of particles being accelerated up to a value equal to the velocity of gas flow, then the maximum relative velocity can be twice that of the gas velocity (see Fig. 1). At all other instants during penetration, the relative velocity between particles and the opposite gas flow are larger than the gas velocity. In impinging stream devices for gas-solid systems, the operating velocity of gas flow inside the accelerating tubes is usually higher than l0 m.s -~ or, sometimes, even above 20 m-s -~. Obviously, the relative velocities in the traditional column devices can never be comparable with that in gas-solid impinging streams. (2) The penetration of particles to and fro between the opposing streams lengthens their residence time in the region active for transfer, i.e. the impingement zone, so that, to an extent, the conditions for enhancing transfer can continue for longer. Elperin [6] observed 5 to 8 times of particle penetration to and fro between the opposing streams in his experiments. For instant processes, such as combustion of powdery coal or fine droplets of oil, such an amplitude of residence time increase is of very great significance. The resulting global behavior of the residence time increase of particles is that the concentration of particles (or droplets) within the impingement zone is much higher than in any other regions. This implies that the impingement zone has much a larger interface area per unit volume for heat and mass transfer. (3) In the impinging streams of gas-liquid systems, high relative velocity between phases and collision between droplets favor surface renewing of droplets, resulting in reduced liquid film resistance and thus increased overall mass transfer coefficient.
6
IMPINGING STREAMS
(4) Impingement between the flows of continuous phase in the opposing streams, plus the oscillation movement of particle penetration, leads to strong mixing in the impingement zone, resulting in homogenization of temperature and composition. In some cases, this favors an increase in the average driving forces of heat and mass transfer and thus promotes the transfer processes. The problems related to mixing will be discussed further in later chapters of this book. It can be seen that impingement between the streams shown in Fig. 1 is of a "soft" nature. As stated above, its flow configuration and impingement action are totally different from the impinging jet impacting on a fixed wall surface, which is a "rigid" impingement.
3 EXPERIMENTAL EVIDENCE FOR ENHANCING TRANSFER As mentioned before, in his early investigation on drying of wet particles, Elperin [6] obtained powerful evidence that impinging streams enhance heat transfer. He determined the heat transfer coefficient is as high as 5800 W.m-2-K-~, while, in comparison, that calculated by the general empirical relationship, assuming the relative velocity to be of the order of the fluidizing velocity, is only about 470 W.m-2-K-~. In addition, the relationships of heat transfer coefficient and pressure drop versus relative velocity obtained by interpretation of experimental data are h c~ u 19 and Ap ~ u 2, respectively; while, by other technical methods, usually h c~ u°~3. This suggests that the employment of impinging streams will yield much higher efficiency. Tamir [5] tested the effectiveness of impinging streams in enhancing heat transfer by introducing a partition between the two opposing streams, which separates the impinging stream dryer into two non-interacting components. The results showed that the partition causes a significant reduction in heat transfer coefficient, h. The values for h in the case without partition are larger than those with partition by 1 to 2 times where the other conditions are the same. In their investigation on circulative impinging stream drying of PVC, Huang et al. [11] measured experimentally the value for the specific volumetric evaporation coefficient to be 16x10 -4 kg-s-~-m-3.K-~, which is about 10 times that in spray dryers, and a conclusion is also obtained similar to that by Elperin according to the heat transfer coefficient predicted from the evaporation coefficient just described above. On the enhancement of mass transfer, Tamir [5] studied the absorption of acetone into water with a similar method, i.e. using a partition. The results they obtained were: under suitable operating conditions and with appropriate structural parameters, the runs without partition yield absorption rates higher than those with partition by over 4 times. The experiments on combustion of powdery coal carried out by Ziv et al. [12] are another instance that shows impinging streams enhancing transfer. They measured the temperature profiles along the direction of flame length in the two cases with and without partition, and the results showed obviously higher temperature profiles in the case without partition than with partition, suggesting that impinging streams enhancing heat and mass transfer leads to stronger combustion.
INTRODUCTION
7
It should be noted that, in all the comparative experiments made by Tamir and his group, the cases without impingement are imitated by using a partition between the two opposing streams. In those cases each stream impinges on one side of the partition, which actually plays the role of a fixed wall surface. Thus, each piece of experimental equipment used with partition is equivalent to two impinging jets. As mentioned in Section 1, impinging jets also enhance transfer between phases very efficiently. Therefore the results of their comparative experiments could not reflect fully the influence of impinging streams on heat and mass transfer. This may account for the fact that the degree of enhancing heat transfer by impinging streams Tamir obtained in the study on drying (1 to 2 times) is much lower than those by Elperin (more than 10 times). There is still much more experimental evidence for impinging streams enhancing transfer. All the evidence, both that mentioned above and that not mentioned, supports the following conclusion: impinging streams are very efficient in enhancing transfer between phases, especially those controlled by diffusion through gas-film. Because transfer phenomena are widely encountered in various processing industries, the method of impinging streams undoubtedly has great potential application.
4 OTHER PERFORMANCES OF IMPINGING STREAMS Impinging streams were first suggested for enhancing transfer between gas and solid phases; however, the results of a large number of investigations have shown that, as well as this effect, the method of impinging streams has many other functions valuable for application; of course, at the same time, it also has some disadvantages, unbeneficial to application [9]. For the processes occurring in liquid phase or multiphase systems with a liquid as the continuous phase, the mixing status has significant effects on the efficiencies. The results of investigations since the 1990s showed that impinging streams have excellent performance for mixing. The most remarkable is that, because of the special flow configuration of the two opposing streams impinging against each other, impinging streams promote micromixing very efficiently [13]. In addition, the results of investigations by the author of the present book show that in liquid-continuous impinging streams there exists a pressure fluctuation of multi frequency in the range of sub-sonic waves, and the maximum amplitude can be as large as over 1 kPa. The details will be discussed in Chapter 11 of this book. Most likely, such pressure fluctuation is one of major reasons for impinging streams promoting micromixing efficiently. Yet, the pressure fluctuation favors kinetic processes, and this has also been proved by experiments. The phenomena of impinging streams promoting micromixing and the existence of the pressure fluctuation in liquid-continuous impinging streams had not been fully considered, and had not even been discovered in the investigations on impinging streams before the early 1990s. Consequently, the application potential of impinging streams for the processes of reactions and precipitation etc. received little attention, although the stagnation jet mixer developed by Brauer [14] was mentioned in Ref. [5]. In practice, since many processes are carried out at the molecular scale, the
IMPINGING STREAMS features of impinging streams promoting micromixing and the existence of pressure fluctuation are of very great value for application. In a number of researches and developments carried out in recent years [15-19], various impinging stream reactors of different structures were used for liquid reactions, reaction-precipitation or reaction crystallization etc. of various systems, and have performed well. Among the applications mentioned, the preparation of nano or sub-micrometer materials by impinging stream reaction-precipitation is an area of great potential. The milling effect resulting from strong collisions between particles in gas-solid impinging streams is another important feature of value for application. The most outstanding advantage of this technology of milling is that no milling material is needed, so that the substance being milled can be protected effectively from pollution. In addition, because milling proceeds in gas flows at high velocity, the phenomenon of overheating is also avoided so that the technology is applicable especially to substances of thermal sensitivity. Impinging stream milling technology was applied industrially as early as the 1970s. A typical example is the Trost Jet Mill [20]. The results of a large number of investigations and applications have shown that "ultrafine" products of the order of sub-micrometer can be produced by such a technology. As stated by the author of this book in the section "Translation Illustration" of the Chinese translation of Ref. [5] (The Chemical Industry Press of China, Beijing, 1996), as a technical method impinging streams cannot be a universal tool. It also has some disadvantages which limit its application. The most obvious problem is the very short residence time of the material in the active region of the impinging stream device. As will be discussed later, in gas-solid impinging streams the average residence time of solid particles is only about 1 s. On the other hand, the flow configuration of an impinging stream device is relatively more complex so that it becomes difficult to arrange a multistage system, such as in column devices. Tamir proposed several structures of multistage impinging stream contactors (refer to Fig. 3.2 in Ref. [5]). However, from the point of view of industrial application, they are obviously impractical. Most processes of industrial interest, such as drying materials containing porous moisture and/or combined water etc., need considerably longer time, even if they are carried out under conditions of significantly enhanced heat and mass transfer. Very short residence time, plus the difficulty of arranging multistage systems, significantly limits the fields in which impinging streams alone can be applied. For the processes restricted by equilibrium, single stage impinging streams can enhance heat and mass transfer to yield higher rates although it is difficult to ensure that the requested processing degree can be achieved. For example, the final absorption fraction or reaction conversion etc. may not achieve the level expected. Nevertheless, impinging streams, as a novel technical method, has a number of superior properties, among which the features of gas-continuous impinging streams enhancing transfer between phases and liquid-continuous impinging streams promoting micromixing have considerable value for application, and have found, and are finding, more and more applications.
INTRODUCTION
5 EXTENSION OF IMPINGING STREAM TECHNOLOGY Figure 1 represents the basic principles of gas-solid impinging streams, and also its essential structure as originally designed. On the basis of the essential structure, various devices can be constructed by extending the idea of impinging streams. Two extension schemes of IS have been proposed' extension of the flow configuration and extension of the phase conditions of the substance systems involved, as described below.
5.1 Extension in flow configuration Starting with the elements necessarily included, the concept of impinging streams can be extended to include various flow configurations. Talnir et al. [5] investigated a number of impinging stream contactors with different flow configurations" and the structures of some of the contactors they studied are shown in Fig. 2. A P
A
P
A------~
A(W) A
A A
P
p+ (P+W) A(W)
l (a)
(b)
A A
(c)
A ~
A A
A
------~ o (d)
P
(c)
Q,~L )
o t~---
(f)
Figure 2 Impinging stream contactors of various configurations [5]. (a) Coaxial-horizontal two impinging streams" (b) Horizontal three impinging streams (c) Coaxial-vertical two impinging streams; (d) Curvilinear two impinging streams; (e) Curvilinear four impinging streams; (f) Four impinging streams. A--air; P--particles: W--water.
10
IMPINGING STREAMS
In addition to those shown in Fig. 2, there are many other different structures. Different impinging stream devices may have different flow configurations, although all of them contain the same essential elements: (1) the streams flow in opposite directions and impinge against each other, and (2) each stream contains at least one continuous phase. Impinging stream equipment contains two types of part: (1) Accelerating tubes, which are also the conduits for feeding fluid of continuous phase; and (2) Equipment body with separate outlet ports for continuous and dispersed phases, respectively. Referring to Tamir's work, the following classification according to various features may be applicable for various impinging stream devices with different flow configurations:
Flow of the continuous phase: Parallel: the streamlines are parallel to the axis of flow, e.g. (a), (b) and (c) in Fig. 2. Rotational: the streamlines are helicoids with respect to the axis of flow, e.g. (d), (e) and (f) in Fig. 2.
Flow of the streams inside the device: Coaxial countercurrent: two streams enter the device in opposite directions along the same axis, and flow as free jets before impingement, e.g. (a) in Fig. 2. Eccentric countercurrent: as above but different flows are not on the same axis, e.g. (b) in Fig. 2. Co-plane-rotational: two streams enter the device tangentially and counter currently with the central lines in the same plane before impingement, and then flow on the wall of the device, with streamlines of a half circle form, e.g. (d) and (f) in Fig. 2. Non-co-plane-rotational: two streams enter the device tangentially and counter currently with the central lines in different planes before impingement, and then flow on the wall of the device, with streamlines of several half circle forms, e.g. (e) in Fig. 2.
Operation modes: Continuous two-side feeding: both phases flow at steady state, and particles are injected into both streams symmetrically; all devices shown in Fig. 2. Continuous one-side feeding: both phases flow at steady state, while particles are injected only into one stream. Semi-batch: only the continuous phase flows at steady state, while particles are circulated inside the device. In addition, according to the feature and number of impingement planes, devices can also be classified as stationary, moving, and multi impingement zone, etc. Readers may refer to Ref. [5].
INTRODUCTION
II
Any modification in flow configuration is aimed at: (1) Producing some advantages in operation or performance; and (2) Making the device more suitable for some specific systems. Very often, the latter is needed in practice, and usually it can be achieved by certain modification; while the former is somewhat complex. It can be seen from the various modifications proposed that, relatively, the common advantage obtained is that the total residence time of particles is lengthened in various degrees. However, essentially, the residence time in the active region could not be lengthened. On the other hand, the total residence time of particles is lengthened in some schemes although a price must be paid: (1) The structure of the equipment must be complicated; (2) The effects of impinging streams described in Section 2 must be weakened; and (3) In the cases of gas-solid impinging streams the resistance of the system must be increased remarkably. So, not every flow configuration in Fig. 2 is of practical significance. More ideal modification schemes with more advantages and fewer disadvantages may possibly be constructed in the future with further investigations into impinging streams. Among various flow configuration schemes, co-axial two impinging streams is the most essential and simplest; while its effects of enhancing transfer between phases and mixing are most significant. On the other hand, this scheme is the key for understanding principles and application of impinging streams. Therefore the discussions in the present book will take this scheme as the major topic.
5.2 Extension in phase conditions Obviously, one of the necessary conditions to carry out impinging streams is that both the opposed streams in impingement must have, at least, one continuous phase. In the impinging streams shown in Fig. 1 the continuous phase is a gas; although a liquid can of course also be the continuous phase. If a liquid is used, the dispersed phase should be a solid or another unmixable liquid. Otherwise, the employment of impinging streams would have less sense. The properties of a liquid are quite different from those of gas. These essential differences must result in different performances of impinging streams with gas and liquid as the continuous phase, respectively. The following facts are clear: (1) Liquid (L) is normally greater in density than gas (G) by three orders of magnitude. (2) L is larger in viscosity than G by two orders, and (3) G has a considerably larger molecular free path , while the molecules of stationary L can only vibrate and/or rotate with extremely small displacement round their balanced positions. Because of these significant differences, the behavioral features of the impinging streams with a liquid as the continuous phase are quite different from those with a gas as the continuous phase. As an example, consider here the case where the dispersed phase is a solid. When a liquid is taken as the continuous phase, the relative velocity cannot be large because the densities of solid and liquid have the same order of magnitude and the friction force between phases is very large. Furthermore, the phenomena of particles penetration to and fro between the opposing streams become non obvious and fine particles tend to follow streamlines. As the result, the enhancement of heat and mass transfer become
12
IMPINGING STREAMS
very weak, as has been proved by experimental data. This aspect will be discussed further later. On the other hand, since liquid is at condensed status and has large density, the interaction between two opposing liquid streams in continuous phase will be much stronger than gas streams, although its operating velocity is usually smaller than the latter. The strong micromixing and pressure fluctuation in impinging streams with liquid as the continuous phase mentioned above would be related closely to such strong interaction between the two opposing liquid streams impinging against each other; these features have great value for application. It can be considered that the extension of continuous phase in impinging streams from only gas to include liquid is progress of major significance which brings great application potential to impinging streams. However, this has unfortunately been ignored for a long time. Since there are significant differences of properties and performances between gasand liquid-continuous impinging streams, the two kinds of impinging streams will be discussed separately in this book. In addition, the method of impinging streams can also be used for systems of single phase, such as gas-gas and liquid-liquid impinging streams etc. In fact, single phase impinging streams have great value for practical application in mixing, gas combustion, etc.
6 APPLICATION STATUS OF IMPINGING STREAMS AND LOOKING AHEAD As mentioned above, the Koppers-Totzek gasifier of powdery coal [4, 5], the Stagnation jet mixer [14] and the Trost jet mill [20] are practical examples of the successful application of impinging streams. Apart from these, very few industrial applications of impinging streams had been by the end of the last century, [9, 21 ]. The following may account for the fact that the application of impinging streams has progressed so slowly:
(1)
Incorrect selection of application objectives decentralized time and efforts of investigations. Guided by the understanding "almost any process in chemical engineering can be carried out" [5], investigations extended over almost all the unit operations in chemical engineering, even including those controlled by internal diffusion so that, essentially, impinging streams cannot play any role, such as calcination of phosphate rock etc. That is, researchers did not focus on the cases where impinging streams were likely to be applied successfully. (2) The engineering problems that are normally encountered in practical application did not receive enough attention, and thus appropriate and feasible solutions were not found. As a result, few complete set technologies have been provided for industry.
INTRODUCTION
13
(3) Investigation on impinging streams with a liquid as the continuous phase started very early although, the perfect features and the application potential of liquidcontinuous impinging streams had been ignored for so long that both quantity and depth of investigations and development in this area were insufficient.
In fact, as described in the last section, the method of impinging streams has outstanding advantages and, simultaneously, intrinsic disadvantages. It can never be expected to become a universal tool. The proper selection of application objectives based on an understanding of the properties of impinging streams, the improvement of its advantages while avoiding the disadvantages, and focusing on solving related engineering problems may be the most important things to push n impinging streams towards industrial application. Fortunately, since the 1990s, technologies employing impinging streams have received increasing attention, and investigations into them have been growing faster than before. It is reasonable to believe, therefore, that more and more technologies applying impinging streams will emerge in various processing industries in the near future. The development of applied technologies of impinging streams has tended to increase significantly in the last 10 years. The following areas may be the most promising for impinging streams application to achieve success: (1)
Preparation of ultra fine powders by reaction-precipitation in impinging streams One of the most important conditions for preparation of ultrafine particles by reaction-precipitation is to create a very high and uniform supersaturation environment for precipitation. The tact that liquid-continuous impinging streams promote micromixing effectively favors such conditions, and so has received much attention in the last ten years and more. Instant reaction- precipitation processes can be carried out in an impinging stream reactor alone. Particularly noteworthy is that such types of reactor can be used for the production of nano materials. Mahajan et al. [22] and Liu et al. [16] studied the rapid precipitation of a number of medicines in two impinging stream reactors to prepare ultrafine products and obtained satisfactory results. By reaction- precipitation in a submerged circulative impinging stream reactor (SCISR), the author of this book obtained a Titania product average-sized 5.68 nm and copper powder sized 5.1 nm, both with very narrow size distribution. The details will be discussed in the relevant chapters of Part II. Essentially, few engineering problems involved in such technologies remain to be solved for successful application.
(2) Impinging stream ~'ombustion The strong micromixing in single phase impinging streams for gas fuel and highly enhanced heat and mass transfer for sprayed liquid fuel or fine powdery coal favor their combustions considerably. The Koppers-Totzek gasifier for powdery coal mentioned above is a typical example of employing impinging streams, and has been proved to be successful. The scheme of multi-frame inclined impingement has been used in some novel cooking stoves. Recent research and developments in the area of combustion have focused on improving the structures of burning
14
IMPINGING STREAMS chambers and the arrangement of burners in order to increase combustion efficiency further.
(3~ Impinging stream drying Drying of solid normal or fine particles is a type of typical process involving parallel heat and mass transfer, and thus is an area where the application of impinging streams could be most promising. In fact, since the 1970s a large number of investigations on this topic have been carried out, and many technologies and related devices have been proposed, as will be described in Chapter 6. However, no essential progress in industrial application has been seen [21] for over four decades. The main reason for this lies in the fact that some of the engineering problems involved had not been solved in the related developments. Most of the particular materials contain both free moisture and combined or in-pore water. The former can be removed instantly under enhanced transfer conditions, provided the particles are not too large; while the latter needs a considerably longer time to be removed because porous diffusion, especially diffusion of the liquor water, is involved. Since it has the intrinsic disadvantage of a very short residence time in the active region, impinging streams alone cannot accomplish the tasks of removing both free moisture and combined or in-pore water; while the design of a multistage impinging stream device would greatly complicate the system and increase its hydraulic resistance. Some researchers have used impinging streams for drying grains. In this case, in addition to the problems above, the energy consumed in accelerating grains would be very large. The author of this book recently developed a circulative impinging stream dryer [11] which, on one hand, utilizes the feature of impinging streams enhancing transfer and, on the other, can provide arbitrary residence time for the material being dried as needed by the arrangement of circulation. It is expected to be applied industrially in the near future. (4) Impinging stream milling As already stated, the most outstanding advantages of impinging stream milling are that it is without milling material and that the milling is carried out in gas streams at high velocity so that the substance being milled can be protected effectively from pollution and overheating. The Trost jet mill was applied successfully as early as the 1970s [20]. The development of applied technologies and devices of impinging stream milling have increased tendency in recent years [23, 24]. This is also an area of application that is currently well to the fore.
(5~ Impinging stream absorption Absorption is typical process involving transfer between phases and so is another area where impinging streams may be applied successfully. However, many systems to be processed by absorption are subject to equilibrium limitations, while to arrange a multistage countercurrent system employing impinging streams, such as in a column device, is very difficult. For such systems the impinging streams method is not a good option. On the other hand, for chemical absorption systems involving fast or instant irreversible reaction(s) in liquid phase, most possibly, an
INTRODUCTION
15
impinging stream device can be successfully applied. A notable objective of application is desulfurization of flue gas from coal burning. As is well known, this is a major problem involving the protection of the human environment. The absorption of sulfur dioxide with Ca(OH)e-water suspension involves irreversible fast reactions in liquid and is a typical case of gas-film diffusion control. Berman et al [25] studied such a process in an impinging stream absorber with three coaxial cylinders. Their results are certainly positive from the point of view of SO~_removal efficiency, but from the standpoint of engineering practice, both the equipment and the system scheme are considerably complex, giving rise to difficulties tk)r industrial application. Further research and development are still needed. Recently, the author of this book developed a novel impinging stream absorption device system with a simple structure and scheme, which has been used for wet desulfurization of flue gas, yielding good results. The details will be discussed in Chapter 7. In addition to the above, it is predicted that impinging streams can also be used for some other processes, such as solvent extraction, emulsification etc., to yield good performances. In the last two decades the application potential of impinging streams has been receiving more and more attention from scientists and engineers and the subsequent research and development has increased significantly, encompassing more and more countries and regions. It can be expected, therefore, that more and more applied technologies of impinging streams will continue to emerge in various processing industries in the near future.
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PART I GAS-CONTINUOUS IMPINGING STREAMS The idea of impinging streams (IS) was originally presented for enhancing transfer processes in gas-solid systems. In the 30 years from 1961 when the concept of IS was presented by Elperin to the mid-1990s, investigations on IS were mainly concentrated on systems with a gas as the continuous phase, while, to an extent, the dispersed phase was extended to include liquid. For processes in multiphase systems, whether the dispersed phase is solid or liquid, the common characteristics of gas-continuous impinging streams (GIS) are: low viscosity of the continuous phase, large density difference between the continuous and dispersed phases, and high operating impinging velocity. These features result in the following phenomena in GIS: strong turbulence in the impingement zone, very large relative velocity between phases, and penetration of particles or droplets in the dispersed phase to and fro between the opposing streams. The latter two can be considered as the special phenomena of GIS. Without question, GIS is one the most effective methods for enhancing transfer between phases to date. Part I of this book focuses on problems relating to gas-continuous impinging streams, including basic regulations, properties, and some of its applications. It is clear that the flow of continuous phase plays a very important role in impinging streams. Part I will start with single-phase impinging streams, because, to a great extent, the flow phenomena in such impinging streams can reflect the flow of continuous phase in multiphase impinging streams. Considering the similarities in movement of liquid and gas, the discussion in this Part will also involve single phase IS of liquid.
17
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-1FLOW OF CONTINUOUS PHASE
As mentioned betbre, the concept of impinging streams (IS) was originally suggested for enhancing heat and mass transfer between a solid and a gas, and the development of the application of impinging streams were long focused mostly on multiphase systems. However, the problems involved in multiphase impinging streams are considerably complex so that it is necessary to study the behaviors of the individual phases separately. In gas-continuous impinging streams, the concentration of dispersed phase, particles or droplets, is usually very low. Particles or droplets have no significant influence on the flow of the continuous phase except that their acceleration consumes part of the dynamic energy of the flow. For investigating the basic regularities of the flow, the behavior of gas-continuous impinging streams may be considered as a simple summation of that of single phase impinging streams plus the movement of dispersed phase without significant deviation. The flow of the continuous phase is the most essential phenomenon, and is the origin of various features of impinging streams. In addition, the method of impinging streams of single phase has a number of practical uses in gas-gas or liquid-liquid mixing, gas jet burning, etc'. Therefore investigation on the flow of continuous phase is of practical significance. The discussions in this chapter relate to single phase impinging streams, the behavior of which may reflect part of those of multiphase impinging streams. Taking into account the similarities of liquid with gas, the discussions will also involve impinging streams of liquid alone: while the differences between gas- and liquidcontinuous impinging streams will be described in detail in the chapters of Part II.
1.1 FLOW CHARACTERISTICS The impingement of two single-phase flows against each other is a very complex and interesting phenomenon and it can be considered that investigations on this phenomenon are not sufficiently up to date. Powell [26] presented a Mirror Image model in 1960, in which impingement between two opposing jets at a distanced of L is considered as equivalent to the impingement of one jet on a flat plane at a distanced of L/2. In other words, all the states in the impingement zone are symmetrical with respect to the impingement plane. Nosseir et al. [27] considered that the concept of mirror
19
20
IMPINGING STREAMS
image may be correct if the jets in impingement are laminar flows; but for turbulent flows, that concept would be questionable. The results of their study on the flow field in the impingement zone showed that pressure fluctuation occurs on the impingement plane, and then fluctuation is enhanced in a feedback mechanism. This pressure fluctuation is very significant and should never be neglected. The results obtained recently by the author of this book show that the pressure fluctuation in liquidcontinuous impinging streams has fundamental influence on micromixing in the impingement zone, and thus on process kinetics. The details will be discussed in Chapter 11. The major disadvantage of the mirror image model is that the interaction between the two opposing streams in impingement was not considered. In fact, even in laminar impinging streams, such an interaction cannot be negligible. Denshchikov et al. [28, 29] studied experimentally the interaction between the opposing flows. Two slit nozzles are mounted opposite each other in a water filled tank sized lx lx0.23 m. The nozzles are in the form of a plane box. The dimensions of the outlets of the nozzles ensure that the jets from them are in two-dimensional flows. Inkstained water streams are ejected from the two nozzles and impinge against each other at the center of the tank. The experimental observations are: the flow direction of a single jet is always stable; when another jet is introduced to impinge against it, both the jets begin to deflect in opposite directions with considerably large amplitude (up to 0.05-0.2 m), and then the directions of deflection periodically vary, as shown in Fig. 1.1, where T is the period of oscillation. Apart from planar deflection, there is also a twisting of the ends of the deflected jets in a vertical direction.
t--O
I=T/4
Figure 1.1 Image of impingement between two-dimensional jets of ink-stained water [29].
FLOW OF CONTINI !OUS PHASF
2i
When the jet on the right-hand side is at the top and that on the left-hand side is at the bottom, the twisting is clockwise; while for opposite positions of the jets, it is anticlockwise. The deflection and twisting of the jets result in the formation of vortices on both sides of the outflow plane, the intensities of which depend on the parameters of the jets in impingement. The researchers considered that the collision between the two jets is associated with retardatioE~ of the liquid, resulting in an increase in pressure in the impingement zone. The experimental measurements were correlated in terms of the period T of the auto oscillation as a flmction of the operation parameters. For two equal jets ejected from the nozzles at the same velocity, the relationship they obtained is
.
T =0.34Re-~(~] ~)45 .
.
.
(1.1)
where --/2
T ,~ = - - - " V
/2 0 d
Re - - -
V
L is the distance between the nozzles, u0 is the velocity of the jet just leaving the nozzles, i . e . , impinging velocity, 6 is the transverse dimension of the nozzles, and v is dynamic viscosity of the fluid. It is clear that such significant deflection as that shown in Fig. 1.1 must be related to the properties of the liquid employed. In comparison with liquid, both the density and viscosity of a gas are much smaller, so that such a strong deflection could not be observed with gas jets. However. in principle, it is possible that such deflection phenomena could occur in gaseous single-phase impinging streams, but the degree of deflection may differ greatly from that of liquor ones. Oren e t a l . [30] also observed deflection of the flows resulting from the interaction between two opposed jets in an investigation on the impingement of cylinder-formed liquid jets with small diameters (3-8 ram) submerged in liquid. The investigation was carried out in an electro-chemical reactor of 0.07 m in diameter and 0.012 m high, and the inlets of the two jets and the two outlets for flows from the reactor are uniformly distributed along the circle and meet at right angles (see Fig. 4.2 in Ref. [5]). According to the structure of the device, most possibly, the mentioned deflection is related to the arrangement of inlets and outlets of the reactor, i . e . , the symmetry of impinging streams is disturbed by the flows toward the outlets. A notable factor is that in the experiments made by both Denshchikov e t a l . [28, 29] and Oren e t a l . [30] the cross-section areas of the jets are considerably smaller. This may be an important reason for the deflection of the jets. Since 1995, the author of the present book has organized a number of investigations, theoretical and experimental, on the properties and application of the submerged circulative impinging stream reactor (SCISR) [9, 13, 15-18, 31]. The flow configuration inside the reactor is two impinging horizontal streams, as shown in Fig.
22
IMPINGING STREAMS
1.2. The propellers on the two sides transport liquid to pass through drawing tubes 60 mm in diameter and impinge against each other at the center. The rotary speed of the propellers ranges from 400 to 1300 rpm, and, correspondingly, the velocity of liquid flow at the outlet of the drawing tube, i.e., the impinging velocity, ranges from 0.18 to 0.6 m.s -~. In experiments with water as the process material, the following global flow phenomena are observed: (1) A strong surge occurs in the impingement zone at high impinging velocity, with the appearance of the flow being significantly different from that outside this zone. (2) The two streams in impingement do not reflect, provided the drawing tubes are mounted co-axially. (3) At very high rotary speed of the propellers, horizontal oscillation movement of the impingement plane along the axis of flow is observed, the maximum displacement of which can be as large as 30-40 mm. The horizontal oscillation obviously results from the pressure fluctuation round the impingement plane. (4) In the range of normal rotary speed of the propellers the horizontal position of the impingement plane is essentially stable, but its micro oscillation can still be observed visually.
Drawing tube
Impingement zone Propeller
Figure 1.2 A simplified view of the flow configuration in the submerged circulative impinging stream reactor. On gaseous impinging streams, Becker et al. [32] made a comparative study of two cases" two opposed free jets impinging against each other and a single jet impinging on a wall plane vertical to the jet axis. The researchers found that in the laminar region the impinging streams and the impinging jet have the same configuration, except for the friction of the wall surface, and the former is equivalent to the combination of the latter with its mirror image on the wall. However, in the turbulent region, a boundary layer forms on the surface of the fixed wall impinged by the jet, which stabilizes the flow and eliminates oscillation on the plane, while in the case of two opposing jets impinging against each other the interaction between the two streams is significant. In addition, Becker et al. mapped the concentration field in impinging streams by marking them with oil condensation smoke from a generator. The results they obtained are" the concentration decreases along the axis towards the impingement plane, and also decreases towards the radial direction. These tendencies are reasonable, because the two jets mix strongly with each other on the impingement plane and mixing is most intensive at the center.
FLOW OF CONTINUOUS PHASE
23
Popiel and Trass [33] studied the flows in impingement of two opposing jets against each other under conditions of low turbulence by the smoke-wire visualization technique. They provided an excellent photograph of the flow behavior of free impinging round jets issuing into ambient air at rest through bell-shaped convergent nozzles, as shown in Fig. 1.3.
Figure 1.3 Streamlines in circle impinging streams [33]. Possibly because the mixing problem involved was not fully considered, the interaction between the two opposing streams in impingement has not received enough attention in earlier investigations. A typical example is the mirror image model proposed by Powell [26], which takes the impingement between two opposing streams as two streams, independent of each other, impinging the same rigid plane from opposite directions. In the investigation carried out by Becket [32] the mixing between the two streams was considered although the target is gaseous impinging streams, for which the mixing problem is very simple and not as important as for liquid systems, because gases have a considerably larger molecular free path. As mentioned in the Introduction, the impingement between two opposing streams is quite different from a single jet impinging on a rigid wall surface. In the impingement of two opposing streams against each other, the interface (the impingement plane) is "soft". On the other hand, the continuous phases of impinging streams of gas and liquid have common properties: they have no fixed forms and certain distances exist between molecules and/or fluid elements. Consequentially, the relative movement between molecules and/or fluid elements can occur with certain magnitude. Of course, there would be a considerable difference between the permitted relative displacements in gas and liquid. In gas-solid impinging streams, solid particles can take the oscillation movement of penetration to and fro between the opposing streams; and even in the single phase impinging streams of gaseous He-He+SiF4, the molecules SiF4 have been found to penetrate into the opposing stream deeply. The penetration phenomenon may also occur theoretically in liquor impinging streams. However, since the densities of both the two streams are large, with very small difference between them, and the friction Ibrce is also very large, the depth of penetration into the opposed stream, if it happened, would be extremely small, so that it is very difficult to observe.
24
IMPINGING STREAMS
In addition, due to the increase in pressure caused by the strong momentum transfer between the streams on the impingement plane, the fluid elements in opposing streams may collide, shear and press each other, resulting in deformation of fluid elements, and finally the elements may break up to reducing segregation scales. This action may be very important and effective for promoting mixing, especially micromixing. Figure 1.4 shows a simplified model describing Element A deforming and breaking-up into Elements A1 and A2 by pressing and shearing after passing through the impingement plane, where the lengths of the arrows represent qualitatively the velocities of the corresponding elements. Since the interaction between elements always occurs mutually, every element could be subject to deformation by shearing and pressing, and/or to breaking up. It is difficult to obtain direct experimental evidence for the model shown in Fig. 1.4, but the model would be reasonable according to the general properties of liquid and the analysis for momentum transfer involved. On the other hand, in general theory it is considered that mixing is caused by flow, turbulence and eddy diffusion etc., yielding reduced segregation scale in the fluid. When the segregation scale is reduced down to the Kolmogoroff micro scale 2, molecular diffusion takes over to achieve a complete homogeneity, i.e. the ideal micromixing status. According to such a theory, the results obtained by the author of this book on the very strong micromixing in the submerged circulative impinging stream reactor (SCISR) [ 13] may be taken as indirect evidence for the model shown in Fig. 1.4. I
|
Impingement plane t = to
Impingement plane t = tl
Impingement plane t=
t2
Figure 1.4 Model for deformation and break-up of fluid element due to pressing and shearing round the impingement plane. Summarizing the results and analysis described above, the following can be generally concluded for the flow characteristics of co-axial single phase impinging streams, at least for those with larger diameter jets: (1) Impingement does not change the axial symmetry of the flows, i.e., each of the streams does not deflect, and the related parameters, such as streamlines, density etc., are kept in axially symmetry; (2) Pressure fluctuation occurs round the impingement plane; and (3) Due to strong momentum transfer, significant interaction between the two streams occurs round the impingement plane, and, in the case of liquid impinging streams, the interactions of collision, sheafing, and pressing etc. between the streams may finally result in reduction of the segregation scale. The characteristics described in Items (2) and (3) above are important effects in micromixing and will be discussed in detail in Part II.
25
FLOW OF CO\r~ftNUOUS PHAS!~
1.2 VELOCITY FIELD IN LAMINAR IMPINGING STREAMS Powell's mirror image model has the major disadvantage that the interaction between the opposing streams in impingement was not taken into account, although it is helpful for understanding some of the global parameters and their profiles in impinging streams, such as velocity field etc; while the model itself is relatively simple.
1.2.1 General equations Based on general principles of fluid dynamics and Powell's mirror image model, Tamir [5] analyzed the velocity field in laminar impinging streams. The case considered was two gaseous coaxial impinging streams, and the distance between the outlets of the two tubes was represented by L. In establishing the model, the following assumptions were made: (1) The fluid has zero viscosity and is incompressible, i.e., ¢tg = 0 and p~ = constant; (2) The flows in impinging streams are at a steady state; (3) The flows are without rotation; and (4) The influence of gravity is negligible. Obviously, the assumption of incompressible fluid greatly simplifies the problem; since pressure drop due to impingement between the opposing streams is usually not large, this assumption would result in no significant error. With the above assumption, the equations for fluid motion can be obtained as [34]
pV ] - u 2
-~ - p [ u x ( v x u ) - - v / ;
(1.2)
The equation for flow without rotation is
Vxu =0
(1.3)
V.u-O
(1.4)
and the continuity equation is
where - v/,
- -vP
+ pg
With the assumptions of fl~ = O, p~ = constant and neglecting the effect of gravity, the motion equation (1.2) is simplified to
OVl-u -~ - - V P /,
(1.5)
26
IMPINGING STREAMS
1.2.2 Planar two-dimensional impinging streams For planar two-dimensional impinging streams, the velocity vector, u, can be represented in the well-known way, as u - iu x
+
jUy
and the equations above can be specified to be as follows.
For motion:
~)u×
~)u~
=
OP
¢)Uy buy ~)p x-"~-X" k - Uay y ~ = - - ay ~
(1.6)
For irrotational flow..
~U x
~Uy
by
Ox
=0
(1.7)
For continuity: Ou
Ou " + Y -0 0x Oy
(1 8)
Using Eqs. (1.6), (1.7) and (1.8), ux, Uy and P can be determined as a function of the coordinates x and y; while it would be more convenient to represent them in terms of the stream function N(x, y) and the velocity potential O(x, y), which are defined by the relationships below.
u
=-~,
~)Y
a0
u -=-~, ~)Y
u
Y - - t - ~OX
Uy -
a0 ~)x
(1.9)
(1 10)
Substituting Eqs. (1.9) and (1.10) into Eqs. (1.7), (1.8) and (1.6) and their integral forms yields
FLOW OF CONTINUOUS PHASE
27
v-~-o
(1.11)
V2¢~-0
(1.12)
Oy
3x O(x, y)
0(v2¢) a(v2~u)
(1.13)
Oy 2
~)x2 and, from Eq. (1.5), we have 2
2
0.5 p(u ~ + u y ) + P - const.
(1.14)
The equations above can be represented in terms of the velocity potential to yield ~p- 0.5M ( - x 2 + y2)
(1.15)
where M is a constant determined from a known velocity at some distance from the origin. From the definitions for the velocity components, Eqs. (1.9) and (1.10), using Eq. (1.13) and considering that d ~r = u~clx- u:dy, one can obtain the expressions for the velocity components below u ~ -
-Mx,
u y -
My
(1.16)
and dv uy . . . . . .
y
dx
X
I t ,,
( ~ - Const..)
(1.17)
The pressure distribution in the flow field is determined by Eqs. (1.14) and (1.16) to be P - P(t - 0 . S p M 2 (x 2 + y2 )
(1.18)
The streamlines in two-dimensional impinging streams are shown in Fig. 1.5, the slope of which demonstrates the direction of flow and is equal to the local velocity vector meeting the relationship b e l o w
28
IMPINGING STREAMS 2 )0.5 _ I1--( bl2x WHy
M (x 2 + y2 )0.5
This relationship indicates that the velocity profile is flat for large x or y; while when x and y are of the same order of magnitude, the velocity is minimal at the centerline and increases with x or y increasing. This is indicated in Fig. 1.5. For the case of neglecting viscosity, these results are reasonable, because no shearing force is imposed on the jets. Integrating Eqs. (1.18) and (1.9) leads to the streamlines equation: xy - gz / M = const.
(1.19)
/ bly--my
xy= gzA~=const.
uy,
N~~
ux=Mx ..........................................
iiiiiiiiiiii!iii t-
x
P u=M(x:+y2)°5
I
Impingement plane Figure 1.5 Streamlines of two dimensional impinging streams.
1.2.3 Axial-symmetric impinging streams For axial-symmetric impinging streams, it is more convenient to use the cylindrical coordinates. The equations above then become the following:
FLOW OF CONTINUOUS PHASE
29
For motion:
I 0(~,E 2~u) + ~2~3E~-/~ r
3(r,:)
,
_
0
(1.20)
r 2 3:,
For irrotational flow with respect to 0 axes: Ou,. , 3z
c)u~') _ 0 Or )0
(1.21)
and tbr continuity: 1 3(ru,. ) + ~ U z _ 0 r Or 3z
(1.22)
where
E-p'-
32~ 3 Or-
"
32g 1 ON + ~ 9 r Or 3z"
(1.23)
3p' u,. - + ~ 3~
(1.24)
The stream function is defined by 3gt Or '
u, =
A specific solution of the equations above is given in terms of the velocity potential 0as 0 - 0.5M (-22. 2 + r 2 )
(1.25)
where M has the same meaning as that in Eq. (1.15); and 0 meets the definition and the restrictive conditions below
II r
u~
=
a~
1 a W"
Or
r 3:
ao ~z
u0 =0
-Mr
13~u
r ar
- -2Mz
(1.26)
IMPINGING STREAMS
30
1.2.4 General three-dimensional impinging streams For general three-dimensional impinging streams it is difficult to define a stream function satisfying the continuity equation. However, if the velocity vector is defined by u = -V¢
(1.27)
then it automatically satisfies irrotational flow defined by Eq. (1.3); and the continuity and momentum equations are given by Eqs. (1.12) and (1.5), respectively. A specific solution for the velocity field satisfying the modified equations above is given by the velocity potential function below:
¢-
(1.28)
0 . 5 M ( - 2 x 2 + z 2 q- r 2 )
and the differentiation of Eq. (1.28) yields the following velocity components: 30
li x
()X
a¢ UY=Oy
-2Mx
(1.29)
= My
3 0 = Mz
Uz = ~ Z
Using the expressions for the velocity components above and Eq. (1.5), one can obtain the expression for the pressure distribution as P = P o - 0 . 5 p M 2 ( 4x2 + y2 + z2)
(1.30)
and that for streamlines as U x "bly "U z -- d x "
dy'dz
- -2x"
y"
z
(1.31)
The results of integrating Eq. (2.30) are y - C l z,
where CI and Cz are integral constants.
xy 2
-- C 2 ,
xz 2 -
C2
(1.32)
FLOW OF CONTINUOUS PHASE
31
1.2.5 Viscous impinging streams In all the above derivations in this section, the influence of viscosity is neglected so that analytical solutions for velocity and pressure profiles can be obtained. When the viscosity of fluid is taken into account, it is difficult to obtain any analytical solution. Kuts and Dolgushev [35] solved numerically the flow field in the impingement of two axial round jets of a viscous impressible liquid ejected at the same velocity from conduits with the same diameter and located very close to each other. The mathematical formulation incorporated the complete Navier-Stokes equations transformed into stream and velocity functions in cylindrical coordinates r and z, with the assumption that the velocity profiles at the entrance and the exit of the conduit were parabolic. The continuity equation is given by Eq. (1.22); and the equations for motion in dimensionless form are:
,~)u~ , .~)u~. . . ~) P u,. Or" +u~ ~)z' ~)z.'
+ 1 1 ~) (r, +~)-u~• R-e-eL-r-7~-Tr'k, Or J ~)z'2
(1.33) , au,___ u,
Or" +
, au,__~= u~
,
a
1
- -'r--;
1 a ( r u,.
a 2u~' + az '2
+--~
where the dimensionless variables are defined as
u,
,
u,. u~
, uz
uz u~
r
,
r R
7
,
z --, R
Re
Ru o v
R is the radius of the conduits; u~ the velocity at the outlet of the conduit, i . e . , the impinging velocity. The velocity components in terms of stream function were given by Eq. (1.26); while the conditions of irrotational flow were determined by Eq. (1.21). The streamlines calculated by Kuts and Dolgushev [35] for some of the operating conditions are given in Fig. 1.6(a) and (b) as functions of dimensionless axial and radial coordinates, where the measure for ~- is from the impingement plane and that for r from the flow axis. A comparison between these results and the data to be introduced later shows that the flow field configuration theoretically calculated above fits the experimental data qualitatively well. The two streamlines figures indicate that the flow in impinging streams has a layered nature and the presence of vortices was not observed. As the Reynolds number, R e , increases, the streamlines are slightly deformed in the direction along the impingement plane. However, this slight difference and the absence of vortices, despite a significant difference in R e between the figures, indicate that the assumption of ignoring the influence of viscosity made before is reasonable indeed; this assumption greatly simplifies the calculation.
32
IMPINGING STREAMS 1.0
.e=,O 'J A
~
/
/
0.5
0.5
0.005 (a)
I
I
0.5 z/R
I
(b)
0.5 z/R
(d)
0.5 r/R
0.04 0.03
-..< (c)
0.5 JR
• 0.02 0.01 0
Figure 1.6 Streamlines (a, b) and pressure profiles in axial and radial directions (c, d).
Figures 1.6(c) and (d) show the pressure profile in the impinging streams. The profile is characterized by the considerable pressure gradients in the direction of the chamber axis, as shown in Fig. 1.6(c), and also in the perpendicular plane, as shown in Fig. 1.6(d). Obviously, in the region a little distance away from the impingement plane, the pressure profile is independent of the impinging distance. A decrease in the impinging distance leads to a more rapid increase in pressure only in the impingement zone. All the above analyses for several specific cases are based on the mirror image model. As mentioned earlier, the major disadvantage of the model is that the interaction between opposing streams in impingement, including momentum transfer and the consequent pressure increase and fluctuation, was not taken into account. In addition, the flows were assumed to be irrotational in the establishment of the models. Naturally, the streamlines calculated do not show vortices existing. Although these results cannot completely reflect the flow characteristics of impinging streams, the information on velocity and pressure profiles that the researchers provided is helpful for understanding the properties and some regularity of impinging streams.
1.3 EXPERIMENTAL RESULTS FOR THE FLOW FIELD IN IMPINGING STREAMS Elperin et al. [3, 36] investigated experimentally the hydrodynamics of co-axial gassolid suspension impinging streams. The dimensions of the device they used are"
FLOW OF CONTINUOUS PHASE
33
diameter of the gas conduit do = 0.05 m; distance between the outlets of the conduits, L, variable in the range of (0.5-8)d0. The air velocity in the conduit was measured with a Pitot tube connected to a micromanometer; while the values and the direction of the velocity and the static pressure in the impingement zone were measured by a threechannel probe attached to a traverse gear. Figure 1.7 demonstrates the distributions of axial and radial velocities, along with the isobars (dotted lines) for the case of L/d = 8. This figure can be related to Fig. 1.3 demonstrating streamlines in impinging streams for reading and analysis. It can be seen that the gas stream leaving the conduit behaves like a free jet flowing into infinite space with a characteristic velocity profile. As it approaches the impingement plane (x/do = 0), the axial velocity profile is deformed and a defined extreme point is observed along the x-axis. This behavior is the result of the hydrodynamic interaction between the opposing streams and the consequent appearance of the radial component of velocity. When the impinging distance becomes even smaller (L/d < 3), the axial velocity profile deforms immediately once the gas flow is ejected from the conduit. v/R
~
I V
~
I
~'"
-3
-2
"
",, ~ - -
-1
0
\...
1
....
2
/
3
Figure 1.7 Velocity and pressure profiles in impinging streams for L/d=8. It can be seen from the isobars (dashed lines) that the highest static pressure appears near the impingement plane. The isobars have complex appearances and assume an ellipsoidal form in the region far from the x-axis. These experimental findings and the symmetry of the flow pattern in impinging streams with respect to the impingement plane are suitable for applying Eq. (1.28) for the pressure distribution in a non-viscous impinging jet far from the x-axis, where the constant pressure surfaces are ellipsoids with the main axis ratio of 0.5:1:1. The maximum pressure is observed at the point x = y=z=O.
34
IMPINGING STREAMS
G)
I
2
~D
¢~
I
I
¢,g3
10
y
c.
P/(0.5pu 0 )
©
0
I
G~
o
-1.0
Ux/U° I
~
-4d
I -2d
I
I 0
I
I 2d
I 4d
Figure 1.8 Variations of dimensionless velocity and pressure along the axis (y/d = 0). Figure 1.8 shows the variation of the dimensionless axial velocity on the flow axis ((y/R - 0) along the radial direction. This variation can be approximately represented by the hyperbolic tangent function below: bl x
e x/d ~
e - x/d
x/d
-x/d
= u o
e
( 1 . 3 4)
+e
The absolute values of the dimensionless velocity vary between 1 and 0. The minus sign in the figure indicates that the velocities are in opposite directions. Figure 1.8 also shows the variation of pressure along the radial direction. The velocity profiles in the x direction shown in this figure are different from those based on the theoretical model in Fig. 1.5. This is because the experimental profiles in the jet are affected by the drag forces of the stagnant atmosphere. The variation of the maximum pressure versus the dimensionless impinging distance L/do at a constant impinging velocity, u0, is shown in Fig. 1.9. It can be seen that, in the range of L/do - 3-8, the maximum pressure at the center of the impingement plane increases moderately as the impinging distance decreases; while, as the impinging distance decreases further, the pressure increases sharply. This is a problem connected with the design of the impinging stream device. When the impinging stream technique is applied to processes such as combustion etc., it is important to determine the characteristics of the variation of the maximal radial velocity. Figure 1.10 gives the experimental results related to this topic. In coaxial two impinging streams, after impingement the two streams are mixed with each other and then turned to be an axial flow. At the point of an axial coordinate y, the axial velocity is affected by two factors: (1) After the flow direction turning, the fluid originally in the region of r < y must flow outward through the point of axial coordinate
FLOW OF CONTINUOUS PHASE
35
v. Therefore the amount of fluid flowing outwards passing through Point y must be increased as v increases. This is a positive factor for increasing radial velocity. (2) The passage area for the radial flow increases as y increases, yielding a negative influence on the radial velocity. As a result of the combined effect of the two factors in contradiction, there must be a m a x i m u m value on the curve describing the relationship of the radial velocity versus the dimensionless radial distance y / R ; the curves in Fig. 1.10 clearly demonstrate such a situation. In the range of L / d - 3-8, the following relationship gives a good approximation to the experimental data:
(1.35) u(}
R
I
I
I
I
I
I
I
I
4 tt~
2
I
i
()
i
I
2
i
4
I
I
i
6
8
L/d
F i g u r e 1.9 Maximum static pressure vs impinging distance at the center of impingement plane.
2.0 S/d
0.5 1.0 I I 3.0 (5 4.0 ...,. 6.0 • 9.0
•
1.6
x
1.2 0.8 0.4 I
0.0 0
1
2
I
3
4
5
6
v/R
F i g u r e 1.10 Variation of maximum radial velocity with radial distance.
36
IMPINGING STREAMS
where R is the radius of the gas conduit. According to Eq. (1.35) the following inference of practical sense can be withdrawn: When y/R - 6, ur/uo = 0.0355, the radial velocity becomes negligible in comparison with the velocity of the gas flow inside the conduit. Therefore the relationship below is suggested as a criterion for the decision of the diameters of gaseous impinging stream device in design:
Did > 6
(1.36)
1.4 TURBULENT IMPINGING STREAMS The theoretical method describing turbulent impinging streams was presented first by Champion and Libby [37]; although it may not be the best, there is nothing better to date. Champion and Libby analyzed both the planar two-dimensional impinging streams and the impingement of two co-axial-cylindrical jets in which the flow is axis symmetrical. Actually, the results they obtained are applicable for both the two cases, provided the two-dimensional coordinates in planar impinging streams are replaced by the cylindrical coordinates. The jets are assumed to be ejected at an initial velocity of u0 along the x-axis, and then expanded rapidly towards the y-direction. The corresponding velocity components are ux and Uy. It is also assumed that the distance between the outlets of the conduits, i.e. the impinging distance, is very small and is equivalent to the diameter of the conduit, d0. The turbulent kinetic energy of each jet is k0, and the mean viscous dissipation is e0. It is then possible to define the following two dimensionless parameters. The first one is the integral scale of turbulence, 10, which is the measure of the turbulent degree and is defined by comparing it to half the separation distance of the jets. This ratio is
1o L/2
k~5/eo L/2
The second parameter is the measure of the turbulence intensity and is defined as
ko/u g . It was found in laboratory experiments that turbulence intensities resulting from a grid or a baffle are such that k o/U2o is of the order of 0.01 and 10/(L~) is of the order of 0.1. The fact that the two parameters are very small forms the basis of an asymptotic analysis of the model. As the quantity k o/U2o approaches zero, the ratio (L / 2) / u 0
flow time
10 / k0°5
turbulence time
is of the order of 1, i.e., it is independent of the limit process. With the two parameters defined above and from a comparison of their orders, it can be considered that the
FLOW OF CONTINUOUS PHASF
37
flows associated with closely spaced jets consist of two regions: one is the outer region between the exit and the neighborhood of the plane containing the stagnation line or point, called the stagnation plane; and the other is a thin layer centered about that plane, commonly called the impingement zone, in which adjustments of various quantities take place on each side. The so-called stagnation plane is the plane containing stagnation points and/or stagnation lines. The two-dimensional flow equations employed by Champion and Libby [37] are just the well known Reynolds stress equations [38], t
...... +_8(u'~u[) Z ~k 8~ 6= ,. , Ok ~.r
+ .8(u'~u,) . . . Ov
. .1 .OP. p Oi
(i
x,y)
(1.37)
where pu[u"k are the Reynolds stresses. The analysis involves solving the partial differential equations that are treated by asymptotic methods to solve for the intensities of the radial and axial shear stresses and for the viscose dissipation. Of principal interest are the dimensionless axial and radial intensities, respectively, defined as G~(()-,,
"~-/k ~,
Gy(()-,~
"2/k o
(1.38)
where g"= x/d is the dimensionless axial distance, and ( = 0 corresponds to the impingement plane. The initial values at ~'= 1, i.e., the exit planes, are determined from experimental data. Kostiuk [39] obtained data suitable for comparison with the theoretical analysis relating to opposing circular jets. The experiments took air as the working fluid. The two impinging jets with an exit diameter of 0.035 m were spaced 0.07 m apart, and the mean velocity at the exit plane of each jet was 9 m.s -~. The turbulence is generated by perforated plates located 0.02 m upstream of the exit planes. The perforated plates have various geometries, but all with a blockage ratio of 50%. A comparison of the dimensionless mean velocity between the results experimentally measured and calculated is shown in Fig. l . l l ; while that of the dimensionless intensity in Fig. 1.12. Note that the calculation based on the theoretical model requires a value for an adjustable parameter. From the figures, it seems that the agreement between the calculated values and the experimentally measured data is excellent. In addition, Fig. 1.12 provides the information of significance below: the intensity of the axial component exceeds that of the radial component. Kostiuk et al. [40] measured experimentally the flow field of the vertical co-axial turbulent impinging streams with a two-component Laser Doppler velocity meter. The opposing gas streams were ejected from two burner nozzles, which were designed to produce a uniform axial velocity profile at their exits. The turbulence in the flow was generated by a perforated plate located at the end of the contraction section in each nozzle. The air velocity at the exit of the nozzle was varied from 4.1 to 11.4 m.s-~; and
38
IMPINGING STREAMS
the distance b e t w e e n the nozzles f r o m 0.02 to 0.103 m. The f o l l o w i n g m a j o r results were obtained:
1.0
0.8 0
E
~) :,/~,-~
0.6
u0, m's -~ d, mm
c'~ ot'~~ a ~'-~ a
V !> []
0.4
<3 O
0.2
I
0
0.2
0.4
8 10 9 6 6 9
55 70 70 70 70 70
h, mm 2 4
2 2 3 3
I
!
0.6
0.8
1.0
~=x/d Figure 1.11 Comparison between calculated non-dimensional axial velocity profile and experimental data.
2.0 h=2 mm
1.6 .p.
(3
1.2
©
~
0.8
(3
"~ 3~ ~ a,~
0.4
~
-~ . . . . ~
0
~_' ~ il ~ - - = .......
" ....
,,~
~
I
I
I
I
0.2
0.4
0.6
0.8
/
1.0
Figure 1.12 Comparison of calculated axial and radial vortex intensities with experimental data.
FLOW OF CONTINUOUS PHASE
39
Under the conditions of turbulence, the time-averaged velocity field is symmetric with respect to the free stagnation plane, provided the flow rates from the two nozzles are equal. The mean axial velocity profile has a similar shape to the curve of ux/uo vs x. The gradient of the time-averaged axial velocity takes the maximum at the stagnation plane, while it approaches zero near the nozzle. The mean velocity field is found to be self-similar for all the nozzle exit velocities, distances between nozzles and turbulence generators tested. This similarity allows the mean axial velocity traverses to be normalized so that all the measured data lie on the same curve. The shape of the curve is similar to that given in Fig. 1.8. The turbulence in the flow is uniform across the jets, and therefore any flame in the flow will experience the same upstream conditions.
This Page Intentionally Left Blank
-2PARTICLE BEHAVIOR As mentioned in the Introduction, the concept of IS was originally presented for gassolid systems. Of course, processing solid-in-gas suspensions is an important area where impinging streams may be applicable. To a considerable extent, the feature of impinging streams enhancing transfer between phases is related to the behavior of particles, especially to their motion. Since the densities of solids and liquids have the same orders of magnitude, the methods for analysis used and the results obtained in this chapter are also suitable, in principle, for liquid-in-gas suspensions. However, when the dispersed phase is a liquid, the problems of re-atomization and coalescence of sprayed liquid droplets possibly occurring during impingement of the two opposing streams must be considered; furthermore, if the liquid is very volatile, vaporization of the liquid and the consequential variation of droplet size should also be taken into account. This chapter starts with the motion of a single particle, which is the basis for understanding the behavior of particles in groups. In a gas-solid suspension being processed by impinging streams, the particles-to-gas volumetric ratio is usually very small, round 10-~-10 -4, so that it can be treated as a thin suspension, neglecting the interaction between particles. For these cases, the analytical results for a single particle can be approximately applied to particles in groups, which is helpful for understanding their motion. The discussions in this chapter are limited to the motion of a single particle in the two cases most essential and most practically applicable: horizontal and vertical co-axial impinging streams.
2.1 MOTION OF A SINGLE PARTICLE IN CO-AXIAL HORIZONTAL IMPINGING STREAMS
2.1.1 Qualitative description In addition to the greatly increased relative velocity between phases, the most important phenomenon in gas-solid suspension impinging streams is that, by inertia, particles can penetrate to and fro between the opposing streams, and thus behave as oscillation movement. This results in lengthened residence time and increased concentration of the particles in the active region, yielding a positive effect on heat and mass transfer between phases. Understanding this behavior by analyzing the motion
41
42
IMPINGING STREAMS
equations for a particle is of great interest. Tamir [5] has analyzed the motion of a single particle, including penetration-oscillation. The penetration-oscillation movement of a single particle in horizontal onedimensional impinging streams is shown in Fig. 2.1, where the coordinate x defines the positive direction of motion or flow.
Ug
'_,,............., , . ~
.....
t
Ug
.....•,-,. ..... -,,L...-,-,,Ji
X
i
b/g
Impingement plane Figure 2.1 Movement of single particle in co-axial horizontal impinging streams.
As can be seen from the figure, the particle enters the gas conduit, which is also called the accelerating tube, with an initial velocity of Up= 0 in a horizontal direction at Point 1, when the relative velocity Ur = u g - Up = Ug; then the particle is accelerated by the gas stream. If the gas conduit is long enough, it will achieve the velocity of the carrying gas, Ug, at Point 2, and then, within the path from Point 2 to Point 3 very close to the impingement plane, the particle moves at a zero relative velocity ur = 0; while if the conduit is not long enough, the particle will reach Point 3 with a velocity of Up,3 smaller than Ug. At Point 4, just after the impingement plane, the particle encounters the opposite gas stream with the velocity (-Ug), when the relative velocity achieves the maximum value of ur = -Ug - ug = -2Ug; or, in the case of the gas conduit being not long enough, u~ = - U g - Up,3, where the minus sign represents the opposite direction of flow. Naturally, such a high relative velocity at that time will strongly enhance transfer between phases. From Point 4 on, the particle decelerates until Point 5, where its velocity vanishes and u~ = -Ug. The distance 4-5 the particle traveled during its first deceleration is the maximum distance of penetration into the opposite stream and is denoted by Xma~.This parameter is of significance for the design of an impinging stream device, and will be discussed further later. After reaching Point 5, the particle is accelerated by the opposite gas stream and moves towards the impingement plane again, and then penetrates into the original stream. When the velocity of the particle in the negative direction vanishes, it is in turn accelerated by the original stream towards the impingement plane; and so on. Because of kinetic energy loss, when the particle
PARTICLE BEHAVIOR
43
reaches the impingement plane again, the absolute value of particle velocity cannot achieve the original one, and the depth of its penetration into the original stream cannot achieve .r,,~,~. So, the oscillation movement of the particle cannot be of equal magnitude, but is damped. After performing damped oscillations several times, the particle completely loses its kinetic energy and drops out of the streams. The following factors may change the oscillation movement of the particle: (1) The friction force of the fluid consumes the kinetic energy of the particle, causing its damped oscillation movement; (2) The action of gravity makes the particle have the tendency to drop out of the impingement zone; whether large or small this tendency depends on the size and density of the particle and the operation impinging velocity, u0; and (3) After impingement the two opposing streams are blended and turned to be a radial flow. The momentum transfer between the particle and the radial flow gives it a radial velocity component, and thus the particle tends to move outwards to leave the impinging stream field. Depending on its position, the particle may obtain a radial velocity at various angles in the range from 0 ° to 360 ° and, consequently, may leave the impingement zone in different directions, as shown in Fig. 2.2.
l
Impingement plane
2O
Figure 2.2 Particles leaving impingement zone in various directions. Obviously, the situations are very complex. To describe every possible condition is very difficult and troublesome; while, on the other hand, it is not necessary to do so from a practical point of view. In the following discussions the motion of the particle will be analyzed based on certain idealized assumptions for the most number of oscillation times possible and the longest residence time possible of the particle in the impingement zone in extreme cases.
2.1.2 Basic relationship for the particle motion As is well known, the regularity of the particle motion is a result of the combined action of various forces. In the stream(s) three forces act on the particle: (1) The field force, which is the gravitational force in irrotational flows; (2) The buoyancy force
44
IMPINGING STREAMS
which acts in the direction parallel but opposite to the direction of gravitational force. According to Archimedes' law, it is equal to the gravitational force acting on the particle minus that acting on the fluid occupying the same volume as the particle but in the direction opposite to gravity; and (3) The drag force that appears whenever there is a relative motion between the particle and the fluid and depends on the relative velocity. The general equation describing the motion of a single particle with the mass mp can be written as
mp--~t -mpg-
Pgg-O.5CDPgAp]up-UgI(Up -Ug)
(2.1)
where the relative or slip velocity between the phases is determined by
I
I /
__
)2
)2 +(b/
--
)2
(2.2)
Equation (2.1) can be simplified or reformed for various specific cases. In the operation of a coaxial horizontal impinging stream device, at first solid particles are usually accelerated by a gas stream inside the gas conduit, i.e. the accelerating tube. The gas velocity in the accelerating tube is generally very high, up to 10-20 m.s -~ or even higher. Depending on the length of the tube, a particle can be accelerated to a velocity of 0.5-0.7 times the gas flow velocity so that its kinetic energy can be considerably large when the particle just enters the impingement zone. Compared with this kinetic energy, the influence of gravity can be negligible. As the particle loses its kinetic energy in the oscillation movement, the influence of gravity becomes more and more significant. However, as mentioned above, what are considered here are the extreme situations. For simplification of the problem, the influence of gravity will not be taken into account for the whole of the process of the particle motion. In addition, since the difference between the densities of gas and solid is very large, the buoyancy force can certainly be negligible. Thus, with the subscript x denoting the motion direction being saved for the onedimensional motion of a single particle, Eq. (2.1) can be simplified to
dup = - 0 . 7 5 C D Pg [Up -Ugl(Up-U g )
~
dt
ppdp
(2.3)
In order to focus attention on the basic feature of the oscillation movement of the particle, the assumption Tamir [5] made that the gas velocity is uniformly distributed in the whole flow field, i.e., Ug = constant, is still employed here. The experimental results obtained by Enyakin [41] illustrated that this assumption can give relatively good approximate results. According to the above analysis for Fig. 2.1, the motion of the particle in the impinging streams is divided into various stages.
45
PARTICLE BEHAVIOR
2.1.3 Solutions of the motion equation for various stages 2.1.3.1 Motion in an accelerating tube One of the necessary conditions for carrying out a process with gas-solid impinging streams is to accelerate particles from zero to a velocity close to that of gas streams before the opposing streams impinge against each other. Therefore a piece of tube is normally set up in an impinging stream device for each stream for particle acceleration. Referring to Fig. 2.1, assume the particle enters the accelerating tube at a zero velocity. From Points 1 to 2, the particle is accelerated by its interaction with the gas stream and moves in the same direction as that of the gas stream. Thus we have
iup-. lThe velocity equation, Eq. (2.3), for the case under consideration becomes dup p~ - - = 0.75C D ~ (u~ - U p ) -
dt
ppdp
where the drag coefficient, C~, depends relationships are well known [20]"
on flow regime,
(2.4)
and the following
Stokes regime, Rep < 0.4" 24 CD = ~
Re P
(2.5)
Transient regime, 0.4 < Rep < 500" 18.5 C o =
~
Re o.5
(2.6)
P
Turbulent regime (Self-similar regime), 500 < Rep < 105: CD -0.44
(2.7)
where the particle's Reynolds number, Rep, is defined as
Rep =
dpu,.pg ~
~g
(2.8)
46
IMPINGING STREAMS
On one hand, solid materials to be processed with an impinging stream device have various sizes, while, on the other hand, the relative velocity between gas and particles varies from time to time in acceleration and deceleration stages of particle motion. Both factors make the value for Rep vary continuously with considerably large amplitude, which may be across various flow regimes. So, the variation of the drag coefficient, CD, in various flow regimes has to be taken into account in the solution of the motion equations for the particle in various stages. Using Eqs. (2.6), (2.7) and (2.8), the equations for particle motion in the various flow regimes become
Stokes regime" dup
dt
= 18
Transient regime: dup
dt
= 13.88
]/g
(2.9)
2 (Ug-Up) = fl
ppdp
0.5 0.5 pg ,t/g
1.5 (Ug - Up)
1.5
ppdp
=f2
(2.10)
Turbulent regime: dup = 0.33 pg
dt
ppdp
(Ug - Up )2 - f3
(2.11)
where f~, f2 and ~ represent the accelerations in the Stokes, transient and turbulent regimes, respectively. Since the gas flow velocity has been assumed to distribute uniformly, ug in all the equations above is a constant. The initial value for the stage considered here is Up =0,
t =0
Provided the velocity of the particle at the exit of the accelerating tube, Up0, is given, by separating variables, Eqs. (2.9), (2.10) and (2.11) can be integrated in the interval of [0, Up0] to obtain the moving time of the particle in the accelerating tube, t~c, as t a c - ~0'p°--L-dup, k - 1 , 2 , 3
fk
(2.12)
It should be noted that, during the movement of the particle with its velocity varying from zero to Up0, the relative velocity varies from zero to (ug- Up0), and it is possible that the flow regime is changed, e.g., from a turbulent regime to a transient and even to a Stokes regime. Therefore, a staged integration is needed for solving Eq. (2.12), i.e., according to the variation in Rep, a suitable expression forfk (k--l, 2, 3) in Eq. (2.12) should be selected for the integral calculation. This is not difficult to work on a computer.
PARTICLE BEHAVIOR
47
On the other hand, the velocity of the particle is defined as dr
(2.13)
up = dt With the definition, we come to dup
dul,
dx
dt
dx
dt
dH p
--H i~
dx
Equation (2.4) can then be rewritten to be dll i~
- 0.75C D
u p d_v
jog p pd p
(u,~ - u p ) 2
(2.14)
Substituting, respectively, the expressions for the drag coefficient, CD, yields the relationships for the corresponding flow regimes. For example, for the Stokes regime, substituting Eq. (2.6) into Eq. (2.14)results in
ul~
du I~
....
dr
18
¢l~
- ~
p pd j-,
(2.15)
(u,,-up)
For the transient regime, using Eq. (2.7) yields () 5
()
dut' - 13.88 p ~ / ~ -1.5 (u ~ - u p )15
up dx
ppdp
(2.16)
while for the turbulent regime, it is dH p
up dr
- 0.33
/O~
~ (u~ - u p ) -
.~
ppdp
(2.17)
Equations (2.15) to (2.17) describe variations in the velocity of the particle with the traveled distance, with the initial condition below . p - 0,
.v - 0
Separating variables and integrating them in the integral of [0, Up0], the distance that the particle has to travel in the accelerating tube, or, in other words, the effective length of the accelerating tube needed, L a~,, can be obtained as
L,,~ - i,'i~'' u~-~up, • ,/'k
k - 1' 2, 3
(2 18)
48
IMPINGING STREAMS
The effective length of the accelerating tube, i.e., the length from the point at which the particles enter to the exit of the gas conduit, is an important structure parameter for impinging stream equipment. Theoretically, particles can be accelerated by gas flow to a velocity close to that of the gas, Ug, provided the accelerating tube is long enough. However, a longer accelerating tube will result in increased hydraulic resistance of the system, and is not practically applicable. On the other hand, too short an accelerating tube would decrease the efficiency of impinging streams. Therefore it is necessary to select a reasonable velocity for particles at the outlet of the accelerating tube. Generally, 50-70% of the gas flow velocity can be chosen for the velocity of particles out of the gas conduit. In other words, the length of the accelerating tube has to ensure that particles reach a velocity Up0 equal to (0.5-0.7)×big.
2. 1.3.2 Motion in the identical velocity stage From Point 2 to Point 3, the particle can be considered to move at a constant velocity which is equal to that of the particle at the outlet of the gas conduit. The motion equation is written as
dup ~=0 dt
(2.19)
The distance the particle travels in this stage is half of the impinging distance, S/2. Consequently, the time of particle motion in this stage, to is
S t c = 2bip0
(2.20)
2.1.3.3 Motion in the first deceleration stage From Point 4 to Point 5 shown in Fig. 2.1, the particle is decelerated due to the resistance of the opposed gas flow, and at Point 5 its velocity vanishes. In this stage the particle moves in the direction opposite to that of gas flow. Therefore we have
lbip--Ugi =U p +big,
Up--big =--(Up +big)
Equation (2.3) for the variation of the particle velocity becomes
dup
dt
=
_0.75C D Pg, '
ppdp
(Up Jr- Ug )2
(2.21)
Similarly, the drag coefficient CD depends on the flow regime. Using the relationships of Eqs. (2.6), (2.7) and (2.8), the motion equations for various flow regimes in this stage, can be written as
PARTICLE BEHAVIOR
49
Stokes regime" dup
#~
~=-18
dt
PPdl~ (up + ug ) = f4
(2.22)
Transient regime"
dup
dt
=-13.88
pO.5#o.5 1.5 (Up +u~) 15 - f5
ppdp
(2.23)
Turbulent regime:
dup = -0.33 p~ (Up + u,,)2 dt ppdp
_ f(~
(2.24)
In the equations above f4, f5 and f, represent the accelerations in Stokes, transient and turbulent regimes, respectively; and, similarly, the gas velocity ug in these equations is a constant. The initial value for this stage is
ttp =
Hp,4,
t -- 0
where Up.4 is the velocity of the particle just arriving at Point 4, and can be taken place by the velocity at Point 2 for good approximation, for which the method of calculation has been described above. Separating the variables involved and integrating Eqs. (2.22) to (2.24) in the interval of [Up~), 0], the motion time of the particle in the first deceleration stage, td, can be obtained as t~ _ l d u t d - f llpl, .fk p'
k - 4, 5, 6
(2.25)
Similar to the motion in the accelerating tube, during the movement of the particle in the first decelerating stage its velocity drops down from Up()to zero, correspondingly, flow regime turning may occur, and so staged integration is also needed for solving Eq. (2.25), while the calculation procedure is the same as for the motion equation of the particle in the accelerating tube. Using the definition for the particle velocity, Eq. (2.13), one can obtain dup
Up dx
=
_0.75C D Pe"-
ppdp
(Up + U,,~ )2
(2.26)
Substituting the expressions for the drag coefficient CD for various flow regimes into Eq. (2.26) yields the relationships for corresponding regimes as follows:
50
IMPINGING STREAMS
Stokes regime: db/p Up dx
~ - - 1 -8
/.tg 2 (Up + Ug) - f4
ppdp
(2.27)
Transient regime: O5 O5 = - 13.88 p g /tg 1.5 (Up +Ug) 1.5 - i s Up dx dup
Dpdp
(2.28)
Turbulent regime: dup
pg
2
Up dx = -0.33 ppdp (Up + Ug) - f6
(2.29)
Equations (2.27) to (2.29) describe the variations of the particle velocity with the distance the particle traveled in various flow regimes, with the following initial condition Up = b/p,4 = b/pO,
X= 0
Separating the variables involved and integrating the equations between Up0 and 0, the distance the particle traveled in the first decelerating stage, i.e., the maximum depth of penetration into the opposed stream, xm,x, can be obtained as Xmax -
io Up ----~Up, k = 4 , 5 , 6 rip() fk
(2.30)
where the functions fk have the same meanings as the corresponding functions in Eq. (2.25). It can be seen from the configuration of impinging streams that the value for Xm,x is of essential guiding sense for the design of co-axial horizontal two impinging stream equipment, mainly for the decision of the impinging distance, S. As mentioned above, during the movement of the particle in this stage, flow regime turning would occur, and so a staged integration is also needed for solving Eq. (2.30).
2.1.3.4 Motion in the first acceleration stage On the arrival of the particle at Point 5, its velocity drops to zero, and then the particle is accelerated by the opposite gas stream and begins its first acceleration motion towards the impingement plane again. In this stage the velocity of the particle is in the same direction as that of the gas flow. Since the gas flow velocity has been assumed to distribute uniformly in the field, the motion equations of the particle in this stage and their solution procedure are the same as those in the stage of particle motion in the accelerating tube. Obviously, this stage will end when the particle arrives at the
PARTICLE BEHAVIOR
51
impingement plane, and then the particle penetrates back into the gas stream it originally existed in. Since its velocity is in the opposite direction to that of the original stream, the particle is impeded and then starts its second deceleration movement. Therefore the distance the particle traveled in the first acceleration stage is equal numerically to the maximum distance of the particle penetration into the opposite gas stream, Xm~x,as described in Section 2.1.3.3. The value for the particle velocity at the end of this stage, up.~,~.~,can be determined by integrating Eqs. (2.15) to (2.17) with respect to x in the interval of [0, .r,,,,~]; and then the motion time of the particle in this stage, t~,c~,can be determined further by integrating Eqs. (2.9) to (2.11) from the value for the particle velocity at Point 5 to up = up,a~. Similarly, during the movement of the particle in this stage, the flow regime may be turned, and so a staged integration is also needed for solving the equations mentioned, with suitably selected equation(s) according to the variation in Rep; the procedures are the same as described above. As mentioned above, in impinging streams, it is possible for the particle to penetrate to and fro many times between the two opposing streams. In other words, it would experience multiple decelerations-accelerations of motion. The time for each deceleration or acceleration motion, td.i or t~,~.~ (i = 1, 2, ...) can be analyzed and calculated with the same principles and procedures as described above. Of course, the distance the particle travels in each of the deceleration stages, Xd.~, can also be calculated; however, those distances lack of significance, except for that in the first deceleration stage, xd.~ = x,,~,. Because the particle must lose part of its kinetic energy during each penetration to and fro between the two opposing streams, its oscillation motion in the impingement zone is with damped amplitude. For example, at the beginning of the first deceleration stage the particle has a velocity of up0; while the velocity of the particle at the beginning of the second deceleration stage must be smaller than up0 because the time or the moving distance tot the particle being accelerated in this stage is shorter than in the last stage, the first acceleration stage; consequently, the depth of particle penetration into the opposing stream must be smaller than x,,,,~; and so forth. According to the above assumption, the particle can penetrate to and fro between the opposing streams, theoretically, for an infinite number of times; but this is, of course, impossible. Once the distance of particle penetration becomes very small, the possibility of the particle penetrating once more into the opposing stream is negligible. Then, the particle gets a radial velocity to leave the impingement zone due to the momentum transfer between the particle and the radial gas flow.
2.1.4 Residence time of the particle in the impingement zone From the point of view of practical application, it is certainly of interest to understand, even primarily, the possible residence time of the particle in the impingement zone by a theoretical analysis. In principle, the total residence time of the particle, t~, can be calculated as the summation of the residence times in all the motion stages:
52
IMPINGING STREAMS tf - Y',td,i + Z tac,i (i -- 1, 2, 3,.-., n)
(2.31)
However, it is somewhat difficult to determine the number of oscillation times, and consequently the total residence time, accurately, because some of the idealized assumptions made above, such as neglecting the effect of gravity and the uniformly distributed gas velocity etc., are not really true. It is not difficult to understand that, in the initial period of the particle motion, when the velocity of the particle is near that of the gas flow, the influence of gravity is small and thus can be neglected; while in the later periods the absolute value of the particle velocity becomes smaller and smaller, and so, in comparison with the momentum of the particle, the effect of gravity becomes non negligible. Furthermore, with regard to the assumption of the uniformly distributed gas velocity, both the interactions between the gas flow and the particles and between the gas flow and the surrounding gas were not completely taken into account. Almost certainly, these excessively idealized assumptions will yield significant deviation from reality. Despite possible error(s), it is still helpful and of significance to get some rough or semi-quantitative understanding about the residence time of particles in the impingement zone from the analysis using the equations given above. The results estimated with Eq. (2.31) can be considered as being about the possible maximum residence time of particles.
2.2 EXPERIMENTAL RESULTS ON THE BEHAVIOR OF A SINGLE PARTICLE IN CO-AXIAL HORIZONTAL TWO-IMPINGING STREAMS It is obviously important to have certain experimental evidences for the corresponding results from theoretical analysis. Some experimental data obtained by several researchers and a comparison between the data and their theoretical results will be introduced below. These analyses mostly have certain disadvantages. The common disadvantage is that the relationships among various variables were analytically calculated separately for various flow regimes; while the turnings of flow regime during the motion of one particle in various stages, e.g., acceleration, deceleration etc., in an impinging stream device were not considered in a combination. Nevertheless, these comparisons are still helpful for understanding the regularities of particle motion. Kuts et al. [35] studied the relationship between the maximum depth of penetration into the opposing stream, Xmax, and the diameter of the particle, dp, in the ranges of d v = 100-1000 gm and Xmax- 0.01-0.11 m; the results are illustrated in Fig. 2.3. The results of the theoretical analysis are: in the Stokes and the transient regimes Xm~ is positively in proportion to dp2 and dp~.5, respectively. It can be seen from Fig. 2.3 that the results theoretically predicted fit the experimental data relatively well.
PARTICLE BEHAVIOR
5?
I
I
0.08
0.04 -
J
I
I
400
800 dp, gm
Figure 2.3 Maximum depth of penetration versus particle diameter: comparison between theoretical (curve) and experimental results. The expressions for x,,,~,xEnyakin [41] derived are similar to those by Tamir [5]" but the constants involved are somewhat different from each other. For the Stokes, the transient and the turbulent regimes, respectively, the expressions Enyakin obtained are as follows. For the Stokes regime" x'nux = 0.0166 pp dt~ P~
ReP
~"~'~ = 0.0245 pp dp p~
Rep0.5
For the transient regime"
For the turbulent regime: r"'~'x = 0.0744 pp dp pg In the investigation by Enyakin the maximum distance of penetration was measured by photographing the trajectories of various kinds of particles in impinging streams; and the results obtained are shown in Fig. 2.4 as the plots of the relative penetration depth x,,Jdp versus Rep. The trajectories photographed proved that the particle does experience damped oscillations.
54
IMPINGING STREAMS
800
400
_
I
0
I
u
o
I
0 0
oyo
o
0
00 -
1200
1600
- y O 0
100 0
I
400
I
I
800
Rep Figure 2.4 Relative penetration depth of particle v s . calculated.
Rep [41]. o measured; - - theoretically
In earlier extensive investigations [3, 42, 43], the local velocities, the average velocity and the residence time of a single particle were measured with isotope Co 6° tracing technique. Since there are turbulent pulsations in the impingement zone, the motion of a single particle is highly stochastic. In order to obtain representative data, the mean velocity was determined by averaging the results of 15 to 25 similar experiments. The motion of Ti, A1 and polystyrene particles in the accelerating tube were studied, and the variations of particle velocities with the distances the particle traveled were determined, with the relative errors < 15%. Figure 2.5 shows part of the results for the relationship of the particle to gas velocity ratio, Up/tUg, versus the accelerating distance, where the theoretically calculated values are not given for simplicity and clarity. From Fig. 2.5 it is clear that if the length of the accelerating tube, lac, is over 1 to 1.2 m, the velocity ratio becomes almost independent of the accelerating length, i.e., the acceleration of particles is terminated within the length of 1 to 1.2 m. The data shown in Fig. 2.5 provide guidance for the design of an impinging stream device, especially for the decision of the length of accelerating tubes. Figure 2.6 shows the relationship between the residence time of a single particle in the impingement zone, tf, and the length of the accelerating tube, lac. As can be seen, in the range of l~ smaller than 1.0, tr increases quickly as l~ increases; while after l~ = 1.0, the increase in tf with l~c is smoothed. The important conclusions that can be drawn from the discussions in this section are: (1) To carry out an impinging stream process it is necessary to accelerate particles by gas flow, i.e., accelerating tubes are needed in a gas-solid impinging stream device; and (2) When the accelerating tube is over 1 to 1.2 m, a further increase in length of the accelerating tubes becomes meaningless for both accelerating particle and lengthening residence time. In the investigations related to gas-solid impinging streams carried out by the author of this book, the dimensions of the experimental devices were determined on the basis of the results introduced above.
PARTICLE BEHAVIOR
55
I
I
I
0.4
0.2
n
I
0
I
0.4
I
0.8
1.2
[ac ~ m
Figure 2.5 Results measured tbr the relationship of velocity ratio versus length of accelerating tube. Material
pp, kg.m --~
d~×l 0 ~, m
u,~, m.s -~
•
Ti-particle
7160
2.25
21.7
•
Al-particle
2790
2.25
21.7
polystyrene
7160
2.25
18.3
0.6
I
I
I
1
0.4
m
I
2 o----
Hg, m/s
0.2
0
1--18.3 2--24.6 I
I
I
0.5
1.0
1.5
lilc ~ m
Figure 2.6 Residence time of particle in the impingement zone vs. the length of accelerating tube.
56
IMPINGING STREAMS
2.3 BEHAVIOR OF A SINGLE PARTICLE IN CO-AXIAL VERTICAL IMPINGING STREAMS 2.3.1 Description of motion phenomena In vertical impinging streams the two streams are flowing vertically. In this case the effect of gravity cannot be neglected. The trajectories of a single particle in the impinging streams are shown in Fig. 2.7. The particle enters the system at Point 1 with a zero velocity. Then it is accelerated by the gas flow and its velocity reaches Up,~mat Point 2. Because of the effect of gravity, its velocity is constantly smaller than that of the gas flow, ug. In this stage the friction force between the particle and the gas flow acts upwards to accelerate the particle; while the gravity partially offsets the accelerating effect. At Point 2 the particle penetrates into the opposing stream and then moves upwards in deceleration from Point 2 to Point 3; while at Point 3 the velocity of the particle becomes zero. In this deceleration stage both the friction force and the gravity act as deceleration effects. Also, at this point the particle turns its direction and, under the double acceleration effects of the friction force and the gravity, moves downwards and then arrives at the impingement plane once again at Point 4. Then the particle decelerates again along the trajectory 4-5, where both drag and gravitational forces are acting in opposition. At Point 5 the velocity of the particle vanishes and it turns its moving direction again. The particle then accelerates along the trajectory 5-6, and the situation is similar to that along path 1-2. At Point 6 the particle is withdrawn from the system.
Y
Figure 2.7 Movement of a single particle in vertical impinging streams.
PARTICLE BEHAVIOR
57
2.3.2 Motion equation and its solution For vertical impinging streams and under the assumptions of neglecting b u o y a n c y force and the particle being spherical in form, the equation for the particle motion, Eq. (3.1), can be rewritten in the generalized form, applicable for various cases, as
du
dt
P = ak i (u,; + bup)" + cg
(2.32)
where the parameters a, b, and c depend on the position of the particle on its trajectory, while n depends on flow regime. The values for these parameters are listed in Table 2.1.
Table 2.1 Values for the parameters in Eq. (2.32) Trajectory in Fig. 2.7
a
b
c
Flow regime
n
1-- 2
1
- 1
- 1
Stokes
1
2-- 3
- 1
1
- 1
transient
1.5
3 -- 4
1
- 1
1
turbulent
2
4--5
-1
1
1
5--6
1
-1
-1
Due to various relationships of the drag coefficient, Parameter k~ in Eq. (2.32) has different definitions for various flow regimes, as follows.
Stokes regime: kj - 1 8
~
(2.33)
dt~Pp
Transient regime k-, - 13.88 -
D0.5. o.5 ~ /J~ 1.5 pj~dp
(2.34)
Turbulent regime kI - 0 . 3 3
p~ p~dp
(2.35)
58
IMPINGING STREAMS
No matter what the flow regime, Eq. (2.32) cannot be solved analytically. Tamir [5] solved it with a numerical method, and the major results obtained are: (1) At Points 2 and 4, just across the impingement plane, the relative velocity between the particle and the gas flow, Ur, varies very significantly. (2) The influences of gas flow velocity on the velocity of the particle in the accelerating stages from Points 1 to 2 and 5 to 6 are quite different from that in the stage from Points 3 to 4. In the latter stage the gravity and the drag force act in the same direction so that the acceleration is very much larger. (3) This difference decreases as the gas velocity increases, because the influence of the gravity becomes smaller in comparison with that of the drag force at higher gas velocity. (4) The results of an analysis of the influence of the particle diameter on the relative velocity between the particle and the gas flow showed that with small diameter of the particle, its velocity quickly approaches the terminal velocity, ut. The smaller the particle diameter, the faster the particle velocity approaches the terminal velocity. (5) The influence of the particle density is similar to that of the diameter, i.e., the smaller the particle density, the faster the particle velocity approaches the terminal velocity.
2.3.3 Terminal velocity As is well known, when the gravity and the drag force acting on the particle are numerically the same but in opposite directions, the relative velocity of the particle with respect to the gas flow will be kept constant; such a relative velocity is called "terminal velocity", denoted as u t - ( U g - Up)t, and is numerically equal to the sedimentation velocity of the particle in a stationary gas. Setting dup/dt = 0 and using Eqs. (2.33) to (2.35) and the data listed in Table 2.1, the terminal velocities for various flow regimes can be directly obtained from Eq. (2.32) as follows:
Stokes regime"
gdppp
(2.36)
b/t - - ~ 18¢/g
Transient regime: _ b/t
5
8pp
5 3 dppg
133
I/ / 1° g 11 lpg
/,tg3
(2.37)
Turbulent regime:
u t - 1.74
I
gdp pp
(2.38)
PARTICLE BEHAVIOR
59
The main significance of the terminal velocity lies in the fact that it defines the minimum gas velocity requested. The particle enters the system at Point 1 at zero velocity. If a gas velocity is employed so that u,. < u~, the particle cannot get into but drops out of the system; if the gas velocity makes the relative velocity equal the terminal velocity, i.e., u , - ut, the particle will move with a constant velocity of zero with respect to the ground, and so will remain at Point 1; while if the gas velocity makes the relative velocity greater than the terminal velocity, i.e., u,. > u~, the particle will be accelerated by the gas flow to the terminal velocity and will then move towards Point 2 at that velocity. Therefore, the terminal velocity is necessary for the determination of the operational range of the gas flow velocity for a specific vertical gas-solid impinging stream system. As an example, let us now examine the specific case of particle motion at 25°C and atmosphere pressure. The related physical properties are: dp= 0.001 m, pp = 1000 kg.m -3, p ~ - 1.145 kg-m -3, and ¢t~- 1.798 Pa-s. The calculated values for the terminal velocity and the operational condition ranges are given in Table 2.2. An important conclusion that can be drawn from the data listed in the fourth column of Table 2.2 is that the Stokes regime cannot exist in co-axial horizontal impinging streams; while other regimes are applicable in this kind of impinging stream. Table 2.2 Operational range for vertical impinging streams Regime
u~, m/s
Conditions for u,.
Special conditions for u,
Stokes
30.3
u~
30.3
transient
4.2
u~
4.2
turbulent
5.1
ut
15.7< ul<3144
2.4 BEHAVIOR OF PARTICLE CROWDS IN IMPINGING STREAMS In the previous sections the motion of a single particle was discussed; and the results obtained may provide certain fundamentals useful for understanding the behavior of particles in impinging streams. In the target systems to be processed by an impinging stream device, a huge amount of particles is usually involved. From the point of view of the concentration of particle(s) in a gas stream, the practical suspensions containing either a single particle or a large amount of particles belong to thin dilution. In many cases, the interaction between particles can be negligible for the latter kind of suspension; however, in certain cases such interactions cannot be neglected. In impinging streams particle crowds exhibit two important contradictory behaviors as follows:
60
IMPINGING STREAMS
(1) The penetration of particles to and fro between the opposing streams causes increased concentration and, consequently, lengthened residence time of particles in the impingement zone. This is the characterizing feature of gas-solid impinging streams. Since the impingement zone is a highly active region for heat and mass transfer, both factors favor the processes to be can'ied out with impinging streams. (2) Collision between particles is unavoidable in the impingement zone. As a result, part of the particles may leave the impingement zone earlier, partially offsetting the favorable effects of the factors mentioned above. In some practical applications, of course, the collisions have positive e f f e c t s - in fact, collision between particles is the major mechanism of grinding and milling by impinging streams. However, in other cases the collision may be harmful, e.g., it may lead to unnecessary grinding of particles or, in the case where liquid is taken as the dispersed phase, to coalescence of droplets, yielding a reduced interface area, etc'. Beyond question, residence time distribution is one of the most important behaviors of particle crowds in impinging streams. This will be discussed in detail in the next chapter as it involves many complex problems.
2.4.1 Distribution of particle concentration in impinging streams Elperin et al. [3, 44] studied the influences of the concentration of particles in the feed suspension on the radial and axial profiles of their velocity and concentration in the impingement zone. The velocity of a single particle was measured using a radiationlabeled particle; while the local concentration of particles was determined by measuring the attenuation in radiation intensity due to penetration through the particles, where the ~-active isotope T12°4 was used as the radiation source. Figure 2.8 shows the experimental results on the concentration profile of sand particles, as a plot of the relative concentration versus the dimensionless radial distance, where the relative concentration of particle, ,6, is presented in terms of the particles to gas volumetric ratio; while 130 is the inlet relative concentration of particles, and x is the distance from the impingement plane. It can be seen in Fig. 2.8 that the maximum concentration is on the axis of the streams, and the radial concentration gradient increases rapidly as the distance from the impingement plane increases. It was mentioned in Chapter 1 that pressure fluctuation occurs round the impingement plane, which causes turbulent pulsations that tend to even out the concentration distribution. This effect diminishes towards longer distance. In addition, the accumulation of particles on the axis may also result from the Magnus effect related to their rotation. At lower gas flow velocity, the pressure fluctuation becomes weaker and, consequently, the tendency to even out the concentration distribution also becomes weaker. Figure 2.9 shows part of the results Elperin et al. obtained for the axial relative concentrations profile of silica gel particles. The major feature of this distribution is that the concentration sharply increases towards the impingement plane and the value near the impingement plane can be as high as 20-28 times the concentration in the feeding stream. This is mainly accounted for by the penetration of particles to and fro
PARTICLE BEHAVIOR
61
between the two opposing streams, i . e . , the oscillation movement. As already mentioned, this is the most important characteristic of gas-solid impinging streams. It is obvious that the more the oscillations, the higher is the concentration of particles round the impingement plane; and, consequently, the longer the residence time of particles in the impingement zone. The tbllowing can also be seen from the results: (1) The concentration of particles round the impingement plane decreases as the inlet concentration, ~ , increases. This may be because higher inlet concentration favors an increased probability of collision between particles, resulting in part of the particles leaving the impingement zone earlier; and (2) Under the same inlet concentration condition, the gas flow velocity does not exhibit a significant effect on the concentration of particles around the impingement plane, although the data on such an influence are not given in Fig. 2.9. Comparison between Curves 4 and 5 in Fig. 2.9 shows that when the inlet concentration fl0 reaches about 0.9x10 -~ m~.m-~, the concentration of particles round the impingement plane does not significantly drop as ~ increases further. The reason for this might be that the decrease in concentration round the impingement plane due to part of the particles dropping out of the impingement zone is smoothed by the increase in the total number of particles entering the system at increased inlet concentration.
~
1.0
I
I
!
I
xld
_ -'-'---,-..
";'-..
0.6
u.= 12.8
~. 0.4
I
-
-,.°2
0.8
1.0
I
m/s
I
''Q 12:72;
-
N
\ I
' 11.88
I
I
¸
. ~
0.8 0.6
-
0.4 0
x/d
ua, m]s
X
- 11.88
"
_
\9.9 20.2
-
I
I
I
I
I
2
4
6
8
10
12
rxlO ~ m Figure 2.8
Radial-relative concentration profile of sand particles. dp=0.67x10-~ m; pp=2650 kg-m ~
62
IMPINGING STREAMS
I
I
I
I
I
28 24 floxlO ~
20 ,\/I
Uo, m/s
0.37
I\\/1[ / 1 ] ///IV 16[-~11[1 2 . , , , , ,/
2 3 4
[/Ig
o.2
0.48 0.75 0.88
15.6 9.9 20.2
-
12I~3
i
4 0
I
I
2
4
I
6
I
I
8
l0
12
x/d Figure 2.9
Axial-relative concentration profile of silica gel particles. dp= 1.23x10 -3 m pp = 880 kg.m-3
An interesting phenomenon of practical significance observed in the same investigation is that in the two cases of the particles being introduced into (a) both the accelerating tubes, termed two-sided feed (TSF), and (b) only one accelerating tube, termed one-sided feed (OSF), the resulting concentration profiles have obviously different shapes under the same other conditions. In the case of one-sided feed the other stream is pure gas. In practical application, the scheme of one-sided feed has the great advantage of convenient operation. The results Elperin et al. obtained are: one-sided feed yields much lower concentrations of particles than those coming from the twosided feed and is non-symmetrical (refer to Fig. 2.10). Tamir [5] considered that, in the one-side feed, inter-particle collisions with those from the opposing stream are absent, and so the particle collision rate is higher for the two-side feed than for the one-side feed, resulting in decreased residence time and decreased penetration distance into the opposing stream and thus a lower concentration of particles round the impingement plane. In fact, under the condition of equal inlet concentration the one-side feed implies directly that the total number of particles is reduced by half; naturally, the concentration of particles round the impingement plane will be lower. This also indicates an insignificant decrease in the residence time due to collision between particles in the case of the one-side feed. With regard to the non-symmetry of the concentration profile, most possibly it is caused by the imbalance of momentums of the two opposing streams in the case of the one-side feed: the total momentum of the pure gas stream must be smaller than that of the particles-in-gas suspension stream under the condition of equal gas velocity.
PARTICLE BEHAVIOR
28
63
I
I
I
I
I
I
I
I
I
I
24 20
/~=xl0 -~" u,~=27.1 m.s -I
-
O two-sided feed • one-sided feed 12 8 4 0
-8
.
-6
!
-4
I
-2
I
0
2
4
6
8
12
14
x/d
Figure 2.10 Axial-relative concentration profile of poppy seeds. dp=0.99× 10--~ m; pp= l 100 kg.m -~
2.4.2 Influence of particle concentration in feed streams As mentioned before, the collisions between particles significantly affect the impinging stream process" while the frequency of the collisions is related to the concentration of particles in the feeding stream. It would be of interest to make certain theoretical predictions for the relationship between the variables mentioned. The analysis presented by Elperin el al. [44] for this purpose is a valuable reference, which is briefly introduced below. Consider the impingement between two opposed particles-gas suspension streams from accelerating tubes of the same diameter. The assumptions made in the establishment of the model are: (1) The streams are symmetrical with respect to both the jet axis and the impingement plane" (2) The gas flow velocity and all the physical properties of gas and solid are kept constant; and (3) The particles beyond collision penetrate into the opposing stream up to x,,,~,~, while any particle will be drawn out of the system immediately once it collides with another particle. Let Z0 denote the initial number of particles crossing the impingement plane and penetrating into the opposing stream at t - 0. Due to collision, the number of particles is reduced by dZ within the time interval dt. The reduction of particle number is assumed to obey the first-order kinetics, that is dZ 1 . . . . Z dt rm
(2.39)
where r,,~ is the mean time between successive collisions. Integration of Eq. (2.39) leads to
64
IMPINGING STREAMS
Z = exp(_ t/Z.m) Zo
(2.40)
where the ratio Z/Zo represents the fraction of particles that do not suffer collision, and can be defined as the probability for collisionless flow, P(t): (2.41)
P(t) = Z / Z o
while [1-P(t)] is the collision probability. It is clear that P(0) = 0; while P(t) decreases with penetration of particles into the opposing stream. For a spherical particle, the effective cross section for collision, or, is o r - Za/p2 If the mean relative velocity of the particle with respect to its neighboring particles is denoted by Up,r, then the collision number along a path of the length equal to Up,r, N, is given by N
-
n m (o'/,/p, r ) -
2
7gnmUp,rd p
(2.42)
where nm is the mean number of particles per unit volume in the system. Then the mean time between successive collisions, ~'m,can be calculated as
m
1
1
N
7£nmUp,rd p
(2.43)
It is obvious that, if all the particles move at the same velocity, then Up, r -" 0 , the mean time between successive collisions ~'m equals infinity, the movement would be collisionless. The larger the value of Up,r, the smaller is the value of Z'm and, consequently, the higher is the collision frequency. In the system under consideration, the total volume of particles, Vp, is much smaller than that of the suspension, V, namely Vp << V. Therefore the averaged particles-to-gas volumetric ratio for calculation can be simplified to _
Vp
V-V
=Vp_nm
V
7c 3
-gd
(2.44)
Solving Eq. (2.44) for nm, substituting the resulting equation into Eq. (2.43); and substituting the further resulting expression into Eq. (2.40) and combining it with Eq. (2.41) yields P(t) - e x p -
dp
(2.45)
PARTICLE BEHAVIOR
65
Equation (2.45) can be used for the estimation of the dependence of the degree of collisionless movement on the inlet concentration of particles. However, the estimation of the values for tim at various probabilities of collisionless movement needs the substitutions of t and Up.,.by some known variables. If it is assumed that the particles have performed one complete oscillation, the particle crossing the impingement plane may collide with the following particles: (1) Particles moving uniformly at a constant velocity in the opposite direction. (2) Particles undergoing their first acceleration and moving in the opposite direction. (3) Particles moving in the same direction but at a lower velocity. These particles are in their second penetration, i.e., in the second deceleration. (4) Particles moving in their second acceleration stage and in the opposite direction. Through a theoretical analysis Elperin obtained the following relationship:
_{ '6° -
dp u p,d IUp,ac 1 In P(t) 3xm~,x[Hp.acl (Hp,acl + b/p()) + 2Up,aclUp0
(2.46)
where ~ is the particles-to-gas volumetric ratio in the feeding suspension; Up.d~ is the mean velocity of particles in the first penetration into the opposing stream, i.e., in their first deceleration stage of motion; Up a~,~ is the mean velocity of particles in their first acceleration; and Up,0is the mean velocity of particles just outside the accelerating tube. Equation (2.46) indicates that ~ is positively proportional to lnP(t). Therefore ]~ must be very small, if collisionless movement is wanted; Vice versa, larger values of/~ would lead to significant collisions. Elperin analyzed the case of the particles-air suspension impinging streams with dp= 2.25x10 -~ m and pp=ll70 kg.m -3 and concluded the following: (1) Collisionless flow can be expected when ]~ < 10-4; and (2) When ]~ > (1-5) x l0 -3, the collisions between particles become significant.
2.4.3 The influence of collision between particles The calculated results described above demonstrate that, in some cases, the collisions between particles cannot be neglected. The results of collision are, firstly, that it shortens the residence time of particles in the impingement zone, partially offsetting the positive effect of impinging streams promoting transfer between phases. Secondly, in gas-solid suspension systems, strong collisions may result in breaking-up of particles; while in a gas-liquid system it causes re-atomization and/or coalescence of liquid droplets. Not all these phenomena are harmful, of course. In fact, the action of particle break-up due to collisions between particles has been successfully used for grinding and milling of solids; while the re-atomization might also be useful on some occasions when further dispersion of liquid may be required. The movement and collisions of particle crowds in impinging streams are very complex and have great randomicity. Many researchers have devoted themselves to the
66
IMPINGING STREAMS
study of these problems for certain methods of qualitative description of such phenomena. Unfortunately, no ideal solution has yet been found. Considering the similarity of the collision between particles in impinging streams and that between molecules in gas molecular motion, Culick [45] and Pai [46] applied a modified Boltzmann equation for solid particles. However, some factors may be unimportant for dynamics of gas molecules; but they have to be considered in the Boltzmann equation for the flow of gas-solid suspension, including the interaction between gas and particles, non-elasticity of the collision between particles, friction problem, and size distribution of particles, etc. If all these factors were introduced into the equation, the mathematical description would become extremely complex, yielding non-linear integral-differential equations that cannot be solved analytically. Kitron et al. [47-49] developed the Monte-Carlo simulation method for solving the Boltzmann equation numerically. Although great simplifications had been made and only a few particles were simulated, the procedure of solution was very complex and tedious; while the results obtained have little practical interest. Probably, this topic calls for further investigation.
-3RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
Any process takes a certain amount of time and the length of the residence time often dictates the occasions when particular equipment or technology can be used. On the other hand, in almost all chemical unit processes the driving forces vary from time to time, and therefore time has the nature of non-equivalence, i.e., an equal time interval yields different, even greatly different, results for the early and later stages of a process. The "result" mentioned here means the processing amount accomplished, such as the increments of reaction conversion, absorption efficiency, moisture removal etc. Normally, these parameters vary as parabolic curves with time. Because of the nature of the non-equivalence of time, in addition to the mean residence time, the residence time distribution (RTD) affects the performance of equipment, and thus receives common attention. For devices employing gas-continuous impinging streams (GIS), the most important residence times and their distributions are those of the dispersed phases, i.e., solid particles or liquid droplets. This is because the materials in dispersed phases are usually the targets of processing, e.g. solid particles to be dried, powdery coal or droplets of liquid fuel in burning, etc; or the substances that play very important roles in the processes, such as catalysts. It call be deduced that, in impinging stream devices, the residence times and their distributions of particles and liquid droplets have similar characteristics; although the numerical characteristics of the distributions such as the mean residence time etc. may have some differences. Therefore, the investigation into residence time distribution of solid particles is of typical significance. In nature, residence time distribution is an important behavior of particle crowds; because of its complexity, it will be subject of locus in this chapter.
3.1 THEORETICAL CONSIDERATION A theoretical analysis is helpful for understanding the basic characteristics of impinging stream processes and the performances of the related devices. In an impinging stream device, where the residence time distribution of particles is most important is in the impingement zone, because this zone is the major active region for heat and mass transfer between phases in such a device. Unfortunately, it is basically
67
68
IMPINGING STREAMS
impossible to measure the residence time and its distribution in this region directly and individually, because the zone has no physical boundary and thus sampling cannot be carried out. Therefore, a modeling-simulation method is needed for investigating this topic, i.e., one uses certain well established theoretical model(s) for simulation and then identifies the reasonableness and feasibility of the model(s) and determines the residence time distribution by comparing the results experimentally obtained with those simulated using the model(s).
3.1.1 Impinging stream device Because of the importance of the residence time distribution in the impingement zone, the basic requirement when designing the experimental equipment is that it should be suitable for understanding the residence time and its distribution of particles in the impingement zone. For this purpose, the rest of the equipment should be as simple as possible. From the point of view of residence time distribution, this means shortening the residence times in the spaces other than the impingement zone as much as possible. Wu and Wu [50, 51 ] studied the residence time distribution of particles in a typical co-axial horizontal gas-solid impinging stream contactor (ISC). The structure of the experimental impinging stream device is shown in Fig. 3.1. The body of the contactor is made of plexiglass. The horizontal cross section of the upper chamber is a square of 0.4x0.4 m, and the bottom is tapered sharply in order to avoid the solid material heaping up. This equipment essentially meets the requirements stated above.
Process particles
Pure process P ~
t Air
Feeder 2
Feeder 1 Airflow 1
Accelerating tube
..... j! !iiii !I !~~i ;'i!~:i!
]
: ii region ::i:::i:: .... on !:/
~, Accelerating Air flow 2 tube
Impingementzone Imp
cles
Figure 3.1 Impinging stream contactor for RTD study.
RESIDENCE TIME OF PARTICLE<+ +~.NDITS DISTRIBUTION
69
3.1.2 Constituents of RTD of particles in the ISC An impinging stream contactor usually contains several flow spaces of different natures, which exhibit not only different flow characteristics but also different contributions to the residence time distribution. In the experimental equipment shown in Fig. 3.1, the particles, as they pass through the contactor, undergo four stages of different characteristics of flow or movement and thus of different residence time distribution properties (as described below), with the time of particles and gas flow motion from the exits of the accelerating tubes to the impingement zone being neglected for the very short distance and very high velocity of the streams.
3.1.2.1 Inside the accelerating tube Inside the accelerating tube the particles are accelerated by the air flow from zero to a certain velocity, usually 60-70% of the gas flow velocity. During acceleration, the relative velocity between the particles and the gas flow is very high. On the other hand, the concentration of particles in the solid-gas suspension to be processed with impinging streams is generally very small, as mentioned above, so that the interaction between particles can be neglected. Therefore the movement of particles can be described approximately with Newton's motion equations for a single particle, as follows: dH
P = 0.75CD[ p" l(u~, - Up)2
(3.1)
dup--0.75CD[ Pa l(ua_Up)2
(3.2)
dt
up dl
ppdp
ppdp
where the drag force coefficient (7'~ can be calculated with the well known relationships for different flow regimes [20], which have been given in Chapter 2 as Eqs. (2.5) to (2.7); while the Reynolds number of particles is defined as
Rep = dPP" ll.,, [u...- . p]
(3.3)
In most cases, the major part of the particles movement in this stage is in the turbulent regime because of the high relative velocity between particles and gas flow, and so this space is also an active region for heat and mass transfer in the impinging stream device. The residence time distribution of particles is related to the properties of the particles and the gas flow, including the size distribution, and the velocity of gas flow and its profile. In practically applicable impinging stream devices, the particles being processed usually have relatively narrower size distribution; the diameter of the tube to particles size ratio, dJd~, is normally very large (>>15); while the gas velocity is high
70
I M P I N G I N G STREAMS
enough (> 10 m.s-J). Under these conditions, the movement of particles inside the accelerating tube can be considered to behave as a plug flow, for which the residence time distribution probability density function, E~c(t), is represented by E.c (t) - 6(t - t.~ )
(3.4)
where the 6-function has the well-known properties below:
f t ~ O, 6 ( t ) - O t-O, 6(t)-oo
and
~+£ 8 (t) d t - 1
The mean residence time of the particles in the accelerating tube can be calculated by the motion equation for a single particle; this procedure was discussed in the last chapter. For instance, for the case where the solid particles are millets with the average diameter dp - 1.6872x 10-3 m, the density pp = 1101 kg.m-3; the length of accelerating tube is L~c = 0.585 m; and the velocity of the air flow inside the accelerating tube, the impinging velocity, u0, at 20°C ranges from 9.48 to 17.36 m.s -~ the calculated
i.e.
mean residence time of particles in the accelerating tube, t,c, is in the range from 0.16 to 0.29 s.
3.1.2.2 In the impingement zone The impingement zone is a small region of a thin cylinder between the exits of the two accelerating tubes without a physical boundary. It is the major active region in the impinging stream device, and therefore the residence time of particles in it is the most important. It was mentioned in Chapter 2 that the penetration of particles to and fro between the opposing streams has an important influence on the residence time. Because the penetration phenomenon is highly random, it is reasonable to assume that the particles in this region are in perfect mixing. This assumption is not only theoretically reasonable but is also supported by experimental evidence. Goldshtein [52] employed a vertical two impinging stream equipment for mixing solid particles. It is difficult to evaluate the practical applicability of the equipment Goldshtein designed, but it is beyond question that the results demonstrate well the performance of the impinging stream device for mixing solid particles. From the assumption of perfect mixing, the corresponding residence time distribution probability density function is well known as
Eim (t) -
1 tim
exp(- J/:
/rim
)
(3.5)
Part of the experimental results on the residence time of a single particle in the impingement zone was given in Fig. 2.6 in Chapter 2. These results, at least, can bound
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
71
the range of the values for the mean residence time of particles in the impingement zone, and thus are useful for understanding the basic natures of IS devices.
3.1.2.3 In the falling down region After impingement, the two originally opposing streams are mixed with each other and combined, and then turned to become a radial flow outwards. As described in Chapter 1, there exist two opposite factors causing the variation of the radial gas flow velocity: (1) The passage area for the radial flow increases as the radial distance, y, increases, yielding a negative influence on the radial velocity; and (2) After the turn of flow direction, the fluid originally in the region of r < y must flow outwards through the point of radial distance y, and so the amount of gas flowing outwards passing through Point y must be increased as y increases, yielding a positive effect on the radial velocity. As the results of the combined effect of the two factors in contradiction, there must be a maximum value on the curve of the radial velocity versus the radial distance y. For the relationship between the radial gas velocity and the radial distance Elperin [3] obtained experimentally the expression below:
u~,-1.5u0 tR)~ exp(-1.52 R )
(3.6)
where R is the radius of the accelerating tube, and u0 is the gas velocity in the accelerating tube, i.e., the impinging velocity. By differentiating Eq. (3.6) with respect to r and setting the derivative to be zero, the maximum value of the radial velocity is found to appear at r = 1.974R, as u~,..,~,X = 0.575u 0
(3.7)
Since most of the particles are thrown out of the impinging streams at that radial coordinate, r = 1.974R can be considered as the pseudo-boundary of the impingement zone. In addition, it is easy to get the relation of u~,,r = 0.004u~, at r = 8R, which is negligible in comparison with u0. If the particles are assumed to leave the impingement zone at r = 1.974R, then the movement of the particles leaving the impingement zone can be assumed to be affected by the gas flow only in the range of the radial coordinate, r, from 1.974R to 8R. The particles may leave the impingement zone in various directions, as shown in Fig. 3.2. To calculate the residence time in the falling down region, only the motion in the vertical direction needs to be considered. If the direction vertically downward is set to be the positive direction for both the vertical velocity of particles, Up,h and the height coordinate, h, and let/6' (0 ° ___fl < 360 °) be the intersection angle between this positive direction and the direction the particles are leaving in, then we have
~=h/cos~
(3.8)
72
IMPINGING STREAMS
r = 1.974R
1 Figure 3.2 Falling down of particles. Particles with initial velocities in different directions will move along different trajectories, and so need different times for falling down to a certain height. The residence time distribution of particles in this stage can be determined with the motion equation, which is obtained from the force balance, while neglecting the buoyant force, as follows: (1) For 0 ° < fl < 90 ° or 270°<]3 < 360 ° and 1.9737Rcosfl < h < 8Rcosfl = 0.75Cf[ ~a ][1.5(Ih13) exp(-1.52 I hi )cos/~b/a duph dt flpdp R cos/3 R COS 1~
-- uph
]2
+ g
(3.9)
(2) For 90°
dUph= - 0 . 7 5 C f [ ~a 1[1.5( Ih 3e x p ( - 1 . 5 2 Ihl dt
ppdp
R cos/3
R COS/~
)cos ~b/a
--//ph
]2 + g
(3.10)
(3) Outside the ranges specified above dttph
dt
=-0.75Cf[
2
Pa ]Uph+g
ppdp
(3.1~)
where the definition of the vertical velocity is known as dh Uph -- dt
(3.12)
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
73
The initial conditions for Eqs. (3.9) to (3.10) are t-0"
Uph--0,
(3.13)
h-l.974Rcosfl
The time needed for particles falling down to the height of h - H depends on the intersection angle ,6, and can be found by integrating Eq. (3.12) as H 1 t - Jl ~)74eco~¢~', dh
(3.14)
b/ph
where up, is determined by an equation properly selected from Eqs. (3.9) to (3.11). Equation (3.14) represents the relationship between the height the particles fall down to, H, and the initial intersection angle of the particles, ,6. In principle this relationship can be expressed by (3.15)
-fit)
where the subscript "t" indicates that the residence time of the particles having the initial intersection angle fl is equal to t. It is clear that the particles with an initial intersection angle o f / 3 = 180 ° need the longest falling time; while those with an angle fl = 0 ° need the shortest falling time. Generally, in the range of 0-180 °, the falling time needed by all the particles with an initial intersection angle of/5' < ~ must be < t. If the particles leaving the impingement zone are distributed uniformly over all the directions and the symmetric nature of the impingement plane is taken into account, the definition of residence time distribution yields directly the expression below: (3.16)
F(t) - ~
180
1.0
I
0.8
"~ 0.6
0.4 0.2
..............
0.0
i
0
i
i
i
i
i
0.1
i
i
i
i
i
0.2
i i 1 " 1
II III IIII III l l i 0.3
0.4
0.5
0.6
t,S Figure 3.3 Sub-RTD of particles calculated for falling region.
74
IMPINGING STREAMS
As an example, for the case where the gas velocity inside the accelerating tube u0 = 11.07 m.s -~ and the solid particles are millets or rape seeds, the sub-distribution of particle residence time in the falling down region calculated for the equipment shown in Fig. 3.1 with the relationships given above is shown in Fig. 3.3, and the mean residence time of particles in this sub-space is t fal- 0.343 s. It can be seen from Fig. 3.3 that the non-uniformity of the residence time distribution in this stage is limited, so that the movement of particles can be assumed to behave as a plug flow without significant error. Thus, the residence time distribution probability density function in the falling down region can be represented by
Efal(t)
-
6(t-tfal )
(3.17)
The mean residence time in the falling down region depends mainly on the structural dimensions, in addition to the properties of the particles and the gas; while is little affected by operation parameters. Therefore the data shown in Fig. 3.3 are applicable for the experimental study to be discussed later in this chapter. Particle
tl°,,
Figure 3.4 Collision of particles on the walls of the tapered bottom of IS device.
3. 1.2.4 In the collision-slipping region Collision on the wall and slipping out of particles in the tapered bottom of the contactor is an additional stage of particles motion. It is possible for the following two situations to occur simultaneously: (1) Elastic collisions on the walls change the directions of particles moving multiply, as shown in Fig. 3.4, where t~, t2, t3, and t4 denote different instants. During the movement from one side wall to the other, particles may collide with others and then change their directions further. Such variation of motion direction must result in significant back-mixing; (2) Non elastic collisions on the walls may also
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
75
occur with part of the particles. In this case particles that drop down to the bottom wall may slip along the incline surface, and, during slipping, the particles may multiply overlap or be inserted in between by the newcomers, leading also to back-mixing. The movement of the particles in this stage is very complex and extremely random, so that to determine accurately the residence time distribution and the mean residence time is difficult, whether by theoretical analysis or experimental measurement. On the other hand, the residence time distribution in this stage is unimportant because this subspace is essentially inert for heat and mass transfer. Considering the presence of significant back-mixing, the flow of the particles in this stage is assumed also to be in perfect mixing, as a first-order approximation, and thus the residence time distribution probability density function is of the form below: E~,~(t) - ~ 1 exp(-t--t/( ) [CS
/
(3.18)
]f C S
where ~., is the mean residence time of particles in the collision-slipping region.
3.1.3 Model for the overall residence time distribution In the impinging stream contactor shown in Fig. 3.1 the particles pass through the four sub-spaces in series described above, and the overall residence time distribution probability density function in the system, E ( t ) , can be written as [53] E ( t ) = E ,~ (t) * Ei,,, (t) * Er~, (t) * E~ (t)
(3.19)
where the symbol "*" denotes the convolution integral defined as Ei(t)*
E i(t)-
I I)Ei(a)Ei(t-a)da-
~DEi(a)Ei(t-a)da
(3.20)
while the relationship between the residence time distribution function, F(t), and the distribution probability density function is well known as F ( t ) - ~D E ( t ) d t
(3.21)
Equation (3.19) can be rearranged to be E(t) = Ei,,, (t) • E~., (t) • E;,~. (t) • Eji, i (t)
(3.22)
E.~c ( t ) * Et~l (t) = d ( t - t .a~.) . d ( t - t f .ai ) = d ( t - t ~c - t f.~l )
(3.23)
where
Note that, for the dynamic response, both t ac and t r~ are transfer lag time, also called pure lag time. For convenience, the total lag time, t~g, is defined as
76
IMPINGING STREAMS
tlag -- t~c + tf~l
(3.24)
Eac (t) * Era1(t) - ~'(t - tlag )
(3.25)
Then Eq. (3.23) can be simplified to
According to the nature of residence time distribution, the following relationship should hold: E(t)
-
0,
F(t)
-
0
for t < tiag
(3.26)
On the other hand, the first convolution integral in Eq. (3.22) can be found by Laplace transformation and its inversion, as
Eim (t) * Ecs (t) - ~_
1
_
tcs --tim
[exp(-t /tc~ ) - exp(-t /tim )] (3.27)
-
1
-
tim --tcs
[exp(-t / tim) - e x p ( - t / tc~)]
{t > tlag)
Substituting Eqs. (3.25) and (3.27) into Eq. (3.19) and solving the resulting expression for the final yield
E(t)
-
1
~_ { e x _p
tcs -- tim
[-(t - flag )/tcs ] - e x p [-(t - t,~g ) / t im ]} (3.28)
~ { e tim --tcs
xp [-(t - flag ) / t im ] - e x p [--(t -/lag )//cs ]} (t-> 'lag )
This is the model for the overall residence time distribution of the particles in the impinging stream contactor under consideration. The model contains several parameters related to equipment structure and operating conditions, i . e . the mean residence times in the four sub-spaces,
t ac, tim, t fal and t cs. Among the four
parameters, the mean residence time in the impingement zone, tim, and that in the collision-slipping region, tcs, are symmetrical parameters, which have the same influence on the overall residence time distribution. It can be seen from Eq. (3.27) or (3.28) that the model requires that tim :/: tcs. If this requirement is not met, the function would have no definition. However, this will not be a problem. It is observed in all
E(t)
the experiments that the relation of t cs > tim always holds.
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
77
The following may be concluded theoretically from the above analysis: (1) The flow of the particles inside the impinging stream contactor has the characteristics of perfect mixing--plug flow in series; (2) Back-mixing in the impingement zone, the major active region for heat and mass transfer, is critical. This may be favorable for some processes; while it is harmful for others since back-mixing usually results in a decreased mean driving force for the processes involved; (3) The existing results of theoretical analysis and some experiments indicate that both the mean residence times of the particles in the major and the secondary active regions, the impingement zone and the space inside the accelerating tube, are very short, totally about 1 s only. This is a great limiting factor for practical application of impinging streams. The primary conclusions above are helpful for understanding possible performances of impinging stream devices.
3.2 METHOD FOR EXPERIMENTAL MEASUREMENT OF PARTICLES' RESIDENCE TIME DISTRIBUTION Theoretical analysis gives some useful information on the characteristics of residence time and its distribution of particles although experimental evidence is always important. It is well known that the measurement of residence time distribution usually employs the dynamic method [54], the so-called input-response technique. However, for measuring RTD of solid particles the input signal is a difficult and troublesome problem. The author of the present book employs an arbitrary known function as the input signal so that this problem is solved. This procedure is also applicable, in principle, to the measurements of RTD of solid materials in other devices.
3.2.1 Input signal In the measurement of residence time distribution, the most convenient and so most widely employed input signals are the impulse and the step change. For a fluid, as continuous phase, whether it is a liquid or a gas, both the signals are simple and convenient to carry out; while for solid particles, as dispersed phase, it is very difficult to input any one of the two signals usually used, and the so-called frequency response [55] is even more difficult. One may consider the facts below: could any conveyer for solid particles or powders be used to provide a step change of tracer to the following equipment under conditions of ensuring stable flow rate? Inputting tracer particles in pulse form is theoretically possible, but the amount of particles to be inputted is a problem. If the amount of tracer particles inputted is as large as is needed, the width of the pulse would be considerably increased, yielding unacceptable experimental errors;
78
IMPINGING STREAMS
while if the amount is small, the randomness of movement of the particles would conceal the regularity of the residence time distribution. Luzzatto et al. [56] studied experimentally the residence time distribution of particles in a co-axial gas-solid two impinging stream reactor with a special ejector to input radioactive-tagged particles to the reactor, and interrelated the data they obtained by Markov chains. In order to keep the flow conditions stable and, at the same time, to ensure that the input signal was close to an impulse, the number of the tagged particles they inputted to the system for each run was only about 200 particles; while, in comparison, the number of process particles in the equipment is almost a million. Since the movement of particles is highly random, one experimental run cannot yield useful data from so small a number of tracer particles, so that the researcher had to repeat the experiment many times under the same conditions to obtain a set of statistically averaged results. The experimental procedure is considerably long-winded, and the interpretation of data with Markov chains is also significantly troublesome. Obviously, if no special restriction is exerted on the input signal, the experimental measurement of residence time distribution of solid particles would become simpler and much more convenient. From observations it is found that, inside the glide tube of a hopper for solid particles feeding, the movement of particles behaves very like a plug flow, as shown in Fig. 3.5. This implies that the hopper can provide the following device with a step change with good approximation. On the other hand, it is also found by experiments that a screw feeder with good shifting performance cannot only control very stable feed flow rate but also give a very stable response to its input of a step change with good reproducibility of data. The observations described above indicate that, with good control, the concentration of the tracer particles in the out stream of the screw feeder can be determined to be a known function of time, and, furthermore, it is feasible to use such a known function as the input signal to the impinging stream equipment to be tested. In this way the experimental procedure can be greatly simplified. Of course, this scheme calls for corresponding mathematical relationship(s) for data interpretation.
Figure 3.5 Plug-flow movement of particles within feeding tube.
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
79
Tracer
P r o c e s s particles
Hopper
b-----q
To a c c e l e r a t i n g tube
Figure 3.6 Feeder to IS equipment. The screw feeder used in the experiments is shown in Fig. 3.6. It is driven by a motor with very stable rotary speed and good speed-shifting performance. Two sets of screw feeders are used in the experimental system, which are mounted oppositely and symmetrically. The outlet of the glide tube of each hopper is connected directly to one particle feeding tube of the impinging stream contactor, i.e., the accelerating tube. In operation, two air streams flowing in the two accelerating tubes accelerate the particles fed by the two screw feeders, respectively, and then the particles-in-air suspension streams are ejected from the accelerating tubes towards the impingement zone. Initially the two feeders are ted with only the process particles B. Under the condition of stable flow, Feeder 1 is inputted with a step change of the tracer particles A; while Feeder 2 is continuously ted with process particles B at a stable flow rate. A certain time later, the particles of Tracer A emerge at the outlet of Feeder 1. The instant the tracer A emerges at the outlet of Feeder 1 is taken as the initial time for the residence time distribution measurement. After that, the concentration of Tracer A in the out stream of Feeder 1 varies with time , and this function of time, ~ ll(t) "-~Ai
'
is the response of Feeder 1 to its
input signal, the step change; it is also the input signal to the following equipment to be tested for RTD, which can be determined to become a known function of time. In order to protect the flow stability from turbulence caused by the input signal, the properties of the tracer used should be as close as possible to those of the process particles. In the investigation carried out by the author of this book the process particles are yellow millets; while purplish-red rape seeds are used as the tracer, the properties of which are very similar those of the millets. The properties of the process and the tracer particles are listed in Table 3.1. The concentration of the tracer is represented in terms of mass fraction, and is measured by manually separating the tracer from the process particles according to the difference in color and weighing the amount of tracer. This is laborious and time-consuming work, but it can yield reliable data. Using the procedure described above, input a step change of the tracer to the hopper and then measure the response of the screw feeder to the step change as a known function of time, which is used as the input signal to the following impinging stream device for the measurement of RTD.
80
IMPINGING STREAMS
Table 3.1 Properties of the materials used for RTD measurement Property Material Pp, kg'm -3
Pb, kg'm -3
dp X 103, m
Millets (B)
1101
661.5
1.6872
Rape seeds (A)
1172
696.5
1.6304
It is clear that the response of the screw feeder to the step change of the tracer depends on the rotary speed of the screw. For convenience, the responses at various rotary speeds are calibrated prior to the experimental measurements as the curves of the tracer concentration versus time, which represent the responses as known functions of time. For each rotary speed, the calibration was carried out two or three times, and the data were averaged. A set of typical data calibrated is shown in Fig. 3.7. It demonstrates that there is good regularity and reproducibility of the data, suggesting the calibrated curves can be used as the input signals to the impinging stream device for RTD measurement to yield sufficiently accurate results.
1.0 O
•E
~0.8
o
~0.6
•-~
~0.4
~D
A
o0.2 O t
0.0 -20
0
20
40
60 I,S
80
100
I
120
t
140
Figure 3.7 Typical response of the screw feeder to a step-change of the tracer concentration (rap = 1.914x10 -4 kg.m-3).
R E S I D E N C E T I M E OF P A R T I C L E S A N D ITS D I S T R I B U T I O N
81
3.2.2 Data interpretation The relationship for the interpretation of data measured for RTD can be easily derived with Laplace transformation [54]. However, the equations so obtained are not convenient for time-dispersed measurement because they still need Laplace transformation to deal with the data. In the following, another relationship will be derived.
re(l)[ a,
.I
Air flow
IS Device
"~1
•[" I
Air flow
mao~mBo(m~o) Out particles Figure 3.8 Principle scheme for measurement of particles RTD. Consider the equipment system schemed in Fig. 3.8, where the subscript B denotes the process particles and A the tracer. Corresponding to the impinging stream equipment with feeding on both sides, the system has two feeds, which are denoted by the superscripts (1) and (2), respectively; while it has only one out stream of particles. For convenience of operation, the tracer A is inputted into one feeding stream,
i.e.,
rt 2)
Ai -- 0. The system is assumed to be operated at a steady state, and only the particles B were fed before, mA~ = 0 for t < 0. From the instant of to = 0, the tracer particles A are fed into the equipment at an amount comparable with the particles B, and then both the particles A and B are fed into the device continuously and simultaneously. In order to keep the momentums of the two streams in balance, the operation of the screw feeders ensures that the total amounts of particles of the two feeding streams are always kept equal,
i.e.,
i.e.,
m~i'
+
m(" Bi
-
m Bi '2)
= mBi,
rn~i'
-- mAi
(3.29)
At the instant t, the residence times of all the particles of the tracer A must be < t; while the particles of B consist of two groups: in the first group all the particles are fed at instant to - 0 or later and so also have residence times < t; while the particles in the second group were fed into the device before to = 0 and so have residence times > t. If the particles in the first group are denoted by the superscript "*", then, from the definition of the residence time distribution function, F, we have [
:f:
Io (mA,, + mBo )dt
f ( t ) - I~(mAo + mBo)dt
IomAodt + Io mBodt f[) mto dt
(3.30)
82
IMPINGING STREAMS
On the other hand, one of the essential conditions for correct measurement of RTD is that the tracer particles have the properties, including RTD characteristics, very close to those of the process particles. This implies that, in the time domain of t _> 0, the residence time distribution functions of particles A and B in the same device should be approximately equal to each other, i.e. F A (t) - F B(t),
t>0
(3.31)
It should be noted that Eq. (3.31) is an approximate relationship. This is because both the individual flow rates of the particles A and B vary with time, although the total flow rate, i.e. the flow rate of A plus that of B, is stable, while the variation of particle flow rate may affect RTD. However, since the variations of the flow rates of particles A and B are not so large, the possible deviation of Function F caused by this factor may be neglected, just like the generalized (dimensionless) residence time distribution function can be extended for application in a certain range [57]. According to the definition of the F-function, the residence time distribution functions of A and B for the case under consideration can be expressed by the corresponding ratios of the amounts of particles coming out from the device to those inputted in the time interval from 0 to t, i.e. ~OmAodt F A (t) - - ~ ~ - lOmAidt (3.32) [OmBodt FB(t ) = . IomBidt Using the relationship represented by Eq. (3.31), Eq. (3.32) yields mA°d--------~t I; = m~3°d-------~t I° J'; mAidt
(3.33)
I;mBi dt
and, from the well-known mathematical theorem of proportion by addition, Eq. (3.33) becomes J'omAodt + ItomBodt
j'omAodt =~ ~omAidt + ]'; mBidt j'; mAidt t
(3.34)
The mass balance at steady state gives t dt + ]'omBidt i0mAodt + J'omBodt- J'0mtodt =!0mAi
(3.35)
83
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
Substituting Eq. (3.35) into Eq. (3.34) and combining the resulting expression with Eq. (3.30) leads to
F(t)- I/~mA°d~t
(3.36)
I;mAi dt
On the other hand, the condition of the total material stably flowing suggests the following relationship holds: into - mAo + rnB{, -2[m~i ) + m~Bii)] - mti- const.
(3.37)
Dividing both the numerator and the denominator in Eq. (3.36) by Eq. (3.37) leads to
[[t (mAo /mto )dt F(t) = " fo (rn {l ' /mti )
(3.38)
while the following relationships are easy to obtain:
mA° = CAo into
and
m~i) mti
(m~
(')+ (2) ) + mBi mBi
1
Therefore Eq. (3.38) becomes
F(t) =
2IoCAo(t)dt
Ill r~til "~Ai (t)dt
(3.39)
or, in the case of time-scattered measurement, ZCAo(ti)Ati F(t) = 2 i Z Cdi ) (ti)All i
(3.40)
In the derivation above, the input signal CAi(t) is an arbitrary function of time, without any restraint condition, but, of course, it should be known; CAo is the response of the impinging stream device to CA~(t) that can be measured by sampling at the outlet of the device.
84
IMPINGING STREAMS
3.3 RELATIONSHIPS FOR FITTING DATA Equation (3.39) or (3.40) can be used simply for the determination of the overall residence time distribution, and further the mean residence time, of the particles in the whole equipment according to the data measured for the concentrations of the tracer in the streams inputted to and coming out from the device. However, as can be seen from the analysis in Section 3.1, the impinging stream equipment includes several flow regions of quite different natures so that the global residence time distribution in the whole device is not helpful for understanding the processing performance of an IS device. A combination of the theoretical model with the relationships for data interpretation makes it possible to obtain some important parameters, such as the mean residence time in the impingement zone etc., by fitting experimental data. For convenience, define the integrals below:
t -(1) I i (t) - ~oC'Ai dt (3.41) I o (t) - Io CAodt Their differentiations are d I i ( t ) _ (-,(1)
dt
"Ai (t) (3.42)
t d I o (t) _ CAo (t) L dt On the other hand, using the expressions defined above, Eq. (3.39) can be rewritten as F(t) -
21° (t) ii(t )
(3.43)
For easier simulation, the equations above are reformed further. Differentiating Eq. (3.43) and rearranging the expression leads to dF(t) = E ( t ) = 2 .-,(1)( t ) i ° (t)] d-----~ 12 (t) [cA° (t)Ii (t)--CAi
(3.44)
Thus, we have the set of differential equations consisting of Eqs. (3.41) and (3.44), including the integral terms, of which the initial conditions are I i (0)
=
0,
I o (0) = 0,
F(0) = 0
(3.45)
RESIDENCE TIME OF PARTICLEq AND ITS DISTRIBUTION
85
This set of equations can be used to determine the residence time distribution function of the particles in the equipment, F(t), according to the data measured for
C~ Ai (t)
and CAo (t). However, for simulating calculation, it is more convenient to use the differential equation related to CA<,, which can be obtained from the modified Eq. (3.43) as 1
l o -2F(t)Ii
(3.46)
Differentiating Eq. (3.46) with respect to t yields 1
CA,, - ~[E(t)Ii + F(t)f~Ali)l
(3.47)
On the other hand, for the device shown in Fig. 3.1, the residence time distribution function F(t) in Eq. (3.47) can be found by integrating Eq. (3.28) with respect to t to be 1
F(t) - ~_ [ t ~ ,_~
exp(t,~g/tc,~ ) - t i m exp(t,ag/Tim )
t cs -- tim
(3.48) -
-tc~ exp(
-- l --/lag
-
_
)+tim exp(
-- I --/lag
_
los
tim
while the differential form of Eq. (3.28) is dE(t) dt
1
1
los -- tim
tim
t --/lag
~[_----exp(-
1
_ ) - _---exp(tim /cs
t - tlag
_ los
)]
(3.49)
Differentiating Eq. (3.47) and combining the resulting expression with Eq. (3.49) leads to dCAo
_
1
d - - - ~ - 2(/c~, - tim) {exp[-(t-tiag)/tim][-lim
-exp[-(t-ti.~g)/tc~][/i+-lc~ tcs
Ii
dCA------L- 2CAi ] dt
+[;c~ e x p ( ~ )-tim exp(t-i~g )1 dCAi tcs tim dt } Its initial condition is
-
dCAi
+ tim ~ -]- d t2CAi
(3.50)
86
IMPINGING STREAMS t < tlag : CAo = 0
(3.51)
The calculations for fitting experimental data can be made by solving Eqs. (3.43) and (3.49) simultaneously with a numerical procedure of trial-and-error to determine the parameters tim and tcs, while tlag is calculated by the motion equations given in Section 3.1.
3.4 MAJOR EXPERIMENTAL RESULTS FOR RTD OF PARTICLES
3.4.1 Measurement of tracer concentration As described before, for all the sets of experimental conditions the response of the screw feeder (1) to the step changes of tracer A have been calibrated prior to the experimental measurements of RTD to get the known functions of time, which are taken as the input signals to the impinging stream device, while the responses of the impinging stream contactor to the input signals are measured at the outlet of the device under corresponding conditions. Although the flow properties of the particular materials are never as good as gas or liquid, most of the experiments yield data with considerably good regularity. Figure 3.9 gives a representative set of measured results. In the fitting calculations below, the data used for CAo are taken from the smoothed curves, just like that given in Fig. 3.9.
0.5
0.4
I
<
[]
-
-
0.3 0.2 0.1 0[
.. re
0
I
I
10
i
i
20
I
I
I
30
40
I
I
I
I
50
t,S
Figure 3.9 Typical data measured for response of IS equipment. mp=5.316×10-3 kg.s -l ua=12.64 m.sq
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
87
3.4.2 Comparison between the results measured and simulated Simulating calculations are carried out for each set of conditions with the relationships derived in Section 3.3. The total lag times, tlag, are calculated with the motion equations according to the corresponding conditions, while the mean residence times in the impingement zone and the colliding-slipping region, tim and tcs, are determined by fitting experimental data with a trial-and-error procedure, i.e., simulations are made with various values for tim and t ~ , and then the optimal values are selected that best fit the data. A comparison between the response of the impinging stream contactor measured and calculated for a typical set of conditions is given in Fig. 3.10. Figure 3.11 shows a comparison between the residence time distribution functions simulated (the curve) under the same conditions as for Fig. 3.10 and the results, as data points, obtained by interrelating the data directly with Eq. (3.39) or (3.40), which are actually the experimentally measured values. Both figures show good agreement between the regularities of the experimental data and the simulated results, indicating that the model equations derived are reasonable and the assumptions on the properties of flow and mixing in various sub-spaces are also reasonable. As predicted from the model, the residence time distribution curve shown in Fig. 3.11 exhibits the feature of perfect mixing with a certain lag time, that is, the feature of plug flow-perfect mixing flow in series. All the experiments yield the results of RTD exhibiting the same feature. For the conditions yielding the results shown in Figs. 3.10 and 3.11, the values for _
the parameters respectively.
t~,~ and tim obtained from best fitting data are 4.0 and 0.9 s,
0.8 0.6 <
0.4 0.2 0.0 0
6
12
18
24
30
36
t,S
Figure 3.10 A comparison between the responses of IS device measured and calculated. [] CA~measured; A CAo measured; ---- CAocalculated. Conditions" mp = 3.016x10 -3 kg.s -I" u~= 14.22 m.s-I" mp/m.~= 0.466; S = 0.08 m.
88
IMPINGING STREAMS
1.0
,,
.~
^
~
A
A
0.8 0.6
0.4 0.2 0.0
•
0
I
I
I
6
,
I
s
I
,
I
12
I
I
•
18 t,S
i
l
|
l
24
|
l
|
30
l
|
|
36
Figure 3.11 Comparison between the F-functions measured and calculated. Conditions: 3.016x10-3 kg's -]" ua = 14.22 m.s-]" mp/ma=0.466; S = 0.08 m.
m p "-
3.4.3 Mean residence times of particles The mean residence times of the particles in various sub-spaces of the impinging stream device are determined by interpretation of the experimental data of 27 sets in total, and the conditions tested range as follows: Particles flow rate: m p = 2.346-9.047x10 -3 kg.s-" Impinging velocity: u0= 11.07-14.22 m.s-1; Mass flow rate ratio of particles to air: mplma-0.466-1.4; and Impinging distance: S = 0.06-0.1 m. The following major results are obtained: (1) The mean residence time of particles in the impingement zone varies in the range of 0.6 < tim _< 1.5 s, mostly in the range of 0.8-1.1 s; (2) The mean residence time in the colliding-slipping region in the tapered bottom of the device ranges from 3.0 to 4.0 s; (3) The total mean residence time of particles in the whole device is between 4.8 and 6.0 s; and (4) No statistically remarkable regularity of the effects of the operating conditions on the mean residence time in the impingement zone, tim, has been found, although the latter is not constant. The results of Item (4) above are somewhat unfortunate, because the parameter tim is the most relevant. The major reason for this may lie in the fact that the operating variables have contradictory influences on tim. For example, an increase in the
RESIDENCE TIME OF PARTICLES AND ITS DISTRIBUTION
89
impinging velocity may enhance the penetration of particles to and fro between the opposing streams, yielding an increased mean residence time in this zone, while, at the same time, it also increases the probability of collision between particles, yielding, inversely, decreased mean residence time in the same zone. In addition, the influences of the operating variables are easily covered by some random variations in the system of particles. All these factors mean that, essentially, no defined regularity on the variation of tim can be determined.
3.5 REMARKS Without doubt, the gas-continuous impinging streams (GIS) method has been proved by a number of investigations to significantly enhance transfer between phases [58-60, etc. ]. This feature gives it good application potential. On the other hand, the residence time of particles and its distribution are obviously important, because they directly affect the processing performance of impinging stream equipment. All the results of investigations to date, both theoretical and experimental, carried out by the author of this book and other researchers indicate that the mean residence time of material in the dispersed phase in the major active region of an IS device, i.e., the impingement zone, is very short, about 1 s only. To a great extent, this fact limits the application of GIS. Indeed, some fast processes, such as combustion of powdery coal or sprayed liquid fuels etc., can be carried out in an individual impinging stream device with very high efficiency. However, many other processes involved in various industries need considerably longer times, even under the conditions of strongly enhanced heat and mass transfer. A typical example is the drying of porous particles such as PVC. This kind of material to be dried usually contains both surface or free moisture and in-porous or combined moisture. The former is easy to remove, while the removal of the latter needs a very long time, say, several tens of minutes, even if the transfer processes involved are strongly enhanced. It is obvious that individual impinging stream equipment is not applicable for such processes. Even those processes needing only several tens of seconds under the conditions of enhanced transfer cannot be carried out with an IS device individually, since, after all, the flow configuration of the impinging stream device is very complex making it difficult, if not impossible, to arrange a multistage countercurrent system. So, very short residence time is a problem that has to be considered in GIS application. Only solid particles' residence time and its distribution are discussed in the present chapter although, because of the similarities of movements of the dispersed pha~es in impinging streams, the results described above are also of referential significance for gas-liquid impinging streams. The differences between the properties of liquids and solids and their influences will be discussed in detail in the relevant chapters. Another problem arising from the results of the investigation on residence time distribution is the strong mixing of the materials in dispersed phase in the impingement zone. The fact that the model of RTD derived above fits well the concentration of the tracer in the out stream of the device indicates that the assumption of perfect mixing of
90
IMPINGING STREAMS
dispersed phase in the impingement zone is reasonable. In addition, the results of other investigations, e.g., those reported by Tamir et al. [61], provided forceful experimental evidence for this argument. For some processes, strong mixing of dispersed phase may have positive significance. However, back-mixing usually implies decreased mean driving force, and thus decreased average rate of the process involved. This may, to an extent, offset the positive effect of IS enhancing transfer. For fast and irreversible reactions the influence of strong back-mixing may be limited, while for those restricted by equilibrium, such as some of the absorption systems etc., it would have significant negative effects on the processes. On the other hand, it is practically impossible to arrange a multistage countercurrent system, like in a tower or column, because of the complexity of the flow configuration of impinging streams. Therefore back-mixing is also a limitation of IS application for some processes, and is thus a notable problem. In the investigation on residence time distribution described in this chapter, some special procedures were employed. In the theoretical analysis, the whole impinging stream contactor is divided into several sub-regions according to the flow properties in them, and the residence time distribution and its constituents in the sub-regions are studied separately. The results are helpful for understanding the performance of an IS device. In the experimental investigation, a known arbitrary function of time is used as the input signal to the device being tested, resulting in a greatly simplified experimental procedure. The results obtained indicate that both the theoretical and experimental methods are successful, and so have general significance for investigations on RTD in other devices, especially for RTD of solid particles.
-4HYDRAULIC RESISTANCE OF IMPINGING STREAM DEVICES
Gas-continuous impinging streams involve flows at high velocity, and so power consumption naturally becomes an important concern [62]. As is well known, the theoretical or minimum work per unit time for fluid transportation is equal to the product of the pressure drop and the volumetric flow rate of the fluid: Wm~n = ~t'V
(4.1)
Because of the very small density of gas, in most processes of practical interest the volumetric flow rates of gases, Vg, are very large. Therefore the hydraulic resistance, or the pressure drop, of the device to be employed has a crucial influence on the total power consumption. Tamir et al. [56, 63-65, 238] studied the total pressure drops over several gas-solid impinging stream contactors with various structures and dimensions. However, the data they measured covered a very large range, from 20 to 3725 Pa, while few generalized results have been obtained for the power needed to operate impinging stream devices. On the other hand, the experimental equipment employed by Tamir et al. contained a number of complex and flow configurations of uncertain necessity, e.g., tangential flow, multi pair and multistage impinging streams, etc., which must yield additional hydraulic resistance; while their sizes were very small. For these reasons, it is difficult to use the results they obtained tot predicting the hydraulic resistance of industrial impinging stream devices, and thus these results had no general significance for equipment design and scale-up. In order to obtain more generalized results that can be used for predicting hydraulic resistance and equipment design, Wu and Wu [66] studied the hydraulic resistance of a horizontal two-impinging stream contactor and examined the total pressure drop across the device and its constituents in detail, paying special attention to the sub-pressure drops caused by the acceleration of particles and impingement between the opposing streams because these two sub-drops are the essential elements that must be present in the operation of any gas-solid impinging stream device.
91
92
IMPINGING STREAMS
4.1 THEORETICAL CONSIDERATION Generally, the hydraulic resistance of a gas-continuous impinging stream device can be considered to result from the following factors: (1) The drag force due to friction between the gas flow passing through the accelerating tube and the wall of the tube; (2) The friction force between the gas flow and the particles in dispersed phase; (3) The impingement between the opposing streams; and (4) Certain structural factors of the equipment involved, such as sudden contraction of flow passage and some other structural parts etc. Examining the individual sub-pressure drops caused by various factors separately is more helpful for understanding the features of impinging stream devices, and can also provide experimental evidence of practical interest for equipment design and scale-up. The most important operating parameter for gas-continuous impinging streams is the velocity of the gas flow at the outlet of the accelerating tube, also called impinging velocity, u0. Therefore, the most convenient approach is to relate all the individual hydraulic resistances to u0.
4.1.1 Flow through the accelerating tubes During the flow of the streams through the accelerating tubes, several factors may lead to pressure drop along the path. These factors include friction between the gas flow and the inside wall of the tube, acceleration of the particles, collisions of particles on the wall and between particles etc. For convenience, the pressure drop through the accelerating tube can be considered to consist of two constituents caused by gas flow and particles, respectively, -Ap,c,~ and -Apac,p, which are discussed separately below.
4.1.1.1 Resistance to the gas flow Similar to the calculation of pipe-resistance, the pressure drop caused by the gas flow passing through the accelerating tube, -Apse,a, can be represented by Lac pa u2 -- APac'a - )~a dac 2
(4.2)
where the friction coefficient, 2a, is a function of both the Reynolds number, Rea, and the relative roughness of the inside wall surface, Edac, and the well-known curves representing the relationship of 2, versus Rea, with e/d~c as the referring parameter, can be found in Ref. [20] or other references. Under isothermal conditions, the gas flow velocity in Eq. (4.2) may theoretically be affected by both the variation of the local pressure and the presence of particles. However, in practical cases the pressure drop caused by the gas flow through the accelerating tube is very small in comparison with the operating pressure so that the latter can be considered as constant, and thus the
HYDRAULIC RESISTANCE OF IMPINGING STREAM DEVICES
93
variation of local pressure is negligible. At the same time, the solid-in-gas suspension to be processed is usually thin dilution, e.g., the particles to gas mass flow rate ratio normally ranges from 0.5 to 2.0 and, correspondingly, the volumetric fraction of particles is very small (< 0.5%). Therefore u0 can be considered as constant, independent of the local pressure and the presence of particles.
4.1.1.2 Pressure drop due to acceleration of particles The pressure drop resulting from the movement of particles can be determined by an energy balance. The overall energy balance round the flow of suspension inside the accelerating tube is written as
0
particles
1
,
1
ma
+--m.u(~ + Pl-----mpoUpo+ 2 p~ 2 air flow
,.
-,
1
-~
~
particles ~.,
,
ma
m.uO+ P2 Pa
(4.3)
air flow
,
input
out'put
Rearranging the expression yields l -
@.~-.p~
-
P,
-
mp
(4.4)
P2 - -2 P. ~ U p o ma
Equation (4.4) defines the minimum pressure drop caused by the particles flowing through the accelerating tube. Actually, it is the pressure drop resulting from the momentum transfer between gas flow and particles, i.e., the acceleration of the particles. In the equation, upo is the average velocity of particles at the outlet of the accelerating tube, which depends on the following: the impinging velocity of gas flow u0, the length of the accelerating tube L,~, the density ratio pJp~, and the mean diameter of particles dp; it can be determined with the motion equations below: du p dt
= 0.75CD [
[3 a
ppdp
du p Pa Up ......... 0.75CD[
ppdp
dl
.~
](u 0 -Up) ~
](u 0 -Up)
(4.5)
2
(4.6)
The initial conditions for the motion equations are t=0:
Up=0,
/=0
(4.7)
The drag force coefficient in Eqs. (4.5) and (4.6), CD, depends on the flow regime, as has been described in Chapter 2; the relationships between CD and the Reynolds
94
IMPINGING STREAMS
number Rep for various flow regimes can be found from common handbooks, e.g., Ref. [20] etc., where Rep is defined as
Rep = dpP~alblO--Up[
(4.8)
On the other hand, the collisions of particles on the inside wall of the tube and between particles also result in a certain pressure drop, which is denoted by -Apac,p2. Some researchers let it be a fraction of the pressure drop caused by pure airflow [67]; however, it would be more convenient and reasonable to correlate it with that resulting from the acceleration of the particles, -Apac,pl, because the factors affecting the two kinds of collision are the same as those for-Ap at,p1. In this way, we have
-- Apac,p2 - a ( - A p a c , p l ) -
1
mp
a ~ p a~HpO
2
(4.9)
ma
where a is a proportional factor. If the combined local resistance coefficient is defined as
(ac,p
=
1+ a
(4.10)
then the combined pressure drop caused by the acceleration and the collisions of the particles on the wall and between them can be represented by 1
mp
2
-- Apac,p - - ( A P a c , p l -t- ~Pac,p2 ) - (ac,p ~ P a ~ u p0 Z ma
(4.11)
Thus, the sub-total pressure drop due to the suspension flow passing through the accelerating tube is written as - A p a c = - ( A p a c , a + Apac, p )
(4.12)
4.1.2 Impingement between opposing streams The gas-solid suspensions ejected from the accelerating tube flow towards the center of the contactor and impinge against each other. Between the exits of the accelerating tubes and the impingement zone nothing produces a pressure drop; or, at least, the pressure drop caused by the flows from the exits of the accelerating tubes to the impingement zone is negligible. It is found from experimental measurements that the pressure drop caused by the impingement between the opposing streams is independent of the presence of particles.
HYDRAULIC RESISTANCE OF IMPINGING STREAM DEVICES
95
To determine this pressure drop by the energy balance round the impingement zone is difficult, because the zone has no physical boundary so that the radial velocity of the gas flow just leaving the zone cannot be determined, even approximately. According to visual observations, it may be considered that the impingement plane plays the role of a turning tube of 90 ° in the sense of yielding a hydraulic resistance to the flow. Thus, the pressure drop across the impingement zone can be represented in terms of the local resistance coefficient, ~m, as ,.) --APim -- ~i,n-2-PaL/(~
(4.13)
The value for ~,~ needs to be determined by experimental measurement.
4.1.3 Resistance due to the structure of the IS device Some of the structural factors, such as the changes in cross section area of flow passage and flow direction etc., may also cause pressure losses. Obviously, these factors depend on the specific structure of the device under consideration and vary from device to device. For convenience, and for more generalization, the resistance resulting from all the structural factors is represented in terms of the combined local resistance coefficient, (~, which is also related to the velocity of the gas flow in the accelerating tube, i.e., the impinging velocity, uo. i.e., o -- APds -- (ds ~ Pa/A(~
z
(4.14)
The individual resistance coefficients caused by various structural factors can be determined by the common methods, e.g., those given in Ref. [20], according to the specific structure of the impinging stream equipment employed. For example, for the contactor shown in Fig. 4.1 in the next section, the two structural factors may cause pressure drops: (1) The flow direction turning by 180 ° caused by the upper damper; and (2) The sudden contraction of flow passage at the outlet of the device. The first factor may be neglected because the airflow velocity in that space is very small (<0.05 m.s-~), while the second factor would yield a pressure drop comparable with that due to the impingement between the opposing streams, and can be generally represented by o
-@~o - Co a-p~.~-o z
(4.15)
where u,,, is the velocity of the gas flow in the outlet tube of the contactor and can be calculated easily from the diameter ratio, while neglecting the influence of the small pressure variation on the gas volume, as
96
IMPINGING STREAMS
/'/ao = 2(@oC)2u0
(4.16)
where do is the inside diameter of the outlet tube of the device and dac is the inside diameter of the accelerating tube. The value for (~o is obtained from the handbook [20] to be 1.0. Thus, Eq. (4.14) can be rewritten as -- Apd s = 2 ( ~ )
4 pa u2
(4.17)
Note that Eq. (4.17) is applicable only for the impinging stream contactor shown in Fig. 4.1 (below).
4.1.4 Overall resistance of the IS contactor Combining the results above, the overall resistance of the impinging stream contactor to the streams can be obtained as --@T -" -(Apac,a q- APac,p q- Apim q- APds )
(4.18)
where-APac,a,-Apac,p,-~im and--Apd s are determined by Eqs. (4.2), (4.11), (4.13) and (4.14), respectively; the values for Aa and ~s involved in these equations can be predicted with the common methods according to the specific structure of the IS device under consideration; while the values for ~'ac,p and ~m remain to be determined by experimental measurement.
4.2 EXPERIMENTAL EQUIPMENT AND PROCEDURE 4.2.1 Experimental equipment To focus attention on the aspects most concerned, the experimental impinging stream contactor used for the investigation of the hydraulic resistance employs the basic flow configuration of horizontal two-impinging streams, with feeding solid particles on both sides. The equipment and its main dimensions are illustrated in Fig. 4.1. Two airflows pass through the accelerating tubes at the same velocity and accelerate the particles fed by Feeders 1 and 2, respectively, to a fraction of the airflow velocity, and then the particles-in-air suspension streams are ejected from the accelerating tubes in opposite directions into the impingement zone to impinge against each other. The two accelerating tubes are the same in both diameter and length; while the impinging distance, S, i.e., the distance between Points B and B' in Fig. 4.1, is variable. The
HYDRAULIC RESISTANCE OF IMPINGING STREAM DEVICES
97
horizontal cross section of the upper plexiglass chamber is a square of 0.4x0.4 m, and the bottom is tapered sharply. The accelerating tubes are made of zinc-plated steel, with a relative roughness of g/d,c = 0.0047. Both the screw feeders 1 and 2 are driven by motors with precisely manipulated rotary speeds so that they supply the impinging stream contactor with particles at stable flow rates. Air out /50
Particles in
Damper
Particles in
!
1 I Feeder, 11
]Feeder 21 A
---~I65I--
~m ~ ~,
/' s20/ I
B " B' [] ." [1 =1
Accelerating tube
S~q
520 ~
=[65 l=
Impingement zone
~ 2 1 Unit" mm Particles out
Figure 4.1 Horizontal two-impinging stream contactor for study on the hydraulic resistance.
4.2.2 Experimental procedure The materials used in the experiments are millets and rapeseeds, the properties of which are identical to those listed in Table 3.1 (Chapter 3). All the experiments were carried out at room temperature, and the operating or structural conditions tested ranged as follows: the velocity of the airflow in the accelerating tube u 0 - 9.48-17.36 -1 m.s , particles to gas mass flow rate ratio mp/m.~- 0.556-1.0, and the dimensionless impinging distance S/d,c = 3.0-6.7. The pressure drops between Points A and C, between Points A' and C, between Points A and B, and between Points A' and B', denoted by-ApAo--ApA'c,--ApAB, and --ApA'B', respectively, are measured with inclined U-shape tubes filled with colored kerosene. The average of--ApAB and--APA.B, is taken as the pressure drop through the accelerating pipes, while that o f - A p a c and-APA,c as the overall pressure drop across the impinging stream contactor,-APT. Consequently, there should be --(/~Pim + @ d s ) -- - - @ T -- (--APac)
(4.19)
98
IMPINGING STREAMS
4.3 MAJOR RESULTS FROM THE EXPERIMENTAL STUDY 4.3.1 Basic characteristics of pressure drop distribution Table 4.1 gives some typical data measured for pressure drops. All the results m e a s u r e d exhibit the same characteristics of pressure drop distribution over the contactor as shown in this table. The following characteristics of the pressure drop distribution over the whole impinging stream contactor can be observed from Table 4.1: (1) The majority of the resistance (> 80%) to the streams, whether pure air or particles-in-air suspensions, occurs in the accelerating tubes, i.e.,-Apa c >_ 0.Sx(--ApT); (2) The presence of particles in the suspension streams does not exhibit any substantial influence on both the pressure drops resulting from i m p i n g e m e n t and caused by structural factors of the IS contactor,-Ap~m and--Apds; while (3) The presence of particles results in a significant increase in the pressure drop through the accelerating tubes.
Table 4.1 Data measured for pressure drop distribution (u0= 14.22 m.s -~, S/dac=4.0) Millets
mp/ma
--APT
Rape seeds
--( Apim+ Apd0
--Apac
--APT
--( Apim+ Apd0
-APac
Pa
Pa
%
Pa
%
Pa
Pa
%
Pa
%
0.0
87.36
15.65
17.9
71.71
82.1
83.38
12.32
14.8
71.06
85.2
0.556
1 3 3 . 6 4 13.69
10.2
119.95
89.8
142.70 15.06
10.5
127.64
89.5
0.588
1 4 0 . 8 8 16.82
11.9
124.06
88.1
144.33 15.58
10.8
128.75
89.2
0.625
1 3 9 . 9 6 15.45
ll.0
124.51
89.0
144.00 14.60
10.1
129.40
89.9
0.667
1 4 0 . 4 8 14.34
10.2
126.14
89.8
147.01 14.80
10.1
132.21
89.9
0.714
1 4 4 . 0 7 14.34
10.0
129.73
90.0
150.53 15.06
10.0
135.47
90.0
0.769
1 4 9 . 2 9 14.67
9.8
134.62
90.2
153.66 14.67
9.5
138.99
90.5
0.833
1 5 0 . 7 3 13.76
9.1
136.97
90.9
163.36 16.75
10.2
146.61
89.8
0.909
1 5 2 . 2 2 13.69
9.0
138.53
91.0
167.34 15.12
9.0
152.22
91.0
1.0
162.32
12.3
142.31
87.7
20.01
HYDRAULIC RESISTANCE OF IMPINGING STREAM DEVICES
99
A very important observation obtained from the characteristics described in Items (1) and (3) above is that the major part of the power needed for operating the impinging stream contactor is being consumed in the acceleration of the particles. It can also be seen that an experimental impinging stream contactor is desirable that has no complicated structural factor yielding additional pressure loss so that it enables one to understand fully the pressure drops caused by the acceleration of particles and the impingement between the opposing streams, which are the essential elements in operating solid-gas impinging stream devices.
4.3.2 Resistance of accelerating tubes to pure air flow The experimental results on the influences of the impinging velocity, u0, and the dimensionless impinging distance, S/d,~, on the pressure drop due to pure airflow passing through the accelerating tube are shown in Fig. 4.2. It can be seen that for the dimensionless impinging distance S/dd~, > 4 -Ap,~.~ is kept essentially constant, and S/d~c exhibits a very small effect; while in the range of S/d,,,c < 4.0, -Ap~c,~ increases slightly as S/d,,,c decreases. In addition, the influence of S/d.~ on-Ap~m exhibits the same tendency, as can be seen in Table 4.2. The latter results are consistent with those obtained by Elperin [3] for the axial profile of static pressure in an impinging stream contactor, which demonstrate that the pressure is essentially constant in the range of S/d.~c > 4 but obviously increases once S/ct.,c < 4. It is also found in the experiments that, when S/d,~, is decreased to a level below 4.0, the pressure fluctuation at the outlet of the accelerating tube is strengthened and the impingement zone becomes unstable. Therefore it can be considered that S/d,~ - 4 should be the operational lower limit for gas-solid impinging streams" and thus the pressure drop due to pure airflow passing through the accelerating tube,-At),~,, can be considered to be independent of the impinging distance in the range of practical interest, i.e., S > 4.0. 120 1 O0
~"
80
.~
6o
I
"¢' 0
X~x
X
X-X--
m~,m_,_..mm--l--m__
40 20 I
2.0
I
3.0
I
I
4.0
I
I
5.0
I
I
6.0
I
I
7.0
I
I
I
8.0
S/d,,~ Figure 4.2 Influences of air velocity and impinging distance on -Ap ...... u0, s.m-l" m9.48 xl 1.06; o12.64; /k 14.22" @15.80; D17.30.
100
IMPINGING STREAMS
On the other hand, Fig. 4.2 illustrates that the velocity of the gas flow in the accelerating tube, u0, has an important influence on-Apse,,; and such influence can also be predicted by Eq. (4.2). 0.05 0.04 ,-~= 0.03 0.02 0.01
1.2
1.5
1.8
2.1
2.4
Re ~x10- 4
Figure 4.3 Friction coefficient of airflow through the accelerating tube. [] from Ref. [20]
A measured.
Using Eq. (4.2), the values for the friction coefficient ~ are obtained by regression of the experimental data, as shown in Fig. 4.3; and, for comparison, those obtained from Ref. [20] are also given in the same figure. As can be seen, with respect to the influence of the Reynolds number on the friction coefficient, both kinds of data exhibit the same tendency; while the values obtained from Ref. [20] are higher systematically than those measured by about 50%. The reason for this is still unclear. By a linearized regression of the experimental data, we come to .n
0 000118
2 a - 0.0213 l•e a"
4.20)
The power of 0.000118 in Eq. (4.20) implies that the effect of Rea on ~ is very small, so that ,;Lacan be considered to be independent of Re~ in the range of practical interest, and a constant of 0.0214 is taken for 2a. Since the pressure drop due to pure airflow passing through the accelerating tube occupies only a very small fraction of the total across the contactor; while the values for 2, obtained from the curves of 2~ versus Rea given in Ref. [20] can also be used directly for calculation without significant error.
4.3.3 Pressure drop due to acceleration and collisions of particles The pressure drop due to acceleration and collisions of the particles on the walls and between them,-Apac,p, can be calculated with Eq. (4.12) according to the data
HYDRAULIC RESISTANCE OF IMr'[ !'q(;ING STREAM DEVICES
101
measured for the combined pressure drop through the accelerating tube, -Apse. It is found from the measured results that both the velocity of particles at the outlet of the accelerating tubes, Up0, and the particles to gas mass flow rate ratio, mp/m~, have essential influences on this pressure drop. The measured data are partly shown in Fig. 4.4.
120 Particles m p/m ~ Millets 0.588 2 Millet,, 0.769 3 Rape seeds 0.588 4 Rape seeds 0.769 !
100
80 & 60 40 20 0
3.0
I
I
I
I
3.5
4.{)
4.5
5.0 /,,to, m - S
I
I
I
5.5
6.0
6.5
7.0
-I
Figure 4.4 Influences of out-velocity of particles and mass flow rate ratio on Apse,p (S/d=4.0).
10
U
I
,
0.5
i
,
i
,
0.6
i
I
I
I
I
0.7
I
I
I
I
I
I
I
0.8
i
,
i
0.9
,
i
,
i
i
1.0
i
,
i
,
1.1
m p/m a
Figure 4.5 Part of the data measured for the local resistance coefficient ~,p (Upo=3.51-6.42 re.s-l); - - averaged over 490 sets of data: + m O r a p e seeds, S/d=4.0; ~[]/~x.omillets, S/d=5.0.
102
IMPINGING STREAMS
The local resistance coefficient (ac,p is calculated for a total of 490 sets of measured data, and the results are given in Fig. 4.5. It is clear from Fig. 4.5 that the data for (ac,p are considerably concentrated: over 85% of the values are in the range 4.3 to 6.2, and the average value is 5.34. This fact indicates that (ac,p can be considered to remain essentially constant and that, consequentially, Eq. (4.11) describes well the pressure drop behavior due to the acceleration and collisions of the particles. Therefore the combined consideration of the two kinds of collisions, i.e., the collisions of particles on the wall and between particles, is reasonable and feasible. It should be noted that, as described by Eqs. (4.5) and (4.6), for a given impinging velocity u0, the velocity of particles at the outlet of the accelerating tube depends on the length of the tube Lac, the particles to gas mass flow rate ratio mp/ma, and the mean diameter of particles dp. Therefore Eq. (4.11) actually involves all the factors affecting-Apac,p. This may account for the phenomenon that no difference between the values for the local resistance coefficient (~c,p measured for millets and rapeseeds, respectively, has been observed. Considering such general applicability, the equation below is recommended for the prediction of the pressure drop due to the movement of particles, including acceleration and collisions:
-- APac, p -
,mp, 2 2.67Pa I,~)UpO
(4.21)
ma
4.3.4 Resistance due to structure of the device As described in Section 4.2, the sub-pressure drop measured between Points B and C or between Points B' and C, -ApBc or -ApB,c, is the sum of the pressure drop across the impingement zone,-Apim, and the structural resistance-Apds, but not any individual one of them. Such a measuring arrangement is not only because the pressure fluctuation frequently occurs round the impingement zone, making accurate measurement difficult, but also because both sub-pressure drops are too small (< 10 Pa) to be measured separately. To obtain the values for these sub-pressure drops, Eq. (4.17) can be specified for the impinging stream contactor studied here as -Apds - 2 (
0.02116 4 0.05
)PaU2=O.O642Pa u2
(4.22)
Thus, for each experimental run, the pressure drop across the impingement zone, -APim, can be obtained from the average value of measured-APBc and-Ap~,c and the value of-Apds calculated with Eq. (4.22) as its indirectly measured value:
--~Pim
1
- -~ [(--@BC) + (-APB'c)I - (-APds)
(4.23)
HYDRAULIC RESISTANCE OF IMPINGING STREAM DEVICES
103
Table 4.2
Typical data measured indirectly for-Api mand (ira
U0, m-s
S/d,a~.= 6.0
-1
S/d~c = 6.7
-Api .... Pa
(,m
-Apim, Pa
~m
9.48
4.90
0.091
4.30
0.079
11.06
6.84
0.093
6.81
0.092
12.64
7.47
0.081
7.48
0.081
14.22
10.76
0.088
11.09
0.091
15.80
13.38
0.089
12.85
0.085
17.36
15.95
0.088
15.12
0.083
By way of illustration, two sets of indirectly measured data for the pressure drop across the impingement zone are listed in Table 4.2. As can be seen from the table, the mass flow rate ratio mp/ma has no effect on -Apim, and the equivalent local resistance coefficient (,m calculated with Eq. (4.13) is essentially kept constant. This implies that the pressure drop across the impingement zone is independent of the presence of particles. The value for (,m averaged over a total of 490 sets of data is equal to 0.096. So, the pressure drop across the impingement zone can be calculated with the relationship below: 9
- Ap~m - 0.048p~u~
(4.24)
4.3.5 Model for the overall pressure drop Combining the results of the theoretical analysis and the experimental investigation above, the overall pressure drop across the whole impinging stream contactor can be represented by rap. 2 2 1 2 La~ p.~uo ~-2.67p,,, ~--)Up0 + 0.048pau 0 +-~ (dsPaU0
__ A p T _ ~,~
" d~,~.
2
Substituting the known factors into the expression above and rearranging the result in the following empirical equation"
-
ApT
-
(0.048 +
~.~ Lac 2 .mp. 2 " + 0 . 5 ( d s )Pab/0 -Jr-2.67p,, ¢--)UpO 2 d~,c ma
(4.25)
104
IMPINGING STREAMS
Equation (4.25) is the overall pressure drop model for the impinging stream contactor. In this equation, the first term on the right-hand side reflects the effect of the gas flow; while the second term reflects that of the particles. It is applicable for calculating the hydraulic resistance of common gas-solid horizontal two-impinging stream equipment with a dimensionless impinging distance of S/d~c > 4. For a given length of the accelerating tube, Lac, density and mean diameter of particles, pp and dp, the out-velocity of particles Up0 in this equation is related to the impinging velocity u0, and can be calculated by integrating Eq. (4.6). In the model equation above, in addition to the dimensions of the device and the properties of the solid particles and the gas involved, only two variable parameters are included: the friction factor 2~ and the structural resistance coefficient ~s. 2~ depends mainly on the roughness of the inside wall of the accelerating tube, and its value can be obtained from common handbooks of chemical engineering. For the zinc-plated tubes used in the experimental study carried out by Wu and Wu [66], the value of 0.0214 can be taken for ~. While the structural resistance coefficient ~s is related closely to the structure of the device to be considered, and its calculation must be aimed at the specific equipment; but all the methods for calculation involved are general. Therefore the empirical model, Eq. (4.25), is universally applicable. A comparison between the data measured for the overall pressure drop across the impinging stream contactor and the corresponding values calculated with Eq. (4.25) is illustrated in Fig. 4.6. Good agreement between the results measured and calculated can be observed clearly, suggesting that the total pressure drop model established is reasonable and feasible for application. In addition, it has the advantages of universal applicability and convenience in calculation. 250
~, g,
200
# I
150
0
1 O0
;>
50 I
0
I
I
I
I
50
I
I
I
I
I
100
i
i
i
i
i
i
i
150
i
i
I
200
I
I
i
i
250
Calculated-ApT , Pa
Figure 4.6 Comparison between measured data and results calculated with Eq. (4.25) for the overall pressure drop cross the IS contactor, u0, m.s-1" n9.48 @ 11.06 e 12.64 A 14.22 x15.60 + 17.36.
HYDRAULIC RESISTANCE OF IMPINGING STREAM DEVICES
105
It is clear from Eq. (4.25) that the velocity of the airflow in the accelerating tube, i.e., the impinging velocity u0, has a fundamental influence on the hydraulic resistance of the impinging stream equipment: the overall pressure drop, -APT, increases in the form of an exponential function as u0 increases, resulting in rapidly increased power consumption. On the other hand, at lower impinging velocity the impingement between the opposing streams could not be efficient so that it is difficult to achieve the goal of enhancing transfer between phases. Therefore the appropriate selection of the value for the impinging velocity is very important in the design and operation of an impinging stream device.
4.4 EVALUATION OF POWER CONSUMPTION AND DISCUSSIONS RELATED TO APPLICATION It has been mentioned before that power consumption in the operation of impinging stream equipment is a problem of great concern, because IS involves gas flow at high velocity. The first conclusion from the investigation on the hydraulic resistance of the impinging stream contactor shown in Fig. 4.1 is that the resistance of the gascontinuous impinging stream device is not large, and, consequently, the power consumption is acceptable, provided the structure of the device is reasonably designed and the material impinging streams being applied for is properly selected, not too heavy and not too large. It is of interest to compare an impinging stream contactor with another type of processing device, although such a comparison may not be easy because of the lack of operation data. With regard to conveying of solid particles by pneumatics, the device most similar to impinging stream equipment is the pneumatic flash dryer, also simply called a flash dryer. Table 4.3 gives a relative comparison between the two kinds of devices, with particles of dp= 0.001 m and pp = 1500 kg.m -3 as the superficial material being processed. As the basis, all the parameters related to the impinging stream contactor are set to be unity. It can be seen in the table that, in common cases, impinging the stream contactor exhibits a much higher efficiency than the flash dryer, occupying less space and have lower power consumption. Generally it can be considered that power consumption should not be a problem in the application of gas-continuous impinging stream devices. On the resistance constitution of the equipment system, the major conclusions that can be drawn from the investigation are: (1) Where millets or rapeseeds are the material to be processed, the power for the operation of the impinging stream contactor is mainly (>80%) consumed in the acceleration of particles; (2) The pressure loss due to the impingement between the opposing streams is independent of the presence of solid particles.
106
IMPINGING STREAMS Table 4.3
A relative comparison between flash dryer and IS contactor dealing with the same amount of solid material Type of equipment
Impinging stream contactor
Flash dryer
1
0.3-0.5
1 (-*2u0)
0.1-0.2 (=ut
Average transfer coefficient
1
0.1-0.2
Length of pneumatic tube
1
510
Effective average residence time
1
3-5
Total pressure drop
1
1-3
Specific power consumption
1
1-3
Operating air velocity Maximum relative velocity between phases
The first conclusion above, i.e., the power being consumed mainly in the acceleration of particles, is of significant interest and so must be fully considered when selecting the materials that the IS is being applied to. It should be noted that in the experiments yielding this conclusion, the sizes of particles, dp, range from 1.6 to 1.7 mm, and the densities of particles, pp, from 1100 to 1200 kg.m-1, i.e., the particles are not very large and relatively slight. Certainly, for smaller and slighter particles, the hydraulic resistance of the IS equipment is never a problem. However, if one uses impinging streams for processing large and heavy particles, the resistance of the system must be greatly increased, because a very high velocity gas flow has to be used for the acceleration of particles. If smaller gas flow velocity was employed, particles would not be fully accelerated so that the enhancement of transfer between phases becomes very weak and the application of impinging streams becomes useless. Therefore, from the point of view of power consumption only, the method of gas-continuous impinging streams is not suitable for processing large, heavy granular materials. The discussion above on the feasibility of IS application only relates to power consumption. In practice, power consumption is not the only factor that has to be considered in the selection of a target system for IS application. As described in the Introduction, in the development of application, the selection of the target system must be based on a complete understanding of the properties of IS, bringing the advantages of IS into play and avoiding its disadvantages; at the same time, one must pay attention to solving the related engineering problems. If this is done, success of application can be expected.
-5INFLUENCE OF IMPINGING STREAMS ON DISPERSITY OF LIQUIDS
In the previous chapters, the materials in the dispersed phases of impinging streams discussed are essentially solid particles, although a few aspects of liquid as dispersed phase were mentioned. Since liquids and solids have densities of the same order of magnitude, quite different from those of gases, the analysis and conclusions described in those chapters, including enhancing transfer between phases, the motion of particles, the residence time and its distribution, and the hydraulic resistance and the related problems, etc., are, in principle, also applicable for the occasions where, instead of solid, liquid is in the dispersed phase without significant deviation. On the other hand, the assembly condition and the inter-molecule force of liquids are quite different from those of solids. In gas-continuous impinging streams with solid and liquid as the dispersed phases, respectively, some different phenomena would occur, which may affect the perlbrmance of impinging streams and thus are of concern. This chapter discusses the problem of liquid dispersion in gas-continuous impinging streams, which is related to the topic mentioned above, and introduces the related results of investigations.
5.1 STATEMENT OF THE PROBLEM Gas-continuous impinging streams with a liquid as the dispersed phase has wide application, such as in the combustion of liquid fuel droplets, absorption, water-spray cooling of air, etc. [9]. In such systems the dispersity of liquids plays a very important role affecting heat and mass transfer rates, because it influences both the interface area and the mean transfer coefficient. Wu et al. [68] investigated the influence of impinging streams on the dispersity of liquid. In a gas-continuous impinging stream device with liquid as the dispersed phase, the liquid is usually atomized into fine droplets with nozzles of an appropriate type, and ejected into gas flows to form droplets-in-gas suspensions before impingement. This can be called the Primary Atomization, and it defines the primary dispersity of liquids. The mechanism of primary atomization and the methods for predicting size distribution (SD) and mean diameter (MD) of the sprayed droplets have been widely reported and some sources of references may be found, e.g., in Ref. [691.
107
108
IMPINGING STREAMS
On the other hand, during impingement between the opposing suspension streams, the following factors may cause variation in the dispersity of the liquid: (1) Collisions between droplets cause coalescence and/or break-up of droplets, the latter being called Secondary Atomization, or, simply, re-atomization; (2) The strong friction force between gas flow and liquid droplets due to very high relative velocity, which can theoretically be as high as twice that of the gas flow [5], i.e., 2u0, may also result in reatomization of droplets. It is obvious that re-atomization yields decrease the mean diameter of the liquid droplets and thus an increased interface area; at the same time, it results in reduced average transfer coefficients, because heat and mass transfer coefficients between gas flow and particle or droplet are in positive correlation with the diameter of the particle or droplet, while coalescence of droplets yields influences opposite to those described above. In their investigation on the absorption of CO2 into NaOH solution, Herskowits et al. [59, 60] determined theoretically the total interface areas and the mass transfer coefficients by comparing the absorption rates with and without reaction in liquid, employing the expression for the enhancement factor due to chemical reaction of second-order kinetics presented by Danckwerts [70]. Based on the modified Boltzmann kinetic equation presented by Culick [71 ], Kitron et al. [72] developed the Monte-Carlo simulation method (MCS) and used it to calculate the droplet sizes, their positions and velocity profile in impinging streams. The interaction between gas flow and droplets, the friction force of gas, non-uniformity of the flow field, and the binary collisions of droplets yielding coalescence and reatomization were taken into account in their calculation. Also, Kitron et al. simulated the absorption processes, based on the investigation by Herskowits et al., in order to understand the influence of collisions on the total interface area, and, on the basis of Faeth's investigation [73], analyzed the influences of the factors similar to those above on the rate of pentane combustion in an impinging stream combustor, and obtained a number of curves describing the time-averaged overall vaporization rates for the cases of flow without collision, with only coalescence of droplets, and with both coalescence and re-atomization, respectively. Unfortunately, few experimental investigations have been reported to date, so that the accuracy of the theoretical results above cannot be verified for lack of experimental evidence. In addition, although many excessive assumptions were made by Kitron et al., the calculations are very troublesome and cumbersome. In order to obtain experimental evidence for understanding the influence of the impingement between the opposing droplets-in-gas suspension streams on the dispersity of liquid, and also to get some practically applicable information for designing and operating impinging stream devices, the experimental investigation described below was carried out.
INFLUENCE OF IMPINGING STREAMS ON DISPERSITY OF LIQUIDS
109
5.2 EXPERIMENTAL EQUIPMENT AND PROCEDURE 5.2.1 Impinging stream device To make sampling easier, the open horizontal co-axial two-impinging stream device is employed in the experiments, as shown in Fig. 5.1.
A ....
1
-...-..:.~.~.~:~.~.S.~.?.....-....
J, 3
3
6
5
6
Figure 5.1 Impinging stream device for the measurement of droplet size distribution, l critical nozzle; 2, 3 rotameter; 4 frame; 5 guider; 6 apron; A-A impingement plane.
Two identical critical nozzles 1 of the Caldyn CSL2 type [59, 60] (the structure of which will be described in Chapter 7 on absorption) are mounted co-axially in opposite directions onto Frames 4 supported by Guider 5, on which the frames can move axially to change the distance between the nozzles, i.e., the impinging distance S. Two streams of water, the process liquid, flow through Rotameters 2 and enter the nozzles on the two sides, respectively, where water is mixed with the compressed air from Rotameters 3 and is atomized to form droplets-in-air suspensions. The two opposing suspension streams ejected from the nozzles flow in opposite directions and impinge against each other on the impingement plane A-A. To avoid possible turbulence of the surrounding air affecting the impinging streams, an Apron 6 is placed just before each frame, and each of the aprons has an opening of a size that just permits the head of the nozzle to cross the apron.
110
IMPINGING STREAMS
5.2.2 Method for measurement of droplet size distribution Various techniques, such as melted wax cooling [74], spray freezing [75], direct high speed photography [76], and Laser Dopler Analysis (LDA) system [77, 78] etc., have been used to measure the sizes and/or size distribution of spray droplets, and each has its own advantages and disadvantages or difficulties for practical application. This investigation was carried out as early as 1995-1996. Mainly because of the lack of advanced laboratory instrumentation, such as LDA, at that time, the procedure of slide-sampling, micro-photographing, and imaging-analysis was used for the measurement of droplet size distribution. On the other hand, slide-sampling, one of the traditional direct methods, has the advantage of a wide measurable range of droplet sizes from 5 to nearly 1000 gm, and thus is suitable for the cases under consideration. Glass slides, size 20x80 ram, are coated with methyl silicone oil to prevent deformation of the collected droplets. In order to minimize the influence of evaporation of water from the droplets on the reality of the data, immediately after each sampling the slide carrying the collected droplets is photographed with an XTT-XF-DF microphotography system. The photos are then analyzed with an imaging analysis system of Type IBAS I-II, produced by Contron Co., Germany, to determine the size distribution and the mean diameter of the corresponding spray droplets.
5.2.3 Arrangement of sampling It is actually very difficult to obtain representative and reproducible data for the size distribution of sprayed droplets, no matter what kind of method is employed for measurement, because the dispersion of liquid by atomization, including re-atomization in impinging streams, is highly random. In the study the following factors are considered carefully in the arrangement of sampling for representative samples and thus statistically trustworthy results: (1) Very fine droplets tend to follow the streamlines. To account for this fact, any possible disturbance of the sampling slides to the flow field must be minimized to prevent the fine droplets escaping from the slides. For this purpose, the number and positions of the sampling slides in each run are arranged according to the moving tendency of the droplets. (2) In the case of multipoint sampling from the same spray, each slide must sample droplets in a suitable amount to avoid failure both in representativity due to fewer droplets and overlapping of the droplets due to an excessive number of droplets, leading to errors in the image analysis. So, the time of exposing each slide to the spray was rigorously controlled at 1-2 s. As stated above, the goal of the investigation is to understand the influence of the impingement on the dispersity of liquid. Therefore it is necessary to measure size distribution and the mean diameter of droplets both before and after the impingement of the opposing streams. In the former case one of the suspension streams is turned off and sampling is made with a single slide at a distance of 0.4 m from the working
INFLUENCE OF IMPINGING STREAMS ON DISPERSITY OF LIQUIDS
111
nozzle, where the suspension jet begins to bend downwards, and below the jet axis by 0.15 m, as shown in Fig. 5.2(a). In the latter case both the nozzles are kept on, and samplings are made for the same spray with three slides in the vertical plane crossing the midpoint between the two nozzles, and the combination of data measured from all the three slides is taken as the results for the run. The three sampling points are so arranged that Slides 2 and 3 are above and below the jets, respectively, and just round the dense region of the impingement zone, while Slide 1 is below Slide 2, as shown in Fig. 2(b), to account for the effect of gravity. In fact it is observed in the experiments that the number of droplets leaving the impingement zone downwards is significantly larger than that upwards, and therefore more droplets have to be sampled under the jets. Since all the three sampling points are far from the center of the impingement zone, the influence of the slides on the flow field may be negligible. 3
nozzle
/\
/1/" \\'\
!
nozzle
400 mm
~-]
slide
2 ",~( /i! i'!
150mm
slide 1 i'i _ _ ~ _ _ (a) Without impingement
(b) After impingement
Figure 5.2 Sampling arrangement. 1, 2, 3: sampling slides.
5.3 MAJOR RESULTS OF THE INVESTIGATION 5.3.1 Size distribution of droplets The measured droplet size distribution is represented in terms of the number frequency function of droplets, ON, which is defined as N(D)
ON -- I(~ N ( D ) d D
N(Di)
_
- Y~ N ( D i )AD i
(5.1)
i
The results of regression analysis of the measured data indicate that the droplet size distributions both before and after the impingement can be expressed by the following equation:
112
IMPINGING STREAMS 0y = exp[-C1 - C2 (In D )3 DAM
C3( D )3 DAM
where Ci (i-1, 2,) are adjustable parameters; while of droplets: DAM -- IO DON(D)dD -
C4 ( D )1.5]
(5.2)
DAM
DAMis the arithmetic
mean diameter
~.DiN(Di)ADi
(5.3)
ZN(Di)AD i
Two sets of typical data experimentally measured are illustrated in Figs. 5.3(a) and (b), in which the curves represent the results calculated by Eq. (5.2) with properly regressed parameters Ci. It should be noted that Ua in the figures is the airflow velocity at the exits of the nozzles; which is quite different from the impinging velocity, u0, in both the nature and the order of magnitude. It can be seen from the figures that the size distributions of spray droplets become narrower after impingement, or, in other words, the droplet sizes become more uniform than before. Such a variation is observed in most of the runs, although some exceptions also appeared (about 10% in all the runs). As is well known, the uniformity of droplet sizes, or inversely, the scattering of the size distribution can be expressed with the parameter "Standard Deviation", or, which is the defined as [Z o =
V
N i ( D i - DAM ~Ni
)2
(5.4)
The larger the value for ~ the greater the scattering of the size distribution, and, correspondingly, the poorer the uniformity of the droplet sizes. Part of the data measured at various air-to-liquid mass flow ratio, ma/mL, are shown in Table 5.1. It can be seen that, normally, the impingement between opposing droplets-in-gas suspension streams makes cr smaller. Only those obtained at ma/mL= 1.34, in the third column of the table, exhibit an exceptional tendency: the values for cr before and after impingement are very close to each other. Table 5.1
Standard deviation of size distribution before and after impingement ma ]mL
0.269
1.340
Case
3.410
5.680
7.848
~ gm
Before impingement
53.05
59.14
42.10
40.20
27.57
After impingement
49.14
59.69
26.77
14.74
16.93
113
INFLUENCE OF IMPINGING STREAMS ON DISPERSITY OF LIQUIDS
0.04
0.03
7
El
zzk
~
2
0.02
0.01
0.0() 0
50
100
150
200
250
300
D, btm (a) u,~=301.6 m.s
~, m~/m,= 0.295. A, 1 without impingement (S=oo); [], 2 after impingement (S=0.4 m)
0.04
7
0.03 E
:zk
~
0.02
0.01 [
0.00
._!
L
50
100
150 D, btm
200
250
300
(b) u~=293.2 m.s -~, mL/m.~,=0.173. A, 1 without impingement (S=oo); U1, 2 after impingement (S=0.09 m) Figure 5.3 Variation of droplet size distribution before and after impingement of suspension streams.
To an extent, the intensity of impingement is dependent on the impinging distance, S, and smaller S implies stronger impingement. Therefore it is also of interest to understand the effect of the impinging distance on the size distribution of droplets. The
114
IMPINGING STREAMS
related results are listed in Table 5.2. As can be seen, mostly, the standard deviation cr decreases as the impinging distance S decreases, although some exceptional data appeared, (the fourth row in Table 5.2). If both the strong randomness of liquid atomization and the non-avoidable difficulty in obtaining representative samples are taken into account, these exceptional data may be ignored, and thus it may be approximately concluded that the impingement between the opposed droplets-in-gas suspension streams plays the role of making the sizes of the droplets uniform, and stronger impingement favors the uniformity of droplet sizes. Table 5.2
Influence of impinging distance on the standard deviation of droplet size distribution S, In
0.25
0.15
0.09
(y, pm
ma[mL
-,,1.0
59.01
32.94
~2.0
20.65
30.13
7.34
37.37
16.90
27.77
The uniformization of droplet sizes implies that, during impingement the larger droplets tend to break up, i.e., re-atomize; while the finer ones tend to coalescence. This is theoretically reasonable. Within the range of smaller distance from the exit of the nozzle to the impingement plane, all the droplets can be considered to move at essentially the same velocity. The larger the droplet, the greater is its dynamic energy. So, shattering collisions tend to occur with large droplets; while bouncing collisions tend to occur with smaller ones, as was observed by Ashgriz and Givi [79] in the case of spray cooling of water. In shattering collisions numerous tiny droplets are expelled radially from the periphery of the interacting droplets, reducing the sizes of the original droplets. On the other hand, the surface tension acting on per unit mass of the larger droplets is smaller, so that the larger droplets are more easily re-atomized than the smaller ones at high relative velocity between gas flow and droplets. Such a uniformization of droplet sizes is a phenomenon of interest for practical application. It suggests that, in gas-liquid processes, the amount of fine droplets entrained by gas flow may be reduced so that the processes may carry on more smoothly. Another interesting result observed is that in the case without impingement, i.e., only primary atomization occurred, the scattering of the droplet size distribution, or, decreases, and, consequentially, the droplet sizes become more uniform, as the gas to liquid mass flow rate ratio, mJmc, increases, as can be seen in the third row of Table 5.1.
INFLUENCE OF IMPINGING STREAMS ON DISPERSITY OF LIQUIDS
115
5.3.2 Mean diameter of droplets For convenience of application, all the experimental data are interrelated in terms of the Sauter mean diameter, i.e., the volume-surface mean diameter, D32, which is defined as Ni Di3AD i D~ = -) ~2 Ni Di- AOi
(5.5)
For the systems of water-air and water-CO2, the Sauter mean diameters of spray droplets and the corresponding size distributions before and after the impingement were measured under a total of 37 sets of operating conditions in the ranges of the air velocity u~,- 110-310 m.s -~, corresponding to R e o - 15000 -70000, the mass flow rate ratio m J m L 0.2-15.0, and the impinging distance S = 0.09-0.4 m. In the data interpretation, various regression equations were tested, and finally the following expression was found to fit the data best and to describe the influences of various factors better: D~2 - K(mL/m.,, )" Re[~ ([Ltm)
(5.6)
where K, a and b are adjustable parameters and Re., is the Reynolds number of airflow at the exit of the nozzle" Re., - d°u"'P'a
(5.7)
tl
where do is the inside diameter of the exiting hole of the nozzle. If the air stream through the nozzle is assumed to be adiabatic flow and the air velocity at the inlet of the nozzle is negligible in comparison with that at the exit, then the air velocity in Eq. (5.7), u,, can be calculated by [69]
u., -
(2
1000 k P2 /,-1 gcRT"l MA k ~ i [ 1 - ( PI )t )/k
(5.8)
where the subscripts 1 and 2 denote the states of airflow at the inlet and the exit of the nozzle, respectively. The applicable values for u~, depend on the pressure ratio P2/P~ and the critical pressure ratio ~ * / P i " the latter is represented by P:,~: 2 k/(k-i) -~-=( ) P, k+l
(5.9)
116
IMPINGING STREAMS
where k is the adiabatic coefficient of the atomizing gas. The following two situations are possible: (1) If P2/PI > P2/1:'1 , u~ can be calculated by Eq. (5.8); and (2) If PffP1 < P 2 / P l , ua should be equal to the local sonic velocity, Uc, which is written as -
(5.~o)
/ ooO gcRr :/M
where MA is molecular mass of the gas, air in this investigation. A quasi-linearized regression is made for the experimental data with Eq. (5.6), which yields K - 3200,
a - 0.09,
b - -0.32
(5.11)
with a complex correlation coefficient of r = 0.7275, which is much greater than the acceptable minimum value, 0.418, for the confidence degree of 1%. A comparison between the measured data and the results calculated by Eq. (5.6), with the values obtained by the regression for the parameters involved, is illustrated in Fig. 5.4. The standard deviation of the calculation is SD=21.06 gm. If the intrinsic difficulties in the measurement of the spray droplet sizes mentioned above are taken into account, then the results shown in Fig. 5.4 indicate that Eq. (5.6) is acceptable for fitting experimental data. The following can be seen from the regressive equation and the results shown in Fig. 5.4: 190 170 150 130
[]
[]
~D
110 ~9
I
90
A
70 50
50
70
90
110
130
150
170
190
Measured D32, m Figure 5.4 Comparison between Sauter mean diameters measured and calculated by Eq. (5.6). A A before impingement; om after impingement; oA water-air system; • A water-CO2 system.
INFLUENCE OF IMPINGING STREAMS ON DISPERSITY OF LIQUIDS
117
(1) The impinging distance, S, is not involved in the regressive equation, Eq. (5.6). Actually, in the interpretation of data, various parameters containing S, such as S itself, do~S, and (l+d,,/S) etc, were tested by introducing them into possible equations. However, no or very weak influences of them on the Sauter mean diameter of droplets, Ds:, were found in every case. Therefore S is cancelled from the final regressive equation. As mentioned above, parameter S affects the intensity of the impingement between the opposing streams to an extent, and S = oo implies no impingement. So, the fact that S does not appear in the regressive equation suggests that the impingement of the suspension streams does not affect the Sauter mean diameter of spray droplets. (2) The negative exponent on the Reynolds number o f - 0 . 3 2 indicates that Reo has a medium effect on the mean diameter. This result is about in accordance with those obtained by Gretzinger and Marshall [80] in their investigation on the external mixing pneumatic spray nozzles. (3) It is observed in the investigation that the liquid to gas mass flow rate ratio, mL/m.,, has little influence on the mean diameter of droplets, as indicated by the exponent of 0.09 on the ratio. In comparison with the results obtained by other researchers [80, 81], the influence exhibited in this investigation is much smaller. The structure of the nozzles of the type Caldyn CSL2 used in the present study being quite different from those used by the mentioned researchers may be the major reason for the difference described above. It should be noted that, generally, the properties of liquid should affect the mean diameter of spray droplets to some extent, both before and after the impingement. In the investigation on the dispersity of liquid in impinging streams described here, however, only water was tested as a process liquid; while other liquids were not. This remains to be studied further.
5.4 CONCLUDING REMARKS Gas-continuous impinging streams have a number of important applications for gasliquid systems, such as combustion of liquid fuel, absorption, etc. All these applications involve transfer processes between phases. A liquid is very close to a solid in density, and thus, as a dispersed phase, would exhibit a number of behaviors in impinging streams similar to those of a solid, such as in the residence time distribution and in the hydraulic resistance, etc. However, because of the difference in their assembly conditions, liquid and solid, as the dispersed phases, have different effects on the performances of gas-continuous impinging streams. From the aspect of macro behaviors, the most important difference is that coalescence and/or breaking up (reatomization) would occur with liquid droplets in impinging streams, but not with solid particles. As a result, the total interface area and the transfer coefficients may change, yielding effects on the transfer processes between phases. Therefore the influence of impinging streams on the dispersity of liquid is a problem that needs to be considered.
118
IMPINGING STREAMS
The author of this book investigated experimentally the influence of the impingement between two opposing liquid droplets-in-gas suspension streams on the dispersity of the liquid in an open device of horizontal two-impinging streams, with the internal mixing nozzles of Type Caldyn CSL2 as the atomizers and water-air and water-CO2 systems as the targets, and with the slide-sampling, micro-photographing and imaging analysis procedure The following results of interest were obtained: (1) The impingement between the two opposing suspension streams makes the sizes of the spray droplets uniform to an extent, yielding narrower size distribution. More intensive impingement favors the uniformization of droplet sizes more effectively. (2) Essentially, the impingement between the two opposing droplets-in-gas suspension streams does not change the mean diameter of the droplets. (3) The Reynolds number of gas flow, Re~, exhibits a medium influence on the Sauter mean diameter of the droplets, both before and after the impingement; while the liquid to gas mass flow rate ratio, mL/ma, affects the same amount very weakly. (4) The Sauter mean diameters of the spray droplets, D32, both before and after the impingement can be correlated and predicted with Eq. (5.6), which gives reasonable and acceptable fitting of the experimental data.
-6IMPINGING STREAM DRYING
6.1 INTRODUCTION In the previous chapters the essential principles and characteristics of gas-continuous impinging streams (GIS) were discussed; while this and subsequent chapters in Part I will focus on the research and development of applied technologies. The contents have been chosen to be as valuable as possible for practical application, while successful or unsuccessful experiences included can be used for reference. As mentioned, like any other technical method, the method of impinging streams (IS) cannot be a universal tool. Oil one hand, IS has the outstanding advantage of significantly enhancing heat and mass transfer between phases; while on the other, it also has its intrinsic faults. From the discussions in the previous chapters, the essential characteristics of gas-continuous impinging streams can be summarized briefly as follows:
(1) The special flow configuration is suitable for processing two-phase or multiphase systems with a gas as the continuous phase;
(2) Its effect on enhancing heat and mass transfer between phases is very significant; (3) There exists strong mixing in the major active region for transfer, i.e., the impingement zone;
(4) The residence time of materials in either the continuous or dispersed phase in the active region is very short, about 1 s only; and (5) In comparison with other existing technologies mainly used for transfer processes, such as plate column etc'., the flow configuration in the devices of impinging streams is much more complex. This increases the difficulty of putting impinging streams into practice, especially in arranging countercurrent multistage systems. Items (1) and (2) above are the most obvious advantages of gas-continuous impinging streams. Since transfer between phases is a problem often encountered in multiphase systems, these advantages provide impinging streams with a wide sphere of application in the chemical, petrochemical, and other processing industries. The characteristic described in item (3) may appear as an advantage on certain occasions, while strong mixing implies serious back-mixing, leading to lower efficiency in many processes.
119
120
IMPINGING STREAMS
The very short residence time of the materials being processed in the major active region is a serious disadvantage or fault of impinging stream technology. In various processing industries, some processes can be carried out within a short time; and, under conditions of strongly enhanced heat and mass transfer, the time interval needed may be shortened further and thus the processes can be carried out with impinging stream technology, yielding significant benefits in reducing the volume of the equipment and reducing power consumption. However, many other processes need considerably longer time, e.g., several tens of minutes, even if under conditions of significantly enhanced transfer, in order to achieve the required processing degree, e.g., certain reaction conversion, absorption or moisture removal efficiency etc., to obtain the specified technological-economic indexes. It is evident that any effective development of impinging streams application must be based on an understanding of its properties. The industrial application of impinging streams can only be developed further when researchers have an in-depth knowledge of the characteristics of IS, including its advantages and limitations, and, on this basis, can bring its superiorities into play and avoid its disadvantages and faults. On the other hand, the industrial application of a new technological method must face many practical engineering problems, in addition to those of the method itself, which must be solved appropriately with the ideas of system engineering. Otherwise, successful application is difficult, even if the technology is very nice. Drying is a typical process involving parallel heat and mass transfer, and is one of the most suitable areas for the application of gas-continuous impinging streams. This has long been one of the hot spots of investigation and so many studies on impinging stream drying have been carried out since the early 1970s [21]. It is true that drying is the unit operation process most involved in the research and development of impinging streams. Many industries involve drying and so many materials, either final or intermediate products, need to be dried. Since the various materials have quite different properties, the impinging stream dryers that have been reported are also of many different types. Although, great efforts have been made in the past over 30 years and more, unfortunately no impinging stream dryer has yet been applied industrially. The lack of in-depth understanding of the properties of impinging streams and the number of unsolved engineering problems encountered during development may account for this slow progress in the industrial application of impinging stream drying. Assimilating positive and negative the experiences obtained in the past, the author of this book has developed the "Circulative Impinging Stream Dryer", an IS device suitable for powdery and/or granular materials [11, 82]. A test with quasi-industrial equipment on a scale of 1000 tones per year has exhibited good performance, and practical application may be expected in the near future. In this chapter, the research and development that has been carried out and the proposed impinging stream drying technologies and devices will be introduced first. By summarizing and analyzing those works, and following the train of thought described above, one may gain some useful understanding. Later, the circulative impinging
IMPINGING STREAM DRYING
121
stream dryer will be discussed in detail, including the basic idea for the dryer design, its working principles, and the major experimental results, etc.
6.2 EARLIER RESEARCH AND DEVELOPMENT Since the 1970s, researchers from different countries have proposed a number of impinging stream drying technologies for drying various materials and this will be discussed briefly below. It should be noted that there are two other kinds of drying technologies incorporating the word "impingement", i.e., "Impingement drying" and "Jet impingement drying". In the tormer, the material to be dried, a solution or suspension, is coated, by ejection and impacting action, on a certain surface vertical to the jet axis and rotating at high speed, where the drying is carried out [83]. In the latter, the jet of drying medium, hot air or super heated steam, impacts continuously the surface of a thin sheet material and dries it. This method is applicable for the drying of paper, tissue, textiles, and films, etc. [62, 84]. For porous materials, jet impingement drying can combine with through-air drying [85], i.e. a part or the whole of the impinged gas flows through the material sheet to increase the drying rate. It may be more suitable to call these two types of drying "impacting drying" and "gas-jet impacting drying", respectively. Their common feature is that the stream impacts a solid surface, and neither involves impingement between opposing streams. As stated in the Section 6.1, they do not belong to the category being discussed in this book.
6.2.1 Impinging stream spray drying Leiner et al. [86] and Elperin et al. [87] proposed an impinging stream spray drying system for aluminum sulfate, as shown in Fig. 6.1. The spray drying is actually carried out in two primary drying chambers with perforated walls placed opposite each other. The hot airflows pass through the perforated walls at high velocity to enter the primary drying chambers in order to avoid the material caking on the walls. The hot airflows contact the spray droplets and partially dry them, and then carry and accelerate the particles flowing through the conduits to enter the secondary drying chamber where the impingement between the opposing hot airflows occurs and drying is finally carried out. It is reported that, since the process is controlled by external diffusion, the action of impinging streams in enhancing transfer is very efficient. The influences of the atomizing pressure, air flow velocity and the initial concentration of aluminum solution on the moisture content of the product were studied. Comparative experiments were also made with and without impingement and the results indicate that impingement can reduce the final moisture content from 18% in the case without impingement to a level below 12%. However, in this comparison the factor below was not considered: if one did not want to employ impinging streams, the structure of the equipment shown in Fig. 6.1 would obviously be unreasonable. In tact, the addition of a secondary drying chamber increases the residence time of the material to a large magnitude. This may be the major reason tor the increase in the degree of drying. No report on the industrial
122
IMPINGING STREAMS
application of the technology and equipment described above has yet been found. The application may be hindered mainly by the considerable complexity of the system. In addition, although the measure of the hot air passing through the perforated walls was employed, possibly the problem of material caking on the walls could not be completely avoided.
t Exhaust Hot air
Primarydrying 2 chamber I /
~
- ' - ........
o,u;i
•"" " '
Hot air
. :i !!....,
!!:::..]-
econ /I
. . . . . . . . . . . . .
V"!':-:<,-~......-'"""~*"4~~ Solution
dryingchamber Figure
Primary drying P chamber ] \
1 Dry salt i
6.1 Impinging stream drying system for aluminum sulfate [86, 87].
Exhaust
Atomizer Secondary air
Heater
Figure
Product
Tangential / nozzles
" Fan
6.2 IS Dryer for microbiological materials [88].
Kuts et al. [88, 89] patented a spray dryer with rotating impinging streams for solutions or suspensions of thermal sensitive materials, as shown in Fig. 6.2. The drying chamber is a horizontal cylinder, 1.2 m in diameter and 4 m long. The main airflows at 150°C enter axially the drying chamber from the two sides with the sprayed
IMPINGING STREAM DRYING
123
slurry concurrently. The gas flow-swirling plates mounted at the entrances bring the gas flows into rotary motions in opposite directions. In order to keep the material from caking on the walls and from being superheated, part of the air is replenished through the tangential nozzles, which forms thin film-like gas flow in the proper direction. Because of the strong evaporation of water, the temperature of the gas flow out of the drying chamber can be reduced by up to 70°C. The evaporation intensity can achieve 0.0078 kgH20.m-3.s -~, equivalent to 28.1 kg.m-~.h -~.
6.2.2 Impinging stream drying of granular materials In all the materials to be dried, granular ones occupy the greater proportion and so the impinging stream drying technology aimed at this kind of material is one of the hot spots of research and development. Tamer et al. carried out wide investigations on the drying of this kind of material in a number of impinging steam dryers with various structures and various numbers of stages; while in China, investigation of this topic started in the mid- 1990s.
6.2.2. 1 Co-axial horizontal impinging stream drying Kitron and Tamir [90] studied the drying processes in a smaller and a larger co-axial two-impinging stream dryer, the structures of which are essentially similar to that shown in Fig. 3.1 (for details readers may refer to Fig. 7.23 in Ref. [5]). The volumes of the drying chambers are 0.65x 10-~ and 2.2×10 -t m ~, respectively, the diameters of the accelerating tubes are identical, 0.02 m, and the dimensionless impinging distance, S/d,,, is adjustable in the range 0.6 to 5.6. There is a restriction in each of the feeding tubes before the position where the particles are fed in, which creates a low pressure zone and prevents the back-pressure effect in order to smooth the particle flow. The drying of two kinds of millets, the regulations and scaling-up of the dryer were studied, and the following main results were obtained: (1) Neither the air flow rate, the volume of the drying chamber, nor the impinging distance affects the values of the heat transfer coefficient; no significant influences of the diameter of the particles or the feeding schemes on the heat transfer coefficient were observed. (2) The effective volume for heat and mass transfer is very small. It only includes the limited space between the exits of the two accelerating tubes; while reducing the total volume of the dryer would result in increased hydraulic resistance of the device. (3) The mass flow rate of particles affects significantly the heat and mass transfer coefficients. The values for the heat transfer coefficient, h, measured range from 120 to 1800 W.m-:.K -j In addition, the influence of the impingement between the opposing
124
IMPINGING STREAMS
streams on heat transfer was studied in another co-axial horizontal two-impinging stream dryer by introducing a partition in the middle of the dryer, which separated the dryer into two non-interacting compartments without impingement between the opposing streams. Again, the equipment used is essentially similar to that shown in Fig. 3.1; but the drying chamber is cylindrical in form and, after impingement, both the gas and particles are discharged from the bottom of the cylinder. The results demonstrate that the use of impinging stream increases the heat transfer coefficient by 0.5 to 3 times and that such influence becomes more significant at higher gas flow velocity, higher particle flow rate and larger hold-up of particles in the dryer. Obviously, this work is limited to fundamental investigation and is considerably far from application. In addition, the basis they used for the comparison between the transfer coefficients of with and without impingement is improper: the partition changes the flow configuration from the impingement between the opposing streams to that of two suspension streams impacting solid surfaces individually so that the process becomes Impacting Drying, as stated above. As mentioned in the Introduction, the latter flow configuration also enhances heat and mass transfer between phases significantly. Therefore the degree of enhancing transfer they determined, 0.5 to 3 times, may be significantly lower than the reality.
6.2.2.2 Co-axial vertical two-impinging stream drying Tamir et al. [91] examined further the effectiveness of impinging streams enhancing heat transfer by comparing the performances of a co-axial vertical two-impinging stream dryer and a spouted-bed dryer. The structure of the former is shown in Fig. 6.3, where the particles are fed on one side. Simply by turning off the upper gas inlet port and feeding the particles through the lower gas-entering tube, the impinging stream dryer becomes a spouted-bed dryer. From the results of the investigation, Tamir et al. concluded that the heat transfer coefficient in the impinging stream dryer (h = 600-1000 W.m-Z.K-j) can be 1.5 to 3.3 times that of the spouted-bed dryer; while the power consumption for gas transportation (theoretically, it equals [volumetric flow rate of gas]x[pressure drop]) for the former is only 10% that of the latter. It is obvious that the hold-up in the spouted bed dryer is much higher than that in the impinging stream dryer. This is the major reason for the significant difference between the power consumptions in the two kinds of dryer. On the other hand, the much higher hold-up must yield much longer residence time and, consequently, quite different processing performance. Therefore these two kinds of dryers are not comparable. Hu et al. [92, 93] studied the drying feature of a vertical impinging stream dryer, including the mean residence time, the moisture removal efficiency, the effects of the initial temperature of the gas and the initial moisture content of the material, etc. The equipment Hu et al. used is different from that shown in Fig. 6.3: a concentric "impingement gave" is added inside the dryer, as shown in Fig. 6.4. The experimental results indicate that the application of the impingement gave may be useful for increasing the mean residence time of the particles in the active region to an extent, while the amplitude that can be increased is limited, and the increase in the residence
IMPINGING STREAM DRYING
125
time must be followed by an increase in the hydraulic resistance. It is reported for millets that the residence time in the impingement gave can be increased to about 10 s. It is obviously impossible to completely dry such a material in one path through the dryer. Therefore Hu et al. also studied an operation scheme with external circulation, i.e., part of the out material enters the dryer again to be dried further. Particles
~
Gas
I
1
Impingement One Particles y
articles
1 Gas Part + Gas
Gas
[
ave
N T
icles + Gas
Left: Figure 6.3 Vertical impinging stream dryer [91]. Right: Figure 6.4 Vertical impinging stream dryer with an impingement cave.
6.2.2.3 Tangential-horizontal impinging stream dryer Tamir et al. [58, 63, 94] studied the drying of granular materials in tangentialhorizontal impinging stream dryers of various flow configurations, such as two-, fourand multistage impinging streams. The tangential-horizontal two-impinging stream dryer is shown briefly in Fig. 6.5. The main dimensions are: the diameter of the accelerating tubes (I.022 m and 0.52 m long; the inner diameter of the outside cylinder 0.107 m and 0.335 m high; the diameter of the inside cylinder 0.062/0.07 m. The drying is carried out mainly in the annular chamber between the external wall of the inner cylinder and the internal wall of the outside cylinder, in which several blades are placed for swirling air to enable a certain increase of the hold-up and thus the residence time, while the blades also result in an increase in the hydraulic resistance. The -3 granular material of dp - 1.9xl() -~ m and p p - 1153 kg.m is dried with hot air at a
126
IMPINGING STREAMS
temperature of 40-45°C. In the range of the impinging velocity from 8.4 to 12.5 m.s -~, the values experimentally measured for the heat transfer coefficient are in the range of 548 to 1035 W.m-2.K -~, the volumetric evaporation coefficient is between 4 and 10 kgH20.s-l.m-3.K -j, the pressure drop across the system 700 to 1550 Pa, the hold-up in the dryer 3.9 to 6.0 kg, and the mean residence time 600 to 1245 s. /
Particles ~
!
1
Air
Air :I "-~1 g.])~_(_i ~.... /ii
A
Accelerating tu~~
1 2,,"~
A
B1
A-A
~ Particles
tt
ii
Figure 6.5 Tangential-horizontal two impinging streams dryer [58].
Air Particles
Figure 6.6 A brief view of tangential four-impinging stream dryer [94].
IMPINGING STREAM DRYING
127
The one-stage tangential horizontal four-impinging stream dryer is constructed by adding opposite to each other a pair of suspension-feeding tubes in the two- impinging stream dryer, as shown in Fig. 6.6 [94]. The results showed that the hold-up is increased to some extent in comparison with that in the two-impinging stream dryer, and thus the residence time is somewhat increased; no significant changes in the heat transfer coefficient and the volumetric evaporation intensity have been observed, indicating that the effective surface area for heat transfer did not increase proportionally as the hold-up increased. The researchers reported data for the pressure drop (the maximum being 600 Pa) even smaller than those obtained in the twoimpinging stream dryer. However, since the two dryers are quite different in size, there is no basis for comparison. From the point of view of practical application, the dryer shown briefly in Fig. 6.6 has excessive complicated flow configuration, and so it may be difficult and troublesome to achieve stable operation. Bar and Tamir [63] studied the two-impinging stream dryer with two pairs of airfeeding tubes, as briefly shown in Fig. 6.7. The purpose of adding the lower two air streams is to increase the hold-up and the mean residence time of the particles in the dryer, and also aims to enhance the turbulence between phases in order to increase the drying intensity. However, the experiments did not show that the secondary air streams increased either the hold-up or the transfer coefficient. On the other hand, the induction of the two secondary air streams results in the greatly increased hydraulic resistance of the system. The pressure drop across the dryer, with two pairs of air-feeding tubes and with a volume treble that of the dryer shown in Fig. 6.6, is as high as 3800 Pa.
~ Particles
Air
B I ~ Primary air flow i b
i
,
I-"
Secondary air flow
t
& Particles
Figure 6.7 A brief view of the tangential impinging stream dryer with two pair gas-feeding [63].
128
IMPINGING STREAMS
In addition to those above, Kitron et al. [95] studied experimentally a tangential horizontal flow multistage dryer. The dryer consists of multiple overlapping annular chambers. Two suspension streams enter the upper annular chamber on one side and impinge against and mix with each other on the other side of the chamber. The combined stream flows through the openings in the partition separating one stage from the other and down to the next annular chamber where the stream is divided into two streams again by a triangular flow splitting device. The two streams flow tangentially and in opposite directions and impinge against and mix with each other on the other side of the annular chamber again, and so on. After several impingements (three times in the experimental equipment used by Kitron et al.), the major part of the gas is discharged to the atmosphere through the central conduit inside the dryer, while a small part of the gas is discharged with the particles together from the cone-shaped bottom of the dryer. The details of the dryer can be found in Ref. [95]. Because of the multilayer structure, the hold-up and the residence time of particles in the dryer are greatly increased in comparison with those in a single stage dryer, while the heat transfer coefficient and the volumetric evaporation intensity are somewhat decreased. No data on the pressure drop across the dryer has been reported, which may be another important problem with the dryer.
6.2.3 Impinging stream drying combinations To solve the problem that the residence time of material in a simple impinging stream device is very short so that drying cannot be carried out individually, in addition to the improvements described above, some technological schemes have been proposed which combine the method of impinging streams with other method; some of these are introduced below.
6.2.3.1 Flash-impinging stream drying The drying of pharmaceuticals was tested at pilot plant scale in the former Soviet Union [96]. The equipment used is actually a combination of a pneumatic flash dryer with an impinging stream dryer, as shown in Fig. 6.8. The wet material delivered by a screw feeder mixes with the hot air at 180 °C in the injector-type chamber, which acts as a small flash dryer, to form a suspension. The suspension is fed into the lower part of a vertical cylindrical dryer through an axially situated tube with a blade-type swirler mounted at the end of the tube. The swirler creates an internal swirl flow in the form of a slight cone with its base directed up towards the dryer outlet. Supplementary air at 150 °C is introduced tangentially at the top of the dryer. The two swirling streams impinge against each other at the middle of the space inside the vertical cylinder, where the material is dried; the mixed stream then leaves the internal swirl at the bottom and moves towards the dryer walls due to centrifugal force and then drops down to the discharge bin. The external swirling of the upper gas stream is returned by means of a deflector plate, and the gas then flows up with the internally swirling air to the exit.
IMPINGING STREAM DRYING
129
The dry-based capacity of the equipment is 1.78 kg.s -~, while the recovery efficiency of the particles is 98-99%. It is reported that the conditions of fluid dynamics are very stable and that the equipment can be operated with a very high concentration of dense phase. It is clear that at least one of the researchers' major intentions in the design of the dryer is to lengthen the residence time of the particles. im
ql Product Flow-swirler Hot (150
eei . ~ Product Hot air (18()'~C) Figure 6.8 Flash-impinging-stream dryer for pharmaceuticals [96].
6.2.3.2 Drying of sewage with high moisture content The process proposed for the drying of industrial sewage of high moisture content is shown in Fig. 6.9 [3, 97]. The system consists of drying in impinging streams and additional processing of the residue by high temperature gases in a pneumatic pipe. The initial moisture content of the sewage in dry-basis is 2.28 to 4.56; while the final one is 0.21 to 0.42 kg.kg -~. The wet material mixes first with the recycled fine powder and is then fed into the impinging stream drying chamber where the majority of the moisture is removed and, at the same time, the material is ground. The resulting gas-solid suspension passes in order the air-lift tube and the centrifugal separator where the material is dried further. The coarse traction separated out is the dry product, while the fine fraction is fed back into the feeder of the raw material as the recycle, which absorbs the moisture in the sewage to avoid caking on the walls. The main structural dimensions and operating parameters of the equipment are: the accelerating tubes 0.3 m in diameter and 0.9 m long: the velocity of gas flows ejected from the nozzles 20 to 25 m.s -~ and the velocity in the accelerating tubes, i.e. the impinging velocity, about 10 m.s -~. The gas from the separator is cleaned of dust and then ejected to the atmosphere. A similar process was introduced in Ref. [89].
130
IMPINGING STREAMS
[-----]
Exhaust
entriuga, _
/separat°r /
Recycling / ~ x N Feed ~ a i r l i f t I ~..._ tubel
Fuel gas
~@. ~' RecYcling ~ ~ Dr3~
Feed ~
~.~
AiriNg(' )~ Combustion chamber
~( )~'~ Air IS dryingchamber Combustion chamber
Figure 6.9 The IS-flash-rotary drying system for slurry [3].
6.2.3.3 Impinging stream mill dryer In the systems introduced above the impinging streams method is combined with other technical methods to carry out the given drying job. In addition to such combinations, it is also possible to utilize the kinetic energy of gas streams at high velocity for milling materials simultaneously with drying. This is actually a multi-functional combination in which the impinging streams play a double role: drying and milling or grinding. A device for the drying and grinding of tail by impinging streams is illustrated in Fig. 6.10 [3]. The raw material is delivered first into the airlift tube and then into the separator to separate it into various fractions of different sizes. The large and wet fraction is guided through a pipeline into the impinging stream dryer, where the solid particles are driven by the gas streams ejected from the nozzles at a velocity of 45 m.s -~ and are accelerated by the gas flows at a velocity of 10 m.s -~ in the channels; and then flow into the milling-drying chamber where the drying and grinding take place simultaneously. Grinding occurs due to interparticle collisions. The gas carrying fine powder flows through the multistage cyclone to remove the fine product, while the gas is cleaned further in a scrubber and then discharged. The initial moisture content of the material being processed in the equipment is 0.333 to 0.37 and the final moisture of the product is 0.002 to 0.005 kg.kg -~ (both on dry-basis); the capacity of the device is 1.83 kg.s -~ (dry product); the heat consumption per kg water evaporated is 3360 to 3800 kJ, and the power consumption per kg product is 60 kWh.
IMPINGING STREAM DRYING
131
! II j
Gas to scrubber
Cyclone
Feed
1
Product
Metering valve
Grinding-- IS drying chamber
Combustion chamber
Figure 6.10 Impinging stream mill dryer for processing granular materials [3]. Tutova et al. [98, 99] reported another kind of impinging stream drying-milling device which utilizes inert particles, as shown in Fig. 6.11. The dryer consists of two impingement chambers situated at both ends of a central drying-milling tube, the reversing flow tube or impact tube which is 1.5 m in length and 0.03 m in diameter. The exits of the impingement chambers to the cyclones are fitted with constraining grids with a mesh dimension of 0.0013 m. The grids prevent large particles from leaving the impingement chambers and increase the mean residence time of smaller particles in the system. The impact tube contains a mixture of steel beads of 0.002 m in diameter and aluminum beads of 0.003 m in diameter in a mass ratio of 1:1 as the milling medium for reduction of particle sizes by interparticle collisions. The wet material is fed on one side into the accelerating tube, and delivered by the hot air flow at high velocity into the impact tube to impinge against the gas-solid suspension entering fiom the opposite end. Due to the action of the switching valve, the impingement occurs periodically in one of the impingement chambers and, consequently, the inert particles and the wet material take the reversing motion between the two impingement chambers at a frequency of 1 to 2 Hz, the latter being dried and ground as it moves in reverse. The main experimental results are: after 4-5 oscillations, equivalent to 10-15 seconds of drying, the moisture content of the material drops from
132
IMPINGING STREAMS
0.179 to 0.00462 kg-kg-~. The dried and ground fine particles are carried by the gas flow through the grids and collected in the separator. The equipment is applicable for drying of some special materials. For example, crystalline lysing lacks thermo-stability and has the tendency to agglomerate, contains a large amount of bounded water (about 17%), and has wide size distribution, while the quality indexes of the product required are: a final moisture content no greater than 0.01 to 0.015 kg.kg -j and a mean diameter of particles no larger than 0.5-0.6 mm. It is therefore a difficult material to dry. It has been reported that the system shown in Fig. 6.11 has successfully solved the problem of crystalline lysing drying. The device was also used for drying wheat grains [100]; Meltser and Tutova [101] analyzed theoretically the drying process of granular materials in the same equipment. t Exhaust gas ImuPbaCt
~
SW~ta~vhi?g
Impingement ~
Accelerating
D-'~ 1~ c// h a m b e r
tube
Feed ]
I
i
\
\ I
I
i i r _1_1_
-I
I
L___I----
I
I
I l°ry I
Hot air
product
Figure 6.11 Impinging stream milling dryer for crystalline lysine [99].
6.2.3.5 Semi-batch impinging stream drying It is clear from the discussions in the previous sections one of the key problems facing researchers is how to lengthen the mean residence time of the material in the dryer. In fact, many granular materials to be dried are porous or contain bounded water. The processes of removing water either in pores or bounded are controlled by internal diffusion and thus need considerably long times, say several tens of minutes. No matter what the device or what its structure or the flow configuration employed, it is almost impossible to accomplish the drying job with a simple impinging stream dryer operated
IMPINGING STREAM DRYING
133
1 14
!
Air + vapor
Jil ii!i
!iii
Air
Figure 6.12 Semi-batch impinging stream dryer [ 102].
continuously. Batch or semi-batch operation is one of the most effective methods, greatly lengthening residence time. Frishman [102] studied the drying of rapeseeds in the equipment shown in Fig. 6.12, which the gas flow passes through in one path while the granular material is retained and circulates. Essentially, it is a combination of impinging streams with spouted bed with a central duct. The lower gas flow at a high velocity draws the particles at the conic bottom of the dryer, carries and accelerates them through the central duct. The suspension ejects from the central duct, and impinges, in the upper space inside the dryer, against the upper gas flow entering from the top of the drying chamber. After impingement the gas with increased moisture exits through the mesh holes of the grid fixed on the upper wall of the drying chamber, while the particles drop down into the annular chamber due to prevention of the grid, move downwards by gravity to the conic bottom of the dryer, and then are drawn into the lower gas flow again. Depending on the times of circulations, the final moisture content of the dry product can achieve any required level. Actually, this is because the residence time can be arbitrarily lengthened in the case of batch operation. Obviously, the device shown in Fig. 6.12 is not suitable for industrial production on a large scale because of its batch operation.
134
IMPINGING STREAMS
6.3 CIRCULATIVE IMPINGING STREAM DRYING On the basis of a thorough understanding of the properties of gas-continuous impinging streams, including its advantages and disadvantages, the circulative impinging stream dryer (CISD) was developed and patented by the author of the present book [11, 103]. Primary tests on a quasi industrial scale have yielded satisfactory results, and it can be expected to be applied industrially in the near future.
6.3.1 Basic ideas for equipment design In view of the fact that most materials needing to be dried are powdery or granular solids, the development of the circulative impinging stream dryer is aimed at this kind of material. For drying, the typical parallel heat and mass transfer process, the feature of impinging streams that significantly enhances transfer, should of course be fully utilized. The water existing in powdery/granular solids can be classified into three types: the surface or free water, the water in pores, and the crystallized (bounded) water. Any material containing water in pores or bounded must simultaneously contain surface or free water. Either surface or free water is easily removable and there are many types of dryer, such as fluidized bed dryers, pneumatic flash dryers etc., capable of meeting the requirement of removing these two kinds of water. However, the removal of water in pores or bounded is difficult, and usually needs a considerably long time. In the past, combined multistage drying systems were usually used for materials containing water in pores or bounded. For example, the drying system traditionally used for PVC is a combination of a pneumatic flash dryer plus a fluidized bed dryer; however, this yields a number of problems: it is a complicated system scheme that occupies a large amount of space, has a higher capital cost, higher energy consumption and presents difficulties in operation and device maintenance. In a common impinging stream device, such as the simple one used by the author in the investigations on the residence time distribution and hydraulic resistance, the mean residence time is very short and so cannot meet the requirement for drying the materials mentioned above. Therefore, a measure is called for that would significantly lengthen the residence time of the material in the dryer, especially in the active region. As can be seen from the discussions above, in all the developments on impinging stream drying technologies this topic receives great attention and, consequently, several measures have been employed for this purpose. However, most of the proposed IS dryers operated continuously can only lengthen the residence time to a limited extent; and some of them involve greatly complicated structure of the equipment, mostly accompanied by significantly increased hydraulic resistance. The dryer shown in Fig. 6.12 can only be operated in batch, and this is its major disadvantage. However, lengthening the residence time by circulating the solid material employed is an effective measure.
IMPINGING STREAM DRYING
135
The goal of the investigation is to fully utilize the superiority of impinging streams in enhancing heat and mass transfer, drawing on the advantages and compensating for the disadvantages of existing technologies, in order to develop simple, highly efficient and energy-saving equipment for drying powdery and/or granular materials that can accomplish the whole of the drying job, including the removal of water in pores or bounded. This is the basic aim when designing the dryer.
6.3.2 Structure and working principles of the dryer The structure of the circulative impinging stream dryer developed is sketched in Fig. 6.13, in which the flow configuration of vertical two impinging streams is employed. The equipment consists of three basic parts mounted co-axially" the body, the upper and the lower accelerating tubes. The body is of a conical/cylindrical shape with different diameters. The lower-middle part of the body is a cylinder with a relatively
I
Hot air I
Exhaust ~
A
, ,
=/IL q:i!::~
(
Wet feed
li!il
Settling chamber
/ :::::
J / /
.
.
.
.
.
.
........-....
Accelerating tube
.
,.,..,,,,.., ..,.......,
,.,..,,,..,/
\
.
.
.
.
.
.
Impingement zone
-.-.......... .............
............. ............. ............ ............
.,.,.,.,.....
"\'\'\,....
i!i!!i!!i!ii!i i!!iiiiiii!ii !iiii~i!iiiii !!!!!i-ii!!! ..,...,.,.., .,...,....,.
............
......-....
i:i:i:-i:i:¸
:::::w:::::! . . . . . . . . . . . . . ............. ............. . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
jY
Annular chamber
:i:!:i:i:~ . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . .
?!!!?i!i ,.,.,.-....,, . . . . . . .
,
I
Dry product Hot air II Figure 6.13 A brief view of the circulative impinging stream dryer (CISD).
136
IMPINGING STREAMS
smaller diameter, the lower end of which is connected to the conical bottom of the dryer, and the latter to the inlet tube for the lower gas flow, Hot air II. The discharging tube for the dried material is fixed on the bottom of the lower cylinder near its vertical wall. The annular chamber formed between the walls of the lower cylinder and the coaxial lower accelerating tube is the space for the movement of the circulative material. The inlet of the lower accelerating tube is distanced by a certain height from the outlet of the entering tube for Hot air II in order to enable the gas flow to draw and carry the particles from the bottom of the annular chamber to enter the lower accelerating tube and to flow upwards. The diameter of the upper body is enlarged gradually to that of the upper cylinder with a top cover. The space inside the upper cylinder/cone, except for that occupied by the upper accelerating tube, is the settling chamber; and the gas exit is settled on the top cover of the dryer body. The upper accelerating tube is also the inlet of the upper gas flow, Hot air I, to which the inlet tube for the material is connected at a certain height over the top cover. The co-axial upper and lower accelerating tubes are spaced apart by a distance S, which is just the impinging distance. To run the dryer, the annular chamber is first filled with dry material and then the gas flows, Hot air I and II, are turned on. The wet material is fed by the screw feeder into the upper accelerating tube and accelerated by Hot air I. The gas-solid suspension formed ejects from the upper accelerating tube and flows towards the impingement zone. At the same time the other gas flow, Hot air II, enters the dryer at the bottom at a mass flow rate about equivalent to that of Hot air I, draws essentially dried particles from the bottom of the annular chamber, and disperses and accelerates them during its flow upwards through the lower accelerating tube to form the upward suspension stream at high velocity. The two suspension streams impinge against each other in the impingement zone, mix and combine with each other, and then become a radially outward flow. The particles in the radial flow drop down, or slip down after colliding on the wall, by gravity into the annular chamber, move slowly towards the bottom, and then are drawn by Hot air II again and circulate inside the dryer, while the gas flows upwards and, after separation from smaller particles by settling, is discharged from the top of the dryer. During operation, the material undergoes two stages of different properties: (1) The fresh material is dried primarily under conditions of very rapid transfer between phases to remove the free and/or surface moisture; and (2) During the downward movement inside the annular chamber and the multiple circulation carried by Hot air II, the particles are finally dried. In the period between the two impingements, i.e., during the downward movement inside the annular chamber, the particle experiences a stage of long residence time but under weak transfer conditions, resulting in effective reduction in the gradients of temperature and/or moisture content inside particles, while later the particles undergo impingement again under the conditions of strongly enhanced transfer. Such a periodical variation of transfer condition favors the removal of the water bounded or in pores. The particles leaving the dryer have generally undergone multiple impingements and a long enough drying time, provided the dryer is reasonably designed, that the required drying degree can be ensured. The dried material is discharged quantitatively through the outlet tube near the side wall of
IMPINGING STREAM DRYING
137
the lower cylinder to keep the level of the material inside the annular chamber from becoming unstable. It can be seen that in the arrangement of flow configuration CISD is somewhat similar to that of the batch IS dryer shown in Fig. 6.12, while the others are quite different. The CISD can be considered to be a combination of a spouted bed with central duct and impinging streams. The lower accelerating tube in CISD is equivalent to the central duct in a spouted bed; but the operating velocity in the former is usually much higher than in the latter. Theoretically, CISD has two advantages in performance: (1) the feature of impinging stream enhancing significantly heat and mass transfer is used to intensify the drying processes, and (2) the total mean residence time can actually be set arbitrarily to meet the requirements of drying various materials, especially those with high moisture content. In addition, in the combination of spouted bed and impinging streams described above, the advantage of lower hydraulic resistance that the spouted bed with central duct intrinsically has is retained in the circulative impinging stream dryer. The research and development of CISD is carried out in two stages: laboratory experiments and pilot plant tests.
6.3.3 Experimental model equipment scheme and procedure 6.3.3.1 Scheme of equipment system Huang et al. [11 ] carried out a modeled experimental study on circulative impinging stream drying with PVC produced by the suspension process as the target material and the scheme of the equipment system used as shown in Fig. 6.14. Two modes for dry product discharging were originally considered: from the bottom of the dryer and by overflow. For this purpose the dryer involved in the system scheme is somewhat different from that shown in Fig. 6.13, i.e., the former includes an arrangement for overflow of the material but the latter does not. If the requirement of drying can be ensured, the mode of product discharging by overflow has the important advantage that the level of the material in the annular chamber becomes very easy to control. For experiments operated continuously in small equipment, like that used in the investigation described here, the wet material feeding is extremely problematic, much more difficult than in an industrial device. Because of the very small diameter of the feeding tube suitable to the dryer, the wet material often clogged up the tube and even locally bonded in it, destroying the condition of stable feeding, and causing the run to be stopped. After a number of improvements and tests, the feeding system briefly shown in Fig. 6.14 was designed which includes a vertical screw feeder and a lower disperser (scatterer). The latter utilizes the impacting effect of the gas flow at high velocity to disperse the wet material and carry it into the upper accelerating tube. This system successfully solves the problem of the continuous feeding of wet material into the dryer at a stable rate, and thus enables the experiments to proceed smoothly.
138
IMPINGING STREAMS
6
7
•
5
: ~4
Figure 6.14 System scheme for CISD experiments. 1 fan; 2 heater; 3 rotameter; 4 CISD; 5 sampling port, 6 screw feeder; 7 scatterer; 8 dust catcher.
6.3.3.2 Major parameters for equipment and system design The main parameters for the experimental equipment and system design are listed below: • • • • • • • •
•
Production of dry product: 0.0033-0.0067 kg.s -~ (12-14 kg.h-~); Gas velocity in the accelerating tube, i.e., the impinging velocity: set as 6 to 9 times the terminal velocity, so that u0 - 2 0 - 3 0 m.s -~" Length of the accelerating tube: calculated with the method described in Chapter 47 L~ - L2 =0.6 m; Diameter of the accelerating tube: do - 0.025 m; Height of the annular chamber: H ~ 0.8L~ = 0.5 m; Diameter of the lower cylinder: calculated according to the required drying time 0.5-1.0 h, the hold-up M ~ 10-12 kg, yielding D - 0.2 m; Impinging distance S - 0.1 Spacing of the inlet of the lower accelerating tube from the exit of the entering tube for Hot air II (simply called the lower spacing, adjustable)" h = 0.02-0.05 m; Temperature of the entering air: it should be much lower than the softening temperature of PVC, take Tg0 - 90-110°C.
m;
IMPINGING STREAM DRYING
139
6.3.3.3 Experimental procedure Dry PVC powder is used as the process material for all the experiments in the first stage of the investigation, the properties of which are the following: /3~ = 472.3 kg.m-~; pp = 946.1 kg-m-~; dp= 1.8x 10-4 m Before each run, the dry PVC powder taken from a factory is mixed with water in a certain proportion; the mixture is sealed with plastic film for over 24 h to restore the PVC's moisture content uniformly. The relationship between the teed rate of the wet PVC and the rotary speed of the screw feeder was calibrated prior to all the experiments. In each run, the annular chamber of the dryer is first filled with a certain amount of dry PVC powder and then the hot airflows I and II are turned on to preheat the system. When the temperature of the material inside the dryer achieves the required value, the screw feeder is turned on and the experimental operation begins. About one hour later, when operation of the system achieves stable conditions, sampling is started and, at the same time, the dry- and wet-bobble temperatures at the inlet and the outlet of the dryer, the flow rates of Hot air I and II, and the rotary speed of the screw feeder are recorded.
6.3.4 Major results of the model experiments
6.3.4.10verafl performance of CISD The experiments of the PVC drying continuous operation yield positive results for the removal of both free and in-pore moisture simultaneously. Some of the typical experimental data obtained is listed in Table 6.1. It can be seen that in the ranges of experimental conditions tested the final moisture contents, xr, can meet the requirements specified by the standard for the product (no greater than 0.4% in wet basis). The initial moisture content of wet PVC, x0, is around 7%, which is significantly lower than that of commercial wet product from the centrifuge (-20%). This is due to the special difficulty in the operation of a small-scale experimental device. Although efforts have been made on numerous occasions to increase the initial moisture content of the wet material, untbrtunately none of them were successful. A further increase in the initial moisture content mainly results in two problems: (1) Very often the feeding system for the wet material is blocked due to clogging so that continuous operation becomes extremely difficult; and (2) the adjustable ranges of the temperature and flow rates of the hot air I, II and the feeding rate of wet material are limited due to the equipment used, including the supplementary ones. At higher initial moisture content, the amount of heat that could be supplied by the hot air flows is not high enough so that the final moisture content of the product increases gradually, and finally stable operation is destroyed due to the difficulty in ejecting the dried material. On the other hand, it is clear that these problems could be easily solved in large-scale industrial equipment, so the problems mentioned should not interfere with the conclusion described above.
140
IMPINGING STREAMS Table 6.1
Typical experimental data of PVC drying (7'8o= 100-1 l0 C°, x0 = 0.07 kg.kg-l)
Impingement distance S, mm
Lower distance h, mm
Impinging velocity u0, m.s - i
Feed rate
1
60
30
19.81
0.00414
0.00325
2
60
30
21.22
0.00414
0.00294
3
60
30
22.64
0.00414
0.00305
4
60
30
24.05
0.00414
0.00271
5
60
30
25.47
0.00414
0.00224
6
60
30
22.64
0.00395
0.00293
7
60
30
22.64
0.00370
0.00290
8
60
30
22.64
0. 00346
0. 00273
9
60
30
22.64
0.00315
0.00256
Run
No
mp, kg's -1
Moisture content of product xf, kg'kg -I
6.3.4.2 Examination of the effect of the internal circulation In order to understand the effect of the circulation of material inside CISD on the drying process, comparative experiments were carried out for the two cases with and without internal circulation of the material. The method for experimental runs without internal circulation is to replace the lower accelerating tube down to the conic bottom of the dryer so that Hot air II cannot draw and carry the material from the bottom of the annular chamber to flow upwards. In this case the drying is carried out mainly during the impingement once between Hot air I carrying the fresh material and pure Hot air II, while the drying inside the annular chamber becomes nothing. The data measured after starting up the system for a short time are shown in Fig. 6.15. It is observed in the case without circulation that the moisture content of the material inside the annular chamber increases gradually, clogging of the material begins at the bottom of the chamber, and the operation becomes more and more difficult. If the annular chamber of the dryer is not filled with dry material at the beginning of the run, the system cannot start up. The results described above indicate that, without the circulation, the requirement for drying PVC cannot be met. In other words, the moisture in the pores of porous material cannot be removed in a simple impinging stream dryer.
IMPINGING STREAM DRYING
141
1.0
0.8
-
S=I00 mm
mp=0.00414 kg-s -~
h=30 mm
u0=25.47 m's -l
-7
0.6
A
•
w
•
¢-i
O
x 0.4
0.2
0.0
__
500
~_
i
L
400
l
~
300
~___L_ 200
l
j
__
lO0
Height from the out port H, mm
Figure 6.15 Results of comparative experiments of with (O) and without (O) circulation.
6.3.4.4 Profile of moisture content of the material along the path It would be of interest to understand the variation of the moisture content of the material inside the annular chamber. The profile of the moisture content of the material along the path of motion inside the annular chamber is measured under conditions of stable operation and with impingement between the streams and the result is illustrated in Fig. 6.16. As can be seen, drying occurs to an extent as the material moves downward, but the rate is limited. The evaporation of water in a small amount results from a small part of Hot air II by-passing the annular chamber. Since the transfer between phases in the chamber is very weak, it is impossible to evaporate water in large amounts. This result is consistent with that predicted. The major phenomenon occurring inside the annular chamber is the moisture inside particles transferring towards to their surface, like the "moderating" in the drying of grains. On the other hand, by multiple circulations the residence time of the material is greatly lengthened so that further drying can be accomplished completely.
142
IMPINGING STREAMS
0.4
S=60 mm
m
h=30 mm
mp=0.00414 kg's-l 7
6o c.!
O
0.3 -
0
o x
0
~)
0.2 --
I
[]
[]
] m i
I
0.1 500
J
400
I
i
300
I
i
200
i
100
i
0
Height from the lower out port, mm Figure 6.16 Profile of moisture content of material in the annular chamber, uo, m's - j O 25.47, • 24.05, [3 22.64, • 21.22.
6.3.5
Influences
of structural
and operating
parameters
6.3.5.1 Criterion for comparison During the investigation, the influences of some structural and operating parameters were examined experimentally. As mentioned above, the drying-moisture removal occurs mainly in the impingement zone. Because the zone has no physical boundary and also the hold-up and the interface area in this zone have great uncertainties, it is difficult to determine the transfer parameters from the experimental data. On the other hand, to achieve the drying process, the other regions in the dryer cannot be lacking among which the annular chamber plays an important role in promoting the internal moisture to move towards the surface, as discussed above. Therefore it is more convenient and may be of more practical interest to employ the parameters based on the total volume of the dryer as the basis for comparison. In the investigation, the volumetric evaporation intensity, Ev, is used as the major comparative criterion, which is defined as E v = ~,q/AHv VD AZlm
kgH20.s -j -m-3.K-t
(6.1)
where q is the heat transfer rate, kJ.s -~, and is calculated by q - mp (Xo - x f ) A H v
and AHv is the vaporization heat of water.
(6.2)
IMPINGING STREAM DRYING
143
It is found by observation that the temperature of the material in the dryer exhibits somewhat different behavior from that of the same material with the two-stage drying mechanism in other common drying processes. In the circulative impinging stream dryer, the variation of PVC temperature is very small and is essentially kept at a constant temperature very near the wet bobble temperature of the hot air. The most likely reason is that the weak transfer condition in the annular chamber promotes the movement of internal moisture towards the surface of the particles so that most of the particles have resumed their wet-surface state when they re-enter the impingement zone. According to this observation, it is assumed, as the first order approximation, that the temperature of the material is kept at the wet-bobble temperature of the hot air throughout the whole drying process. Thus, the logarithm mean temperature of the material can be calculated by
AT~,,~ -
Vg() _ Tgf In[ (T,~/t - T~, ) (T,~r - Tw )]
(6.3)
6.3.5.2 Influence of the impinging distance Some of the results on the influence of the impinging distance S are given in Fig. 6.17 as a plot of the volumetric evaporation intensity, E,, versus the dimensionless impinging distance, ( = S/d,,. It can be seen that in a large range of (variation the volumetric evaporation intensity increases significantly as (decreases. This phenomenon can be explained as follows: at smaller ~" the two opposing jets impinge against each other with small expansion and low velocity decay yielding stronger impingement and penetration of particles to and fro between the opposing streams [5], both of which favor enhancing the evaporation-drying process; on the contrary, at large ( the stream jets have expanded significantly and their velocity has greatly decayed so that the effect of impingement becomes very weak, yielding decreased evaporation intensity. However, in the range of even smaller ~" (3.0-4.0), the variation of Ev with ( i s not obvious. Further extended experiments [104] show that, for ~" < 3.0, E, inversely decreases as S/d,, decreases. It is observed that, with very small (, the concentration of particles round the impingement plane is significantly increased and a parclose-like layer of particles is ~brmed, which impedes the relative movement between phases and the penetration of particles. This may be the major reason ~br the decrease in E,. with ( decreasing. The results suggest that there is an optimal range for the impinging distance. On the other hand, the results of the investigation on the hydraulic resistance of the IS device carried out by Wu et al. [66] indicate that in the range of ~'< 4.0 the pressure drop and fluctuation strengthens sharply as the impinging distance decreases; which are consistent with those from the investigation on the pressure profile in the impingement zone by Elperin [3 ]. These results imply that too small ~'will result in an increase in the instability of the operation. From a combined consideration of the factors mentioned above, the dimensionless impinging distance of ( = 4.0 is recommended as being optimal.
144
IMPINGING STREAMS
20.0 u0=25.47
m . s -t
h=30 mm -7
16.0
5' -7
zz ¢--i
2z ×
12.0
8o 4.0
I
3.0
3.5
i
I
4.0
i
!
4.5
i
5.0
(=S/do Figure 6.17 Influence of impingement distance on the volumetric evaporation intensity, mp, kg-s-~" .00.00414; 00.00395; [] 0.00370; ~. 0.00346; A0.00315.
6.3.5.3 Effect of the lower spacing The results on the influence of the lower spacing, h, on the volumetric evaporation intensity are illustrated in Fig. 6.18. The influence seems very complex, but it must be closely related to the amount of particles drawn and carried by Hot air II, mpr. The results Wu et al [ 105] previously obtained in an experimental study showed that, in the case without obvious by-passing of lower airflow through the annular chamber, the circulative mass flow rate of particles increases as the lower spacing increases. As can be seen in Fig. 6.18, for the dimensionless lower spacing in the range of ~:= h/do < 1.5, the volumetric evaporation intensity Ev increases rapidly as ~: increases. The reason is the increase in the circulative mass flow rate of particles, resulting in increased concentration of particles in the impingement zone, i.e., the enhanced transfer between phases is mainly due to the increase in interface area in the active region. However, as ~:increases further, too many particles are carried by airflow into the impingement zone to form a parclose-like layer of particles, as in the case of too small an impinging distance, leading to obviously decreased volumetric evaporation intensity, Ev. The phenomenon that, after~:> 1.75, Ev somewhat increases as ~: increases is still difficult to explain precisely. The most likely reason is that part of hot air II by-passes through the annular chamber at large lower spacing, resulting in the circulative flow rate of particles decreasing inversely. From the data shown in Fig. 6.18, the dimensionless lower spacing of ~:~ 1.5 is recommended as optimal.
IMPINGING STREAM DRYING
145
16.00 ~g.s -7
-1
14.00
~.
E -7 TO
:~'
~at)
12.00
% X
>
10.00
8.00 0.00
.
.
.
.
1.25
, 1.50
1.75
2.00
2.25
~=h/d Figure 6.18 Influence of the lower distance, u0, m's-~- O 25.47" • 24.05; D 22.64; • 21.22.
6.3.5.4 Influence of impinging velocity The experimental results on the influence of the operating velocity of gas flows in the accelerating tube, i.e. the impinging velocity u0, on the volumetric evaporation intensity at various feeding rates of the material are shown in Fig. 6.19. All the experimental curves exhibit linear relationship. It may be considered that any impinging velocity in the whole range of 20-25 m.s -~ tested is feasible for operation. Of course, larger impinging velocity implies increased power consumption. The decision on the operating impinging velocity for a practical IS dryer depends on the balance between the enhancement of transfer processes and the power consumption. The influence of the impinging velocity on the hydraulic resistance of an IS device has been described in detail in Chapter 4.
6.3.5.5 Influence of the feed rate of material The results on the influence of the feed rate of wet PVC on the volumetric evaporation intensity are shown in Fig. 6.20. In the range of the feed rate, rap, tested the volumetric evaporation intensity E, increases linearly as mp increases. This is because of the increase in the surface area of wet particles as the feed rate increases. This does not, of course, imply that the feed rate can be increased infinitely. Increase in the feed rate directly suggests increased capacity of the dryer, while an impinging stream dryer has a finite capacity. At certain flow rate and temperature of hot air, there is an increase in the degrees of both the drop in temperature and the rise in humidity of the drying medium, resulting in decreased driving forces for heat and mass transfer so that the drying rate decreases and the moisture content of the product increases.
146
I M P I N G I N G STREAMS
20 S=85 mm
7
h=30 mm
16 E -7r.g3
6
A
(',4
12
×
I
18
i
I
20
i
I
22 u0, m-s
24
26
-1
F i g u r e 6.19 Influence of impinging velocity on the volumetric evaporation intensity, mp, kg.s-l: C) 0.00414; • 0.004395;/~ 0.00370; • 0.00346; A 0.00315.
9.5 S=85 mm, 9.0 -
7
h=35 mm
-
©
(3
8.5 -
8.0 ×
7.5
7.0 0.0034
I
I 0.0036
I
I 0.0038
I
I 0.0040
i 0.0042
mr,, kg/s F i g u r e 6.20 Influence of feeding rate. u0, m's-~: © 25.47; • 24.05" A 22.64; A 21.22.
IMPINGING STREAM DRYING
147
0.32
0.30
w
7
0.28
E
r-i
3.0 0.26 -
0.24 0.0028
O/
Uo= 22.64 m's-1
I
0.0032
i
I
0.0036
I
I
0.004
I
0.0044
mp, kg.s -j Figure 6.21 Influence of feed rate on moisture content of product. A set of data on the variation of moisture content of the product with the feed rate under typical operating conditions is given in Fig. 6.21. In critical cases the moisture content may be above the value specified; and in more serious cases the material in the annular chamber may completely loss its ability to flow, the system is blocked out, and finally the operation is destroyed.
6.3.5.6 Study of the arrangement of the product discharge position In the equipment shown in Fig. 6.14 two possible options for the product discharge were considered. It is reasonable that researchers would want to employ the scheme of overflow through the upper outlet. If such a scheme was feasible, the control of the material level in the annular chamber, which has an important effect on the stable operation of the device, would become very easy and convenient. The premise for putting such a scheme into practice is that the particles in the impinged suspension can be classified by gravity upon their moisture contents. That is, only the particles with moisture content lower than that specified can be carried by the radial gas flow to fly and to drop down into the discharging area outside the annular chamber; while those with higher moisture contents would drop down into the annular chamber to circulate again. The basis for this idea is that water in a porous particle does not change the diameter of the particle, so that the particles with lower moisture content have smaller density, undergo less influence of gravity, and thus can fly for a longer path.
148
IMPINGING STREAMS
Unfortunately, the experiments yield negative results: with the upper discharging port, the out-particles are obviously finer in size, the out rate is unstable, and the moisture content of the out material is higher and higher as the process proceeds. Even without analysis, it can be judged that the moisture content of the product does not meet that required. The following theoretical analysis is carried out to check the negative results above. The following assumptions are made: (1) The particles fly radially by the drag force of the gas flow and, at the same time, drop down by gravity; (2) Since the suspension is thin dilution, the interparticle action can be negligible and so the motion equations for a single particle are used; and (3) The influence of buoyancy is neglected. According to Elperin et al. [6], after impingement between the opposing streams, the radial velocity of the gas flow decays following the relationship below:
Uar = 1.5
Ua
exp - 1.52
(6.4)
With Eq. (6.4), the radial distance of the position where the maximum radial velocity of gas flow appears, rma~, can be calculated to be 1.974 times the diameter of the accelerating tube, Rac. Assume the particles are carried out of the impingement zone at r = 1.974Rac with an initial radial velocity of zero. From the force balance, the movement equations of sphere particles after leaving the impingement zone can be obtained as Horizontal direction:
d/Apr = --O.5CDrPaAp I.pr -/AarI(.pr -/gar)
mp dt
(6.5)
Vertical direction:
dupz dt
[ -mp -° oz a plUpI l.pz
(6.6)
The initial conditions of Eqs. (6.5) and (6.6) are T = 0: z = 0, b/pz- 0; ~ 1.974R, /'/pr-" 0
(6.7)
where the radial and axial drag coefficients are calculated by 24 CD~ = ~ ,
Repr
CD~ = ~
24
Repz
(6.8)
The mass of a particle is related to its diameter dp and moisture content x: Ic dp3pp ( l + x ) mp - -~
(6.9)
IMPINGING STREAM DRYING
149
where pp is the density of a dry particle. The relationship between the particle velocity and the distance the particle travels is well known as dr
it~, . . . .
dr
dz
,
Up~ = - - -
(6.10)
dt
Thus, the trajectories of particles with various diameters and moisture contents, after leaving the impingement zone, can be determined by solving simultaneously Eqs. (6.5). (6.6) and (6.10) with the initial conditions, Eq. (6.7). Because of gravity, the particles drop down as they fly outwards with gas flow so that they move along parabolas. It is clear that this flight cannot continue infinitely. Once a particle drops down to the surface of the material bed, its flight must stop. In other words, the vertical distance between the impingement plane and the surface of the material bed is the limitation of the particle's flight; and the difference between the radial distances traveled by various particles at the end of their flight is the possible maximum separating distance. The results calculated for the flight trajectories of particles with various diameters and moisture contents are shown in Fig. 6.22. The figure indicates that the radial gas flow exhibits certain classification effect for particles with various diameters. However, the moisture content of the particle has almost no effect on the flying distance. These theoretical results illustrate that the arrangement of the upper overflow discharging port is totally unfeasible. 0.()0
-5 0.04
6 7
0.08
dpx 104, m -
0.12 _
0.16
--
1
3.0
2 3 4 5
2.7 2.4 2.1 1.8
6
1.8
7 8
1.8 1.8
1
0.20 0.00
().()4
X~
0 0 0 0.06 0.04 0.02 0 J
] 0.08
I
[ 0. i 2
112
314 0.16
I 0.20
F, ITI
Figure 6.22 Result calculated for the flying trajectories of particles of various diameters and moisture contents.
150
IMPINGING STREAMS
6.3.6 A brief introduction to the quasi industrial test The results of the model experimental investigation indicate that the essential structure of the circulative impinging stream dryer designed by the author of this book is feasible for porous powdery and/or granular materials such as PVC produced by the suspension process. Its performance has achieved the expected goal: The feature of impinging streams enhancing transfer is fully utilized; while the residence time can be arbitrarily set so that the water, both free and in pores or bounded, can be removed in one device simultaneously. It can be expected that the development of its industrial application would yield obvious benefits in simplifying the system scheme and energy saving etc. In comparison with some advanced dryers, such as the whirlwind dryer, it has the additional significant advantage that the major part of the product is discharged from the bottom of the dryer, yielding a greatly reduced load for the dust collection system. On the basis of the investigation above, a quasi industrial test was carried out, as described briefly below. The equipment and system for the test were originally designed for PVC from the suspension process. Because of a requirement from the factory, the target material was then changed, and the new one was the highly chlorinated PV. The wet feed was the crystalline of highly chlorinated PV from the centrifuge containing large amounts of crystalline water and its moisture content was as high as 54 to 60% in wet basis. The other physical properties of the target material were: The mean diameter of particles dr= 350 ~tm, the density of particles pv= 920 kg.m -3, and the bulk density of dry product pb = 464 kg.m -3. The main parameters for the design were as follows: • Capacity of dry product: 0.139 kg.h -1 (1000 t-y-1) -1 • Velocity of gas flow in the accelerating tube: u0 = 25 m.s • Length of the accelerating tube: Lac - 0.6 m • Diameter of the accelerating tube: do= 0.15 m • Height of the annular chamber (cylinder): H = 0.7 m • Diameter of the annular chamber: D = 0.75 m • Impinging distance: S ~ 0.45 m (adjustable in a certain range) • Lower spacing: h ~ 0.225 m (adjustable in a certain range) • Temperature of hot air: Tg0 = 130-140°C The equipment system scheme is essentially the same as that shown in Fig. 6.14; but with two differences: (1) The orifice plates are used for metering airflow rates; and (2) Since the equipment is much larger than that used in the model investigation and therefore the feeding rate is much larger, the screw feeder for wet material feeding in the quasi industrial test is not as complex as that shown in Fig. 6.14 in structure, but is a common one. The primary operations of the equipment system yield positive results: (1) With the highly chlorinated PV crystalline from the centrifuge containing about 60% water as the feeding wet material, the final moisture content of the product from continuous
IMPINGING STREAM DRYING
151
operation of the system can achieve the specified index, i.e. no larger than 0.6%. (2) Although the capacity of the equipment cannot achieve 1000 tonne per year, as designed for PVC, for the change in the target material, its ability in terms of water evaporation exceeds the designed index by 20 to 30%. Development is still continuing; and it can be expected to be put into production in the near future.
6.4 CONCLUDING REMARKS Gas-continuous impinging streams (GIS) is a very effective technical method for enhancing heat and mass transfer between phases. Drying, being a typical process of parallel heat and mass transfer, is one of the areas where GIS can be expected to be applied successfully. Much research and development has taken place on the application of GIS in drying and many IS dryers with various structures and working principles have been proposed. Unfortunately, no successful application of GIS for drying in industry has yet been reported. Although many factors may account for this slow progress, the two most important problems that need attention are: (1) There may not be a thorough enough understanding of the properties of impinging streams, its advantages and disadvantages, resulting in improper selection of target materials and, consequently, giving unexpected results; and (2) Some engineering problems related directly to the application have not been properly solved so that no complete set of technologies can be provided to industries, delaying industrial application of GIS in the field of drying. Without any doubt, the application of GIS in the drying field is of great potential and it is reasonable to expect that drying technologies employing GIS will appear in various industries in the coming years.
This Page Intentionally Left Blank
-7IMPINGING STREAM ABSORPTION
7.1 ADAPTABILITY OF IMPINGING STREAMS FOR GAS-LIQUID REACTION SYSTEMS lkbsorption is another important unit operation process involved in the chemical, petrochemical, and a number of other processing industries. Since the discovery of the major effect of chemical reaction in promoting absorption, especially after the establishment of the systematical analysis method by Denckwerts [70], chemical absorption has been more widely applied industrially. It is clear that whenever an absorption treatment is needed in the processing industries, people always prefer where possible to employ chemical absorption. Of course, in the application of impinging streams in the area of absorption, what we are most concerned with is its application in chemical absorption too. This is the focal point of the discussion in this chapter. Chemical absorption is a kind of gas-liquid reaction and must involve transfer between phases, and thus is one of the important areas where, it is hoped, impinging streams can be applied successfully On the other hand, as mentioned before, the method of gas-continuous impinging streams has the outstanding advantage of significantly enhancing transfer between phases, while, at the same time it has the intrinsic disadvantages of very short residence time in the active region and relatively complex flow configuration, so that it cannot be applied for every gas-liquid reaction or chemical absorption system. Any suitable technology and/or equipment for a chemical absorption system is closely related to the nature of the reaction(s) involved in liquid [57]. To achieve success, the appropriate selection of the target systems of IS application must be made according to such nature, combined with the properties of impinging streams. According to Denckwerts [70]. the nature of a gas-liquid reaction system can be characterized by the parameter M, which is defined as: m =
Possible maximum reaction rate in liquid film Possible maximum rate of mass transfer through liquid film
(7.1)
Parameter M ha~ different definitions for different types of reactions, and various definitions can be found in ReE [57] or other textbooks and monographs on chemical ~eaction engineering.
154
IMPINGING STREAMS Table 7.1
Features of reactions in liquid phase at various values for M Range of M value
Feature of reaction
81 vs 6L *
Reaction region
~/M >>3
instant
61 ~ 0
at the interface
~/m > 3
fast
81 < ~
inside liquid film
x/M -- 1
middle
81 -- ~
up to liquid film
x/M <0.3
slow
61 > ~
over liquid film
x/M <<0.3
extremely slow
81 >> ~
whole liquid phase
* ~--thickness of the reaction layer; ~--thickness of liquid film The characteristics of liquid reaction with various values for M are listed in Table 7.1. It can be seen that, in the last two cases, i.e. ~ <0.3, the processes are controlled by reaction kinetics and so enhancement of transfer becomes of no use; in the medium case of ~ ~ 1, both diffusion and reaction kinetics affect the overall rate of the process, a measure of enhancing transfer may have a certain positive effect; while in the former two cases, i.e. x/M >3, the reaction(s) in liquid proceed fast, the global processes are controlled by diffusion, and thus the measure of enhancing transfer will play a positive key role. Although the parameter M does not involve diffusion through the gas film directly, it has important referential value for the selection of the target system for impinging streams application in the area of absorption, because the diffusion resistance of a gas film has an order of magnitude comparative to that of liquid a film in most systems of practical interest. It can be concluded from the simple analysis above that impinging streams can only be used for gas-liquid reaction or chemical absorption systems involving fast reaction(s) in liquid for success. Concerning the flow configurations, it is clear that impinging streams with gas as the continuous phase is most suitable for chemical absorption, while, with liquid as the continuous phase, impinging streams cannot generally give a perfect performance. In an absorption process with gas-continuous impinging streams, the liquid is usually atomized into fine droplets. For the absorptions involving fast reaction(s) in liquid, the atomization of liquid provides a large interface area for transfer between gas and liquid. It can be expected that with the effect of impinging streams enhancing transfer, chemical absorption processes can be greatly intensified. On the other hand, the flow configuration of impinging streams is relatively complicated so that it is difficult and usually unfeasible to arrange a countercurrent multistage system, such as in a column device. In addition, there is strong mixing in the active region of an impinging stream device. Both factors are unfavorable for absorption systems involving reversible reaction(s) in liquid, even if they are fast ones.
IMPINGING STREAM ABSORPTION
155
These systems are subject to the limitations of equilibrium, and thus it is difficult to achieve the higher absorption efficiency or conversion required in a single stage impinging stream device with strong mixing. From the discussions above, the following general principle for selection of target systems for IS application can be concluded: the gas-continuous impinging streams method is especially applicable to gas-liquid reaction or chemical absorption systems involving fast-irreversible reaction(s) in liquid.
7.2 EARLIER INVESTIGATIONS Investigations on impinging stream absorption began in the mid-1980s and were mainly concentrated in Israel. Up to the 1990s the work carried out was essentially on the fundamentals and focused mainly on analyzing and verifying the enhancement of transfer by impinging streams and searching for related experimental evidence. Few researchers gave detailed consideration to the feasibility or even the possibility of impinging streams application for the target systems so that, essentially, the results obtained cannot be taken as a basis for further development. However, those works provided some referential values for later investigations.
7.2.1 Models for absorption enhancement In impinging streams with gas as the continuous phase, there exist all the factors enhancing transfer mentioned in the previous discussions related to gas-solid systems. In addition to these, Tamir [5] considered that the following factors may further enhance transfer in gas-liquid impinging streams:
(1) Because of the collisions between droplets and the shearing effect of the gas flow carrying the droplets, the original droplets may be re-atomized, yielding an increased interface area. (2) The collision, deformation of droplets, effects of shearing force between droplets and gas flow, and surface tension are factors that cause circulation of liquid at the surface of and inside droplets, favoring surface renewing and, consequently, promoting transfer between phases. However, as described in Chapter 5, because the aggregation status of liquid is different from that of solid, both re-atomization and coalescence are possible in gasliquid impinging streams with liquid as the dispersed phase. This introduces some complicated uncertainty factors. According to the results obtained, it is uncertain whether droplet re-atomization increases the interface area for transfer, because coalescence of fine droplets decreasing the interface area counteracts or even exceeds the positive effect of re-atomization. More possibly, a negative effect may be obtained, i.e., the interface area may be reduced to a certain degree.
156
IMPINGING STREAMS
Tamir [5] analyzed the effects of impinging streams enhancing physical and chemical absorption processes. To describe the enhancement of absorption, the following two enhancements were defined to account for the two factors: oscillation movement and re-atomization-coalescence of droplets, respectively E1=
Mass absorbed in the presence of oscillations Mass absorbed in the absence of oscillations
(7.2)
and E2 -
Mass absorbed in the presence of re - atomization or coalescence Mass absorbed in the absence of re- atomization or coalescence
(7.3)
Based mainly on the analytical results for single particle motion in impinging streams, Tamir derived a number of expressions for the two parameters for various flow regimes in the two cases with and without chemical reaction, in which the parameters such as the droplet size, the motion times of a particle in the accelerating and decelerating stages, particle to gas velocity ratio at the outlet of the accelerating tube, etc. were involved (see Eqs. 11.2 to 11.25 in Ref. [5]). Unfortunately, those models may contain a number of defects. Firstly, the influence of the relative velocity between phases on transfer coefficients has not been considered, while such an influence is just the most intrinsic reason for impinging streams enhancing transfer processes. Secondly, the assumption on the re-atomization and coalescence of droplets is short of both theoretical accordance and experimental evidence. These, plus the randomness of both dispersity and motion of droplets, make the models generally less general meaningful and difficult to apply. The following fact might be of interest: Tamir carried out a number of experimental studies in order to verify the effect of impinging streams enhancing absorption processes [106-108], while the results were essentially independent of the models mentioned above.
7.2.2 Absorption equipments Generally, absorption equipment with impinging streams includes two essential elements" the atomizer and the absorption chamber.
7.2.2.1 Atomizers For absorption processes carried out in impinging streams with gas as the continuous phase, atomization of the liquid is an essential operation. In earlier investigations on impinging stream absorption, all the liquids were atomized by pneumatic nozzles. The nozzles used were mainly of two types, the first being shown in Fig. 7.1a. With this atomizer, the gas contacts and mixes with the liquid outside the nozzle and is generally called the "external mixing nozzle" in industry, although Tamir and co-workers called it the "no-mixing nozzle". The second one is the internal mixing nozzle of Caldun CSL2 type, i.e. the so-called Critical nozzle [59-00, 109], which has a special structure,
IMPINGING STREAM ABSORPTION
157
as shown in Fig. 7.lb. This special structure ensures that the gas contacts and mixes fully with the liquid inside the nozzle, while yielding larger resistance as well. From the point of view of transfer, the mixing chamber inside the nozzle is also a considerably active region, in which the fresh gas contacts the fresh liquid, both with large driving forces for transfer, so that a significant part of the absorption must be carried out in it. Unfortunately, to date no data on the states of gas and liquid at the outlet of the nozzle have been measured so that the contribution of the mixing chamber to absorption cannot be determined. Gas
.... Liquid[ ,'-~.... ,%%%
f i
(a)
External mixing type
Liquid
(b) Internal mixing type Figure 7.1 Pneumatic atomizers. In general, a pneumatic nozzle can produce sprays of fine droplets to provide a large interface area for heat and mass transfer; but the power consumption for atomization is very high. In some cases, e.g., when it is used in technical equipment for environmental protection to remove harmful gases, its high power consumption may become a significant economic problem.
7.2.2.2 Structure of absorption equipment In earlier investigations aimed mainly at verifying the enhancement effects of impinging streams, the common horizontal two impinging streams was mostly employed, although some other flow configurations were also sometimes used.
158
IMPINGING STREAMS
~ Liquid
~ Liquid
~ Gas
~ Gas
(b) With a partition
(a) Normal
• Liquid
Liquid
~ Gas
Cas
i'-
....
r
....
I
'
"
I
(c) With concentric nozzles
(d) With eccentric nozzles
Figure 7.2 Two-impinging stream absorber for demonstrating enhancement. The structures of the experimental equipment used in the investigations by Tamir et al. are shown in Fig. 7.2. Figure 7.2(a) shows the common co-axial two impinging
stream absorber, in which the major part of the gas does not pass through the nozzles but enters the absorption chamber concurrently with the sprayed liquid. The absorber shown in Fig. 7.2(b) has a partition at the middle to separate the two opposing streams and the rest is completely the same as shown in Fig. 7.2(a). The researchers aimed to verify the enhancing effects of impinging streams for absorption by comparative experiments carried out in the absorbers shown in Figs. 7.2(a) and 7.2(b). The absorber shown in Fig. 7.2(c) also has the flow configuration of co-axial two impinging streams, but all the feeding gas is used as the atomization medium; it is called "concentric nozzles". In that shown in Fig. 7.2(d) all the feeding gas is also used as the atomization medium, but the two nozzles are placed eccentrically and so the device is called "eccentric nozzles". Tamir et al. [109] also studied an impinging stream absorber operated in bubbling mode, as shown in Fig. 7.3. The absorber takes liquid as the continuous phase while gas is dispersed in liquid, so it actually belongs to liquid-continuous impinging streams (LIS). The experimental results obtained showed that this flow configuration exhibits a higher absorption rate than that shown in Fig. 7.2(a). Combining them with the results
IMPINGING STREAM ABSORPTION
159
from an investigation on micromixing carried out by the author of this book [110], it might be considered that the major reason for increasing absorption rate in the absorber shown in Fig. 7.3 is that strong micromixing in the continuous liquid phase yields a decreased resistance of liquid side. Obviously, this mechanism is different from that of gas-continuous impinging streams enhancing transfer. This flow configuration is of interest for the gas-liquid reaction processes for which the bubble-bed reactor is suitable, but the other aspects of the reactor have not yet been evaluated. Bubble /
G/2
G/2
n
W/2
,-
8888-/00888 o O8o8oO O8o
_
o o° o
W/2
Figure 7.3 Bubbling impinging stream absorber.
T
G
L #-
1 4
L
Figure 7.4 Impinging stream loop reactor. Probably inspired by the fact that the Jet Loop Reactor has successfully been applied in industry, Gaddis and Vogelpohl [111] proposed an impinging stream loop reactor, as shown in Fig. 7.4. It seems that their main purpose is to lengthen residence times in the reaction vessel. The principles of the reactor's operation are somewhat similar to those of the Air Left Reactor (ALR). The only difference lies in the fact that
160
IMPINGING STREAMS
it also employs a flow configuration of impinging streams. During operation, two gas flows are fed into the nozzles on two sides through separate tubes, and then mix with the liquid flowing through the nozzles at their exits. By strong shearing force the gas is dispersed into liquid to form two gas-liquid two-phase streams of lower density, and the latter flow upwards at considerably high velocity, sucking the liquid or gas-liquid mixture from the main tube of the reactor to cause internal circulation inside the reactor. Considering the flow and mixing status, it can be seen that the liquid is not a unique continuous phase in the mixture before impingement, while the gas may also be such a phase or, at least, not a fine-dispersed phase. This is the specialty of this reactor. The difficulty in operation of the reactor is that the impinging velocity cannot achieve higher levels and so the effect of impinging streams enhancing transfer is limited. Using a cycling pump may be a possible way of solving this problem and, in fact, the researchers have done this, but it led to an increased consumption of energy and a more complicated system. In addition, Ponikarov et al. [112] studied more special impinging stream absorption equipment. It employs the flow configuration of rotating impinging streams and the impingement occurs in a collision chamber of half-circle form. This equipment appears to be of less practical interest and so will not be discussed further here.
7.2.3 Major results of the investigations The following systems were studied by the researchers mentioned above: absorption of CO2, acetone and ammonia into water, and absorption of CO2 into NaOH solution. According to the nature of the reactions in the liquid phase involved, single stage impinging streams is only applicable for the absorption of CO2 into NaOH solution; while from the point of view of economics, using NaOH to absorb CO2 is generally unfeasible. The investigations on impinging stream absorption of the systems above therefore have little practical interest. In fact, earlier investigations on this topic focused mainly on verification of impinging streams enhancing absorption and obtaining fundamental data, while little attention was paid to feasibility of application. The following are the major results obtained.
7.2.3.1 Experimental evidence for IS enhancing absorption The results of the comparative experiments on the absorption of C O 2 and acetone [107, 108] into water are: the absorption rate in the absorber without partition shown in Fig. 7.2(a) is higher than that in the absorber with a partition shown in Fig. 7.2(b), indicating the enhancing effect of impinging streams. However, as mentioned in the Introduction, the basic standard they used for comparison is unreasonable. In the absorber with a medium partition, the flow configuration becomes two jets impinging fixed wall surfaces separately, and the latter enhance transfer significantly, too. In other words, the intrinsic enhancing effect of IS should be stronger than that reported. For example, the impinging jet on a fixed wall may enhance the absorption rate by x times
IMPINGING STREAM ABSORPTION
161
than normal, and IS enhance by y times than the impinging jet on a fixed wall; so the total number of the times enhanced intrinsically by IS than the normal should be xxy.
7.2.3.1 Mass transfer coefficient Obviously, mass transfer coefficient is a topic of general interest. Tamir, Herskowits et al. [59, 106, 107, 109] studied experimentally the absorption of CO2 and acetone into water in a two impinging stream absorber operated in various modes with various atomizers. The data they measured for the volumetric mass transfer coefficient are listed in Table 7.2, which are representative among earlier investigations. As mentioned, from the point of view of practical application, impinging streams is not suitable for the systems given in Table 7.2. On the other hand, the absorption processes for which impinging streams is applicable normally involve fast reaction(s) in liquid and thus are controlled by gas-film diffusion. Therefore the most important should be the gas-film mass transfer coefficient, kG, which is absent in the table.
Table 7.2
Volumetric mass transfer coefficient in two impinging stream absorber
Nozzle type
Operation mode
Volumetric mass transfer coefficient, s-~
Ref.
CO2--H20
External mixing
Spray
0.041 < kLa <0.37
[ 106]
CO2--H20
Critical nozzle
Spray
0.176< kka <0.611
[59]
CO2--H20
Critical nozzle
Bubbling
0.411 < kka <0.951
[59]
Acetone -- H20
External mixing
Spray
0.07< Koa <24
[107]
Acetone--H:O
Critical nozzle
Spray
0.9< Kca <9.8
[109]
System
7.2.3.3 Others The influences of the liquid and gas flow rates, the diameter of the absorption chamber, the distance between nozzles, and the flow configuration on absorption rate were studied by the researchers mentioned above. These will not be discussed in detail here because of the length limitation of the chapter; for the details, the reader may refer to the original references as cited in the text above. It should be noted, however, that in all the investigations above, the data for mass transfer coefficients are always correlated with the gas and/or liquid flow rates, but not with the impinging velocity, u0, although the latter is the operation parameter extremely important in every impinging stream device.
162
IMPINGING STREAMS
7.3 WET DESULFURIZATION OF FLUE GAS (I) GENERAL CONSIDERATIONS Air pollution caused by SO2 in flue gas from coal- and oil-burning is a global issue and thus receives more and more attention and the governments of many countries have passed environmental protection legislation which requires that air pollution problems be solved. In the past few decades, a number of schemes have been proposed for this purpose, such as fuel pretreatment, concurrent burning and adsorption, flue gas post treatment, etc. [113-115]. Among them, flue gas desulfurization (FGD) is the most reasonable from the points of view of both technology and economics, making it the most practically applicable. Different categories of processes, such as dry, semidry, and wet processes, have been developed for FGD; wet processes are the most promising due to their lower operating cost, higher efficiency, more stable operation, and because there is no secondary dust pollution of the environment, and therefore are more widely applied. Since the goal of a wet-process of FGD is for environmental protection only, with no value-added product, cost and performance are the most important factors to be considered and both are closely related to the type of absorber. Therefore, the search for high-performance absorption equipment is of great importance. Several types of equipment, e.g., packed column [116], spray tower [117], and swirling flow plate column [ 118] etc., were tested or even employed industrially, but none of them can be considered as ideal. On the other hand, for many years, no investigations had been reported on the application of a method effectively enhancing transfer, like gascontinuous impinging streams, for FGD. By the end of the 1990s, Tamir organized studies on this topic and recently the author of the present book carried out a test of FGD by impinging stream absorption on a small pilot plant scale. It can be considered that a gas-continuous impinging stream device is the best option for wet FGD; on the other hand, from the point of view of IS application, taking FGD as the applied objective is also the most reasonable selection. Many kinds of absorbents have been tested for wet FGD. From the points of view of a low-cost, easily available raw material with high absorption efficiency and, consequently, a smaller device requirement and lower operation costs, hydrated lime may be the best. In the wet GFD process by absorption with Ca(OH)z-in-water suspension as the absorbent, the following reactions are generally considered to occur in the liquid phase: Dissolution and dissociation of SO2: SO2 ( G ) = SO2 (L)
(7.4)
SO 2 + H20 - HzSO 3 = H + + HSO 3 - 2H + + SO z-
(7.5)
Dissolution and dissociation of the absorbent: Ca(OH)2(S) - Ca(OH)2(L) - Ca 2+ + 2OH-
(7.6)
IMPINGING STREAM ABSORPTION
163
Reaction to yield products of desulfurization: Ca ~+ +SO~- + 0 . 5 H 2 0 - C a S O 3 •0.5H20 $
(7.7)
CaSO3-0.5H20 + 0.502 + 1.5H20 - CaSO4.2H20$
(7.8)
In the system under consideration, the main reaction in liquid, Reaction (7-7), is certainly instantaneous and irreversible, while the mid-reactant Ca 2+ needs to be supplied continuously by the dissolution of Ca(OH)2, which might not be fast enough. Scala [119] and Scala et al. [120] modeled the general case of absorption with instantaneous reaction in a droplet with sparingly soluble fines and, as an example, simulated numerically the absorption of SO2 into an aqueous slurry droplet of Ca(OH)~ at ambient conditions in spray dry desulfurization. The results they obtained showed a significant concentration gradient of Ca(OH)2 inside the droplet, although they were supported by no experimental evidence. The significant gradient implies that absorption would be affected by dissolution of Ca(OH)2 and/or diffusion of Ca 2+ through liquid film to some extent. However, some assumptions Scala made are invalid, at least, ~br the case to be studied, as described below. On the other hand, Scala's approach is classified as a semidry-process of FGD, which normally exhibits lower sulfur-removal efficiency and much higher consumption of Ca(OH)z, although it may save some water to a certain degree. This may be why this method could not be as widely applied in industries as wet processes. The case Scala studied is quite different from that in the wet FGD by absorption with Ca(OH)2 suspension as the absorbent. In the latter process, in order to ensure full contact between gas and liquor suspension, it is important to employ a considerably high liquid to gas flow rate ratio, around 1.0×10 -2 m~.m-l, yielding very low concentration of Ca(OH)2 required for the absorbent. For example, for the representative case where SO~ concentration in flue gas is around 5000 mg.m -~ the concentration of Ca(OH)2 in the absorbent required for a suitable Ca/S mole ratio is only about 0.7 to 0.8% and the density of the suspension is as small as 1050 kg.m -3, suggesting a very thin dilution. In comparison, in the typical case considered by Scala, the density of the suspension was as high as 2200 kg-m -~. So, in the wet FGD with a thin Ca(OH)2 suspension as the absorbent, the following can be expected: (1) The amount of SO2 dissolved in the liquid phase per unit Ca(OH)2 in the same time interval is much larger than in Scala's case, especially at higher pH. (2) The reaction between dissolved SO2 and Ca(OHh must promote dissolution of Ca(OH)2. In fact, Ca(OH)2 has a non-zero solubility in water (about 0.0017 kg per kg water at room temperature), and the fast reaction must: move the equilibrium of Ca(OH)2 dissolution, promoting the latter to dissolve. (3) The circulation motion inside the droplets at very high relative velocity in gas-continuous impinging streams can never be negligible [5], which also promotes dissolution of Ca(OH)2. (4) In the Ca(OH)2 suspension newly prepared by digestion of CaO with water, which is normally used for absorption, the Ca(OH)2 particles newly formed are very fine (sized less than 1 ~tm) and their surface is more active, resulting in their much easier dissolution and thus much easier reaction with
164
IMPINGING STREAMS
S O 2. Therefore the system can be considered to meet the criteria of ~ > 3, and the absorption is assumed to be totally controlled by diffusion of SO2 through gas-film, as was done by Berman et al. [25]. Another feature of the system under consideration is that solid products are formed during the process. This may give rise to some engineering problems which must be addressed.
7.4 WET DESULFURIZATION OF FLUE GAS (11) INVESTIGATIONS IN ISRAEL 7.4.1 Experimental equipment and procedure As stated above, the key elements of a gas-continuous impinging stream absorber are the absorption chamber and the atomizers. The structure of the experimental absorber used by Berman et al. [25] is shown in Fig. 7.5 in which Reactor 1 is made of organic glass, with a structure of co-axial cylinders. There can be three cylinders at most which are called, in order from outward, Cylinders I, II, and III. There are several passage holes for flue gas and absorbent suspension from one cylinder to the next appropriately arranged on the walls of the cylinders. The structure of the atomizing nozzles used is shown in Fig. 7.6. Pure air is used in the nozzles for the primary atomization and the atomized absorbent mixes with the flue gas after exiting from the nozzles. A replaceable nut (5) makes it possible to change the diameter of the flue gas exit, d, within the range 5 to 10 mm, and thus to vary the velocity of the mixture. According to its working principles, the nozzle shown in Fig. 7.5 should be classified as a three-flow internal-external mixing nozzle; Berman and co-workers called it a "complete external mixing type". The experimental absorption process proceeds as follows: The pseudo flue gas consisting of SO2, N2 and CO2 enters the nozzle and then ejects from the flue gas chamber, while the absorbent, Ca(OH)2 suspension, is atomized by the pure air flow to form a spray and is then ejected from the nozzle. After mixing of the absorbent spray with the flue gas, the two-phase streams formed on two sides enter the reactor, where the impingement between the opposing streams and the reactions take place. The mixture containing both tiny droplets and gas is thrown onto the wall of Cylinder I to form a turbulent film moving along the cylinder wall. The flue gas and absorbent leave Cylinder I through the passage holes to Cylinder II, and then to Cylinder III where absorption takes place continuously in the films on the walls of the cylinders. After Cylinder III, the liquid is collected in the funnel under Reactor 1, while gas is discharged after further treatment. Samplings of the flue gas are made at the inlet and outlet of the absorption chamber and the contents of SO2 in the samples are determined by titration with Ba(C104)2.3H20, with an accuracy of + 3-5%. The reproducibility of data on sulfur-removal efficiency is within the range +_4% to 6%.
IMPINGING STREAM ABSOR PT~o'N
165 l Flue gas
C~eaned gas
Sampling
j4
[-
2 I I~ __~
~
~1
I 2
I
Reacted suspension
Absorbent
t
Pure air
Figure 7.5 Experimental equipment used by Berman et al for FGD [25]. 1-reactor; 2-spray nozzles" 3-collection chamber; 4-fabric filter.
Atomizing air
Flue gas
.....
0a
•"'"".............................................,
~
Absorbent
......
II ,..,..,~y.............. . . 9:~';.'.....'X..'.'..'.5"Z,';-".." .... ..
.
.
.
.
.
.
.
..,::.::~
.
"/";4-';, \
•3
'~ 4
Figure 7.6 Spray nozzle used by Berman et al [26]. l-inner nozzle; 2-chamber for atomizing air; 3-chamber for flue gas feeding; 4-sealing; 5-replaceable nut.
166
IMPINGING STREAMS
According to the structure of the equipment shown in Fig. 7.5, absorption can take place in the following three regions: (1) Round the outlets of the nozzles. This region has a very short residence time although; its contribution to absorption is not negligible because both the driving force for the reaction(s) and the relative velocity between phases are considerably large. (2) The impingement zone inside Cylinder I, i.e., the major active region of enhanced transfer. (3) On the walls of the cylinders. This region has a small specific interface area and a significantly smaller driving force for both reaction and mass transfer for much lower concentration of SO2 in flue gas left after absorption in the previous two regions, but the residence time in it is very long compared with the previous regions. It may be considered from the analysis described above that the contributions of the three regions to absorption are comparable. In other words, it is uncertain whether the effect of impinging streams is the lynchpin or not for carrying out absorption in such a device.
7.4.2 Major Results One of the major results Berman et al. obtained is that the mole ratio of Ca in the absorbent to S in the flue gas has the most important effect. The experimental results for the influence of the Ca/S ratio on the sulfur-removal efficiency, r/s, at different concentrations of CO2 in the flue gas are shown in Fig. 7.7. These data were obtained in a reactor with two co-axial cylinders; the experimental conditions were: flue gas flow rate V~ = 0.001 m3.s-1, diameter of Cylinder I in the reactor D~ = 0.06 m, diameter of flue gas exit of the nozzle d - 10 mm, clearance between Cylinders I and II A2 = 5 mm. The results in Fig. 7.7 show the significant influence of CO2 in flue gas on the sulfur-removal efficiency r/s. The reason for this is clear: CO2 reacts with the absorbent Ca(OH)z, too: Ca(OH)2
+ C02
--
CaCO3 + HzO
and thus results in an additional consumption of Ca. As observed, the dependence of r/s on Ca/S exhibits asymptotic behavior. This is consistent with the results obtained by Yoon et al. [121], Sahar et al. [122] and Newton et al. [123]. On the other hand, the asymptotical tendency of r/s variation implies that, when the ratio of Ca/S is over a certain critical value, (Ca/S)c, the sulfur-removal efficiency r/s is kept constant, independent of Ca/S. This is a very important conclusion of practical interest. Further, Berman et al. [25] determined the values for (Ca/S)c at various concentration of CO2 in flue gas, as follows: Cco2 - 0, (Ca/S) c - 1.6; Coo 2 - 10%, (Ca/S) c - 1.8; Cco2 - 15%, (Ca/S)c - 2.0
IMPINGING STREAM ABSORPTION
167
100 •
~-
80
60 0.8
A
R
s/
• CO2=0 • CO2=10% i CO2= 15%
[ 1.0
I 1.2
I 1.4
I 1.6
[ 1.8
[ 2.0
2.2
Ca/S, mol.mol -~ Figure 7.7 Dependence of absorption efficiency of SO~ on the ratio Ca/S for different concentrations of CO2 in flue gas [26].
In addition, the influences of some of the operation conditions and the structural parameters of the reactor were also studied. The major results they obtained were as follows" (1) In the range of the flue gas flow rate from 0.5x 10-3 to 1.5xl 0 -3 m3.s-~, the sulfurremoval efficiency obtained in the reactor of a single cylinder drops slightly as the flue gas flow rate increases, while in the reactor of two co-axial cylinders this efficiency is essentially independent of the flue gas flow rate. (2) In the range of 7 to 20 m.s -~ of the flue gas flow exit velocity from the nozzles, the sulfur-removal efficiency increases as the velocity increases, but the tendency to increase gradually becomes weaker. (3) In the reactor with two co-axial cylinders the sulfur-removal efficiency increases slightly as the diameter of Cylinder I, Dj, increases, while in the reactor of a single cylinder D~ has no significant influence on the sulfur-removal efficiency. (4) In the comparative experiments the reactor with two co-axial cylinders exhibits obviously higher sulfur-removal efficiency than that of a single cylinder.
The results described in item (4) demonstrate that in the absorption equipment used by Berman et al. the third region mentioned in Section 7.4.1, i.e. the walls of the cylinders, makes a significant contribution to the absorption. The researchers interpreted the mass transfer coefficient with the mass transfer model for absorption, and the resulting empirical relationships represent the Sherwood number Sh as a function of Reynolds number Re. For details, the reader may refer to Ref. [25]. It is clear that the items included in the relationships they obtained are related closely to the reactor structure.
168
IMPINGING STREAMS The optimal operating conditions determined by Berman et al. are as follows: Number of cylinders in the reactor Flue gas exit velocity from the nozzles Consumption of absorbent suspension per unit flue gas Ca/S Ratio (for CO210% in flue gas) Flow rate ratio of atomizing air to flue gas Specific productivity of the reactor Hydraulic resistance of the reactor Desulfurization efficiency, r/s Degree of absorbent utilization
2 10-15 m.s -1 0.25× 10-3 m3-m-3 1.8 mol.mol -l 0.08 m3.m-3 1.1×104 m3.m-3.h-~ 1.1 kPa 94-97% 52-54%
It can be considered that the experimental data obtained by Berman et al. are good. However, there are a number of disadvantages and difficulties in the scheme from the point of view of engineering application. Firstly, the multi co-axial cylinder reactor contains a number of internal parts, leading to a much larger surface area of the walls on which the newly precipitated solid may cake, and also increased difficulty in cleaning the fouling unavoidably formed during operation. Secondly, in this approach the use of pneumatic nozzles of "external mixing" type for atomization, with compressed air as the primary atomizing agent, results not only in a greatly increased gas load of the system but also a complicated scheme. In addition, for a coal burning power plant, e.g., generating around 4000 m3.h-~ of flue gas per MW, it is impossible to have the entire flue gas passing through the nozzles. Even if a large number of nozzles are used, there would be a great deal of difficulty in designing and manufacturing nozzles to meet such a requirement. Also, a pneumatic nozzle is one of the more energy-consuming types of atomizer. In addition to the works introduced above, Berman et al. also studied the absorption of SO2 with limestone and NaOH-enhanced limestone suspensions as the absorbents [124] in a reactor of the same type as that shown in Fig. 7.5. The main results obtained include the following: (1) The major factor determining FGD process is absorbent type and the ratio of (absorbent/S). (2) For the absorbent Ca(OH)2 or Mg(OH)2, the degree of absorbent grinding must correspond to a specific surface area not smaller then 10-12 mZ.g-j (measured by BET). (3) At a constant ratio (absorbentJS) - 1.6-1.8 and with the condition 3 above, the sulfur-removal efficiency, r]s, may reach 94-96%. In this case, the flue gas flow rate, SO2 concentration in it and the specific suspension consumption, in a range greater than 0.2×10 -3 m3.m-3, do not influence the desulfurization process. (4) The employment of CaCO3 as absorbent reduces r]s up to 74% for CaJS - 1.6; when CaJS = 3.8 r]s - 84%. The addition of NaOH in the range of NalCa - 0.075-0.1 mol.mol -~ increases r]s up to 99%.
IMPINGING STREAM ABSORPTION
169
7.5 WET DESULFURIZATION OF FLUE GAS (111) INVESTIGATIONS IN CHINA Based on investigations into understanding the nature of impinging streams, the Impinging Stream Gas-Liquid Reactor has been developed by the author of the present book and co-workers [125]. It is mainly suitable for systems involving fast reaction(s) in liquid. In order to develop application, it is employed for the wet desulfurization of flue gas, and tests at small pilot plant scale have yielded results [ 126, 127]. The method of scaling-up and a number of engineering problems that may possibly be encountered in practical application have been carefully considered in the equipment design and investigation. It can be expected to be applied industrially in the very near future.
7.5.1 Experimental Equipment The adaptability of impinging streams to absorption systems and the features of the reactions involved in the wet desulfurization of flue gas with Ca(OH)2-suspension as the absorbent have been discussed in detail in Section 7.3. It is certain that the application of impinging streams for wet desulfurization is a good option.
7.5.1.1 Consideration of the problems related to equipment design Besides the appropriate selection of reactor type, the two problems can be highlighted: (1) Atomization of Ca(OH)2 suspension" and (2) Caking on the walls and cleaning. In the operation of gas-continuous impinging stream absorption system, the liquid first needs to be atomized into the gas flows to form droplets-in-gas suspension streams. In earlier investigations [5] all the atomizers used were pneumatic, although they had various structures. The use of a pneumatic nozzle in such equipment leads to high power consumption, giving rise to economic problems. In fact, among various methods of atomization, the pneumatic one exhibits the highest energy consumption. Also, if additional compressed air is used as the primary atomization agent, the gas load of the system is greatly increased and the system becomes more complex. In the investigation led by the author of this book centrifugal pressure nozzles are used for this purpose" these are discussed in detail later. Another characteristic of the system under consideration is the formation of the solid products CaSO~.0.5H20 and CaSO4.2H20 making fouling of newly formed solid particles on the walls unavoidable. In order to lessen caking and make foul cleaning easier, there should be as few internal parts as possible inside the absorption equipment, and they should be as simple as possible in form.
7.5. 1.1 Conditions for experimental equipment design To obtain results of interest for the design and scaling-up of industrial devices, experiments were carried out at small pilot plant scale. The conditions for the equipment and system design are listed in Table 7.3.
170
IMPINGING STREAMS 7.3
Table
Basic conditions for reactor and system design Atmosphere
Temperature of flue gas Flow rate of flue gas, m3.h-1
300-500
Content of SO2 in gas, mg.m-3
1200-3000
Operating pressure
Atmosphere
Concentration of Ca(OH)2 in absorbent, kg.kg-1
0.01-0.05
7 . 5 . 1 . 2 Atomizers
In the absorption device the eddy pressure nozzles patented by Wu et al. [128] are used for atomizing the Ca(OH)2 suspension, the structure of which is shown in Fig. 7.8. Two essential conditions for a pressure nozzle operation are liquid or aqua suspension flow rotating and ejecting at very high velocity (several decades m.s-~), so this type of nozzle is usually called a Centrifugal Pressure Nozzle. According to common empirical data, the power consumption for a centrifugal pressure nozzle is only about 7-8% that for a pneumatic nozzle. Furthermore, the nozzle shown in Fig. 7.8 requires about 2030% less energy than common pressure nozzles. This is because it employs a highly efficient flow rotating chamber of half empty spherical shape and has a compact structure with a very short passageway for flows at high pressure. It has been used in a number of engineering projects, such as spray granulation in fluidized beds etc., and has performed well. The nozzle head can now be made of abrasion resistant materials such as high-grade ceramics so that the problem of the orifice wall being abraded has been effectively solved.
Body Conduit ~/plate
,
~ ~
Figure
2
~ /
~J" ~
~,
Pressing //cover
ozz,e
~K-N\\\~
~ 7.8 The eddy pressure nozzle.
he ad
Orifice
171
IMPINGING STREAM ABSORPTION
7.5.1.3 Structure of GIS gas-liquid reactor and its improvement The gas-continuous impinging stream gas-liquid reactor for the experiments of wet desulfurization of flue gas employs the flow configuration of horizontal coaxial two impinging streams, as shown in Fig. 7.9. Two eddy pressure nozzles are mounted co-axially inside the gas conduits on two sides of the absorption chamber. The pseudo flue gas is divided into two equal streams, which carry the atomized fine liquid droplets and flow through the two gas conduits, respectively, at essentially the same rate, and then eject from the conduits and impinge against each other at the center of the chamber to form a highly turbulent region, the impingement zone, where the major absorption is carried out. After absorption of SO2, the solid phase inside the droplets mostly becomes fine particles of CaSO3.0.5H20 and CaSO4-2H~O. Part of the droplets drop down by gravity, and are then discharged through the outlet tube at the bottom of the absorption chamber as the waste liquor (for which further treatment is not considered here because there successful methods already exist). A liquid -sealing mechanism is arranged to prevent shortcut of flue gas to atmosphere. After the primary separation of droplets by gravity, the gas flows upwards and is further separated from the residual droplets by the damper. The gas is then discharged from the top of the chamber to the atmosphere. _
Cleaned gas Damper
~ Absorbent
~~Absorption ~ - ~ ~
I ' ~ ' - ' - D. ~ ~ I
I I
All
Flue gas
Absorbent
:'i'ii ~ l l _
;:1
/Gas-conduit
Vtue
gas
~ ~ . ~
Impingement [ zone
~
~ T I !/,~/
......,t B ~ S '
~J
chamber
J
Nozzle
T
Waste T tluid
Figure 7.9 A brief view of GIS gas-liquid reactor. The major dimensions of the reactor designed according to the conditions listed in Table 7.3 are" diameter of the absorption chamber D , - 700 mm, height of the cylinder vessel H~- 950 mm, diameter of the gas conduits do- 80 mm, height of the gas conduit
172
IMPINGING STREAMS
axis from lower edge of the cylinder vessel h = 400 mm, and impinging distance S is adjustable from 4d0 to 6d0. The major feature of the reactor shown in Fig. 7.9 is its simple structure with few internal parts. This not only reduces the walls on which the solid particles may cake, but also favors increasing the time-averaged effective interface area for transfer because the droplets can essentially keep their fine-dispersed condition during the whole period of contact with the gas flow. For flue gas containing ashes it has the additional function of wet dust-removal. In the primary design, following that of the solid-gas impinging stream contactor used in earlier investigations on RTD and hydraulic resistance etc., the two nozzles are placed in the middle of the gas conduits (Point A, represented by the solid lines in Fig. 7.9), i.e., there is a piece of conduit downstream each nozzle for acceleration of droplets by the gas stream. The experiments show that the cross section areas of the sprays do not match those of the conduits: the former are significantly larger than the latter. As the result a considerably large part of the liquid is sprayed onto the inside wall of the conduit to form a liquid film flowing outwards along the wall and then dropping down to the bottom of the absorption chamber, as shown in Fig. 7.10. This part of the liquid has a very small specific surface area so that, essentially, no reaction with the gaseous reactant can be expected, and is thus wasted. On the other hand, it is found from both experiments and theoretical analysis that, unlike solid particle feeding, all the droplets ejected from the pressure nozzles have considerably high axial velocity and so do not need to be accelerated further by air stream. Therefore the nozzles are replaced to the outlets of gas conduits (Point B, represented by the dashed lines in Fig. 7.9), i.e., the piece of conduit originally designed for the acceleration of droplets is deleted. The experiments show that such a modification is efficient, and the details will be described later.
0
Figure 7.10 Coalescence of spray droplets on the wall of conduit and the consequential flow.
7.5.2 Experimental scheme and procedure Commonly, flue gas consists of N2, NOx, C02, 802 and air. In the process of wet desulfurization with hydrated lime as the absorbent, N2 is inert, and the amount of Ca(OH)2 consumed in its reaction with NOx can be neglected; while both CO2 and SO2 react significantly with Ca(OH)2. However, what we are mostly concerned with is the
IMPINGING STREAM ABSORPTION
173
situation of SO2-removal. On the other hand, it is well known from the existing results of investigations made with other equipment [25, 121-123] that, in the case of CO2 existing in flue gas, the sulfur-removal efficiency can be kept constant, provided the mole ratio Ca/S is greater than its critical value (Ca/S)c and, as mentioned above, Berman et al. have determined the critical values for various concentrations of CO~ in flue gas. So, the existence of CO2 in flue gas is not a problem for the application of wet desulfurization technology. In addition, because the concentration of Ca(OHh in the suspension required is normally very low, less than 1%, the temperature of flue gas would drop to the wet bubble temperature once the flue gas contacts spray droplets of the absorbent suspension due to the evaporation of water. So, this temperature can be considered as having no effect on the reactions in liquid. In order to focus on the major topic of desulfurization, the pseudo flue gas used in the experiments is prepared with pure SO2 and air at room temperature, 25-30°C during the run of most experiments. The Ca(OH)2 suspension is prepared just before each run with commercial lime, with CaO content k 95%, and water, and the concentration of Ca in the form of Ca(OH)2 in the prepared suspension for each run is chemically analyzed. The experimental system scheme is shown in Fig. 7.11. _
l 21]
Cleaned gas
Air 5~ .
3
SO~
I,
" - " - - - ~ A' [
~
I Discharging
Figure 7.11 System scheme for FGD experiments, l-lime emulsion tank; 2-metric pump; 3-fun; 4-bomb of SO2; 5-rotameter; 6-mixer; 7-GIS absorber; 8-nozzle. Air and gaseous SO2 in the required ratio enter Mixer 6 to mix fully with each other, and the resulting pseudo flue gas is divided into two equal streams to enter Absorber 7. The air flow rate is adjusted by a butterfly valve in the pipeline and measured with a Pitot tube-pressure difference meter and that of SO2 by the rotameter 5. The total gas flow rate is also monitored by a wind velocity meter of DF-3 type at the gas outlet of the reactor. For each run, gas-samplings are made at both inlet and outlet of the reactor, and the SO2 concentrations in the samples are measured with the Iodine-quantitative method, a standard and authentic method of determining the integral amount of SO2 absorbed in the reactor.
174
IMPINGING STREAMS Nozzle [ ~ ~-~
.....
Detector
::i!iiii!:. Laser ::!!!!iiiii!!i)!:: beam Spray
Figure 7.12 Measurement of sizes and size distribution of spray droplets. Due to the difficulty of measuring the sizes of the spray droplets inside the reactor during operation, the sizes were measured by simulating the in situ conditions, using a laser particle measuring instrument of FAM type developed by Shanghai University of Technology. A scheme of the measurement is shown in Fig. 7.12. To obtain representative samples, the laser beam is arranged to enter and pass through the spray at multipoint, and the data are averaged. The system's hydraulic resistance, mainly the absorption device, is a very important parameter affecting the economic index of the technology because a huge amount of flue gas is emitted from the power plant. To evaluate the hydraulic resistance of the equipment, the pressure drop Ap, between Points A and B shown in Fig. 7.11 is measured in each run. From experience, liquid to gas volumetric flow ratio, VL/Vc, must be an important operation variable for the process under investigation so experimental runs were operated at such defined ratios, while the concentration of Ca(OH)2 in the absorbent is determined by the required Ca/S mole ratio. A diaphragm metric pump is used for liquor transportation through the nozzles to spray. The atomizing pressure has a significant effect on both the amount of liquid sprayed and the size of the droplets. Since higher pressure will result in greater power consumption and because the nozzles used in the study are of high efficiency, a relatively low atomizing pressure of Pat = 1.0 MPa was selected, and for most of the experiments the system was operated at this pressure level. The experiments show that at this pressure absorbent can be atomized to good dispersity.
7.5.3 Data interpretation 7. 5.3. 1 Essential assumptions In addition to sulfur-removal efficiency, mass transfer coefficient is an important parameter of concern. Because of the difficulty in determining of the interface area, it is difficult to determine the gas-film transfer coefficient, kc, so that few data for it have been reported to date, whereas the volumetric one, kca, is easier to get. The current project also aims to obtain the gas-film transfer coefficient. It is difficult to measure individual parameters locally during operation. The only option is to determine the parameters involved by interpreting the global data from the
IMPINGING STREAM ABSORPTION
175
measurements at the inlet and outlet of the device. To do so, the following assumptions are made" (1) Based on the considerations described in Section 7.3, the absorption in the wet desulfurization process with dilute Ca(OH)2-water suspension as the absorbent is assumed to be governed by the diffusion of SO2 through the gas film, and so the equilibrium concentration of SO2 at the interface is equal to zero" (2) The space inside the cylinder and under the damper is the effective region for absorption, in which both gas and spray droplets are distributed uniformly and are in ideal mixing. From the existing results, the assumption of ideal mixing is reasonable for gas. However, the droplets' distribution deviates from uniformity, and the density in the impingement zone, the active region for gas-liquid contact and mass transfer, is higher [5]. However, since this region has no physical boundary and the non-uniformity varies from operation to operation, generalization is difficult. On the other hand, values for transfer coefficient based on the whole effective volume, VR, with uniformly distributed gas and droplets are safer and thus more meaningful from the point of view of application; (3) The spray droplets have the same mean residence time in the effective volume as that of gas. The essence of this assumption is that the spray keeps its dispersion status and there is no surface area loss during contact with the gas until absorption is essentially accomplished. For the device shown in Fig. 7.9, this assumption is relatively reasonable because few internal parts are employed so that the absorption process can be considered to be carried out before the droplets collide on any wall" (4) The specific surface area for transfer, a, within the effective region is calculated from the Sauter mean diameter of spray droplets, which remains constant throughout the process. Because of liquid properties, both re-atomization and coalescence of droplets are possible during impingement between the opposing droplets-in-gas suspension streams. The results Wu et al. [129] obtained showed the following: large droplets tend to break-up (re-atomize) and small ones tend to coalescence so that the size distribution becomes narrower, while the average diameter of the droplets remains mainly unchanged after impingement. This is the experimental evidence for this assumption.
7.5.3.2 Mass transfer model and its solution According to Assumption (1) above, the absorption flux is obtained from the wellknown mass transfer model, as N s - k(~ (Csc - 0 ) -
kc, CsG
(7.9)
The absorption rate per unit volume is N s - k GaCsC
(7.1o)
176
IMPINGING STREAMS
and that in the whole effective volume is (7.11)
M s =Vkk~aCs~
According to Assumption (4) above, the specific interface area calculated from the Sauter mean diameter of spray droplets, a, is kept constant. Thus, the integral amount of SO2 absorbed within the residence time of the gas and droplets in the effective volume of the reactor, tf, can be obtained as A M s - VRkGa ~of Cs~dt
- VRkGaCsG,lmt
f
(7.12)
Consequently, the volumetric mass transfer coefficient is obtained from the logarithmically averaged data measured for the variation in SO2 concentration before and after absorption, Csc,~m, as kGa -
AM S VRCsG,lmtf
(7.13)
where the time of gas-liquid contact, tf, is calculated according to Assumption (3) above, by tf = VR / V c
(7.14)
and the specific interface area, a, is calculated from the Sauter mean diameter of the spray droplets dp" VLtf 2 6VLtf a - (1 / 6)~dp3 7Cdp = dp
(7.15)
In addition, the efficiency of sulfur-removal, r/s, is defined with the results measured for gas composition as
r/s =
VG(Cs,in -Cs,ou t ) VGCs,in
(7.16)
7.5.4 Results and discussion 7. 5.4.1 Sizes of spray droplets
The dispersity of the absorbent has a fundamental effect on the reactions involved. As the basis for the determination of mass transfer coefficient, the Sauter mean diameters
IMPINGING STREAM ABSORPTION
177
of spray droplets of absorbent with various Ca(OH)2 concentrations atomized at a fixed pressure of 1.0 MPa are measured in the equipment shown in Fig. 7.12, and the results listed in Table 7.4. As can be seen, the concentration exhibits a certain influence on the mean diameter. Since only the Sauter mean diameter is needed for calculating the specific interface area of the droplets, the data on size distribution are not given here. Table 7.4
Sauter mean diameters of spray droplets with various concentrations of Ca(OH)2 (atomizing pressure P~t = 1.0 MPa) Concentration of Ca(OH)> kg/m~
3.5
4.4
5.3
6.2
7.1
8.0
8.8
10.0
28.86
30.58
31.07
32.60
36.46
36.62
37.56
37.75
Sauter mean diameter of spray droplets dp, gm
7.5.4.2 Overall performance of the reactor As mentioned, the aim of the study is to develop desulfurization equipment of industrial interest so understanding its general performance is important. Several sets of typical operation data measured under stable operations are listed in Table 7.5. The comparable data are: depending on coal type, SO2 content in flue gas ranges from 1400 to 11400 m g / m ~, while the permitted discharge level in China is normally 1200 mg/m 3. The data show that the designed equipment exhibits satisfactory global performance and meets the requirements for desulfurization by wet process. Under moderate operation conditions, the content of SO2 in the cleaned gas can achieve a much lower level than that permitted. Even if the mole ratio of Ca/S is as low as 1.0, a sulfurremoval efficiency of nearly 90% can be achieved (see the fourth row in Table 7.5)" while the pressure drop across the reactor is very small, ca. 400 Pa only. Table 7.5
Typical operation data of FGD system Vo m3"s-~
Cso.i,, mg "m-3
VL/Vo
Ca/S
uo
At
Cso.out
qs
Ap
L.m-~
mol.mol-~
m.s -j
s
mg.m-3
%
Pa
0.08
3200
0.84
1.4
7.0
3.02
240
92.5
405
0.08
2400
0.84
1.4
7.0
3.02
132
94.5
405
0.08
2000
1.20
1.4
7.0
3.02
80
96.0
405
0.08
2400
0.85
1.0
3.02
3.02
276
88.5
405
0.10
2000
0.65
1.4
10.2
2.42
91.8
91.8
380
178
IMPINGING STREAMS
7.5.4.3 Influence of liquid~gas ratio on sulfur-removal efficiency For absorption controlled by diffusion through a gas film, it is necessary to provide a large enough interface area; the interface area depends on the liquid to gas flow rate ratio, VL/V~, essentially for a defined dispersity of the absorbent. On the other hand, increasing VL/V~ must lead to increased power consumption so it is important to optimize the flow rate ratio.. The experimental results on the influence of VL/Vc on the sulfur-removal efficiency are shown in Fig. 7.13. To keep the conditions of atomization essentially the same, all the experiments in this set were carried out at a fixed volumetric liquid flow rate, VL; while the gas flow rate, V6, for each run varies according to the required liquid to gas flow rate ratio and, simultaneously, the corresponding concentration of Ca(OH)2 was used to keep the ratio of Ca/S the same as 1.4. As can be seen in Fig. 7.13, in the range of VL/V6
96 94
j ~ J
92 90 88 86 84
0.4
0.6
0.8
1.0
1.2
VL/VG×103, m3.m -3 Figure 7.13 Influence of liquid/gas flow rate ratio on the efficiency of sulfur-removal (VL=0.067 m3.s-1" Ca/S=I.4; CsG,in=2000 mg.m-3; Pat=l.0 MPa).
IMPINGING STREAM ABSORPTION
179
Z5.4.4 Influence of Ca/S mole ratio The results on the influence of the Ca/S mole ratio on sulfur-removal efficiency are given in Fig. 7.14. The experiments were carried out at a fixed liquid to gas flow rate ratio of 0.85x10 -3 m3.m -3, and the required Ca/S mole ratio was met by varying the concentration of absorbent Ca(OH)~. The results show that, in the range of Ca/S >1.0 the sulfur-removal efficiency slightly increases as Ca/S increases, but the magnitude of variation is not large. As is well known, the reaction between SO2 and Ca(OH)2 in liquid is instantaneous and irreversible. If there is no limitation, the reaction proceeds fast and quantitatively to achieve high sulfur-removal. The fact of the very small influence of Ca/S mole ratio on the sulfur-removal efficiency in the range of Ca/S > 1.0 suggests s very small influence of Ca(OH)? dissolution in comparison with diffusion through a gas-film. So, the assumption of diffusion through gas-film control is relatively reasonable. 100 95 90 85 80 75
I
1.0
,
I
1.2
l
I
1.4
,
I
1.6
,
I
1.8
~
I
2.0
Ca/S, mol-mol-~
Figure 7.14 Influence of Ca/S mole ratio on the efficiency of sulfur-removal (VLIVo = ~ m-~; Csoi, 0.85x10 -~ m-" • . = 2400mg.m -~" V(; = 0.08 m3•s-l "u0=7.0m. s-l ,Ca/S=l.4mol.mol -~" P, = 1.0 MPa). It should be noted that when qs achieves about 90% the concentration of SO2 in gas would be low enough (around 240 mg.m -3 in this set of experiments). So, it can be considered that for the pseudo flue gas with a composition similar to that used in the experiments, without any other active component like CO2, Ca/S = 1.0 is the most reasonable mole ratio, while for practical flue gas, an appropriately higher Ca/S should be employed to account for the additional consumption of Ca by CO2. The data shown in Fig. 7.14 are somewhat different from those by Berman et al. [25] who reported a sulfur-removal efficiency of about 80% under the condition of CO2 - 0 and at Ca/S - 1.0 (see Fig. 4 in Ref. [25] or Fig. 7.7 in this book). Possibly, this difference results from different absorber structures. In that used by Berman et al. spray droplets may collide rapidly on the wall, resulting in a significantly decreased effective interfacial area.
180
IMPINGING STREAMS
7.5.4.5 Influence of S02 concentration in feeding gas The experimental results for the effect of S O 2 concentration in feeding gas, Cso,i,, on the sulfur-removal efficiency, r/s, are shown in Fig. 7.15. The figure demonstrates that r/s drops continuously as Csc,in increases. However, this does not imply a decrease in the absorption rate. In fact, this rate is a monotonously increasing function of Csc,i,, while the tendency of decrease in r/s results from the fact that an increased amount of SO2 needs to be absorbed. These data suggest that, in the case of very high CsG,in, an improvement in operating conditions, e.g., increasing VL/Vc, is needed in order to achieve higher sulfur-removal efficiency. 100
96
m
92
88 ,
,,
I
,,,,,
I
1600
800
,
2400
,
I
,
I
3200
J,
,l
4000
CmG,in, mg-m
4800
-3
Figure 7.15 Sulfur-removal efficiency vs. SO2 concentration in feeding gas (VL/VG= 0.85×10 -3 m3"m-3; VG= 0.08 m3.s-1" u0 = 7.0 m.s-1' CaJS = 1.6 mol.mol-l" Pat = 1.0 MPa).
0.05 0.04
S 6
0.03 ,I
0.02
"
•
"
1
l
I
0.01
0.00
i
1600
I
,I
I
,
i
2400
I
i
i
i
i
.i
3200
i
.l
I
•
4000 -3
Csai~, mg.m
Figure 7.16 Influence of S O 2 concentration in feed gas on gas-film mass transfer coefficient. (The experimental conditions are the same as for Fig. 7.15).
IMPINGING STREAM ABSORPTION
181
It is interesting to understand the influence of SO2 concentration in the feeding gas on the gas film mass transfer coefficient. A set of values for kG calculated with the model equations derived in Section 7.5.3.2 from the data yielded in Fig. 7.15 is plotted in Fig. 7.16 as ko vs Cs.~n. It seems that the value for kG at Cso,~n = 1600 mg.m -3 exhibits larger deviation from the average. The most likely reason for this phenomenon may be that, at that concentration of SO2 in the feeding gas, the content of SO2 in the out gas, after absorption, becomes too low, only about 30 mg.m -~, making it difficult to measure accurately. This value is therefore questionable and may be considered as resulting from experimental error. If this point is ignored, it can be considered that the value for ko is kept constant, independent of the concentration of SO2 in the feeding gas. The fact that the gas-film mass transfer coefficient is independent of the SO2 concentration further shows that the absorption under consideration is controlled by diffusion through gas film and the calculation method derived in Section 7.5.3 is feasible.
7.5.4.6 Influence of device design To examine the reasonableness of the device structure, the effects of impinging distance, S, and placement of the nozzles are studied in certain ranges.
94
~. 93
92
1
J I
4.0
,
I
4.2
~
I
4.4
,~
I
4.6
,
I
4.8
,
l
5.0
S/4, Figure 7.17 Influence of dimensionless impinging distance on sulfur-removal efficiency (Vo = 0.08 m~.s-j" VdVo = 0.84x10 -~ m~.m-~" Ca/S = 1.4 tool.tool -J, u0 = 7.0 m.s-1" Cso.in= 3200 mg'm-3; Pat = 1.0 MPa). The results on the influence of the impinging distance are shown in Fig. 7.17 as a plot of r/s versus the dimensionless impinging distance, S/do. In the range tested the sulfur-removal efficiency decreases continuously as S/do reduces. The most likely reason is that at smaller impinging distance the concentration of droplets in the impingement zone increases, giving enhanced collision between droplets and an increased tendency of the droplets to coalescence, thus reducing the interface area. On
182
IMPINGING STREAMS
the other hand, the results on the hydraulic resistance of the impinging stream contactor obtained by Wu and Wu [66] showed that, when the dimensionless impinging distance reduced to a level below 4.0, the resistance of the equipment increased sharply as the impinging distance reduced further. Therefore, for both higher absorption efficiency and lower energy consumption, S/do = 4.0 can be considered as the operational minimum value. The results of the comparative experiments on the influence of nozzle placement are listed in Table 7.6. Obviously, all four sets of experiments yielded the same results: the sulfur-removal efficiencies with the nozzle placed at Point A were lower than those with the nozzle placed at Point B represented in Fig. 7.9. The reason for this was discussed in Section 7.5.1.3. These results show that the repositioning of the nozzles from Point A to Point B is reasonable and efficient. Table 7.6
Sulfur-removal efficiencies with nozzles at different positions
No
WG
CsG,i n
VLX 103
m3"s-I
mg'm-3
m3"s-1
1 2 3 4
2000 0.08
2200 2400 2600
Ca/S m°l'm°l-~
Uo m's-1
1.0 0.0667
1.2 1.4 1.6
7.0
rls, % Position A Position B 95.0
97.5
94.0
96.2
92.7
94.5
91.2
92.0
Z5.4.7 Interpretation of mass transfer coefficient The traditional method for interpretation of mass transfer coefficient is based on the well known n-law, in which the influences of the related parameters are represented by expressions including some dimensionless numbers, such as Re, etc. However, because of the impingement between the opposing streams, the phenomena involved in impinging streams are much more complex. In fact it has been found in the investigation on micromixing carried out by Wu et al. [ 110] that the Reynolds number Re cannot be used as a criterion for scaling-up. In all the investigations on impinging streams by the author of this book no n-law-based method could be used. In addition, the mass transfer coefficient generally depends on the relative velocity between phases, ur. However, the movement of droplets during operation of the impinging stream device is quite complex. Possibly, the droplets may penetrate to and fro between the opposing streams so that the movement is multistage, including acceleration and deceleration, and the relative velocity we are concerned with becomes a variable, varying with time, its variation regularity varying from operation to
IMPINGING STREAM ABSORPTION
183
operation, making it difficult to determine a representative time-averaged value for it. In the present study the gas-film mass transfer coefficient, kc, is correlated to the impinging velocity, uo, i.e., the velocity of the gas flow at the exit of the conduit. Using the model equations described above and from the experimental data yielding Fig. 7.13, the calculated data are given in Fig. 7.18 as the plot of kc versus u0. By regression, the experimental data are fitted to represent the relationship between gasfilm mass transfer coefficient and impinging velocity by k G - 2.9×10-4ui) 75821
(7.17)
with the standard deviation of SD - 2.45x10 -4 m.s -~. The exponent of 1.75821 suggests that u0 has an essential effect on kc, and is thus a very important operating parameter. The volumetric mass transfer coefficient, kca, is also calculated from the same set of experimental data as those in Fig. 7.18, and the results are shown in Fig. 7.19. Distinguishing from that for k~; versus u0, the relationship between kca and u0 is essentially linear. The experiments in this set were carried out at a fixed liquid flow rate, VL. So, higher impinging velocity implies larger gas flow rate and thus shorter main residence time in the effective volume of the reactor, i.e., the hold-up of the droplets inside the effective volume becomes smaller. On the other hand, at a fixed Ca/S ratio, for larger gas flow rates an absorbent containing a higher concentration of Ca(OH)2 has to be used, yielding larger mean size of droplets, as shown in Table 7.4 and consequently a smaller interface area per unit mass of absorbent, i.e., in addition to its strong positive influence on k(;, the increase in impinging velocity has a double negative effect on the specific interface area, therefore u0 affecting kc and kca to different degrees is reasonable.
0.04
-7
0.03
0.02
0.01
J I
,
I
6
8
i
I
10
J
I
12 b/(), m . s
i
-1
I
14
,
I
16
18
Figure 7.18 Influence of impinging velocity on kc (VL= 0.067 m~.s-r- Ca/S =1.4; CsG.in= 2000 m
g-m-~~; P~,t= 1.0 MPa).
184
IMPINGING STREAMS 1.1 1.0
"7
0.9 []
0.8 0.7 0.6 0.5
i
4
I
6
i
I
i
I
8
i
10
I
12
i
I
t
14
I
16
18
Uo, m's -1 Figure 7.19 Influence of impinging velocity on k~a (VL = 0.067 m3.s-1" CaJS =1.4; mg.m-3; Pat = 1.0 MPa).
CSG,i n "-
2000
It can be seen in Figs 7.18 and 7.19 that in range of the impinging velocity u0 from 5.53 t:o 16.62 m.s -~ the values for volumetric mass transfer coefficient and the gas-film one ranged from 0.577 to 1.037 s-~ and 0.00641 to 0.0416 m.s -~, respectively, showing clearly the effect of impinging streams enhancing transfer between phases. As mentioned in the Introduction, the Rotating Packed Bed (RPB, also called HIGEE), proposed in the late-1960s, a little later than IS, is another kind of technology that significantly enhances transfer between phases. It has received some attention in recent years and a number of investigations have been carried out into it. It is interesting to compare transfer performances of impinging stream gas-liquid reactor (ISGLR) and RPB. A comparison between the values measured in ISGLR and RPB for the volumetric mass transfer coefficients is given in Table 7.7. Table 7.7 Comparison between volumetric mass transfer coefficients measured in ISGLR and RPB Device
System
Rotary seed, rpm
KLa or KGa, s-1
Reference
ISGLR
SO2-Ca(OH)2
0
0.58-1.04
[127]
RPB
CO2-NaOH
~ 1400
0.21-0.44
[ 131 ]
RPB
CO2-NaOH
1400
0.35
[ 132]
RPB
CO2-NaOH
1550
1.14
[133, 134]
RPB
CO2-NaOH
1369
0.66
[ 135]
IMPINGING ,£TRFAM ABR~R PTtON
185
According to the data listed in Table 7.7 it can be considered that the volumetric mass transfer coefficient in ISGLR is about equivalent to those in RPB, i.e., the two kinds of device have roughly the same ability to enhance transfer. On the other hand, in operation, the main body of impinging stream equipment is motionless, while that of RPB rotates at very high speed, and consequenty needs higher capital cost, power consumption and maintenance cost, and, possibly, has a shorter lifetime.
7.5.4.8 Resistance of the reactor As mentioned above, hydraulic resistance is an important factor to be considered in the decision of impinging velocity. The data on the influence of impinging velocity on the pressure drop across the reactor measured at various liquid flow rates are given in Fig. 7.20. The following can be seen from the figure: (i) The liquid flow rate, VL, exhibits a middle effect on the pressure drop Ap; (2) In the ranges of impinging velocity and liquid flow rate tested all the pressure drops measured are small, around 400 Pa only; (3) In comparison, the resistance of the impinging stream gas-liquid reactor used in the present study is larger than that of the gas-solid impinging stream contactor by about 40%, as reported in Ref. [66]. This may be accounted for by the fact that the interaction between gas and liquid is stronger than that between gas and solid. Similar phenomena were also observed in other equipment for gas-liquid contact, such as wet-wall tower, packed column etc., i.e., the resistance of the tower or column with liquid sprinkling is always larger than that of a dry tower or column; and (4) The variation in resistance with the impinging velocity exhibits different tendencies: before u0 ~ l0 m-s -~ Ap increases rapidly with u0. while after this value the variation in Ap is smoothed. The reason for this is not clear yet. Most likely, a shift of flow pattern occurs at this point. 500
v
400
©
/•
300
I
v L, in3 . h-' ;
0.20 0.24 0.26 0.28
200 .
v 100 ~ _ _ _
I
4
.
l ..........
6
J. . . .
8
I
,
!0
u 0, m . s
I
!2
-I
,
I
14
,
[
~
16
~
_
18
Figure 7.20 Resistances of the reacto~ at various impinging velocity.
186
IMPINGING STREAMS
To balance the opposing effects of larger mass transfer coefficient and lower energy consumption, according to the data in Figs. 7.18 or 7.19 and 7.20, a range of 10 to 15 m.s -1 is taken as the optimal window for the impinging velocity, u0. This yields a sufficiently large operating elasticity.
7.5.5 Conclusions An investigation was made to evaluate flue gas desulfurization of (FGD) by absorption in a gas-continuous impinging stream gas-liquid reactor developed recently for systems involving fast reaction(s) in liquid. The following were concluded: (1) The new reactor has good global performance for FGD. Under moderate operation conditions, the content of SO2 in the cleaned gas can achieve a level much lower than that permitted; (2) The following optimal or feasible conditions and structural parameters were determined: liquid to gas flow rate ratio VL/Vc = 0.85-1.0x10 -3 m3-m-3; impinging velocity u0 = 10-15 m-s-~; dimensionless impinging distance S/do > 4; molar ratio Ca/S = 1.0 for pseudo flue gas without CO2; the nozzles were mounted at the outlets of the gas conduits; (3) The gas-film mass transfer coefficient, kG, was determined based on the Sauter mean diameter of spray droplets. The results show essentially no influence of initial concentration of SO2 on kc, suggesting that the process is controlled by diffusion through gas film and that the method proposed for the determination of k~ is feasible; (4) The data on the relationship between impinging velocity and gas-film mass transfer coefficient were fitted by k~ - 2 . 9 × 1 0 - 4 u ; 75821 , with the standard deviation SD = 2.45×10 -4 m-s -1, implying u0 is a strong effecting variable, and thus a very important operation variable; (5) With the impinging velocity u0 ranging from 5.53 to 16.62 m-s -1, the measured volumetric mass transfer coefficient k~a is in the range 0.577 to 1.037 s-~ and k~ from 0.00641 to 0.0416 m.s -~, showing clearly the effect of gas-continuous impinging streams enhancing mass transfer; (6) The impinging stream gas-liquid reactor has low hydraulic resistance. In the range of operation conditions tested, the pressure drop across the reactor, Ap, is round 400 Pa only.
7.6 DESIGN OF A DEVICE FOR LARGE GAS FLOW RATES The results described in Section 7.5 illustrate that the application of gas-continuous impinging stream gas-liquid reactor for the wet desulfurization of flue gas performs to good effect. Gas-liquid reaction is a large category of important reactions involved in many processing industries, among which many systems involve fast or instantaneous
IMPINGING STREAM ABSORPTION
187
reaction(s) in liquid phase. One can say almost with certainty that gas-continuous impinging streams (GIS) will find more and more important applications in the field of absorption. The major advantages of GIS over other methods for the wet desulfurization of flue gas are its high sulfur-removal efficiency, very large volumetric mass transfer coefficient which necessitates only a small device, and relatively small resistance to the streams. Therefore the application of GIS for wet FGD can be expected to yield great economic and social benefits. However, for such a purpose the related engineering problems need to be solved. From the point of view of practical application, the major nature of the wet desulfurization of flue gas lies in the fact that the flue gas to be treated has extremely large flow rates and, consequently, the amount of the absorbent to be atomized is also very large. For coal burning power plants, e.g., around 4000 m3.h-~ of flue gas per MW is generated, and a coal-burning power station with a capacity of 400 MW will exhaust over 1.6 Mm3.h -~ Correspondingly, the amount of atomized absorbent suspension needed by a wet FGD system for such a station would be around 1.6 km~.h-j. Although the GIS gas-liquid reactor shown in Fig. 7.9 performed well in the test on a small pilot plant scale, its structure is not suitable for treating extremely large amounts of gas, like practical flue gas. An extremely large-volume reactor would be needed and the amount of absorbent to be atomized would make both design and manufacture of the equipment extremely difficult, if not totally impossible. The Combined Multifunctional Impinging Stream Gas-Liquid Reactor was designed [130] to make this technology suitable for processing huge amounts of gas. The reactor's structure is shown in Fig. 7.21 with a vertical view in Fig. 7.22. It employs multiple groups of flow configurations of horizontal-coaxial four impinging streams. The reactor consists of two major parts: the tower body (1) and several groups of impinging stream components I, II, etc, mounted inside the tower body (1) at various altitudes. The tower body (1) is a vertical cylinder. Near the top of the cylinder a mesh can be installed as a foam remover 2. The cylinder has a top cover (in pan or conical shape) connected to the gas exhaust port (3), and a liquid discharge port (4) near its bottom. All groups of impinging stream components, I, II, etc. are identical in size and working principles (Fig. 7.21 shows only Groups I and II as examples). For each group there are four gas conduits (5). At the outlet of each gas conduit, one nozzle (6), or a set of nozzles is installed for atomizing liquid, either pure or containing solid particles. Liquid or solid-in-liquid suspension is supplied to nozzle (6) through the high pressure liquor feed pipeline (7). Above the four conduits, a droplets-removal damper (8), in pan or conical form, is placed. The damper (8), tower body (1), and either the bottom of tower (1) or the damper (8) below the four conduits in the upper group, form a subchamber for absorption. For each group, the four gas conduits are divided into two subgroups, with two conduits in each sub-group. The two conduits in each sub-group are placed coaxially, with the exits of the conduits facing each other. The axes of the two sub-groups of the conduits are perpendicular to each other.
188
IMPINGING STREAMS
2
1 i
II !~i/ , i,,
7 ~id
feed ~
Gas~//d, 6J
/~
J
Impingement zone
Liquidfeed
:i'/ ......... oii '
~
d
.L ~ ~
7
d°ut ~ 6 ~
4
Figure 7.21 Combined multifunctional impinging stream gas-liquid reactor. 1-tower; 2-screen foam-remover; 3-gas outlet tube; 4-liquid outlet tube; 5-gas conduit; 6-eddy pressure nozzle; 7liquid feeding tube; 8-damper.
IGF
1 3
GF Figure 7.22 Vertical view of the combined multifunctionai impinging stream gas-liquid reactor. 1-tower; 3-gas outlet tube; 5-gas conduit, 7-liquid feeding tube.
IMPINGING STREAM ABSORPTION
189
The distances from the outlets of the conduits to the center where the two axes meet are equal (See Fig. 7.22), and the distance between the exits of the two conduits in each sub-group is the "impinging distance". Depending on the requirement of the amount of liquid or suspension to be used, one nozzle or a set of nozzles can be installed inside each conduit with the exit(s) towards the same direction as the outlet of the conduit, i.e., towards the center of the absorption chamber. All groups of the impinging stream component I, II, etc. are operated concurrently. For Nozzle (3), a pressure atomization nozzle, also called the centrifugal pressure nozzle, can be employed. It is desirable to use the eddy pressure nozzle (Chinese patent, ZL00230305.1). The latter has a higher flow-rotating efficiency, and thus requires less energy to atomize the liquid or solid-in-liquid suspension. The purpose of using the screen foam-remover is to separate the gas flowing upwards from the foam or fine droplets carried by the gas flow. Such a remover has a high efficiency of foam-removal and gives small hydraulic resistance. However, if solid particles are present in the gas or liquid, or there is solid product from the chemical reaction, the solid particles may clog the mesh of the foam remover (2), resulting in blockage of the gas flow channels and increased hydraulic resistance, and also making it difficult to clean up. In this case, the foam remover screen can be replaced by an internal wet cyclone, as shown in Fig. 7.23. The wet cyclone can also achieve very high separation efficiency (>99%), but its hydraulic resistance is larger than that of the foam remover screen by about 600-800 Pa. If the requirement for dust removal is not high, it is recommended that neither foam remover screen nor the internal cyclone be used. It can be seen that in the combined multifunctional impinging stream gas-liquid reactor shown in Figs. 7.21 and 7.23, the working principles and action in enhancing transfer between phases for each sub chamber of absorption are the same as those of the reactor shown in Fig. 7.10. The first difference between the two reactors is that a pair of impinging streams is added in the direction perpendicular to the flow axis of the original pair of impinging streams. As a result, the utilization factor of the space inside the sub chamber is increased. The major active region in impinging stream equipment (the impingement zone) is only a thin layer with a diameter 8-10 times that of the gas conduit and a thickness of about a quarter to a half of the impinging distance, so when only one pair of impinging streams is used the utilization factor of the space is very low. The addition of another pair of impinging streams approximately doubles the utilization factor of the space and doubles the gas flow rate that can be processed in the sub chamber for absorption. Another feature of the combined multifunctional impinging stream gas-liquid reactor is the employment of multiple groups of impinging stream components placed one above the other, with each group including four streams. This arrangement further increases several-told the gas flow rate that can be processed in the reactor, while all the groups are operated individually without disturbing each other. The arrangement slightly increases the hydraulic resistance of the system. Compared with the existing equipment for wet FGD, the resistance of the reactor shown in Fig. 7.21 or 7.23 is still much smaller.
190
IMPINGING STREAMS GO
J
/3
9
1
::-~'-:: ,,i,\;~:: II
.:::t::..
liiiiiiiii ',:[
'.L'
I
Figure 7.23 Combined multifunctional impinging stream gas-liquid reactor with the screen foam remover (2) in Fig. 7.21 replaced by the internal cyclone (9) in this figure. According to the data for volumetric mass transfer coefficient measured in the device on a small pilot plant scale, for a certain load of flue gas to be processed, the required total volume of the reactor under consideration would be very small, only about 1/3 that of existing wet FGD equipment. In addition, the arrangement of the internal wet cyclone shown in Fig. 7.23 enables the reactor to have simultaneously high ash-removal efficiency. The reactor is especially suitable for the wet desulfurization of flue gas with hydrated lime or dilute ammonia solution as the absorbent. The design of the large-scale reactor suitable for a power station has now been accomplished and is expected to be applied industrially in the very near future.
-8IMPINGING STREAMS COMBUSTION AND GRINDING
The combustion of powdery coals and sprayed liquid fuels and the grinding and milling of solid materials are fields in which gas-continuous impinging streams have been successfully applied for many years, and so are of industrial significance. Although the author of this book has not carried out any experimental work in these two areas, because of their importance, this chapter has been included to keep the integrality of the book as one specially discussing impinging streams. Consequently, the chapter mainly introduces the reader to some representative results obtained by other researchers although, from time to time, the author's own opinion is included.
8.1 MODELS FOR PARTICLES AND DROPLETS COMBUSTION In principle, gas-continuous impinging streams (GIS) can be applied for the combustion of gases, powdery solids and sprayed liquids. Since gas-combustion is relatively simple and the process is essentially independent of the major feature of GIS, i.e., that it significantly enhances heat and mass transfer between phases, the discussions in this chapter will focus on the combustion of the latter two kinds of fuels.
8.1.1 Evaporation-burning equations for a single droplet Chemically, combustion is a violent oxidation reaction, normally occurring at high temperature. According to the Arrhenius relationship, high temperature defines the kinetics feature of extremely fast reaction(s) and, consequently, any transfer between phases must be the governing factor, so that impinging streams should be applicable. In order to increase transfer rates, a liquid fuel must first be atomized into fine droplets to create a large interface area, no matter what kind of burner is being employed. The behavior of a single droplet during burning is the foundation for understanding and analyzing the process. Barnard et al. [136] proposed that, if a liquid droplet exceeds some critical size but less than about one millimeter in diameter, the combustion takes the form of a spherical diffusion flame round the droplet and the burning rate is determined by the vaporization from the surface of the droplet. The fact
191
192
IMPINGING STREAMS
that burning is closely related to vaporization is the feature of the combustion of sprayed liquid fuel which distinguishes it from that of powdery solid fuel. Let us consider the symmetrical burning of a spherical droplet with the radius rp in surroundings without convection. Assume that there is an infinitely thin flame zone from the surface of the droplet to the radial distance rf~ [ 137], which is much larger than the radius of the droplet, rp. The heat released from the burning is conducted back to the surface to evaporate liquid fuel for combustion. Because the reaction is extremely fast, there exists no oxidant in the range of rp< r < rn; while no fuel vapor is available at r > rn. At a quasi steady state the mass flux through the spherical surface with the radius r (>rp), Mfv, can be obtained with Fick' s law as
Mfv
_
-pgDfv
dmfv
dr,
(8.~)
+ m f v M
where mfv is the mass fraction of the fuel vapor at radius r. In surroundings without convection the flux of inert gas Mg = 0, and so the total flux M = Mfv + Mg = Mrv. Thus, at the external surface of the droplet, Eq. (8.1) becomes [ dmfv dr ]s +mfv,sMfv,s
Mfv's =-pgDfv
(8.2)
From the continuity equation at the quasi steady state, we have
(8.3)
Mfv,s 4~p2 - Mfv 4err 2
Substituting Eqs. (8.1) and (8.2) into Eq. (8.3) and integrating the resulting equation from rp to rn gives the flux of fuel-burning, Mfv,~, as Pg Dfv M fv,s= ~ l n [ 1 rp
+
mfv,s - mfv 1 - mfv,s
]
(8.4)
On the other hand, the diameter of the droplet reduces continuously as the vaporization-burning proceeds. The variation of the droplet radius with time can be determined from the mass balance round the droplet itself: d 4 3 4~p2. dt (-3 n'rp PL) = M fv,s
(8.5)
Simplification of the equation results in drp
Mfv,s m
dt
PL
(8.6)
IMPINGING STREAMS COMBUSTION AND GRINDING
193
Substituting Eq. (8.4) into Eq. (8.6) and integrating the resulting equation from t=0 to t leads to r~,9
__
2 rpo - 2t P~DJv l n [ l PL
mt\1,s
mfv
]
(8.7)
1 - mt.v,,~
Let the final radius of the droplet be r p - 0, the time for complete burning of the droplet, tb, can be obtained to be 7)
r~)PL
tb =
(8.8)
2p,gDfv ln[1 - mtv'~ - mtv ] 1- mr-v,s To eliminate the variable mfv, one can use the relationship of mass balance around the neighborhood of the burning zone: (8.9)
Mmtv 110 - - M oomox oo
where O represents the mass of oxidant needed for burning unit mass of fuel, and the subscript oo denotes the states in the bulk gas outside the burning zone. Utilizing Eq. (8.9), Eq. (8.4) becomes pg Dfv M t~,~= ~ l n ( 1 5'
+
mfv,s - mfv//O 1 - mt~'.s
)
(8.10)
For the calculation of tb the data for the mass fraction of fuel vapor at the surface of the droplet, mr,., is needed. From the heat balance one can obtain ~g Mt.v,~=~ln(l+ Cpgrpo
Cpg(T~ - L ) + A H c mo× ~ / O - O s / M fv,s
)
(8.11)
where Q~ is the heat flux through the gas surrounding the droplet to evaporate liquid fuel of the mass M~-v~.If without the heating process of the droplet before burning, i.e., the droplet enters the system just at the boiling point of the fuel, then the amount of heat Q~ meets -Q~/Mt~,~AH v
(8.12)
The amounts of T~ m,,x.~ in Eq. (8.11) are normally known; while the results from analysis indicate that the temperature at the surface of the droplet is always just under
194
IMPINGING STREAMS
the boiling point of the fuel, TBp. Taking the place of Ts with TBp, then the flux of combustion can be calculated with Eq. (8.11). Using the heat balance, the expression for the time needed for complete combustion of the droplet under consideration is rewritten as
rp2oPL
tb =
(~
2 28 ln[1 + Cpg
C0g
(8.13)
- TBp~ + a / - / c / O
]
zM-/v
It is clear that intensifying combustion is just to shorten the time needed for complete combustion tb. It can be seen from Eq. (8.13) that tb is positively proportional to the square of the radius of the droplet, rp. Therefore the most effective measure for the intensification of combustion is full atomization of the liquid fuel to increase its dispersity, i.e. reducing the size of the droplet. For heavier oils fine atomization is of more importance because of their large densities PL, as can be predicted by Eq. (8.13). In practice, the combustion of atomized liquid fuel is a very complex process. The model introduced briefly above cannot be considered as completely consummated. In the derivation of Eq. (8.13) several assumptions were made, which might be unreasonable. For example, pressure is assumed to have no effect; while actually this parameter does have an effect. As the pressure approaches the critical value its influence becomes very significant. At the extremes, i.e. when the critical pressure is arrived at, the vaporization heat of the fuel becomes zero. For combustion at high pressure, Spalding [137] proposed a theoretical method for the prediction of the burning rate; and later Rosner [138] and Dominicis [139] improved the theory. In addition, You [ 140] summarized the models on the influences of the boundaries on the two sides of the interface and the internal circulation on the evaporation rate proposed by Prakash et al. [141 ], which indicates that the internal circulation promotes heating of the droplet and increases the evaporating rate during heating.
8.1.2 Burning equations for a single particle The combustion of a fine particle of solid fuel has many identical features to that of a liquid droplet; but with important differences. One of the differences is that, in addition to no evaporation existing, the combustion of particles of solid fuel is frequently controlled by chemical kinetics, i.e., diffusion-governing is not the only possibility. Let us consider the burning of an ideal spherical particle in static gas. The oxidant diffuses to the surface of the particle to react with the carbon: C + 02---~C02, while the latter diffuses out from the surface of the particle. The combustion heat is transferred to the surrounding gas partially by convection and partially by radiation. The following assumptions were made in the modeling: (1) The process is at a pseudo steady state. (2) The temperature the highest at the surface, and continuously drops down outwards from the surface of the particle; and the concentration of oxidant is highest in the bulk
IMPINGING STREAMS COMBUSTION AND GRINDING
195
gas and continuously drops down towards the particle surface, while the concentration of the product of burning exhibits the opposite profile. (3) All the transport properties of gas are uniform. The diffusion equation for the oxidant round the surface of the particle is written as
M°x" - -P~D°~
dm°x )
dr
+ M~m°x
(8.14)
s
Similarly, since M~.~= 0, there would be M s - Mo,<~+ M,;,~- Mo,,.,~. The burning rate per unit surface area can be represented by
M,,,,,~ - -Me.sO
(8.15)
The minus sign in the equation indicates that the motions of carbon and oxidant are in opposite directions. Using the continuity equation, Eq. (8.3), and Eq. (8.13) and with the known concentration of oxidant at the surface of the particle, mo.... the integration between r = rp0 and r - oo results in the equation for the flux of burning to be
Mox.~ - ~P~D°" ln(1 + m°x'~ - m°x'~ ) rp()
(8.16)
O + mox,s
Since carbon is reactive, namely, m ..... ---. 0 and mox.~- 0.232 or even less, as well as O = 32/16 - 2.67/br stoichiometric combustion of carbon in oxygen. Thus, Eq. (8.16) is simplified to
M,,x.~
=
PgD°x moxoo ( ) r~o 0
(8 ~7)
The problem needing to be considered further is the case where the resistances of the reaction kinetics and the diffusion are comparable. The oxidation rate of carbon per unit surface area is represented as [138]
M ~., - k~Pmo,
(8.~8)
Expanding the logarithmic term in Eq. (8.16), simplifying, and using Eq. (8.18) leads to
k,Pmo,, ~ e x p ( - E / R T , )
mox~ - m{,x,S
p~D,, x
0 + mox,s
(8.19)
Solving Eq. (8.19) for mox.Jmo~J, and substituting the resulting expression into Eq. (8.16) yields the equation for the flux of burning as
196
IMPINGING STREAMS
M c,s =
m o x , oo
(8.20)
Rsr + Rdiff
where the resistances of the surface reaction and the diffusion, Rsr and Rdiff, are defined, respectively, as
Rsr
m
1 ksPexp(-E/RTs)
.
'
Rdiff =rp0[O+ m°×'s ] pgDo x
(8.21)
The following can be concluded from Eq. (8.21): (1) A large particle has very large diffusion resistance; and (2) An increase in Ts yields reduced reaction resistance. The inference made by Tamir in Ref. [5] on the influence of Ts is incorrect. With the method similar to that used in the last section, the expression for the time of complete burning can be derived to be tb = P_____~__s[ rp0 + mox,s ksP exp(- E/RTs)
0.5 r20 (O + mox,s)
]
(8.22)
pgDox
It can be seen that, for the process controlled by diffusion, the complete burning time tb is positively proportional to the square of the initial radius of the particle, rp2 . For estimation the following empirical relationship may be useful" t b -- 108 rp2
(8.23)
while for the process controlled by reaction kinetics, the complete burning time is positively proportional to the initial radius of the particle, rp0. In the establishment of the models above for liquid droplets and solid particles, a number of assumptions were made, or, in other words, a number of influencing factors were ignored. Therefore it might not be expected to obtain exact results from the calculation with these models. However, they are a useful guide for analysis and understanding the problems involved.
8.2 INTENSIFICATION OF COMBUSTION PROCESSES DUE TO IMPINGING STREAMS From the basic properties of gas-continuous impinging streams and the burning models derived in the last section, impinging streams intensify the combustion processes of atomized liquid and powdery solid fuels by the following mechanisms:
(1) Enhancingtransfer betweenphases: As mentioned above, the combustion of both droplets of liquid fuel and particles of solid fuel are multiphase reactions normally
IMPINGING STREAMS COMBUSTION AND GRINDING
197
occurring at high temperature. Certainly, the feature of gas-continuous impinging streams enhancing transfer between phases can promote and speed up the processes. (2) Breaking up droplets or particles: The models above indicate that the complete burning times of both a single droplet and a single particle under the condition of diffusion control are positively proportional to the square of the radius of the droplet or particle. Under common conditions it is possible for both re-atomization and coalescence to occur with sprayed droplets in impinging streams and the mean diameter of the droplets is kept essentially constant [129], while for the combustion of liquid fuel occurring at high temperature coalescence is obviously impossible because of the existence of the diffusion flame layer round the surface of the droplets. As the result, the mean diameter of the droplets must be reduced not only continuously by evaporation-burning but also by re-atomization. For the combustion of powdery solid fuel, the intrinsic action of gas-continuous impinging streams grinding or milling particles that has been widely applied in industry must yield a positive influence on combustion; such action is particularly effective in destroying the ash-layer on the surface of burning particles to eliminate internal diffusion resistance. (3) Changing trajectou: The motion of particles has been discussed in Chapter 2, and the conclusions made are also applicable for liquid droplets. By inertia, droplets or particles penetrate to and fro between the opposing streams, and at the instant they penetrate into the opposing stream the relative velocity reaches an extremely high value yielding greatly increased transfer coefficients; simultaneously, the residence time of particles/droplets in the active region is lengthened to some extent, favoring the combustion being accomplished in a smaller space. (4) Intensifying circulation inside droplets': The shearing effect of gas flow in impinging streams would intensify the internal circulation of droplets. The analyzed results by Prakash et al. [141] indicated that the internal circulation shortens the heating time of droplets newly entering the system and increases the evaporation rate during heating. Many experimental evidences have been obtained for impinging streams intensifying combustion processes. For example, the results obtained by Enyakin [142] and Enyakin-Dvoretskii [143] from their investigations on the combustion of high sulfurcontent oil showed that the oil can burn in impinging streams at a very low excess air factor (as low as 1.02), the number of burners can be decreased, and the length of flame is obviously reduced. The latter result implies that a smaller combustion chamber can be used. From the aspect of modeling, Tamir [5] proposed a model for impinging streams enhancing combustion. The major substance of the model is the use of movement equations for the analysis of the enhancement due to the oscillations of particles penetrating to and fro between the opposing streams, although some other enhancing factors that existed were not taken into account and the model lacked any supporting experimental evidence.
198
IMPINGING STREAMS
On the other hand, impinging streams may possibly also have unfavorable influences on the combustion of atomized liquid fuel, e.g., the action of the shearing force may cause deformation of the droplets, resulting in increased drag coefficient and, consequently, decreased relative velocity between phases. Enyakin [144] analyzed this influence by introducing a parameter accounting for the increase in drag coefficient due to the deformation of the droplets and the results obtained showed a significant influence.
8.3 IMPINGING STREAM COMBUSTORS 8.3.1 Furnaces for gas and liquid fuels Many kinds of impinging stream furnaces have been used for gas or liquid fuels. Figure 8.1 shows a representative impinging stream boiler of the type BKZ- 320-140GM developed earlier, in the combustion chamber of which eight burners in four opposing pairs are placed horizontally at identical heights. The boiler produces 320 tonnes per hour of superheated steam at 14 MPa and 570 °C, with a volumetric heat release of 267 kW-m-3. Enyakin et al. [ 143, 144] examined the performances of the boiler by burning gas and high-sulfur black oil in the range of production from 240 to 327 tonnes steam per hour. The data of composition and properties of the oil were: moisture = 0.8-9.1%, S = 3-4%, H - 9.7-10.6%, N + O 2 = 0.18-0.83%, C = 79-86%, kinetic viscosity v= 7-8x10 -5 m2-s-1, density PL = 979--990 kg.m -3, heat release AH= 37000-39800 kJ-kg -1. The properties of gases were: Type I: CH4 = 34.3%, C 2 H 6 - 31.35%, C3H8 = 1 0 . 2 5 % , C4H10 = 0 . 6 5 % , N2 "- 22.15%, O2 = 0.65%, AH = 40360-41403 kJ.m-l; while Type II: CH4 = 89.6%, C2H6 = 2.85%, C3H8 = 0.1%, C4H10 = 0.052%, N2 = 4.3%, O 2 -3.1%, AH = 33336 -33901 kJ-m-~. The major results from the oil burning tests were: (1) In all the operations with air excess or> 1.045, the heat losses due to chemically incomplete combustion were virtually zero. For co= 1.02 the identical heat losses were about 0.15%. (2) The heat losses due to mechanically incomplete combustion were less than 0.1%, provided or> 1.025, while only for o~< 1.007 this term of heat loss becomes larger, about 0.26%. (3) Soot particles were not observed, even if co= 1.007. (4) The content of SO3 in the flue gases with o~< 1.035 was less than 0.002%. This implies that the boiler of type BKZ-320- 140GM equipped with impinging stream burners provides complete combustion with a negligible corrosion rate of the heated surfaces. (5) The flame length for o~> 1.04 was approximately 5 to 6 m, while 8 to 9 m for o~= 1.01-1.03. The flame length is defined as the sum of the distances from the face
IMPINGING STREAMS COMBUSTION AND GRINDING
199
of the burners to the impingement plane and from the level of the burner axis to the end of the visible flame along the vertical direction. (6) The employment of vortex-type burners is not advisable with impinging streams. The swirling of the flow leads to significant opening out of the fuel spray, causing the axial velocities of gas flow droplets to decrease very rapidly and thus weakening the impingement between the opposing streams, resulting in a weakened transfer enhancement effect.
=
"~lgfj/Jffffffff./fff~
~"
A-A
Figure 8.1 BKZ-320-140GM impinging stream steam boiler. 1 burner; 2 combustion space; 3 secondary combustion space; 4 steam superheating; 5 water economizer. The main results of the gas burning tests were: incomplete combustion was observed for o:' < 1.03, and the corresponding heat losses were 0.126%. The results described above indicate that impinging streams enable oils or gases to burn highly efficiently, with negligible heat loss at low air excess factor, uniform heating of surfaces, and negligible corrosion of the heated surfaces by the products burned.
8.3.2 Koppers-Totzek gasifier for powdery coals Coal-gasification is a very important process involved in the production of chemical fertilizers, energy and many intermediate chemical products. As has been mentioned, the Koppers-Totzek coal-gasifier is an instance of the earliest application of impinging streams, applied industrially as early as 1952 [4, 145] although the term "Impinging Streams" was not used at that time. The structure of the gasifier is briefly shown in Fig. 8.2. Powdery coal sized about 100 gm is fed into the oxygen and steam streams; the two opposing streams co-axially enter the combustion chamber operated at a pressure very close to atmosphere pressure. The gasification reaction occurs round the impinge-
200
IMPINGING STREAMS
ment plane at about 2000°C where the reaction time is very short, about 1 s only. The carbon is burned almost completely and tars are gasified along with the carbon. Since the temperature is very high, the ash in the coal becomes molten and more than a half of it is removed at the bottom of the equipment, where it solidifies in the water bath. The raw gas stream leaving the reactor at about 1500 °C is quenched to solidify the fly ash particles first before they enter the heat-recovery system, then is separated from about 95% of the ash particles in a set of cyclones and is finally washed in a Venturi scriber to remove the ash further. Because of extremely high combustion temperature, the CO content is very high while those of CO2 and hydrocarbons are very low in the gaseous product. A typical analysis of the raw gas is listed in Table 8.1. The coal consumption of the Koppers-Totzek gasifier with four burners is 700 tonnes per day, and the gas production under normal conditions is 55000 m3.h-~. Raw gas
. oo~
Feed water
O~OOO
g~[
[~~
Quenchin water .11
A--A Fly ash
[I-
A -~ ~ i i : i ~
--~ apowdery coal
OZiZZZZZ::~ '1' Water bath
[Oxygen+ steam
Slag
Figure 8.2 Koppers-Totzek powdery coal gasifier. Table 8.1
Composition of raw gases from Koppers-Totzek process CO2
CO
He
Inert
H2S
COS
CO+H2
6.37
64.93
26.94
0.92
1.16
0.13
91.42
IMPINGING STREAMS COMBUSTION AND GRINDING
201
The Koppers-Totzek coal-gasification process has the following main advantages: (1) It is most versatile for processing a wide variety of coals; (2) No tars in the gaseous product; (3) The gasifier is simple in structure and easy to maintain; (4) The capability of raising a large amount of stream; (5) The reduction of possible deposits on the reactor walls; and (6) The capacity of the gasifier can be increased simply by increasing the number of impinging stream pairs, and the design is convenient. In addition to the two combustors introduced briefly above, researches and development on the application of gas-continuous impinging streams to combustion has been carried out since the 1970s, e.g., the investigations made by Goldberg and Essenhigh [146], Ziv et al. [12], Goldman [147-149] and Liu et al. [150]. Mostly, these works involve experimental studies and model analyses, and mainly aimed at the improvements of combustor structure and the arrangement of burners in order to increase combustion efficiency. Most of the experimental combustors employed were very small and their application seems far from practical.
8.4 IMPINGING STREAM GRINDING Grinding or milling is also a field in which impinging streams earlier applied successfully in industries, as well as the well-known Trost Jet mill, has been collected into the standard tool-book [20] as early as the 1970s, although somewhat later than the Koppers-Totzek coal gasifier. Similarly, the term Impinging Streams was not included in the names of the mills initially employing the impinging stream method. The advantages of the use of gas-solid impinging streams for grinding and milling lie in the fact that the equipment is without a milling part so that possible damage to thermally sensitive materials due to the heat generated by friction can be effectively avoided; thus the method is applicable even for explosive materials and the possible pollution of product by the grinding materials is eliminated as well. Because of these outstanding advantages, there has been an increase in related research and development, e.g., the works by Zhang et al [151,152]. It should be noted that the impingement or impacting of a solid-gas suspension stream on a solid surface can also result in the effect of grinding and milling. However, as mentioned in Section 8.1, flow configuration is beyond the scope of this book. In solid-gas impinging streams, grinding or milling is carried out mainly by the collisions between particles accelerated fully by the gas flows [152]. In order to increase grinding efficiency, in addition to enhancing collisions, it is also important to have a reasonable and effective classification arrangement for removing the particles that are fine enough at a suitable time. In principle, the major differences between various mills are their different schemes for the arrangement for impingement, collision, and classification. Many types of grinding or milling machines have been proposed. This section introduces some representative ones developed and applied earlier.
202
IMPINGING STREAMS i Feed
Prod
r
ct
...........F ! u i d i z e d
~..~
Air
~
bea .....A i r
(a) Front-sectional view
(b) Cross-sectional view
Figure 8.3 Fluidized-bed jet mill. Figure 8.3 shows the fluidized-bed jet mill. The feed of material is from the top directly into the fluidized bed and not through the nozzles. The particles are fluidized by the gas stream created by the opposed nozzles in the lower section of the mill. The gas stream accelerates the particles to high velocities. Size reduction occurs due to collisions between particles and collisions of particles on the chamber walls. The expanded particles-gas stream flows upwards into the high-efficiency deflector-wheel air classifier in the upper section of the mill. Coarse particles are deflected back to the fluidized bed by the wheel, while those of appropriate sizes determined by the classifier exit through the fine classifier outlet at the top of the mill as the product. The fluidized bed jet mill consumes some 15-20% less energy than the jet mills discussed below. It is suitable for fine and ultrafine grinding of most free-flowing materials, including thermal-sensitive products, fluorescent powders, ceramics, abrasives etc. Another kind of impinging stream milling device is commonly called an opposed jet mill. The coarse particles to be ground are first fed into gas flows and then, after mixing with air, delivered into the mill to impinge against the opposing particles-gas stream. The grinding action is achieved by interparticle collisions. Since the residence time is very short, not all particles are fully ground through a single path so it is necessary to carry out a classification operation and to return the oversized particles for further grinding. Figure 8.4 shows the Trost Jet Mill from Colt Industries. In principle, it is an opposed jet mill and its grinding action is achieved by interparticle collisions in impinging streams. It has a centrifugal classifier, the grinding capacity can be 1 to 2300 kg per hour and the airflow rate varies from 0.2 to 28 m3.mim -~. Figure 8.5 shows the Majac Jet Pulverizer from Donaldson Company, a milling device whose operation is also based on the principle of opposing jet impingement. It employs a mechanical classifier. A screw or other feeder discharges the material to be pulverized into the impact zone or into the classifier, depending on the feed material. The impinged gas and powder enter the upper mechanical classifier. The oversized material flows downwards through an annular space against the elutriating air, through two down-comer legs to the nozzles where it is accelerated by the opposing gas streams
IMPINGING STREAMS COMBUSTION AND GRINDING
203
of high velocity and the streams impinge against each other, causing interparticle collisions and thus grinding. The fineness of the product is controlled primarily by adjusting the speed of the classifier and the amount of fan air delivered to the classifier. While other effects can also be achieved by changing the nozzle pressure, distance between nozzles of the gun barrels, and position of the classifier disk. Depending on the capacity and the required fineness of the product, the pulverizers operate on quantities of compressed air ranging from 0.6 to 13 m3.min -~. Laree
particles
Discharge s i z e s)
....;~i...7;ii
............................... ( v ari able :,~-~..~~:,~;~i ~........ •...............
:, f:/.: y
Classification chamber
*~
.... ~.a~:=~,~'.... :~, Fine pamcle :;::!~ ~ :~. separation
j
7{ ;i;i7:: i i
Material input . ',,:'%; {',:~'%~,,/:',,:Up
Li
stack
~,;
~i
i
;.{, '{,
.owo,,.ok
~
P tube (variable sizes i
Impact chamber
0 tube
Figure 8.4 Trost jet mill.
Product
;~-
4 .atena, --~ feeding Fan air
v
t
Compressed air, steam or gas
z~: < < . ~
Pulverizing region
{:[;. [ii" .......... ...
Figure 8.5 Majac jet pulverizer.
204
IMPINGING STREAMS
In most applications, the economics of the use of this type of jet pulverizer becomes attractive in the range of product fineness of 98% through 200 mesh or finer; as finer products are required, this equipment becomes increasingly attractive.
P A R T II LIQUID-CONTINUOUS IMPINGING STREAMS
In various process industries, including the chemical and petrochemical industries, many processes are carried out in liquid phase or multiphase with liquid as the continuous phase, most of these processes involving chemical reaction(s). Since liquid is a condensed condition, the movement of the molecules is greatly restricted. For processes occurring on a molecular scale, the status of micromixing becomes extremely important. With the discovery in the 1990s of its excellent feature of efficiently enhancing micromixing, investigations on impinging streams were diverted to liquidcontinuous impinging streams (LIS). Because the properties of liquids are essentially different from those of gases, impinging streams with liquid and gas as the continuous phases exhibit totally different performances and it is therefore necessary to discuss them separately. Part II focuses on liquid-continuous impinging streams and related problems, including the features of LIS that efficiently promote micromixing, pressure fluctuation phenomena in LIS, promotion of kinetic processes by LIS, and the application of LIS in the preparation of ultrafine particles, e t c . Finally, this we will introduce some important research and development on LIS devices and look forward to the future applications of LIS.
205
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-9DIFFERENCES BETWEEN PROPERTIES OF CONTINUOUS PHASES AND CLASSIFICATION OF IMPINGING STREAMS
9.1 PROGRESS OF INVESTIGATION ON LIQUID-CONTINUOUS IMPINGING STREAMS Investigations into liquid-continuous impinging streams (LIS) actually started as early as the late-1950s. However, initial studies focused mainly on the fluid dynamics, e.g., Powell's mirror model [26] etc; while the term Impinging Streams was not used at that time. Investigations into LIS application in various unit operations in chemical engineering essentially began in the mid-1970s when the concept of impinging streams was presented by Elperin [6]. When the investigations aimed at application development began, there had for a long time been a lack of in-depth understanding of the features of LIS. The ideas guiding the investigations were simply connected to gascontinuous impinging streams and the investigations were still aimed at the enhancement of transfer between phases. Unfortunately, the goal of enhancing transfer is incorrect or, at least, meaningless/'or LIS, as will be discussed later in this chapter. The discovery that liquid-continuous impinging streams promote micromixing very efficiently is the most important progress in the field of IS since the 1990s. It extends greatly the fields in which impinging streams can be expected to be applied effectively, especially those systems involving chemical reaction(s). However, except for combustion of powdery coals and sprayed liquid fuels involving extremely fast reactions, no instance of IS having been applied successfully for any chemical reaction system has been reported, although the monograph by Tamir [5] was entitled "Impinging Stream Reactors". Without doubt, the discovery of the effectiveness of impinging streams for chemical reaction systems is the most important impelling factor for the investigations on LIS, because, alter all, chemical reaction devices are the core of and key to chemical production. With the discovery of the excellent micromixing nature of LIS as the juncture, the emphasis of investigation into impinging streams switched to liquid-continuous impinging streams. Although the statistics are possibly incomplete, in the first stage of IS investigation, i.e., the early 1960s to mid-1970s, the number of papers published on LIS was less than 10% of the total related to IS. In the second stage, from the mid207
208
IMPINGING STREAMS
1970s to the beginning of the 1990s the number was about 20 to 30% of the total, while in the third stage, from 1991 to the early 2000s, a little over a decade, this proportion has increased rapidly to about 60%. In addition, in the latest stage of IS investigation a number of achievements of great potential for application have appeared, such as for the preparation of nano materials by reaction precipitation, some useful properties of impinging streams such as strong pressure fluctuation in liquid-continuous impinging stream were found, and the R&D on LIS devices has progressed significantly. Therefore the switch of emphasis of investigations into IS described above may also indicate that the industrial application of impinging streams may advance rapidly in the coming years.
9.2 DIFFERENCES BETWEEN PROPERTIES OF CONTINUOUS PHASES AND THEIR INFLUENCES ON THE PERFORMANCE OF IMPINGING STREAMS 9.2.1 Differences between properties of liquid and gas Because the composition of liquids is quite different from those of gases, their physical properties are also essentially different. For carrying out impinging streams, the following differences are important: (1) Liquid normally has a higher density than gas by about three orders of magnitude; (2) Liquid has a higher viscosity than gas by approximately two orders of magnitude; and (3) From the point of view of the molecular motion theory, gas has a considerably large molecular free path, while the molecules of stationary liquid can only vibrate and/or rotate round their balanced position with extremely small displacements. These differences must lead to great variation in the performance of impinging streams.
9.2.2 Influences of property differences on the performance of impinging streams The significant differences between the properties of liquids and gases result in totally different performances of impinging streams with liquid and gas as the continuous phase, as discussed below.
9.2.2.1 Transfer coefficients between phases As mentioned earlier, in gas-continuous impinging streams heat and mass transfer between phases are enhanced efficiently mainly by the following factors: (1) Very high relative velocity between phases round the impingement plane, even higher than in common devices by several tens of times; (2) Oscillation movement of particles or
DIFFERENCES BETWEEN PROPERTIES OF CONTINUOUS PHASES
209
droplets penetrating to and fro between the opposing streams; and (3) Very strong turbulence in the impingement zone. When a liquid is used as the continuous phase, the dispersed phase should usually be solid or insoluble liquid; while using a gas as the dispersed phase might be an illadvised option with less practical sense. Whether the dispersed phase is liquid or solid, the particles or droplets will very quickly achieve an identical or nearly identical velocity to that of the continuous phase, namely the liquid, and then follow the streamlines, no matter what their initial velocity, because of the very large friction tbrce between phases and the very small difference between the densities of the liquid and the solid or another liquid. Therefore it is impossible for any larger relative velocity to appear in a liquid-continuous impinging stream device during operation. Secondly, because of very large drag force and very small density difference, the tendency of the particles or droplets penetrating into the opposing stream by inertia is extremely weak, and cannot even be observed. In tact, to date no such phenomenon in LIS has been reported. With LIS, there is very strong turbulence in the impingement zone. However, because of the absence, or extremely weak influence, of the former two factors, both of which are major factors enhancing transfer, no significant enhancement of transfer between phases can be expected in liquid-continuous impinging streams. In fact, the experimental data obtained to date are in agreement with the inference mentioned above. For example, Tamir et al. [ 153, 154] studied the dissolution of urea in water in two liquid-continuous impinging stream devices of different structures; the values tot the mass transfer coefficient were 1.33×10 -4 and 1.60×10 -4 m.s -~, respectively; while Davies [155] studied the dissolution of sodium chloride and potassium sulphate in water in a traditional stirred tank reactor and obtained the data of 1.0×10 -4 and 1.26 m.s -~, respectively, for the mass transfer coefficient for the two substances. The differences in physical properties seem favorable for the dissolution of urea. By comparing these data, less enhancement of mass transfer is observed; although, probably because of their partiality, Tamir et al. considered IS to be superior to other equipment.
9.2.2.2 Interactions between opposing streams An LIS device is usually operated at lower impinging velocity, u0, than GIS although, the momentum transfer between the opposing streams in LIS is much stronger than in GIS because of the very high density of the fluid in the continuous phase. To understanding this more deeply, a semi quantitative analysis is made below. For convenience, the subscripts G and L are used to denote the operating parameters and the properties of the gas and the liquid, respectively, of the continuous phases and it is assumed that the devices of LIS and GIS under consideration have the same cross section area of the drawing tube in the LIS device or the accelerating tube in the GIS device, i.e., A~. = A~; = A; and assume P l . -- 1 0 3 j O G '
hi()L -
0.
lu0~~
(9.1)
210
IMPINGING STREAMS
According to Newton's law, the momentum transferred per unit time in the LIS device, ME, is M L - mLUOL -- VLIOLUOL -- ULAPLUOL -- APLU20L
(9.2)
M G - ApGuZG
(9.3)
M L _ APLUZL _ A × I O 3 p G ( O . l u o G ) 2 _ l O A p G U0G 2 -- 10M o
(9.4)
and, that in GIS is
The use of Eq. (9.1) yields
i.e., the intensity of momentum transfer between the opposing streams in LIS is larger than in GIS by about 10 times. Under conditions of high intensity of momentum transfer, together with the dense assembly condition of liquid molecules, much stronger interaction must occur between the opposing streams impinging against each other, including collisions between fluid elements and/or molecules, mutual pressing and shearing, etc. The intensive momentum transfer and the strong interaction between the opposing streams lead to the global results below:
(1) S t r o n g m i c r o m i x i n g . In fact, Wu et al. [110] and Xiao [13] have found that LIS has the excellent feature of promoting micromixing very efficiently. (2) P r e s s u r e f l u c t u a t i o n : The results of an investigation carried out by Sun et al. [156] showed that, under normal operating conditions of the submerged circulative impinging stream reactor (SCISR), there is considerably strong pressure fluctuation, the major frequencies of which are under about 1 kHz and the maximum amplitude can be as large as 1.4 kPa. Obviously, the two phenomena mentioned above are closely related to each other. The pressure fluctuation implies that the fluid elements and/or molecules vibrate and, correspondingly, part of the kinetic energy of the streams is converted into vibration energy. On the other hand, micromixing occurs at the molecular scale and must be related to the motions of fluid elements and/or molecules. For liquid phase processes occurring at the molecular scale, or multiphase processes with a liquid as the continuous phase, both micromixing and pressure fluctuation have significant positive effects and thus have significant value for application. In practice, the results of a number of investigations have shown excellent performance and application value, such as promoting process kinetics and favoring the preparation of ultrafine particles, etc.
The related substances will be discussed in detail in the following chapters.
DIFFERENCES BETWEEN PROPERTIES OF CONTINUOUS PHASES
211
9.3 SUPPLEMENTARY CLASSIFICATION OF IMPINGING STREAMS Classification is very useful method for scientific understanding. In his monograph [5], Tamir worked out a classification of impinging stream reactors, the major part of which was introduced in Section 5.1 of the Introduction to the present book. This classification is helpful for giving an overview of the field of impinging streams, although some flow configurations included in it are of no practical meaning. It is a little unfortunate that Tamir did not include continuous phase; however, based on the investigations before the early 1990s, this was reasonable. From the discussions above we can see that, because of the significant differences between the properties of liquid and gas, the two kinds of impinging stream with liquid and gas as the continuous phase, respectively, give totally different performances. From the point of view of taxonomy, it is necessary to distinguish them. Therefore a supplementary classification should be made according to continuous phase, i.e., impinging streams should also be classified according to the continuous phase involved in: • Gas-continuous impinging streams (GIS); and • Liquid-continuous impinging streams (LIS).
This Page Intentionally Left Blank
-10MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
10.1 MACROMIXING AND MICROMIXING Mixing between different substances is widely involved in various processing industries and its importance is recognized. The discussion in the present chapter, to some extent, overcomes some macromixing problems from the point of view of the mutual dependence between macro- and micro-mixing; but those aspects independent of micromixing will not be discussed here. According to the properties of the substances, most of the materials involved in processing industries cannot undergo micromixing, i.e. mixing on a molecular scale. The mixing between insoluble liquids, e.g. emulsion, also does not involve micromixing, because the dispersity of any liquid can never achieve the molecular scale level. It is very easy to achieve good micromixing conditions with gases because they have a very large molecular free path, but an investigation on micromixing of gases is of less practical interest. Therefore micromixing of practical interest only occurs between two or more mutual-soluble liquids; the discussions in the present chapter focus on this aspect. Mixing between mutual-soluble liquids is usually carried out by the action of certain external forces, e.g. agitation. Mixing occurs at various scales and the segregation scale varies with time. An idealized model of the mixing between a black liquid and a white liquid is illustrated in Fig. 10.1.
(a) Beginning
(b) Later
(c) Further later
(d) Final
Figure 10.1 Mixing between miscible black and white liquids. It is absolutely necessary to distinguish macro- and micro-mixing from the point of view of their significantly different influences on processes. It may be considered simply that macromixing is the mixing between fluid elements, while micromixing is that between molecules. Macro- and micro-mixing characterize different mixing scales, but are closely related to each other. Macromixing occurs by flow, turbulence, and
214
IMPINGING STREAMS
eddy diffusion, while simultaneously, the segregation scale is reduced. When the segregation scale is reduced to the so-called Kolmogoroff microscale 2, molecular diffusion takes over, crossing the length 2, to achieve complete homogeneity. So, molecular diffusion is the ultimate mechanism of micromixing [22]. The Kolmogoroff length 2 is the minimum size of eddies and, according to the existing theory, depends only on the energy dissipation per unit mass involved, ~.
/~- (V3/E) 1/4
(10.1)
The status of micromixing is described by the parameter "characteristic time constant for micromixing", tM, which can be simply called micromixing time. It represents the time needed to achieve completely uniform micromixing, and is correlated to the microscale 2 by tM =
(0.52) 2 ~ D
(10.2)
where D is the diffusivity. Micromixing can occur at any instant during the mixing operation; but the extent of uniformity varies. For example, even for two large lumps, micromixing also occurs at the interface between the two lumps; but the extent of mixing is very low. Ideal mixing, i.e., a completely uniform status, can only be considered to have been achieved when there is no segregation scale larger than 2 anywhere in the whole of the mixing volume. It is clear that for the processes carried out on a molecular scale, e.g. chemical reactions etc., good micromixing is essential for the molecules to contact the others effectively only under such conditions. The obvious inference that can be made is that both macro- and micro-mixing are closely related to the flow configuration in the device considered.
10.2 METHODS FOR INVESTIGATION OF MIXING PROBLEMS 10.2.1 Macromixing For a long time, the distinction between macro- and micro-mixing was not fully considered, the general concept being just of "mixing". The major parameter examined in investigations on mixing is the Mixing Time, i.e. the time required to achieve full mixing. Although various methods are used for experimental measurement, such as decolorizing, photography, and electro-conductivity etc.; the electro-conductivity method is relatively traditional, feasible, and convenient for most systems. The experimental procedure is as follows: Input a certain amount of tracer, mostly KC1 solution in impulse, into a device filled with process liquid and then immediately measure and record the variations of electro-conductivity at various positions inside the device. The electro-conductivity inside the device gradually approaches uniformity
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
215
with the elapse of time. When the relative deviation of the value measured at any point from the average over the device is less than +5% and, at the same time, the variation of the measured value with time at any point is also less than +5%, a uniform mixing is considered to have been achieved and the time interval from t = 0 to that instant is taken as the mixing time, t,,,.
10.2.2 Micromixing A number of experimental methods and mathematical models for micromixing have been proposed to date. Danckwerts [ 157] first noted the influence of incomplete mixing in homogeneous reaction systems, and proposed the concept "segregation". He suggested using the time-varying parameter "intensity of segregation" to describe micromixing phenomena. The intensity of segregation at an instant was defined as the ratio of the mean square deviation of the composition from the average at that instant to the same deviation at the initial time. From its physical meaning, the parameter can be used as a measure of micromixing. However, the measurement of the two quantities mentioned above must be carried out on the molecular scale and is thus difficult. Following the train of thought of the intensity of segregation, many researchers made great efforts for a long time to solve the problems in experimental measurement. Although various methods, such as electro-conductivity [158], fluorescence [159, 160] etc., have been proposed, it is still difficult to collect correct data on the molecular scale, and thus the micromixing status cannot be reflected correctly. Bourne et al. [161-163] presented a chemical method for micromixing measurement which has been widely accepted and applied. The core of the method is the use of the following parallel-competition reactions as the detection system: A+B~ R
Rate constant k~
(~o.3) R + B -+ S
Rate constant k 2 << k~
where B is the limiting component, i.e., Component A is always excess. Since the second reaction producing S is much slower than the first one (by several thousand times), its pseudo first order reaction time constant 1/(k2CBo) can be taken as the characteristic reaction time constant of the system, &, where CB0 is the initial concentration of B after mixing of the fed reactants. If micromixing is ideal, the reactions would stop quickly without Product S forming, because B is exhausted due the very fast reaction between A and B. On the contrary, if any S forms, there must exist segregation of components, resulting in poor contact between A and B, and, consequently, the reaction between R and B producing S occurs. Therefore the relative amount of S formed can be used to characterize micromixing. For this purpose the selectivity of S, Xs, is defined as
216
IMPINGING STREAMS 2Cs X s = 2C s + C R
(10.4)
Parameter Xs is also called the "segregation index", because its value characterizes simultaneously the extent to which the components are segregated. Theoretically, it is impossible for S to form so that Xs should be zero, provided micromixing is good. On the other hand, if Xs has a non-zero value, segregation must exist. The larger the value for Xs, the higher is the segregation degree. It is clear that Xs is affected by the status of micromixing only in the case where micromixing needs a longer time, i.e., in the range of tM > tR. If tM < tR, Xs will theoretically be zero. However, it cannot actually be zero but a smaller constant, depending on a number of factors, e.g., measuring errors, purities of the reagents and even process water, and operating conditions, etc. Even so, Xs can be used as a guide for understanding micromixing. The relationship measured for Xs versus a certain governing variable, which will be described later, is normally an asymptotic curve: initially Xs decreases when the variable increases; and then gradually becomes a small constant. Thus, one can determine the condition for tM = tR by finding the critical point at which Xs turns from a varying value to a small constant on the measured curve of Xs versus the governing variable. This is the procedure of bounding the micromixing time tM with the characteristic reaction time constant, tR, of the system described by Eq. (10-3). It is the case that the micromixing time, tM, cannot be measured directly, but is bounded by tR. The micromixing time tM is an important parameter that has a practical value for understanding micromixing quantitatively; while Xs can only be used for relative comparison qualitatively, although it can be a guide. The main advantage of Bourne's method is its ability to determine experimentally the values for tv. Of course the experimental procedure for the determination of tM is very troublesome. In order to determine one value for tM, a large set of experiments has to be carried out to yield a curve of tM versus the governing variable. This is probably why a number of the results of investigations reported, e.g. in Refs. [164-166], did not include the data of tM but only analysis and evaluation of micromixing with the segregation index Xs as the criterion. In addition, Fournier et al [167,168] and Villermaux et al [169] proposed several new chemical systems for investigation on micromixing, including that employing the reducing reaction of iodic acid etc.
10.3 FLOW AND MACROMIXING IN SCISR 10.3.1 Design ideas and basic structure of SCISR The submerged circulative impinging stream reactor (SCISR) [15] has been developed by the author of this book on the basis of a number of investigations, some of which were supported by The National Natural Science Foundation of China. It is a new type
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
217
of reactor suitable for the systems of liquid reaction(s) or multiphase reaction(s) with liquid as the continuous phase. Originally, it was designed mainly for preparation of ultrafine powders by reactionmprecipitation. The basic ideas for the design are: (1) Utilizing the feature that impinging streams promote micromixing to create very high and uniform supersaturation for precipitation; (2) Providing longer mean residence times to meet the requirements of most reaction systems. The structure of an SCISR is briefly shown in Fig. 10.2. It consists of two essential elements: impingement between opposing streams and circulative flow. The latter makes it possible to set the mean residence time arbitrarily. The working principles are as follows: The fluids on both sides of the reactor, liquids or liquor suspensions, are pushed forward by the two propellers mounted co-axially to flow through the drawing tubes and impinge against each other at the center of the reactor to form a strongly turbulent impingement zone; The impinged streams flow outward radially, pass through the annular chambers between the walls of the reactor and the drawing tubes to return to the two sides, respectively. Then the fluids are pushed forward by the two propellers flowing through the drawing tubes and impinging against each other again, and so on, to circulate inside the reactor. The reactor can be operated continuously or in batch mode. When operated continuously, the feed fluid(s) enter(s) the reactor at the inlet(s) of the drawing tube(s) on one side or two sides, according to the specific conditions; the part of the fluid over a certain level overflows trom the reactor through the overflow outlet 4 (see Fig. 10.2) to keep the level inside the reactor stable.
I f |
v_
4
~:."""-[] •
3
_
(10 I
2
Figure 10.2 A brief view of the SCISR. 1-drawing tube; 2-propeller; 3-impingement zone; 4overflow outlet.
218
IMPINGING STREAMS
10.3.2 Macromixing time For a full understanding of the nature of the reactor, the flow and mixing status in an SCISR are studied first [ 16, 31 ]. Comparative experiments were made between an SCISR with an effective volume of 3.6x10 -3 m3and a traditional stirred tank reactor (STR) with an effective volume of 0.6xl0 -:~ m 3 for the macromixing times. The electro-conductivity input--response technique was employed, with KC1 solution as the tracer, for the measurement of macromixing time. The experimental measurements were carried out with pure water and a solution of glycerin in water as the process liquids, respectively; and the criterion described in Section 10.2.1 was used to determine macromixing times. The results obtained are shown in Fig. 10.3. As can be seen, under conditions of identical specific effective power, Perf, the mixing time in SCISR is always longer than in STR by 20 to 80%; where the specific effective power is defined as
Pelf = ( P - Po)/VR
(10.5)
where P is the power practically inputted into the filled reactor by the agitator or propellers; while P0 is that inputted into the empty reactor. The comparative results indicated by the data given in Fig. 10.3 are completely opposite to those of Brauer [14] who reported much shorter mixing times in the stagnation jet mixer than in STR. Although the stagnation jet mixer lacks some of the key elements needed for the reactor, it has a flow configuration very similar to that of the SCISR. This phenomenon is unexpected, but the multiply repeated experiments verify that the data are true. 40.0
mlk
Reactor
30.0
Material*
•
SCISR
[]
STR
PW PW
•
SCISR
GW
STR
GW
20.0 a. 10.0
0.0
I
0.8
I
1.2
I
I
I
n
'
i
I
~
1.6 2.0 2.4 Mixing timet .... s
t
2.8
Figure 10.3 Mixing time measured in SCISR and STR (PW: pure water; GW: solution of glycerin in water).
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
219
The results of a theoretical analysis indicate that the longer macromixing time in the SCISR results from its special flow configuration, because in the SCISR there are flow regions without mixing.
10.3.3 Flow configuration and residence time distribution It can be seen in Fig. 10.2 that the flow in the SCISR is in mirror symmetry with respect to the impingement plane, and also approximately in axial symmetry so it is enough to take a half volume of the SCISR, either on the right or the left side, into account in the analysis. According to the flow characteristics, the effective volume of the SCISR consists of three regions of different mixing properties: (1) The impingement zone; (2) The space inside the drawing tubes; and (3) The annular chamber between the wall of the reactor and the drawing tube. The flow in Region (1) can be assumed to behave as perfect mixing because of the impingement between the opposite steams; while those in Regions (2) and (3) can be considered as plug flow because of higher velocity. There is an additional region round the inlet of the drawing tube where the behavior of the flow is highly complicated due to the strong turbulence caused by the propeller and so is very difficult to determine. On the other hand, the residence time of the stream in that region is extremely short, less than 1% of the total. If this tact is taken into account, the influence of the complex flow in that region on the overall RTD may be ignored and thus the flow in SCISR can be represented by the simplified flow model shown in Fig. 10.4, for which the residence time distribution density probability function (RTP) [ 170], E(t), can be represented by
E(t) - -
1
R
Ej(t)+~E2(t)+...+ R+ 1 (R + 1)2
R i-I
(R + 1);
Ei(t)+...
(10.6) -
Ri-I
~ E , ( t ) • ~(R + 1)'
with El(t)-
E()(t),
E2(t)-
E l ( t ) * E ( ) ( t ) - E()', ...
(10.7) Ei(t) - E,_j (t) * E()(t) - Eli i, ...
where R is the circulative ratio, E()(t)is the RTP of the materials passing through the system once, while Ei(t)is the RTP of the materials circulating i times through the system. The symbol "*" denotes the convolution integral defined as
220
IMPINGING STREAMS
Vo
I Plugflow I I re on
Perfect nixing
[
re, on
Figure 10.4 Simplified model for the flow in SCISR.
f (t)* g(t) - j o f ( u ) g ( t - u ) d u - j o f ( t - u ) g ( u ) d u
(10.8)
while ",i" the (i)th-order convolution integral: E0' - E 0 , E 0 , . . . , E 0 w
w-
j
(10.9)
iEo
For the system shown in Fig. 10.4 E0 can be determined by Eo(t) - E m (t) * Ep (t)
(10.10)
and the density probability function for the perfect mixing and the plug flow regions,
Em and Ep, are well known, as E m( t )
--
-~1 exp(-t / i m), Om
E p ( t ) - 8(t-[,p)
(10.11)
where i m and ip represent the mean residence times in the complete mixing and the plug flow regions, respectively. Substituting Eq, (10.6) into Eq. (10.10) yields E o (t) - [__~1exp(-t/O m )] * [ 6 ( t - Op)]
Om
(10.12)
1 exp[-(t-Op)/Om]- ~ce'/em exp(-t / 0 m)
em
em
It should be noted that both the process fluid and the tracer are fed at the inlet of the drawing tube and that the volume of the drawing tubes occupies about one fifth, so there must be E 0 - 0 for t < iv / 5 . If the volumetric fraction of the plug flow region is denoted by fm, then we have
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
i,,, = .l,,, V~, /V() ,
/-~, - (1- f,,, )vR /u(,
221 (10.13)
Eq. (10.12) then becomes 0, Eo(t )
=
V()c(
1 • f,,~
t < 0.2ip
)1i,,,
(10.14)
- V()
e x p [ f , , V R ,],
.f,,,VR
, > 0.2{p
Define the convenient constants A and B as
V()
Vt)e(i-f"')lf'"
A --
~
B
L,,VR
-
(10.15)
~
L,,V~
Eq. (10.14) can be simplified into
0~
E()(t)- ] t
t < 0.2{p (10.16)
lit, A ~,-
t > 0.2ip
and its Laplace transform is written as L[Et,(')]- E/,(s) - A 1 - ~ s+B
(10.17)
Using Eq. (10.17) and the convolution integral law, the Laplace transform of each of the sub-probability density functions in Eq. (10.6) can be obtained to be
(s+B) ~
(i = 1, 2,..., n)
(10.18)
Substituting Eq. (10.18) into Eq. (10.6) and rearranging the result in -E(s)-
" AiR '-I 1 Z - - - i-I (R + 1) i (s + B) i
(10.19)
The inversion of Eq. (10.19) yields E ( t ) - e -B'
ZI
a~ Ri-I
i-i ( i - 1 ) ! ( R + l ) i
[ i-I ],
(10.20)
222
IMPINGING STREAMS
It is noted that the right-hand side of Eq. (10.20) is just the series expansion of an exponential function. Therefore the overall residence time distribution probability density function in the SCISR is obtained to be A AR E ( t ) - +----~ R exp[(R +
1
_ B)t]
'
(t > 0.2t-p ) -
( 10.21)
The residence time distribution model above includes the two parameters f , and R. The volume fraction of the impingement zone is originally determined according to the visual observation of the flow appearance to be fm = 0.186, while the value for it determined later from an analysis for the fluctuation intensity in the investigation on pressure fluctuation in the SCISR, of which the details will be discussed in the next chapter, is very close to the value mentioned above. This value is therefore confirmed. The circulative ratio, R, is obtained by fitting data; and then the probability density function E ( t ) can be determined. It is notable that the values for R obtained by fitting data, i.e. indirectly measured values, and those obtained by the determination of the impinging velocity, u0, made later are in good agreement. Such agreement indicates the reasonableness and feasibility of the theoretical and experimental methods employed in several investigations carried out by the author of this book. Using the electro-conductivity input-response technique still with a KC1 solution as the tracer, the residence time distributions are measured at various rotary propeller speeds and the data are illustrated in Fig. 10.5; for comparison the results calculated with the model equation Eq. (10.21), with the values for R obtained by fitting data are given in the same figure as the curves. The curves coincide well with the experimental data as can be observed in the figures, implying that the theoretical model is reasonable. The important conclusions that can be drawn from the discussion are: the SCISR has the flow configuration of circulative perfect mixing-plug flow in series; and the region without mixing occupies a large volume fraction, >80%. Therefore the longer mixing time in the SCISR obtained in the last section is reasonable.
10.4 MICROMIXING IN SCISR The submerged circulative impinging stream reactor (SCISR) has exhibited excellent performances in a number of applied investigations, especially in the preparation of ultrafine powders, suggesting that SCISR has a good micromixing capability. In order to gain a more quantitative understanding, an investigation on micromixing was carried out experimentally, as described below.
10.4.1 Experimental equipment and procedure The experimental SCISR is the same as that used for the measurements of macromixing and residence time distribution, as shown in Fig. 10.2, while its major dimensions are shown in Fig. 10.6 and the equipment system scheme is illustrated in Fig. 10.7.
M I C R O M I X I N G IN L I Q U I D - C O N T I N U O U S I M P I N G I N G S T R E A M S
223
0.3
V~x ! ()~', m ~.s i
1,o
~.
4.17:
0.2
3 0.1
0.0
A
()
1200
2400 t/S
A I A
A | A
3600
4800
(a) N =700 rpm ().4
0.3
,.....
N, rpm I
•
900
2
•
I 2()()
3 •
I 5()()
().2
().1
I
)
~4()
~
I
~
48()
" I~"-~T'w'~-
~,
72()
~-
-~'-~ --~
960
I 2()()
t/s tb) ' ~ , - 8 . 3 3 x 1 0 ~ m~.s i F i g u r e 10.5 Comparison between results measured and calculated for RTP in an SCISR.
10 10 i
_____.,
7() - - - ~
256
F i g u r e 10.6 Dimensions of SCISR (in mm).
224
IMPINGING STREAMS
~
~
1
~...j2 vA 3
B
I
_L I r,1 i
Figure 10.7 Scheme of the experimental system.l-tank; 2-rotameter; 3-overflow outlet; 4-valve; 5-propeller.
10.4.2 Governing variable and its experimental measurement Essentially, the micromixing in an SCISR is caused by the impingement between the opposing streams. Correspondingly, the impinging velocity, i.e., the velocity of the liquor streams at the exits of the drawing tubes, u0, should be the governing variable. In all the investigations carried out in the past the segregation index, Xs, was correlated to the Reynolds number Re. This is questionable, because Re cannot be used as the criterion for scaling-up the process and the related device. As is well known, Re is defined as Re = d ° p f u ° ltl
(10.22)
If Re was kept constant in scaling-up, a very small impinging velocity u0 would be employed for a large system with a large diameter drawing tube do. On the other hand, the kinetic energy for the impingement is supplied by the streams, which are positively proportional to u 2 . Therefore a very small impinging velocity directly implies very small kinetic energy of the streams and consequently very small energy dissipation per unit mass and thus very poor micromixing. This is a typical practical instance indicating that the traditional method based on mlaw for scaling-up in chemical engineering is inapplicable not only for most of the systems involving chemical reactions but also for some physical processes, like impinging streams. In view of the extreme importance of the impinging velocity, u0, in the implementation of impinging stream technology, in all the investigations worked by the author of this book u0 is employed as the governing variable, as was done in the chapter on absorption. Since the drawing tube in the experimental reactor is very short, the measurement of the velocity in the tube, i.e., the impinging velocity u0, is very difficult. Xiao [13]
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
225
measured the mean velocity in the drawing tube with the electro-conductivity inputresponse technique with KC1 as the tracer. Since the stream passing through the tube needs a time, the response of the out stream to the input signal has the feature of pure lag, i.e., velocity-time lag. A typical response curve automatically recorded is shown in Fig. 10.8.
° Iiiiiiiiiiiiiiiiiiiiiiiiiiiiiiiiiii i!!ii!!!ii!~ Lp ~ii!!!!i!!iiiiiiiiiiiiiiii!ii!iii!!iiiiiiiii!ii!!!i!i ...i...i...L.':...::..2..i...............":...i...i...i.tS:...i._::...i..L..i...i...i...L..::...i...i...i.. Figure 10.8 Response curve recorded in the measurement of impinging velocity. Because the moving velocity of the recording paper is constant, the lag time can be determined according to the length the paper travels up to the instant the response just emerges, Lp, provided that the instant the tracer was inputted, to, was exactly labeled. Then the mean velocity can be calculated by u{} - L d / tl~
(10.23)
where Ld is the length of the drawing tube. The results measured are given in Fig. 10.9, in which the regression curve is used for determining the relationship between the rotary speed of the propeller, N, and the impinging velocity later. It should be noted that the curve is not generally applicable. Even for another reactor with the same dimensions, its applicability is uncertain, because the flow rate transported by the propeller depends not only on the rotary speed of the propeller but also on the shape of the paddles and the gap width between the propeller and the drawing tube, while errors made in mechanical manufacture are unavoidable. 0.8 T
c~
0.6 0.4 0.2 0.0 3(X) 5(X) 700
900 ll00 1300 1500 N, r.min J
Figure 10.9 hnpinging velocity measured at different propeller rotary speeds.
226
IMPINGING STREAMS
An important piece of information that can be obtained from the measurement is that the impinging velocity is very small under normal conditions of SCISR operation, only of the order of magnitude of 10-1 m.s -~, implying very low power consumption, although the SCISR gives an excellent performance of promoting micromixing efficiently.
10.4.3 Experimental procedure According to the works by Bourne et al. [ 162, 163], the azo coupling reactions between o~-naphthol (A) and diazotized sulfanilic acid (B) to produce monoazo dye (R) and bisazo dye (S) are used as the competitive-consecutive (series-parallel) reaction scheme for the determination of micromixing, for which the rate constants have been determined [163], and their values at 298 K are k~ = 7.3x 103 m3.mol -~-s-I k2 = 2.08 m3.mol-l.s -!
kl/k2 = 3500 The experiments are carried out in continuous operation mode. In all the runs the initial concentration ratio of A to B, CAJCBo, is kept constant at 1.4 in order to ensure B being the limiting component. Samplings are made at the overflow outlet of the reactor, and the samples are analyzed with a spectrophotometer of Type 722S. The absorbance of each sample was measured at an interval of 10 nm in the wave length range from 460 to 600 nm, and the concentrations CR and Cs in the sample mixtures are determined by a linear regression with the specific molar absorbance measured. Then the values for the segregation index, Xs, are calculated by Eq. (10-4) with the determined values for CR and Cs.
10.4.4 Major results for micromixing 10.4.4.1 Influence of feed flow rate ratio The results of the influence of the feed flow rate ratio of Reactant A to Reactant B, VA/VB- tZ, on the segregation index Xs at the impinging velocity of u 0 - 0.16 m.s -~ are demonstrated in Fig. 10.10. It is interesting that tz exhibits a significant effect on Xs. As can be seen in this figure, in the range of tz< 5 Xs < 0.03, suggesting good micromixing, while outside this range the micromixing becomes poor. Liu et al. [166] also obtained similar data in their investigation on micromixing in a rotating packed bed. This phenomenon should result from a macromixing condition. To enable B to react completely with A yielding a smaller segregation index, it must be ensured that B fully contacts A. It is clear that such a condition would be achieved more easily when the feed rate ratio o(is smaller. If this ratio is large, say tz = 15, it would be difficult for
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
227
the molecules of B to contact all the fluid elements of A, even if the macromixing is very strong. This may account for the phenomenon exhibited in Fig. 10.10. Figure 10.1 1 illustrates the variation of Xs with the impinging velocity at a given feed rate ratio of c~= 1.0. Even at the lowest impinging velocity tested, u()= 0.16 m.s -~, the value for Xs is only 0.013, while the corresponding conversion of B is 0.987, suggesting Xs is low enough that complete uniform mixing can be considered to have been achieved. The following can be concluded fi-om a combined examination of Figs. 10.10 and 10.11: (1) Generally the micromixing in the SCISR is very good even at an impinging velocity as low as 0.16 re.s-J; and (2) Micromixing is restricted by macromixing to an extent. However, Figure 10.11 indicates that at lower teed rate ratio c~, the determination of micromixing time becomes difficult when all the values for Xs are very small. On the other hand, the case of large teed rate ratio must be encountered in practice. To get representative data that are safe for application, all the other experiments are carried out at a fixed teed rate ratio of c~= 15. 0.12 0.08 0.04
O.(X)
I
I
I
()
I
5
I
I
I
1()
I
15
m
20
Figure 10.10 Effect of the flow rate ratio a'on X s. u()=0.16 m.s j" CB0=2.5 mol.m -(''
V\=8.33×10 ~ mS.s 1
0.10 0.08
:<
~c
0.06 0.04 0.02 ,
0.1
i
0.2
i m .... i ....
0.3
;--m
~---
q ....
0.4 l,l(),m - s
)
0.5
.... ~
j - - - ~ m
0.6
0.7
I
Figure 10.11 Variation of X~ with tt()at a'=l.0. C~()=2.5 lnol.m-(~; V:\=8.33x10 -~ m~.s i
228
IMPINGING STREAMS 0.16 0.12 ~< 0.08
-=
1 =:
0.04 0.00
I
0.15
I
I
I
I
I
I
i
i
i
0 . 2 5 0 . 3 5 0 . 4 5 0 . 5 5 0.65 Uo
Figure 10o12 Determination of micromixing time. CB0=2.5 mol'm-3; VA=8.33x10-~ m3.s-~" a~=15; T=298 K.
10.4.4.2 Determination of the micromixing time As mentioned above, to determine one value for tM needs an experimental curve of X~ versus u0 for the case where the impinging velocity u0 is taken as the governing variable. The plot of a set of typical data is indicated in Fig. 10.12. From the curve the critical point at which Xs turns from a varying value to a small constant is found at u0 = 0.184 m/s. With the known reaction rate constant at 298 K, the micromixing time tM can be calculated as tM - t R -
1
-
k2CBo
1
2.08x2.5
= 0.192 s - 192 ms
In comparison, the micromixing time in a general stirred tank reactor (STR) ranges from 500 to 1000 ms. If it is noted that the impingement zone in the SCISR occupies only a small volumetric fraction (less than 20% of the total), the value of 192 ms for the micromixing time in the SCISR indicates clearly that impinging streams do promote micromixing very efficiently.
10.4.4.3 Influence of the impinging velocity on micromixing An inference drawn from Fig. 10.12 is that at an impinging velocity higher than 0.184 m.s -j, the micromixing time should be shorter than 192 ms. To understand how u0 affects tM, various conditions yielding different characteristic reaction time constants, tR, have to be created. Because of the difficulty in preparing a high concentration solution of B, four sets of different T and CB0 combinations are used, each set corresponding to one value of tR, as listed in Table 10.1. Each set of experimental measurements yields one curve of X~ versus u0 and corresponds to one value of tR. All the curves have a shape similar to that shown in Fig. 10.12, but various curves have somewhat different asymptotic constants of Xs. It seems that, at higher concentrations of solution B, macromixing has a greater influence on micromixing.
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
229
Table 10.1
Experimental conditions for determination of tM VS U0 Set No
T, K
k_~,m".mol-J.s J
CB,~, mol'm -~
1
298
2.08
2.5
2
303
2.63
2.8
3
308
3.30
3.2
4
308
3.30
3.5
The data determined for the micromixing time, tM, from various curves are illustrated in Fig. 10.13, which meet approximately the relationship below: i M o~ ull 1-5
(10.24)
The reason for the decrease in the micromixing time with the increase in impinging velocity is clear: a higher impinging velocity implies an increased energy dissipation rate and thus more efficient micromixing.
10.4.5 Comparison between micromixing performances of SCISR and STR It is of interest to compare experimentally the micromixing performance of an SCISR with that of an STR. The effective volumes of the two reactors used in the comparative experiments are: SCISR 3 . 6 x l O ~ m -~ and STR 0.6x10 -:~ m ~. Since the STR has no parameters corresponding to the impinging velocity in the SCISR, u0, the specific effective input power, P~,~, and the rotary speeds of the propellers in the SCISR and the paddles in the STR are taken as the basis for comparison, respectively, where P~j is defined in the same way as described above. All the other operating conditions in both the reactors are the same. The data measured tot the segregation indexes in the SCISR and the STR are shown in Figs, 10.14 and 1{}.15, respectively. It is very clear that the micromixing performance of the SCISR is significantly better than that of the STR. 250 200 fi 150
z--
~ 100 3
50 l
0.15
l
l
i
l
l
0.2
l
l
l
i
i
l
l
4 l
i
0.25 l!(),
m.s
i
i
0.3
l
l
l
l
0.35
-.]
Figure 10.13 Influence of impinging velocity on micromixing time at (~=15
230
IMPINGING STREAMS 0.040 0.030 0.020 0.010 0.000 0.2
0.6
1.0
1.4
P~f.f,
1.8
2.2
k W . m -3
Figure 10.14 Influence of specific effective power on Xs in SCISR and STR. • - - 0 . 6 x 1 0 -3 m ~
STR; • -- 3.6x 10-3 m 3 SCISR.
0.04 0.03 0.02 0.01 0.00 300
600
900
1200
1500
Rotary speed N, r.min-~
Figure 10.15 Segregation index at various rotary speed of propellers or paddles in SCISR and
STR. 4 - - 0 . 6 x 1 0 -3 m 3 STR; o--3.6x10 -3 m 3 SCISR.
10.4.6 Comparison between measured and theoretically predicted results for micromixing time According to the existing theoretical model c o m m o n l y used, micromixing depends essentially on the energy dissipation rate per unit mass, e. In an S C I S R the energy c o n s u m e d in micromixing is supplied by the dynamic energy of the opposing streams impinging against each other. In operation the two streams have the same velocity, and the energy dissipation rate can be written as
P
-
2 x - -1 m u o2 2
where m is the mass of each stream:
-
mu
:
o
(10.25)
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
231 (10.26)
m - AdPl.u o
Combining Eq. (10.25) with Eq. (10.26) yields 3
(10.27)
P - AdPl.u o
The micromixing occurs mainly in the impingement zone [31]. If the volumetric fraction of the impingement zone is denoted by f,,, we arrive at
v,,, - .f ,,, vR
(10.28)
Thus the energy dissipation rate per unit mass, & can be determined by
~" .
.
P
v,,~p,
.
.
AdU(3)
(10.29)
,f mVR
Substituting Eq. (10.29) into Eq. (10.1) and rearranging the resulting expression in the Kolmogoroff length to be A _
l(/Zt, PI" )3 ] 1 / 4 _ [
fmVR (
g
Ad
/"/f
) 3 ] 1/4
(10.30)
PlU()
Combining Eq. (10.30) with Eq. (10.2) leads to 0.25 f,,1Vlq ,uj- .~ I/', 0.25 fmV~, 1/2 -J.5 tv=--[ ( ) l .... [ (,uj)3] uo D
Ad
pr.uo
D
Ad
PI
(10.31)
where the volumetric fraction of the impingement zone has been determined to be f,, = 0.186" while all the other structural or property parameters are known. The micromixing time calculated for the four sets of conditions yielding the data listed in Table 10.1 are given in the fourth column of Table 10.2; and the corresponding data measured experimentally are listed in the fifth column of the same table for comparison. Table 10.2
Calculated and experimentally measured micromixing time No
T, K
~t,,,m/s
lx,,.,,~,ms
l~,,.,.,~,,ms
tv~,Jt~l.~.,p
1
298
().184
594
192
3.09
2
303
0.245
285
136
2.10
3
308
0.255
205
95
2.16
4
308
().32(~
142
87
1.63
232
IMPINGING STREAMS
It can be seen from the data listed in Table 10.2 that the values calculated with the model above for the micromixing time are systematically longer than those measured by approximately 2 to 3 times, suggesting the micromixing performance of the SCISR is far superior to that predicted by the existing theory of micromixing. Such deviations cannot be explained by measuring error because all the experimental operations were carried out with great care. The only reasonable explanation is that the existing theory of micromixing lacks generality and is not applicable to certain process systems including liquid-continuous impinging streams. As has been mentioned above, the flow configuration in the SCISR is quite different from that in the traditional the STR. In the SCISR a number of phenomena favor the occurrence of micromixing, such as the strong interaction between the opposing streams including collisions, shearing, pressing etc., and also pressure fluctuation [156, 171]. It is most likely that a combination of all these factors promote micromixing. On the other hand, the existing theory of micromixing was based on the data from experiments carried out in the STR, and, naturally, the phenomena mentioned above had not been taken into account. So, the deviations indicated in Table 10.2 are not difficult to comprehend. Although certain factors existing in the SCISR have been found to promote micromixing efficiently, unfortunately their individual contributions cannot yet be described quantitatively and further investigations are called for.
10.4.7 Relationship between macro- and micro-mixing In Section 10.1 the general relationship between macro- and micro-mixing was discussed from the point of view of the mechanism in which mixing occurs. Put simply, micromixing between two or more mutual-soluble liquids occurs first through macromixing. Seemingly the inference would hold that the micromixing time should be longer than the macromixing time, i.e. tM > tin. However, this inference is suitable only for the case where macro- and micro- mixing occur and proceed continuously in the same specified space. Comparing the data shown in Figs. 10.3 and 10.13 it follows that the macromixing time, tin, is longer than the micromixing time, tM, in the same SCISR by about one order of magnitude. However, this is not in contradiction with the mechanism of mixing described in Section 10.1. In consistence with the measuring method, the macromixing time tm indicates the time at which the bulk of the fluid inside the reactor has achieved the overall condition of uniform macromixing. The fact of longer macromixing time is due to the special flow configuration of circulative perfect mixing-plug flow in series in the SCISR. The region in the SCISR without mixing occupies a major part of the total volume (over 80%) but does not play any role in mixing, while the effective mixing essentially occurs only in the impingement zone. The fact that the micromixing time in the SCISR is very short indicates only that it is very easy for the impingement zone to reach the state complete micromixing so that Reactant B is consumed immediately due to the reaction once it contacts Reactant A of excess, yielding a very low segregation index Xs. This situation does not imply that the whole of the reactor
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
233
has achieved a condition of molecular uniformity. In fact, according to the arrangement of flows, there must be a difference between the compositions of the streams, at least, those inside and outside the drawing tubes. For example, the concentrations of A and B in the inside stream must be higher than in the outside stream, while the concentration of P in the inside stream must be lower than that in the outside stream. In any case, the efficient micromixing in the impingement zone is a very important and useful feature of liquid-continuous impinging streams. An understanding of the relationship described above may be helpful for further development of SCISR application.
10.5 MICROMIXING IN IMPINGING STREAM REACTOR WITHOUT CIRCULATION In addition to those described in the previous sections, other investigations on micromixing in LIS have also been reported, e.g., in Refs. [22, 164]. The work to be introduced below is a representative one. As part of the fundamental work for the preparation of ultrafine medicines, Mahajan et al. [22] studied micromixing in the two-jets mixer. The two impinging jets (TIJ) mixer Mahajan et al. used as the reactor is briefly shown in Fig. 10.16. The structure of the TIJ mixer is similar to the device designed by Midler et al. I172]. The device is without material circulation, and can be operated continuously in two modes: free and submerged jet impingement. In the former mode the level inside the reactor is just like that shown in Fig. 10.16, while in the latter the level is increased to be over the jetting tubes. In the case of free jet impingement, two opposing coplanar jets of small diameter and with slight inclination impinge against each other to form a liquid film that eventually disintegrates into small droplets. The size of the film increases with jet Reynolds number, and the film itself remains vertical as long as both jets have equal momentum at the point of impingement. To investigate a possible scale-up criterion for the TIJ mixer, the jet nozzle diameter was changed using HPLC tubing of 0.5, 1 and 2 mm; while the length-to-diameter ratio was kept constant at 10.
Figure 10.16 Two impinging jets (TIJ) mixer.
234
IMPINGING STREAMS
The reaction system, the experiment procedure, and the analytical method used for the determination of micromixing in the TIJ mixer are the same as those described in the last section of this book; but Mahajan et al. correlated their experimental data not with impinging velocity u0 but with the jet Reynolds number Re. Also, the researchers employed the measure of increasing both the initial concentration CB0 and the reaction temperature to raise the sensitivity of the procedure. The characteristic reaction time constant tR- 200 ms at 25 °C and CB0 - 2.5 mM, while tR- 65 ms at 35 °C and CB0 = 4.7 mM, which can be used to bound the micromixing times, tM, no greater than them, respectively. The major results they obtained are as follows: (1) The results obtained under the conditions of tR- 200 ms indicate that, in the range of Re > 700, both submerged and non-submerged operation modes can yield a micromixing time less than 200 ms; while at smaller Re the micromixing becomes no good. (2) The effect of mixer geometry, mainly the nozzle diameter, on micromixing was studied in the submerged operation mode by comparing the measured values for the selectivity of S, Xs. The results showed that, for a given Re, the micromixing quality decreases with increasing jet diameter; or a higher Re number is required to maintain the same micromixing at larger jet diameter. From the comparison between the critical (break) points on the X s m R e curves it follows that, as the Reynolds number decreases, the micromixing becomes poorer and tM increases. Yet, in the submerged jet operation the macromixing associated with the entrainment of surrounding liquid by the jets introduces an additional factor, depending upon geometry and hydrodynamics, which affects the selectivity curve as a function of Re. (3) The influence of the impinging distance was also studied under the identical conditions of tR = 200 ms and submerged jet operation. At a given Re, Xs increases with increasing impinging distance; however, the dependence is very weak over the range of the distance tested. For practical purposes therefore the micromixing in the TIJ mixer can be assumed to be independent of the impinging distance, provided the two jets do impinge upon each other. (4) The interpretation of the data for the micromixing time results in the relationship below: d, tM ~ if-X=) iNe
i.5
1
.-----?j
(10.32)
//0
In addition, Mahajan et al. compared the correlations obtained by various researchers in mixers of different flow regimes and different geometries.
MICROMIXING IN LIQUID-CONTINUOUS IMPINGING STREAMS
235
10.6 COMPARISON BETWEEN THE INVESTIGATIONS ON MICROMIXING IN LIS AS CONCLUDING REMARKS According to the classification, both the reactor used by Mahajan et al. and that described in Section 10.4 belong to the liquid-continuous impinging streams type of device. A comparison between the two investigations, including their substances and results, is very interesting. From the comparison the following are of significance:
(1) Both investigations employ the same reaction system for the determination of micromixing and the results from both investigations demonstrate that liquidcontinuous impinging streams has the excellent feature of promoting micromixing efficiently. (2) The dependences of the micromixing time upon the impinging velocity, tM versus u~, from both investigations are the same. The two results verify each other, suggesting the relationship of t M o< u~ a5 does reflect the objective regularity of the micromixing phenomena in liquid-continuous impinging streams.
(3) The methods for interpreting the data used in the two investigations are different. Employing the traditional method based on the re-law, Mahajan et al. correlated the micromixing time tM with the Reynolds number Re, while the author of the present book disregarded the traditional method and, from a qualitative analysis of the energy source for micromixing, took the impinging velocity u0 as the governing variable to correlate the data for micromixing and, furthermore, took it as the basic parameter for scaling-up the process and equipment. As a result, the method based on the ~-law encountered difficulties. Thereupon, Mahajan et al. proposed that for a larger jet diameter a larger Re number is required to keep the micromixi~'~g quality the same, while this causes the Reynolds number Re to completely lose its meaning as a criterion for scaling-up. As discussed above, if the Re number is taken as the criterion for scaling-up, i.e., keeping it identical in scaling-up, then, according to the definition of Re, one has to use a very small impinging velocity for a large system with a large jet diameter. On the other hand, the energy required for micromixing is supplied by the streams impinging upon each other, which is positively proportional to the square of the impinging velocity, so a very small impinging velocity in a large system must result in extremely poor micromixing. Therefore the results Mahajan et al, obtained verify the inference that the Re number cannot be used as the criterion for scaling-up, but the impinging velocity u~ can. (4) The reactors used in the two investigations have different flow configurations and may be applicable in quite different areas. The essential difference between the two reactors is that they are with or without circulation of material. In the reactor with internal circulation the mean residence time can be arbitrarily set to meet the requirements of most reaction systems, while the reactor without material circulation has very short residence times so that its application is limited to certain special systems only. In addition, the flow configuration of circulative perfect mixing--plug flow in series brings some useful and special functions. For example, in the preparation of ultrafine products by reaction--precipitation, after
236
IMPINGING STREAMS
nucleation in the impingement zone the flow of suspension with very low supersaturation passing through the region without mixing favors the deactivation of the surface of newly formed fine particles, and in crystallization the same flow configuration favors the growth of crystals to yield large sized crystalline, etc. (5) Neither investigation yielded a satisfactorily quantitative description for either the mechanism or the regularity of liquid-continuous impinging streams promoting efficient micromixing. This still seems to be a problem that needs further in-depth investigation. (6) The shortest micromixing times bounded in the two investigations are 87 and 65 ms, respectively, although these would not be the limitations of liquid-continuous impinging streams promoting micromixing. For example, with an improved structure the submerged circulative impinging stream reactor may be operated at higher impinging velocity, and thus the micromixing time may be even shorter. From the point of view of raising the performance of the reactor, this is the direction for further improvement in which great efforts should be made. In both investigations the azo coupling reactions between cz-naphthol (A) and diazotized sulfanilic acid (B) to produce monoazo dye (R) and bisazo dye (S) are used as the competitive-consecutive (series-parallel) reaction scheme for the determination of micromixing. Theoretically, increasing the initial concentration of Reactant B and/or the reaction temperature can yield a shorter characteristic reaction time constant tR for bounding even shorter micromixing time tM. However, as observed in the experiments carried out by the author of this book and colleagues, further increases in the initial concentration of B and the reaction temperature can cause extreme difficulties in the experimental operation and result in confusing data, and thus are not practically feasible. A number of researchers have carried out investigations into developing new methods for the determination of micromixing. In addition to those by Fournier et al. [167, 168] and Villerrmaux et al. [169], Kling and Mewes [173] proposed a two-tone laser induced fluorescent method for the measurement of macro- and micro-mixing, which can determine the local segregation intensities but cannot determine the micromixing time. Therefore the development of a new experimental method that can determine the micromixing time, preferably a method of online fast-measuring, is another problem for investigations into micromixing that needs solving urgently.
-11
-
PRESSURE FLUCTUATION IN THE SUBMERGED CIRCULATIVE IMPINGING STREAM
REACTOR
11.1 INVESTIGATION METHOD OF PRESSURE FLUCTUATION 11.1.1 Meaning of pressure fluctuation As described in Chapter 9, the major characteristic of liquid-continuous impinging streams (LIS) is that a very strong interaction occurs due to the impingement upon each other of opposing streams with very high momentums. According to the general principles of physics and hydrodynamics, one of the global representations of such strong interactions is pressure fluctuation; once generated, the pressure fluctuation will transmit in various directions in the tbrm of longitudinal waves, with fluid as the medium. The generation of pressure fluctuation implies the following: (1)Vibration occurs with molecules and/or fluid elements and thus their modes of motion change; such change must affect micromixing because the latter is closely related to the motions of molecules and/or fluid elements; (2) As the vibration is generated, energy conversion must occur and at least part of the dynamic energy of the flow is converted to vibration energy; and (3) As the conversion of energy occurs, it is likely that the energy distribution over the molecules may change randomly and, as a result, part of the molecules may get more energy to achieve higher levels meeting the requirement for inducing reaction(s), so that the process kinetics may be enhanced. In fact, experimental evidence supporting this inference has been obtained, as will be described in the next chapter. Therefore pressure fluctuation is one of the most important characteristics of liquid-continuous impinging streams, and investigations into pressure fluctuation is of interest for understanding the performances of LIS and also for further development of its application.
237
238
IMPINGING STREAMS
11.1.2 Investigation method of pressure fluctuation The most important characteristics involved in pressure fluctuation are the frequency and the amplitude and must therefore be the major focus of investigation into pressure fluctuation. In addition, it is necessary to understand the distributions of the frequency and the amplitude of fluctuation over the space inside the reactor and the influences of some factors on them. The amplitude mentioned above indicates the extent of the pressure fluctuation; actually it characterizes the intensity of the fluctuation. For convenience, the term "fluctuation intensity" is also used in this chapter as it means the same as "amplitude". In principle, pressure fluctuation signals can be determined directly by suitable probes. However, the device used for the measurement must meet two essential requirements: (1) the frequency response of the instrumentation, including the probes and recorder etc., must match the range of signals to be measured, and (2) the influence of the probes on the flow field in the space to be detected must be minimized as possible. Otherwise, reality or even a near reality cannot be achieved. As mentioned in the previous chapter, the most important operating parameter for an SCISR is the impinging velocity, u0. It defines the kinetic energy possessed by the streams taking part in the impingement and thus must have an essential effect on the pressure fluctuation. It has been found by Wu et al. [ 110] that the Reynolds number based on the classical n-law cannot be used for interpreting data for micromixing in an SCISR, and the impinging velocity is chosen to be the basic parameter for the design and scaling-up of LIS devices. It is clear that u0 is more directly and closely related to the pressure fluctuation, and so it is also used as the basic affecting variable for investigation and interpretation in this work. Two methods were employed for the substantial studies mentioned above, as described below.
11.1.2.1 Statistical analysis of fluctuation intensity The instantaneous pressure at a certain point in the reactor can be considered as the sum of the average pressure and the pulsant one. To describe the intensity of fluctuation, the standard deviation of the instantaneous pressure, PSD, is used as the index. Higher PSD suggests stronger pulsant. It is calculated by (11.1) where ~ is the average pressure: _
1
n
P=-- ZPj n i=1
(11.2)
n is the number of sampling times at the corresponding point, n = 132000 in the investigation, while pj is the instant pressure obtained in the j-th time of measurement.
PRESSURE FLUCTUATION IN THE SCISR
239
11.1.2.2 Analysis for power spectrum The basic mathematical method for power spectrum analysis is the Fourier transformation. By the way, transient fluctuation can be expressed as the sum of the number of simple harmonic waves, which is helpful for understanding fluctuation. A frequency spectrum analysis for pressure signals can yield a profile of the frequencies and that of the amplitude along the frequencies. The basic equation of Fourier transformation can be expressed as
H ( f ) - ~_~ t)(t)e -i7~~ dt
(11.3)
and, for a time-discrete system, it call be rewritten as ,,_i
2x-
i--~
n
H(.[],)- ~ P iexp(-i
n
jk),
n
---
(11.4)
where i - ~ / - 1 • #1 is the number of measuring time points, n - 132000 in the present study" pj the measured value at point j; H(fk) is the value for H at k after transformation, and the frequency at k is f,, -
km
(11.5)
11
where m is the number of sampling times in one second, i.e., the sampling frequency, m = 120000 in the present investigation. The power spectrum describes the frequency composition of the fluctuation, and, for the fluctuation signals, it reflects the distribution of wave energy along frequencies. It can be expressed square
root
ogqH S)l)'
ill
spectrum
three forms" the mean square spectrum,
(RMS),
H ( f ) I , and the
logarithmic
IH S)I',
the mean
power
spectrum,
In calculations using Fourier transformation, the frequency-truncation of
measured signals may lead to "leakage". In order to reduce such leakage, the Haining window function is used, which is defined as 1 w(t) . . . .
2
1
2
2xt cos~,0
r<
< t < T~.
(11.6)
where T~. is the truncating interval, i.e., the length of the window function, T c - m/n here. Using the window function, the Fourier transformation for time-discrete system becomes
H(,/I, )
_ '~
- w ( J ) P i exp(-
i_-=()
HI
i27c 17
kj)
(11.7)
240
IMPINGING STREAMS
11.2 EXPERIMENTAL EQUIPMENT AND PROCEDURE 11.2.1 Experimental equipment The system of experimental equipment for measuring pressure fluctuation is illustrated in Fig. 11.1. The structure of the experimental SCISR is the same as that shown in Fig. 10.2; for convenience of operation it is without cover. The propellers are driven by two motors of stable speeds and with a stepless-speed-adjustable device. Highly accurate micro pressure probes of Model XCQ-062 made by the Kulite Corporation are used for measuring the pressure fluctuation. The outside diameter of the probes is 1.6 mm, which is very small in comparison with the dimensions of the reactor so that their influence on the flow field can be negligible. In addition, the inherent frequency of the probes is 330 kHz, much higher than those to be measured, and thus complete fluctuation signals can be detected without any loss.
2
l L'° Figure 11.1 System scheme for measurement of pressure fluctuation. 1-SCISR; 2-pressure probe; 3-pressure transducer; 4-A/D PCI boards; 5-computer.
0.3
0.2 & m-
0.1 0.0 200
.m
400
~-~
•
600
800
1000
N, r-min -1 Figure 11.2 Relationship between uo and N.
1200
PRESSURE FLUCTUATION IN THE SCISR
241
The experimental measurements are carried out with water as the process liquid. The dynamic measuring system of Model DSTS10 developed by Beijing Aerospace University is employed for signal transformation and recording. Its modulator consists of an exciting power source for transducer, amplifier, filter, zero-setting circuit, conditioner, and PCI card. The signals detected by the probes are transmitted to a computer through two PCI boards, each of which contains five individual paths. After filtering, the amplified signals are transmitted to the sampling card and recorded in it for further treatment.
11.2.2 Measurement and control of the impinging velocity Similar to the case of the investigation on micromixing, the impinging velocity cannot be adjusted and controlled directly, but is done by changing the rotary speed of the propellers, N. Prior to all the measurements the curve describing the relationship between u~ and N was calibrated with the same method as that used in Ref. [110], and the results are shown in Fig. 11.2, in which the curve is, in turn, used for conversion between the rotary speed and the impinging velocity in the data treatment. The curve in Fig. 11.2 is essentially the same as that shown in Fig. 10.9; but with some differences in specific data. The existence of such differences is natural, because the shape of the propeller paddle and particularly the width of the gap between the paddle of the propeller and the drawing tube have a fundamental influence on the flow rate drawn by the propeller, while errors in mechanical manufacture are also unavoidable..
11.2.3 Arrangement of measuring points and sampling frequency In order to understand the profile of the pressure fluctuation over the volumetric space inside the reactor, multipoint measurement is carried out in each run with five probes. The measuring points are arranged according to the coordinate system shown in Fig. 11.3, where the x-z~ plane is the impingement plane, x-y is the horizontal plane, and y-~ the vertical plane; the values of the coordinates are in mm. The flow inside the SCISR is considered to have an approximately axial symmetry. For convenience, part of the data are interrelated in a pillar coordinate system, and the radial coordinate, r, is determined by F--
X - q-- .7-
The sampling frequency employed is f - 120 kHz, and the number of samples sampled at each measuring point is 132000. To ensure that complete signals can be detected, the frequency-cutting circuit in the instrumentation was turned off throughout all the measurements.
242
IMPINGING STREAMS
iI ,
/
t
-
-
17-\i ',2" 0
I/
tl7" \1
\ _
Y
Figure 11.3 Coordinate axes for the arrangement of measuring points.
11.2.4 Pre-treatment of the experimental data
11.2.4.1 Error point wiping During measurements, interferences are unavoidably generated by power source, equipment, etc., giving error points in the detected signals. To obtain the actual pressure fluctuation caused by impingement between the opposing streams, these points must be wiped. An abecedarian analysis showed that the probability density of the pressure fluctuation signals detected meets about the Gaussian distribution. So, these error points can be wiped off simply by the analysis of probability density, which was done for each run before data interpretation.
11.2.4.2 Blank experiments It was also found in the experiments that noise can be detected, even if the probes were put into stationary water. As a result, in every effective measurement, what is obtained is the sum of real signal and noise. To solve this problem, a blank measurement is made after each run, and the noises are deducted from the measured values. This cannot be done in the time domain but in the frequency domain, because the fluctuation has time-randomness.
11.3 EXPERIMENTAL RESULTS AND DISCUSSION 11.3.1 Intensive fluctuation region Pressure fluctuation has the effect of promoting chemical reaction(s). Therefore the spatial profile of the intensive fluctuation region inside the reactor is significant for
PRESSURE FLUCTUATION IN THE SCISR
243
understanding the performance of the reactor and for further application development. The pressure fluctuation is caused by the impingement between the opposing streams" the impingement occurs first round the impingement plane and the maximum velocity appears on the flow axis. Therefore it is reasonable to infer that the intensive fluctuation region should be concentrated around the impingement plane, especially around the flow axis, i.e., the most intensive fluctuation should appear at Point (0, 0, 0). However, the experimental results are completely unexpected. The profiles of PSD on the horizontal and the vertical planes measured at an impinging velocity of u{}- 0.1 m-s - I are shown in Figs. 1 1.4 and 11.5" and those on the impingement plane and its parallel planes in Fig. 1 1.6.
0.08 0.06 a., 0.04 0.02 -3
-2
-1
0
1
2
3
y, c m
Figure 11.4 Profile of fluctuation intensity on x-y plane at uo = 0.1 1 m/s.
0.2 %
a.,
0.1 0.0-3
--2
-1
0
1
2
3
~', c m
Figure 11.5 Profile of fluctuation intensity on x-z plane at u{}= 0.1 1 m/s.
244
IMPINGING STREAMS
0.064 2~ a.,
0.060
0.056
-4
-2
0
x, cm
(y-o)
0.04 0.03 0.02 0.01 -4
-2
0
2
4
x, cm (y = -2) Figure 11.6 Profiles of fluctuation intensity on planes parallel to x - z plane at u0 = 0.11 m ' s
-l
A combined consideration of the figures above reveals that the intensive fluctuation region is located neither around the impingement plane nor on the flow axis. It extends from the outlet of the drawing tube towards the impingement plane, with increasing diameter, and takes the form of a couple of truncated cones with an empty core, as shown in Figl 11.7; the distribution of the most intensive fluctuation points along the y-axis is given in Fig. 11.8. As can be expected, the intensive fluctuation region is symmetrical with respect to the impingement plane, provided the two streams leaving the drawing tubes are at the same velocity. This is simply because the two sides of the reactor have the same structure. In addition, the figures also show that the region is essentially symmetrical about the flow axis, suggesting that the above assumption is reasonable.
PRESSURE FLUCTUATION IN THF SCISR
245
Drawingtubei,
,,
Lwingtube
"
2/
Figure 11.7 Profile of intensive fluctuation region in the space between drawing tubes.
4
/• j
\
i
3
...... f ' • J
-3
-2
-1
0
1
2
3
y, cm Figure 11.8 Profile of strongest fluctuation points along y-axis at u0 = 0.10 m-s-~. Multiply repeated experiments have been carried out in order to verify the spatial profile of the intensive fluctuation region described above; all the results obtained are the same as those given in Fig. 11.7 and are essentially independent of the impinging velocity utt. The reason for the unexpected phenomena is still as yet unclear. It is almost certain that such a profile of the intensive fluctuation region is related to the aggregation status of the liquid, the special flow configuration in the SCISR, and, possibly, also to eddy generation in the flow. The problems seem complex and further investigations are necessary.
11.3.2 Volumetric distribution of fluctuation intensity Chemical reactions always proceed in the space of the reactor volume. Therefore, knowledge of the volumetric distribution of pressure fluctuation intensity is helpful in understanding the reactor's performance. Because of the axial symmetry of the
246
IMPINGING STREAMS
fluctuation, it can be considered that all points in the same circle with a radius of r on a plane vertical to the y-axis have the same intensity, for which the value is denoted by PsD(Y, r) and is averaged over all these points. From the measured data, the volumetric distribution of the intensity at different u0 can be calculated by integration; the results are shown in Fig. 11.9 as the plots offv versus PsD, where fv is the volumetric fraction of the region with intensity less than PsD(Y, r), f v - V p s D / V t , Vt is the total volume between the outlets of the two drawing tubes, and VpsD is the volume with intensity lower than PsD(Y, r) in the same space. For the SCISR used, V t - 771 cm 3, which is about 25% of the total effective volume of the reactor. As can be seen from Fig. 11.9, for u0 < 0.1 m.s -~, no significantly intensive region is observed. The volumetric fraction of the intensive region increases as u0 increases, while the volume with an intensity between 0.1 and 0.2 kPa occupies only about 10% in the range of u0 < 0.2 m-s -~.
1.0
.20
0.8 0.6 0.4 0.2 0.0 0.00
0.05
O.10
PF, kPa
O.15
0.20
Figure 11.9 Volumetric fraction of fluctuation intensity lower than PsD at different impinging velocities.
11.3.3 Definition of the impingement zone The impingement zone is the major active region in the SCISR, and so is of interest to define the region. However, it is somewhat difficult to define directly because it is without physical boundary. Wu et al. [31] made a definition by observing flow appearance and taking the region with turbulent waves at the surface as the impinging zone, which is centered and surrounded by two planes vertical to the flow axis and the inside wall of the reactor. The distance between the mentioned planes is about 1/2 that between the outlets of the drawing tubes, i.e., the impinging distance S. The volume of impingement zone thus defined is about 0.186 of the total of the reactor. On the other hand, the experiments above yield a profile of the intensive fluctuation region as shown in Fig. 11.7. It is considered that the fluctuation results from the impingement between the two opposing streams and that the fluctuation intensity
PRESSURE FLUCTUATION IN THE SCISR
247
characterizes the interaction between the streams to a considerable extent. Therefore, it may be more reasonable to define the impingement zone according to Fig. 11.7. From the data, the region almost around the flow axis inside the truncated cones is not an intensive region. However, from the flow configuration, the interaction of the opposing streams inside this region, including mixing, pressing, and shearing etc., should be significantly strong. Therefore, the impingement zone should be the whole of the region surrounded by the external surface of the truncated cones. The volumetric fraction so calculated is about 0.18 to 0.2, which is approximately equal to that reported in Ref. [311.
11.3.4 Influence of the impinging velocity on fluctuation intensity The influences of impinging velocity are studied by examining two parameters: the intensity at the most intensive point and the integral intensity of the intensive region.
11.3.4.1 Influence on intensity at the most intensive point By studying the multipoint measurements, it is found that Points (4, 1, 0) and (3, 2, 0) and their symmetrical ones exhibit the most intensive fluctuation. Figure 11.10 shows the effect of u{} on the intensity at Point (-3, -2, 0). As can be seen, the fluctuation intensity at that point increases linearly as u0 increases. The reason for this is clear: higher u0 simply implies greater dynamic energy than can be provided by the streams. Under the condition of u0 = 0.202 m.s -~, the fluctuation intensity PSD at the most intensive point can be as high as {}.278 kPa. Furthermore, according to the Gaussian distribution, the maximum should be about six times the standard deviation, implying that the maximum amplitude of fluctuation should be about 1.67 kPa, while a value practically measured for the maximum amplitude is 1.58 kPa, so the maximum amplitude of the pressure fluctuation in SCISR may be figured to be 1.6 kPa. 0.30 0.25 va
n.
....P
0.20
a., o.15 0.10
// IVj
0.05 0.00 0.00
•
m//
0.05
/J
/
£
/,
/
/
/. j/
//
•
J
•
O.10 u{},m-s
O.15
0.20
1
Figure 11.10 Influence of u{}on fluctuation intensity at Point (-3,-2, 0).
248
IMPINGING STREAMS
11.3.4.2 Influence on the integral intensity As can be seen from the results given above, the intensive fluctuation region is concentrated mainly in the cylinder between the outlets of the drawing tubes. For convenience, let us define the integral intensity of the intensive region, denoted by Ip, as
Ip - ~-3 ~PsD (y, r)2zcrdrdy
(11.8)
where PsD(Y, r) represents the averaged fluctuation intensity of all the points in the circle with a radius of r on the plane vertical to the y-axis. The integral intensity defined by Eq. (11.8), Ip, does not have the unit of kPa and so has a different meaning from the fluctuation intensity. However, its unit kPa.m 3 is the same as that of work or energy, and may reflect the total energy of the fluctuation so that it may be related to the mechanical energy consumption. Hence, it is interesting to examine the influence of the impinging velocity on this quantity. Substituting the values calculated at various u0 for PsD into Eq. (11.8) and integrating the resulting expression yields the relationship between Ip and u0, as illustrated in Fig. 11.11. Again, the integral intensity increases linearly with the increase in impinging velocity u0. This fact reflects clearly the qualitative relationship of energy conversion in the pressure fluctuation caused by the impingement between the opposing streams: the larger the impinging velocity, the larger is the kinetic energy of the streams taking part in impingement; as a result, the larger is the integral intensity Ip, i.e., the larger is the total energy of fluctuation.
1.4
1.2 1.0 ~.-
0.8
x
0.6
0
0.4 0.2 0.0 0.00
me
0.05
0.10
uo, m.s
-1
0.15
0.20
Figure 11.11 Influence of u0 on the integral intensity of intensive fluctuation region.
PRESSURE FLUCTUATION IN THE SCISR
249
11.3.5 Power spectrum analysis for pressure fluctuation For the frequency characteristics of the fluctuation, a power spectrum analysis is made. Figure 11.12 shows the power spectrum at several points with u o - 0.11 m.s -~" while Fig. 11.13 gives those at Point ( - 4 , - 1 , 0) at various impinging velocities. It is obvious from the figures that the power density is higher in the low-frequency range. It decreases as frequency increases, while it is kept essentially constant after a certain critical frequency. Although the critical frequency varies with operating impinging velocity (see Fig. 11.13), all the values are iess than 1000 Hz. As can be seen from the variation tendency, a power function of u~ approximately holds in the range ~,f frequency lower than the critical value, while over this range the power density is essentially independent of frequency. The major energy of the fluctuation is concentrated in the r a n g e ,)f lower frequency, but is less in the high frequency' , a n g e . There are some peaks in the aco~tstic wave range, but all are lower. It can also be seen that the forms of all the power spectrums measured :,t various impinging velocities and at various points are similar, implying that all the peaks at whatever frequency i~crease simultaneously as u0 increases.
O.Ol 0.1
I
1E-3 1E-4 __
0.1 0 .0 1 iE 3
1E-4
(o.o,o)
rO
t, rO ~
0.1
lrO
! 000
,
,. ..,
,..
0,1 0.01 1E-3 IE-4 0.01 1E-3 1E-4
500
{,-
500
i
,
-
i
! 500
2000
,
,
1500
!O00
i500
'
,
3000
,
,
2000
i
500
,
,
1000
2500 ,
~
2500
i
,
,
3500
i
2500
4000
, P,
,
3000
i
2000
3500
,
3000
,
4000 .- 2 , ( ! )
I
,
3500
i
4000
I
)
. 0
.
. 1000
500
.
.
. 2000
1500
.
.
.
2500
3000
3500
4000
f, Hz
Figure 11.12 Power spectrum of fluctuation at different points (uo= 0.11 re.s-l).
011
0 01
[ .i
~
=
0.1
~
1E-30"01
-30
I
Uo=0.06m-s ! '
i 500
J
'
I 1000
'
I 1500
,
i 2000
'
i 2500
I~ i
,
i
.
500 •
'
I
.
i
,
] 0001500
!
'
2000
i
i
500
'
~950{7"}
i 3500
,
! 4000
i
'
3000
i
'
~500
i
~(.)00
uo=O. 17 m. s -t
~.,.
' ![l'rr11~ !
'
-I
0.01
0
i 3000
uo=O.10m.,s
0
! E 3
'
i
1000
'
I
1500
'
I
2000
'
I
2500
'
I
3000
'
i
3500
,
I
4000
f, Hz Figure 11.13 Power spectr',,: <:~"fl~cw:at{cn et ~Mr.~ ,'z --1 O) ar~,:!Jiffereqt m:
250
IMPINGING STREAMS
It can be concluded from the figures that the pressure fluctuation due to the impingement between the opposing streams in the SCISR has the charactistics of wide spectra and multi-frequency, and the fluctuation energy is concentrated mainly in a range of frequency no greater than 1000 Hz, with weaker fluctuation in the acoustic wave range. From the results described in Sections 11.3.4 and 11.3.5 it follows that the impinging velocity, u0, has a fundamental effect on both the fluctuation intensity and the main frequency range. Therefore increasing u0 should be an effective measure for enhancing pressure fluctuation. Unfortunately, the operating u0 is limited by the design of the existing SCISR. Too high a velocity will lead to very strong turbulent waves at the liquid surface and unstable operation. The results above indicate the direction for further improvement of the reactor design. The pressure fluctuation in liquid-continuous impinging streams must affect the micromixing status and, consequently, it favors reaction kinetics. The substances related to this topic will be discussed in the following chapters.
11.4 CONCLUSIONS AND DISCUSSION The pressure fluctuation in the submerged circulative impinging stream reactor (SCISR) during operation is investigated experimentally by online measurement of the instantaneous pressure signals using micro pressure probes that respond well high frequency, and with water as the process liquid. The following can be concluded: (1) There is a considerably strong pressure fluctuation in the SCISR during operation, of which the maximum standard deviation of instantaneous pressure can achieve 200-300 Pa, implying that the maximum amplitude can be as high as about 1.6 kPa. (2) The intensive fluctuation region is concentrated between the outlets of the drawing tubes and around the flow axis, and takes the form of a couple of truncated cones with an empty core and with coinciding bottoms. The region is symmetrical with respect to both the impingement plane and the axis of the streams and is essentially independent of u0. The space surrounded by the external surface of the cones is defined as the impingement zone. (3) Using multipoint measurement, the points of the most intensive fluctuation are ascertained, and the spatial integral intensities of fluctuation between the outlets of the two drawing tubes are calculated. The results of examining the influences of the impinging velocity, u0, on these two amounts indicate that both increase linearly with u0 increasing. (4) The results of power spectrum analysis show that the major fluctuation is concentrated in the range of frequencies < 1000 Hz; while at higher frequency range the fluctuation has lower energy and there are some lower peaks in the acoustic wave range. (5) The impinging velocity, u0, exhibits significant effects on both the intensity of fluctuation and its major frequency range. Both increase as u0 increases.
PRESSURE FLUCTUATION IN THE SCISR
251
The major contribution of this investigation is that the existence of the very strong pressure fluctuation due to the impingement between the opposing streams is verified with conclusive experimental evidence; in addition, the basic regularities of the influence of the impinging velocity on the fluctuation are determined, i.e., both the intensity of the fluctuation and its main frequencies increase as the impinging velocity increases. However, the results obtained so far can only indicate the relationship of energy conversion qualitatively, but cannot describe the relationship quantitatively. In addition, because of the limitation of the reactor design, the range of the impinging velocities tested was not wide enough so that the data at higher impinging velocity is lacking, while such data are necessary for understanding the complete regularity. The vertical circulative impinging stream reactor (VCISR), especially the VCISR of type II to be discussed later in this book, can be operated at higher impinging velocity and thus may be applicable for investigations on this topic. The pressure fluctuation must affect the condition of the micromixing in the device and thus promote process kinetics. The results to be introduced in the next chapter will provide the experimental evidence for this topic. Unfortunately, a quantitative description for such influences cannot as yet be made because of the complexity of the problems involved, and further investigations are certainly needed in order to make the relationships clear, particularly the influence of pressure fluctuation on micromixing. Therefore the pressure fluctuation in liquid-continuous impinging streams remains a problem needing further in-depth investigation.
This Page Intentionally Left Blank
-12INFLUENCE OF LIQUID-CONTINUOUS IMPINGING STREAMS ON PROCESS KINETICS
12.1 QUALITATIVE ANALYSIS FOR THE INFLUENCES OF PRESSURE FLUCTUATION AND MICROMIXING As explained in Chapters 10 and 11, the existence of pressure fluctuation and the promotion of micromixing are the major features of liquid-continuous impinging streams. For processes occurring on the molecular scale in liquid or multiphase systems with a liquid as the continuous phase, these features have important implications. One of their valuable applications is the promotion of process kinetics. In investigations related to the kinetics of processes in liquid phase or multiphase systems with a liquid as the continuous phase, the assumption generally made to date is that of the complete mixing in liquid phase. Although the influence of incomplete mixing on the reactions in homogeneous systems had been presented by Danckwerts [157] as early as 1958, and Li and Chen et al. [174-178] also studied this problem, in almost all the investigations aimed at the establishment of kinetic model(s) the traditional stirred tank reactor (STR) and the complete mixing assumption were used. However, as described in Chapter 10, the micromixing condition in an STR is never ideal. According to the molecular collision theory of kinetics, only collisions between molecules with high enough energy can result in reaction(s). In other words, a chemical reaction occurs only when the two conditions are met: (1) the molecules to take part in the reaction collide with each other: and (2) the colliding molecules have a sufficiently high energy level. The first condition requires good micromixing conditions, i.e., good mixing on the molecular scale to enable molecules to make contact with each other. Gases easily achieve a condition of completely unilbrm mixing as the molecules can easily contact each other because they have a very large molecular free path. However liquids do not achieve such uniform mixing so easily because they are in a condensed condition and so the molecules in stationary liquid can only vibrate and/or rotate around their balanced position with extremely small displacement; to a large extent, their mixing status depends on the flow configuration inside the processing device.
253
254
IMPINGING STREAMS
As mentioned in Chapter 10, it is necessary to make distinction between macro- and micro-mixing according to the scales the mixing occurs on. For processes carried out on the molecular scale, the condition of micromixing, i.e. mixing on the molecular scale, has a fundamental influence on the efficiency. "Poor" micromixing is characterized by the existence of the segregation scale greater than Kolmogoroff micro length 2. Since fluid elements tend to follow the streamlines, it is difficult for molecules in segregated different elements to contact each other. In this case, the assumption of completely uniform mixing must lead to deviation from reality. In other words, poor micromixing restricts the probability of collision between the molecules taking part in the desired reaction, while, on the other hand, effectively promoted micromixing favors the increase in the probability of collision between molecules. The liquid-continuous impinging stream device has been proved to have the feature of strong micromixing so that, most likely, it can create such a positive effect. Secondly, it was described in the previous chapter that is pressure fluctuation in the operation of a liquid-continuous impinging stream device. The existence of pressure fluctuation implies the following:
(1) A considerable part of the fluid elements and molecules change their modes of motion. In microscope, pressure fluctuation resulted from vibrations of fluid elements and/or molecules. The changes of the modes of motion may promote micromixing, because micromixing is closely related to the motion of fluid elements and/or molecules; (2) The changes of modes of motion also indicate the occurrence of energy conversion, e.g., part of the dynamic energy of the flow is converted to vibration energy; and (3) During the conversion of the energy form, most possibly the energy distribution over the molecules changes so that part of the molecules may get more energy to achieve a higher level and thus more molecules can induce a reaction when they contact the other(s) by collision. The reasonable inference is that strong micromixing and pressure fluctuation promote process kinetics by the mechanisms of increasing both the probability of collision and the effective collisions. Or, more generally, process kinetics not only depends on the nature of the substance system involved and the operating conditions but also relates to the flow configuration in the processing device employed.
12.2 CRYSTAL-GROWTH KINETICS OF DI-SODIUM PHOSPHATE 12.2.1 Basic principles 12.2. 1.1 General principles Crystallization from solution can occur only in supersaturated solution. Under lower supersaturation nucleation cannot occur spontaneously; spontaneous nucleation
INFLUENCE OF LIQUID-CONTINUOUS IS ON PROCESS KINETICS
255
happens only when the supersaturation achieves a certain level. The concentration at which spontaneous nucleation begins is known as the super solubility of the crystalline substance under consideration. The region between the curves of solubility and super solubility is called the metastable region of the solution. More than 70 years ago, Ding et al. [179] found that in practice there are two super solubility curves: the nucleationstarting curve and the quantity nucleation curve. It is clear that the experiments for measuring the crystal-growth rate must be carried out in the metastable region and under the nucleation-starting curve to avoid errors resulting from nucleation. It is generally considered that crystal-growth includes two steps: (1) diffusion of the solute from the solution bulk towards the surface of the crystals, and (2) crystallizing reaction on the surface, the kinetics of which are represented, respectively, by Diffusion o f solute" dm ---
dt
DA
d
(c-c,)
(12.1)
Crystallizing reaction on the surface" dm __ = k,A(C, -Co) dt
(12.2)
The overall rate of the process is expressed by dm
--=
dt
K A ( C - C o)
(12.3)
where the overall crystal-growth rate coefficient, K, is determined by K-
1
1/k,+a/D
(12.4)
In practice, the kinetics of crystallization is even more complex; some different relationships were introduced in Ref. [180]. For the system to be investigated, since disodium phosphate has a very good crystallization nature, few complicated factors would be involved, and Eq. (12-4) should be valid. On the other hand, as an implication, the equation for the diffusion rate based on Fick's law includes the assumption of the solute concentration in the liquid bulk being completely uniform, which is actually difficult to realize and thus may yield a deviation from reality. The poorer the micromixing, the larger would be the deviation. Therefore the crystal-growth rate coefficients measured in different devices with different micromixing conditions may be different from each other.
256
IMPINGING STREAMS
12.2. 1.2 Interpretation of experimental data In order to avoid other complicated factors concealing the major problems being examined, the overall crystal-growth rate equation, Eq. (12.3), is assumed to be valid for the system under consideration and is used as the basis for the interpretation of data. The experiments are carried out as follows: add a known amount of crystal seeds narrowly screened into the solution with a supersaturation rigorously controlled to grow for a certain time interval; weigh the amount of the grown crystals; and then determine the overall crystal-growth rate coefficient according to the difference between the masses of the crystals before and after growth. The crystals of NazHPO4 are of non-spherical form and so the volumetric and surface shape-coefficients, (Pv and ~,, should be introduced for correction. The total mass of the crystals, m, and their total surface area, A, are calculated by 3
m - N(Pvd~P p ,
A - N(flsd p
(12.5)
The total number of crystals tested, N, is calculated from the initial mass, m0, and the initial mean diameter, dp0, as N --
m°
3 (,Ovdp0Pp
(12.6)
where dp0 is the Sauter mean diameter, i.e., the volumetric-surface mean diameter, of the initial crystalline particles, and is calculated as the arithmetic mean value of the hole sizes of the meshes on the upper and lower screens. At an instant, the relationship between the total mass and the mean diameter of the crystals is determined by dp - [
m ]1/3 N (PvPp
(12.7)
Substituting Eqs. (12.6) and (12.7) into Eq. (12.5) leads to
A -
('°s
mZ/3m 2/3
(12.8)
~vdpoPp
Further, substituting Eq. (12-8) into Eq. (12-3), integrating the resulting equation between t = 0 and t = tr and rearranging yields the expression for the overall crystalgrowth rate coefficient as K =
3(m~/3 - mU3)(,OvPpdpo
(,OsmU3ACmtf
(12.9)
where the average supersaturation, ACre, is determined according to the concentrations of the solute in the solutions at the inlet and the outlet of the crystallizer. In fact, the crystallization process is operated such that the difference between the two
INFLUENCE OF LIQUID-CONTINUOUS IS ON PROCESS KINETICS
257
concentrations is very small so that the average value is accurate enough. For crystalline di-sodium phosphate, the statistically determined values for the volumetric and the surface shape coefficients are (,0,- 1.3 and ~ , - 7.3, respectively.
12.2.2 Experimental investigation 12.2.2.1 Testing materials In order to obtain results of more practically applicable value, in the investigation by Wu et al. [181] the commercial Na2HPO4.2H20 crystalline and the corresponding mother liquor are taken directly from a plant producing inorganic salts as the testing materials, the chemical compositions of which are illustrated in Table 12.1. Table 12.1
Compositions of the samples used in the experiments Composition, %mass Phase pH
Na2HPO4
Cl-
SO 4-~
F-
NaOH
Na~CO~
Crystals
8.51
43.02
1.13
0.29
0.0006
0.638
0.638
Mother liquor
6.36
33.96
2.58
0.05
0.0011
1.0
1.0
12.2.2.2 The metastable region of di-sodium phosphate solution To ensure that all the overall crystal-growth rate coefficients are measured under the conditions without nucleation, the metastable region of the solution has to be determined first and therefore the solubility and super solubility need to be measured. For the measurement of solubility the traditional dissolution-equilibrium procedure is employed, i.e., put both the crystals and the mother liquor into a container submerged in a water bath continuously stirred, control the temperature inside the container rigorously at a given value, with the fluctuation no greater than _+0.I°C. When the dissolution equilibrium is achieved, measure the concentration of the solute in the liquid phase as the solubility at the temperature given. The equilibrium condition is judged by the criterion that the relative deviation of the values obtained in, at least, three times of adjacent measurements is not greater than 1%o. The method used for the measurement of the metastable region is also the traditional one, i.e., using the light scattered by the newly formed fine crystals to detect the nucleation status. The experimental equipment is shown in Fig. 12.1. The water bath tank is surrounded with a light-proof black screen, with two windows vertical to each other with dimension of 2×3 cm for light beam casting and visual observation, respectively. To parallel the industrial condition of induced nucleation, all the measurements are carried out under conditions with the existence crystal seeds.
258
IMPINGING STREAMS Cold water
,le,
O
earn
)
Water bath
Figure 12.1 Equipment for the measurement of super solubility. A beam of parallel light from the focus lamp passes through the glass crystallizer filled with the NazHPO4 solution of known concentration. Under the condition of continuously stirring, the cold water is inputted into the water bath to lower its temperature and, further, to cool the solution inside the crystallizer at the cooling rate of about-0.1 °C.min-~; the temperature of the solution is measured with an accurate thermometer with a scale of 0.1°C. When the temperature of the solution achieves the dissolution-equilibrium one determined by the solubility curve, add the crystal seeds of known amount to the solution, continue to cool it, observe the solution through the window with care, and record the temperatures of nucleation starting and in quantity, respectively. The amounts of crystal seeds added in each run are 0.15-0.3 g, with sizes of 0.42- 1.60 mm. The results of a checking calculation show that the relative variation of the solution concentration due to the growth of the crystal seeds is only about 1%, and so the experimental accuracy can be ensured. Under each set of conditions the experimental operation is repeated at least six times. The results experimentally measured for the metastable region of the commercial Na2HPO4 solution are given in Fig. 12.2, where the data points for nucleation starting and in quantity are the average values of the six measurements. The reproducibility of the observed data is very good and the maximum relative-error, i.e., the ratio of the difference between the measured value with the maximum error and the average to the width of the metastable region, is no greater than 3%. As can be seen in the figure, with the substance studied there are two super solubility curves, i.e., curves of nucleation starting (Curve 2) and nucleation in quantity (Curve 3). The regularity of the variation of the super solubility with temperature is very good so it can be considered that the metastable region determined by the three sets of data shown in Fig. 12.2 is accurate enough. According to the data shown in Fig. 12.2, the experiments for measuring the overall crystal-growth rate coefficient should be operated between Curves 1 and 2.
INFLUENCE OF LIQUID-CONTINUOUS IS ON PROCESS KINETICS
259
50
........................................................................................................................................................................................................................... i 3
40
Q ZZ e-i
Z
d
2
i
30
20
10
20
25
3O
35
40
T ,"C
Figure 12.2 Metastable region of commercial Na2HPO4 solution, o-solubility;/X-nucleation starting D-nucleation in quantity.
12.2.2.3 Crystal-growth rate of di-sodium phosphate For comparison, the experiments for measuring the overall crystal-growth rate coefficient are carried out in an impinging stream crystallizer (ISC) and a fluidized bed crystallizer (FBC). The experimental system scheme is indicated in Fig. 12.3, where the impinging stream crystallizer (ISC) is of the same structure and dimensions as that of the submerged circulative impinging stream reactor shown in Fig. 10.2 but the drawing shown in Fig. 12.3 is the side view. Also, the ISC is without the top cover. The Na2HPO4 solution of known concentration enters the upper tank 1 and is cooled to a certain temperature by indirect heat exchanging with the cooling water passing through the coiler; then flows continuously into ISC 2 through the inlet tube, the exit of which is at the inlet of the drawing tube: the excess solution overflows from the ISC through the overflow port on the side-wall of the ISC. The temperature of the solution inside the ISC is controlled by heat exchanging with the water from the thermostat to achieve the given supersaturation. When both the flow and the temperature achieve their stable states, the narrow- screened Na2HPO4 crystals of the amount m0 are added to the ISC and grow in the solution inside it for a certain interval of time, and then the whole of the suspension is removed from the ISC and filtered immediately while keeping warm. The crystals are washed with alcohol and dried at low temperature, and then weighed to obtain the final amount of the crystals, mr.
260
IMPINGING STREAMS
IT
Cooling water
1
Recycling water :Overflow
% - I . ~ ~ 1
I
_ .,.,,>'~
/
2
Figure 12.3 Scheme of experimental system with impinging stream crystallizer. 1-cooling tank; 2-ISC; 3-thermometer; 4-thermostat.
The temperature of the solution inside the ISC is measured with an accurate thermometer with a scale of 0.1°C so that the accuracy can be _+0.05 °C. The solution out of the ISC is sampled at the outlet of the overflow port and the sample is analyzed chemically for the concentration of Na2HPO4. The average of the concentrations of Na2HPO4 in the inlet and the outlet solutions is taken as the mean concentration of the solution inside the ISC; and its corresponding super- saturation is then determined as the difference between the concentration and the solubility at the operating temperature, ACre, while the latter is obtained from Curve 1 in Fig. 12.2. Since the amount of the solute consumed in the growth of the crystals is only about 1% of the total amount in the solution passing through the ISC, implying a very small variation of the concentration, an arithmetic mean value of the concentrations in the inlet and the outlet solutions is accurate enough for calculation. The experiments are repeated 2-3 times for each set of conditions, and the values measured for the overall crystal-growth rate coefficient are averaged. The time interval for the growth in each run ranges from 1200 to 2400 s. To ensure that the supersaturation is stable, both the amount of the crystal seeds added and the flow rate of the solution passing through the ISC are very small in comparison with the content in the ISC and, consequently, the overflow rate is also very small, so that no crystals being carried by the overflow solution are observed, i.e., no loss of the crystals needs to be considered and the measurement accuracy for the amount of the crystals after growth can be ensured to meet the requirement.
INFLUENCE OF LIQUID-CONTINUOUS IS ON PROCESS KINETICS
261
The structure of a experimental fluidized bed crystallizer (FBC) is shown in Fig. 12.4, where the crystallizer is actually a universal equipment for the measurement of crystal-growth rate. The solution enters the FBC at its bottom, and leaves the FBC by overflow. All the other parts of the experimental system are the same as shown in Fig. 12.3, and so are not shown in Fig. 12.4. The operation procedure for the FBC is the same as for the ISC. For convenience of comparison, the corresponding conditions, temperature and concentration of the solution, operated in the ISC and the FBC are rigorously controlled to be the same, with the deviation of the operating temperature no greater than 0.05 °C.
Overflow
Na2HPO4 solution
Recycling water
Figure 12.4 Fluidized-bed crystallizer (FBC). The values measured in the ISC for the overall crystal-growth rate coefficient of Na2HPO4 are listed in Table 12.2. As can be seen, the reproducibility of the data is in a reasonable range and that of most data is very good. A similar table summarizing the experimental data measured in the FBC, with the same mean diameters of crystal seeds, dp~, and under the same operating temperatures, T, was also obtained; but it is not given here. A comparison between the overall crystal-growth rate coefficients obtained from different crystallizers is of interes!, The data measured in the ISC and FBC are listed in Table 12.3. It is obvious that the values measured in the impinging stream crystallizer for the overall crystal-growth rate coefficient, K~s, are greater systematically than those measured in the fluidized bed crystallizer, KFB, by about 15 to 20%. Nothing can account for such differences, except that the strong micromixing and the pressure fluctuation occurring in liquid-continuous impinging streams promote process kinetics. The impinging stream crystallizer used in the experiments is of a horizontal structure
262
IMPINGING STREAMS
and so its operating impinging velocity, u0, is limited, because higher u0 will lead to liquid splashing and thus unstable operation. Therefore it is not sure w h e t h e r the crystal-growth rate can be increased further by increasing the impinging velocity. Table 12.2
Overall crystal-growth rate coefficient of Na2HPO4 in SCISR
doox 104
mox 10 3
m
kg
T, °C
ACnl,kg.m -3
tf, s
kg
Measured
Average
2.51
0.7802 0.7803
32.7
3.98 4.67
1800 2100
1.4041 1.6382
6.79 6.44
6.61
2.51
0.8010 0.8008
35.9
4.98 5.39
2400 2400
2.4892 2.7594
8.65 8.88
8.77
2.51
0.5715 0.4509
38.6
6.21 7.87
1800 1800
2.1931 2.3024
11.39 11.46
11.43
3.48
0.5610 0.5614
33.9
7.58 13.4
1800 1800
1.3243 2.1316
7.58 7.25
7.42
3.84
0.7401 0.7403
36.5
7.37 9.24
1800 1800
2.1415 2.1972
9.99 8.20
9.10
3.84
0.8121 0.8118
38.7
7.73 8.39
1800 1800
3.5173 3.6787
14.13 13.53
13.83
5.22
0.5428 0.7217
34.2
9.56 10.25
1800 1800
1.2459 1.7409
8.68 8.65
8.67
Conditions
0.6924 5.22
1800
1.6895
13.28
8.12 11.36 13.04 9.31 11.18 8.37
1800 2400 1800 1800 2400 2400
2.1329 3.6147 4.2943 3.9875 5.3435 1.0877
14.63 12.60 16.29 18.98 16.86 8.54
37.0
7.21 5.76 6.64 8.49
2400 2400 1800 2100
1.1422 0.9868 1.3417 1.9875
10.74 13.47 16.12 17.49
38.8
7.92 8.52
1800 1500
2.2521 2.1972
20.49 22.24
8.74
2100
1.7322
13.78
7.59 5.92 7.12 9.45
1800 2100 1800 1800
1.2355 1.3687 1.9611 2.0349
10.29 13.90 17.88 16.16
7.67
1800
1.5851
25.61
8.29 9.78
1500 1200
1.7799 1.4219
27.16 23.80
37.2
33.6
7.36
0.5425 0.5419 0.6222 0.6227
7.36
0.7507 0.7500
7.36
39.6
0.7602 8.95
8.95
0.7611 0.7624 0.9229 0.8410
33.4
36.7
0.5301 8.95
0.6209 0.5810
Klsx 10 6, m.s -j
6.78
0.6918 0.6931 0.8001 0.8415 0.8409 0.5434
5.22
mfx 10 3
39.0
13.50
17.38
10.92
16.81 21.37 12.66
17.02 25.52
INFLUENCE OF LIQUID-CONTINUOUS IS ON PROCESS KINETICS
263
T a b l e 12.3
Comparison between overall crystal-growth rate coefficients measured in the ISC and the FBC, respectively
dpoX 104, m
T, °C
3.48
5.22
7.36
8.95
K[s/KvB
Kls
KI:B
6.61
5.16
1.281
35.9
8.77
7.54
1 163
38.6
11.43
9.53
1 199
33.9
7.42
6.30
1 178
32.7 2.51
Averaged Kx I 0 f', m.s -]
36.5
9.10
7.86
1 158
38.7
13.83
11.58
1 194
34.2
8.67
7.30
1 188
37.2
13.50
10.50
1.286
39.6
17.38
14.70
1.182
33.6
10.92
8.25
1.169
37.0
16.81
14.35
1.171
38.8
21.37
19.15
1.116
33.4
12.66
9.69
1.307
36.7
17.02
14.54
1.|71
39.0
25.52
22.23
1.148
It is also interesting to examine the influence of temperature on the crystal-growth rate. For this purpose the generalized Arrhenius relationship below is used: K - K ° e x p ( - E / RT) The mean observed active energies obtained by regression of the experimental data for various crystal seeds with different mean diameters and in different crystallizers are listed in Table 12.4, where, similarly, the subscripts IS and FB denote the parameters in the impinging stream crystallizer and the fluidized bed crystallizer, respectively. T a b l e 12.4
Observed active energies for Na2tlPO4 crystal-growth in the ISC and the FBC, respectively, obtained by fitting data dptl× 104, m
Eis, kJ'mol -]
EFB, kJ'mol -j
2.51
73.87
81.38
3.48
106.02
103.21
5.22
103.31
104.27
7.36
103.44
128.86
8.95
104.01
101.06
264
IMPINGING STREAMS
The following can be observed from the table:
(1) In the small range of the crystal sizes, the observed active energy for NaePO4 crystal-growth is related to the crystal size. As can be seen the values for the active energy with the initial mean size of crystals of 0.251 mm in both the ISC and the FBC are obviously smaller, while with larger crystals the active energy remains essentially constant, about 104 kJ-mol -j, independent of the crystal size. Wu [180] observed similar phenomena in an investigation on the crystal-growth rate of Na3PO4. As has been mentioned, crystal growth involves two steps: the diffusion of solute molecules and the crystallizing reaction at the surface. So the overall process may be affected by both diffusion and reaction kinetics. The data listed in Table 12.4 indicate that with crystals initially sized 0.251 mm the influence of the diffusion is more significant, while with those initially sized 0.348 mm or larger the process is mainly affected by reaction kinetics. Since the difference between the densities of the liquid and the crystals is relatively small and the viscosity of the mother liquor is very large, the fine crystals tend to follow the streamlines, yielding small relative velocity between the crystals and the mother liquor and thus reduced liquid-film transfer coefficient. Globally, this shows an increased influence of the diffusion of solute molecules through the liquid-film with smaller sized crystals. Therefore the phenomena described above are reasonable. (2) Except for a few questionable data, the values for the observed active energy measured in the two crystallizers of different types, E~s and EFB, show little difference and can be considered to be more or less identical. On the other hand, the values measured in the impinging stream crystallizer for the overall crystalgrowth rate coefficient,/(is, are obviously and systematically larger than those in the fluidized bed crystallizer, KFB. Therefore it can be affirmed without the need for further analysis that, with the observed frequency factors, there must 0
0
be K~s > K w . From the point of view of the molecular collision theory, this suggests that more effective collisions occur in the ISC. Therefore the results given in Table 12.4 support the inference made in Section 12.1, i.e., the strong pressure fluctuation in liquid-continuous impinging streams causes a considerable part of the molecules that obtain more vibration energy converted from the flow dynamic energy to achieve higher level. In addition, the good micromixing in the LIS creates favorable conditions for the increase in collision probability. Thus, the two factors result in a larger observed frequency factor for the crystallization 0 system, K~s. The most important result obtained from the above investigation is the experimental evidence that the crystal-growth rate in the impinging stream crystallizer is higher than that in the fluidized bed crystallizer. More generally, the results certify that the crystalgrowth rate not only depends on the nature of the substance system involved and the operating conditions but is also related to the flow configuration in the device used, and thus indicates a new possible way for the enhancement of crystallization.
INFLUENCE OF LIQUID-CONTINUOUS IS ON PROCESS KINETICS
265
12.3 KINETICS OF ETHYL ACETATE SAPONIFICATION 12.3.1 Chemical reaction and experimental method For further verification that the LIS promotes process kinetics, Wu e t al. [181] studied comparatively the kinetics of ethyl acetate saponification in the submerged circulative impinging stream reactor (SCISR) and a stirred tank reactor (STR), respectively. The chemical reaction is represented by CH3COOC2Hs(A) + NaOH(B)
-
CH3OONa(R) + C2HsOH(S)
(12.10)
Chemically, the system has a very stable kinetic nature and is without by-reaction and so the kinetics data are easy to obtain. Therefore it is very suitable for the goal of the present investigation. The experiments are carried out in the SCISR with an effective volume of 3.6xl 0 -3 m 3, the structure of which is the same as that shown in Fig. 10.2, and a traditional stirred tank reactor (STR) of 0.6x10 -~ In 3 in the effective volume with dampers" both the reactors are operated in batch mode and the temperature of the reaction mixture is rigorously controlled with an accuracy of +0.05°C during the reaction in the same way as for the investigation on the crystal-growth kinetics described in the previous section. The impinging velocity in the SCISR operation is controlled at about 0.26 m-s -j, corresponding to the propeller rotary speed of 900-1000 rpm, while the STR is operated under the condition of fully agitating, which is determined by visual observation, and the rotary speed of the paddle is about 1200 rpm. As the reaction proceeds, the NaOH in the reaction mixture is consumed gradually so that the electroconductivity of the mixture drops continuously. Because the electro-conductivity of the mixture is a monodrome/'unction of the concentration of NaOH, CB, the variation of CB during the reaction can be determined continuously by measuring the electroconductivity of the mixture with an online probe, while that of the ethyl acetate is easily determined as a function of time by a molecular mass balance. The data are correlated with the well known kinetics equation below rj\ - k C A C B
(12.1 1)
12.3.2 Major results The results of correlating the experimental data measured in the range of 25-45 °C are given in Table 12.5, where the subscripts IS and ST denote the parameters determined in the SCISR and STR, respectively. The results reflected by the data in Table 12.5 are similar to those on the crystalgrowth kinetics of Na2HPO4, i.e., the values for the reaction rate constant measured in the SCISR, k~s, are systematically higher than those measured in the STR, ksT, by about 20%.
266
IMPINGING STREAMS Table 12.5
Comparative data for the rate constants in the reaction kinetics equation (12.11)
T,°C
Rate constant, m3.kmol-J.s-J
Rate constant ratio
kis
ksT
kls / ksv
25.0
0.175
0.137
1.28
35.0
0.239
0.196
1.22
45.0
0.758
0.651
1.16
The interpretation the experimental data obtained in the SCISR and STR, respectively, with the Arrhenius relationship yield the measured values for the active energies in the two reactors being E~s = 57.5 kJ-mol -~ and Esv = 60.1 kJ-mol -~. The difference between the two values is small, and is actually within the scope of experimental error so that they can be considered to be identical. These results lead to the conclusion similar to that obtained in the investigation on the crystallization kinetics of Na2HPO4, i.e., with the values for the frequency factor of the reaction represented by Eq. (12-10) measured in the two reactors, there must be k0,~s> k0,sv.
12.4 CONCLUDING REMARKS Comparative investigations were made for the crystal-growth kinetics of Na2HPO4 in the impinging stream crystallizer (ISC) and the fluidized bed crystallizer (FBC) and the kinetics of ethyl acetate saponification in the submerged circulative impinging stream reactor (SCISR) and the stirred tank reactor (STR) and the following was concluded: (1) In the ranges of the operating conditions tested, the overall crystal- growth rate coefficient of Na2HPO4 measured in ISC, K~s, is higher systematically than that measured in FBC, KVB, by 15 to 20%, while the reaction rate constant of ethyl acetate saponification measured in the SCISR, k~s, is larger systematically than that measured in the STR, ksf, by about 20%. (2) The results of correlating the data on the overall crystal-growth rate coefficient of NazHPO4 with the generalized Arrhenius relationship indicate that the values for the observed active energy obtained in both the ISC and the FBC are essentially identical. With the crystal seeds sized 0.348 mm or larger, the observed active energy is determined to be about 104 kJ.mol -~, while with the smallest crystal seeds tested, average-size 0.251 mm, the observed active energy is obviously smaller than with the larger crystal seeds. (3) The regression of the data on the kinetics of ethyl acetate saponification in the range of reaction temperature from 25 to 45 °C yields the active energy in the
INFLUENCE OF LIQUID-CONTINUOUS IS ON PROCESS KINETICS
267
SCISR E~s = 57.5 kJ-mol -~ and that in the STR EST = 60.1 kJ-mol -~. The difference between the two values is small and is within the scope of experimental error so they can be considered as identical. (4) The results of the two kinetics investigations support the theoretical inference that the efficient micromixing and the considerably strong pressure fluctuation in liquid-continuous impinging streams promote process kinetics. Obviously, the results described above are of significance. They have a certain theoretical accordance and indicate reasonable directions for the further development of LIS application. As described above, the major reasons of LIS promoting process kinetics may be that the effective micromixing increases the collision probability between molecules and that the considerably strong pressure fluctuation changes the energy distribution over the molecules, causing the part of the molecules that obtains more energy to achieve the higher level required for inducing reaction. However, the theoretical analysis above is only an inference and, limited by the current technological conditions, no direct experimental evidence can as yet be provided. Therefore, it seems that the results reported in this chapter also raise a number of new problems that require further investigation, two of which are:
(1) Influences of micromixing and pressure .fluctuation: What are the mechanisms of their effects? How can direct experimental evidence be obtained for these effects? How can these influences be quantitatively described? How can their individual contributions to the global influence be determined? (2) Real kinetics dam: To date, almost all the kinetics data on reaction systems in liquid phase or multiphase with liquid as the continuous phase have been measured in traditional stirred tank reactors. From the results reported in this chapter, it is likely that significant deviations exist in the existing kinetics data. On the other hand, the LIS device cannot yet be considered as absolutely ideal for kinetics investigation, not least because its micromixing time, tM, is not zero. What then is the ideal equipment and conditions for obtaining real kinetics data? All these topics, valuable both academically and from an application standpoint, remain to be further investigated in depth.
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-13PREPARATION OF ULTRAFINE POWDERS BY REACTION-PRECIPITATION IN IMPINGING STREAMS I: "ULTRAFINE" WHITE CARBON BLACK
Because of the change in degree of dispersal, ultrafine particles have a number of excellent characteristics that are quite different from those of corresponding solids in normal conditions" hence they have several valuable application features and so have received much attention. Consequently, the preparation technologies for ultrafine powders have become one of the hot spots of investigation. In many countries nanotechnology is becoming one of the key technologies for the 21 st century. Scientifically, the term "ultrafine powder" or "ultrafine particles" is used to describe solid products in which the particle sizes are no greater than 100 rim. The "ultrafine" white carbon black to be discussed in this chapter is the product of particles of a smaller size than those in common products, i.e., "ultrafine" here is not a scientific but a commercial term. The ultrafine powders in the scientific sense, e.g., nano copper, nano TiO2 and nano hydroxyapatite, and related topics will be discussed in later chapters. Nevertheless, the principles involved in the preparation of "ultrafine" white carbon black by impinging stream reaction-precipitation are very similar to those involved in the preparation of the nano powders mentioned above. Therefore this topic is discussed here under the overall title "Preparation of ultrafine powders". Because of their important application value, much research and development on the preparation technologies of ultrafine powders has been carried out in the last twenty years and more, and hundreds of preparation methods have been proposed. Since they are not the major topic of this book, neither a description of the classification of the methods nor an introduction to the details of the various methods will be covered here. On the other hand, reaction-precipitation methods generally have a number of advantages such as lower cost, moderate operating conditions, lower equipment requirements, convenience of operation, and normally yield good-performing products etc." thus they occupy an important position among the various methods. The liquid-continuous impinging stream (LIS) device has the features of efficient micromixing and strong pressure fluctuation and, being the major equipment in the preparation of ultrafine powders by reaction-precipitation, it has exhibited excellent performances in a number of application cases. However, of the various preparation
269
270
IMPINGING STREAMS
methods, the LIS device is only applicable for the reaction-precipitation process and not for others. It can be considered therefore that the LIS device has no superiority over other preparation technologies.
13.1 ADAPTABILITY OF LIQUID-CONTINUOUS IMPINGING STREAMS FOR PREPARATION OF ULTRAFINE POWDERS Essentially, the key process in the preparation of ultrafine powders by reactionprecipitation is crystallization from a solution. As mentioned in the previous chapter, crystallization from a solution includes two steps: nucleation and crystal-growth. Both can occur only in a supersaturated solution and spontaneous nucleation can occur only when the concentration of the solute in the solution is over the super solubility of the substance involved. The rate equation for nucleation derived from the principles of thermodynamics is represented by [182] I -16~rcr3M2 1 Np - Kp exp 3R3T3p2(lnS) 2
(13.1)
where K r, is the nucleation rate constant; S, is the relative supersaturation, S = C/Cs, and Cs is the solubility of the solute. For practical application, Eq. (13.1) is not convenient; and the empirical equation below is usually employed: Np = KpAC t'
(13.2)
where AC is the supersaturation, AC = C - Cs. It is obvious that both nucleation and crystal growth consume the solute in the solution, resulting in a decrease in supersaturation. Therefore there is competition between nucleation and crystal growth for the solute in the solution. From Eq. (13.2), the first premise for preparation of ultrafine powders by precipitation is that the crystallization must be carried out at very high supersaturation in order that nucleation in quantity can occur instantaneously to yield a large amount of crystal nucleus, so that the nucleus cannot grow significantly for lack of the solute and thus remains fine. In the preparation of ultrafine powders by reaction-precipitation, supersaturation is produced by the certain chemical reaction(s). This implies that, chemically, the necessary conditions for the preparation of ultrafine particles by reaction-precipitation are: (1) The reaction producing the solid product must be very fast so that the solute can be produced in quantity instantaneously to achieve the very high supersaturation expected; and (2) The solid product to be prepared in the form of ultrafine powder should generally have a very low solubility. The solution of the substance with larger solubility usually has a wide metastable region so that it is not easy to achieve very high supersaturation rapidly by chemical reaction. Secondly, the solid precipitating out from a solution, i.e., crystallization, must be carried out in an environment of high and uniform supersaturation otherwise the newly
PREPARATION OF "ULTRAFINE" WHITE CARBON BLACK
271
formed nucleus may grow quickly at low supersaturation leading to a product with a wide size distribution, even if the mean supersaturation is very high, because the existence of a local concentration difference is unavoidable under not completely uniform conditions. In the case of product having a wide size distribution, the mean size of the particles in the product is usually larger, because the large particles have a very large "weight" in the weighted mean size. Summarizing the discussions above, the necessary conditions for the preparation of ultrafine powders by reaction-precipitation may be concluded to be the following: (1) Very fast reaction that can yield instantaneously a quantity of the substance to be precipitated; (2) Very low solubility of the substance to be prepared in the form of ultrafine particles: (3) Extremely high supersaturation; and (4) Very uniform supersaturation. As has been seen before, liquid-continuous impinging streams (LIS) have the excellent features of efficient micromixing and very strong pressure fluctuation. These characteristics promote the reaction(s) in liquid phase in the preparation of ultrafine powders to proceed more rapidly to yield instantaneously high supersaturation on one hand; and the strong micromixing can ensure very uniform supersaturation in the solution, on the other. Therefore it can be expected that a finer product with a narrower size distribution can be prepared with an LIS device as the reaction-precipitation equipment. The results of several investigations have proved that the LIS device is one of most advantageous technical equipment for preparation of ultrafine products by reaction-precipitation.
13.2 PROPERTIES OF WHITE CARBON BLACK AND CHEMICAL REACTIONS IN ITS PREPARATION BY PRECIPITATION PROCESSES "White carbon black" is the global title for synthetic hydrated silica and silicate, scientifically termed "hydrated silicon dioxide" (SiO2.nH20), and is an important additive of rubber etc'. [1831]. Because of its white color and properties comparable with carbon black, it is termed "White Carbon Black", for which the Chinese quality standard is given in Table 13.1. Both its application value and the commercial price of white carbon black are closely related to its particle size [184], although the size index had not been included in the standard. For example, the products sized about 30 to 50 lam can be used only in production of general rubbers etc: while the "ultrafine" products sized less than 10 lam can be used as a reinforcing agent in high-grade rubber products such as meridian tyres, insulation and piezoelectric rubbers etc, because of their large specific surface area, strong linking ability, good dispersity, and good optic and mechanical properties; they are also widely used in the production of accurate foundry, high grade fillings, coatings, precise ceramics, compounded materials, and light-guide fibres etc. Accordingly, the market price tot "'ultrafine" white carbon black can be over ten times higher than the common ones.
272
IMPINGING STREAMS Table 13.1
Chinese National Standard of White Carbon Black (GB 10517-89) Item Content of SiO2, % Color Residual on mesh of 45 ~tm, %
Index > 90 Prior or equal to the standard sample < 0.5
Heating loss, %
4.0-8.0
Calcinations loss, % pH
< 7.0 5.0-8.0
Total Cu, mg.kg-j
< 30
Total Fe, mg.kg-~
< 1000
Total Mn, mg.kg-~
< 50
DBP adsorption, cm~. g-~
2.00-3.50
In the profession, people are used to using size to describe the dimension of the assembled particles, but not the primary particles. In the discussions in this chapter, this convention will be followed. There are many industrial methods for the production of white carbon black [184], e.g., the gaseous phase processes, the carbonization method, etc., among which the liquid phase processes, especially the precipitation methods, are most widely employed. The liquid phase processes cannot yield so fine a product with such high activity as those produced by gaseous phase processes; however their products can meet many requirements with very low costs and therefore have received wide attention. Among the various processes producing white carbon black of various grades, the precipitation methods occupy the dominant position. However, the common precipitation process is unable to produce a product with very fine particles, e.g., no greater than 10 ~tm. Because of the attraction of the high application value and high commercial price, a number of investigations and developments had been made in modifying the precipitation process in order to produce ultrafine white carbon black, giving rise to several processes, such as the collosol process, the gelation process, and the twice collosol process etc. The Degussa co. in Germany and the Silica co. in Japan have successfully produced products with assembled particles no greater than 10 ~tm by the process of twice collosol. However, these improved processes usually have the disadvantages and/or difficulties of complicated system schemes, harsh requirements of operation and control, and higher cost etc. It is of interest therefore to develop a simpler process with greater operating flexibility. Chen[ 16] studied experimentally the preparation of "ultrafine" white carbon black by the common (one-step) precipitation process with the submerged circulative impinging stream reactor (SCISR) developed by Wu [15] as the reaction-precipitation equipment and obtained satisfactory results.
PREPARATION OF "ULTRAFINE" WHITE CARBON BLACK
273
In the common precipitation process, the double decomposition reaction between an inorganic acid and the sodium silicate is employed to yield hydrated silica: Na20.mSiO2 + 2H + + aq = 2Na + + mSiO2.nH20$ + aq
(13.1)
As in the preparation of most ultrafine powders, an ageing stage of a certain timeinterval under the condition of continuous stirring is needed for the reacted mixture for deactivation of the surface of fine particles in order to get a stable product consisting of assembled particles. It has been mentioned before that, in addition to the effective micromixing and the strong pressure fluctuation, the SCISR has the special flow configuration of perfect mixing-plug flow in series. With the SCISR as the reactionprecipitation equipment, nucleation in quantity occurs essentially in the impingement zone, while, after nucleation in quantity, the deactivation of the surface of the newly formed fine particles may occur to an extent during the solution of very low supersaturation after nucleation, carrying the solid particles flowing through the regions essentially without mixing, the annular chamber between the drawing tube and the reactor wall and inside the drawing tubes restraining coalescence of the particles. On the other hand, the time that the suspension flows through the regions without mixing is much longer than that through the impingement zone, and thus the ageing can be carried out partially during the flow so that the time for ageing after reaction may be greatly reduced. A large amount of water is usually contained in the dictyo-structure of the hydrated silica precipitated, including both free and combined moisture. The precipitate is separated out by filtering and washed to remove impurities, mainly Na and acid radical ions. The cake is made into slurry again by stirring and the latter is then spray-dried to yield the powdery product of white carbon black. The inorganic acid used for the preparation of white carbon black can be any of the hydrochloric, sulfuric or nitric acids; that mostly used is sulfuric acid and it is also used in the present investigation as the reagent or the precipitant.
13.3 EXPERIMENTAL EQUIPMENT AND PROCEDURE 13.3.1 Experimental equipment In the investigation on preparation of "ultrafine" white carbon black the submerged circulative impinging stream reactor (SCISR) is used as the reaction- precipitation equipment, the structure of which is the same as that shown in Fig. 10.2; it also has the same effective volume of 3.6×10 -3 m 3, but the top cover is not used for convenience because the process is carried out at room temperature and under atmospheric pressure. For further understanding the performance of the SCISR by comparison, the preparation experiments are also carried out simultaneously in a stirred tank reactor (STR) with an effective volume of 0.6×10 -3 m 3, the structure of which is indicated in Fig. 13.1. In order to mimic industrial conditions, the STR is equipped with three dampers distributed uniformly along the circle; the stirrer is a flat paddle.
274
IMPINGING STREAMS
per
Figure 13.1 Stirred tank reactor for comparative experiments.
13.2.2 Experimental procedure The testing materials used in the investigation are commercial sodium silicate with the modulus of 3.3 and sulfuric acid. The following substances are studied: (1) experiments for optimizing the operating conditions for the SCISR in semi-batch operation mode; (2) experiments in the SCISR operated continuously under the optimized conditions; (3) comparative experiments for the SCISR and the STR; and (4) a study of the final treatment of the precipitate from the continuous operations of the SCISR. In semi-batch operation, the SCISR is first filled with a solution of sodium silicate with certain concentration, and then a sulfuric acid solution of a given concentration is dripped at a certain rate into the reactor to react with the sodium silicate at a controlled temperature. The reaction continues for a certain interval of time after the dripping has finished. Stirring is then stopped for ageing of the precipitate for a term, and then the precipitate is sampled and the sample is measured with a laser particle-measuring instrument of FAM type to obtain the sizes and size distribution of the particles in the wet product. The experiments of continuous operation are carried out under the optimal conditions determined by the semi-batch operation experiments. Na2SiO3 and H2SO4 solutions of given concentrations are fed into the reactor at the inlets of the drawing tubes on the two sides of the SCISR and react with each other, while the reacted mixture over the settled level overflows from the reactor through the overflow port on the upper side wall. When the operation achieves a steady state, the overflow slurry is collected in a container for ageing, and then the precipitate is separated from the liquid, sampled and measured for the size distribution. Since the investigation is aimed at the preparation of ultrafine product, the measuring terms are limited to the size and size distribution of particles in the reacted precipitate in most of the experimental runs.
PREPARATION OF "ULTRAFINE" WHITE CARBON BLACK
275
The comparative experiments are carried out only in semi-batch operation mode; the experimental conditions and the operation procedure are identical for both the SCISR and the STR, and the specific effective power inputs for the two reactors are controlled rigorously to be identical. The experimental run in the study on the final treatment of the precipitate is carried out in the SCISR operated continuously under the optimized conditions for about 48 h to accumulate a large enough amount of product for spray drying. The whole of the reacted precipitate separated out is fully mixed and is then used for the spray drying experiment.
13.4 RESULTS AND DISCUSSIONS 13.4.1 Semi-batch operation The major goal of the semi-batch experiments with the SCISR as the reactionprecipitation equipment is to determine the optimal conditions. An experimental technique of normal design is employed and a total of eight influencing factors are examined. The normal design conditions are listed in Table 13.2 where the range of ageing time is determined primarily by searching experiments. The number of temperature levels tested is only two, because complete gelation happens in the reaction mixture so that the operation is destroyed even at 60°C, and so any test at higher temperature becomes meaningless. Table 13.2
Normal design for the experimental conditions Level
Operation variable* A
B
C
D**
E
F
G
H
1
38
600
8
1
30
300
0
1800
2
60
900
9
2
60
3
--
1200
10
3
120
600
1
3600
900
2
7200
* Operating variables: A--reaction temperature, °C; B--rotary speed of propellers, rpm; C--concentration of H2SO4, kmol'm-~; D--feeding position; E--time of feeding, s; F--reacting time after feeding; G .... amount of dispersion agent, g; H--time for maturation, s. ** Feeding position: 1--inlet of drawing tube, 2--outlet of drawing tube, 3 --center of the reactor. The results of the normal-designed experiments are listed in Table 13.3, where I, II, and III denote the summations of the Sauter mean diameters, d32, at Levels 1, 2, and 3, respectively; and R is the extreme difference at a certain level. From Table 13.3 it follows that the order of influence of various factors on the mean size of the particles in the precipitate is D>C>B>F>H>G>E>A. The influences of the latter three factors, G, E, and A, can be considered as very weak, while the most serious factors leading to gelation of the reaction mixture are, in order, A and G, suggesting the reaction temperature cannot be too high and the amount of dispersion agent cannot be too large.
276
IMPINGING STREAMS
Table 13.3 Results of the normal designed experiments Variable
A
B
C
D
Run NO
E
F
G
H
Level
Results Status
1
1
1
1
3
2
2
1
2
1
1
1
3
1
3
1
2
4
1
1
2
5
1
2
6
1
7
2
d32,gm
1
2
thicken
1.455
1
2
1
normal
1.168
3
3
3
3
thicken
1.481
2
1
2
3
1
normal
1.471
2
3
3
1
1
3
normal
1.452
3
2
1
2
3
2
2
normal
1.337
1
1
3
1
3
1
3
2
thicken
1.360
8
1
2
3
2
2
3
1
1
normal
1.297
9
1
3
3
3
1
2
2
3
normal
1.667
10
2
1
1
1
1
3
1
3
gelation
1.248
11
2
2
1
2
3
2
2
2
gelation
1.360
12
2
3
1
3
2
1
3
1
gelation
1.399
13
2
1
2
3
3
3
2
1
gelation
1.463
14
2
2
2
1
2
2
3
3
gelation
1.451
15
2
3
2
2
1
1
1
2
gelation
1.430
16
2
1
3
2
2
1
2
3
gelation
1.541
17
2
2
3
3
1
3
3
2
gelation
1.420
18
2
3
3
1
3
2
1
1
gelation
1.471
I
12.688
8.538
8.111
8.035
8.404
8.35
8.353
8.269
II
12.783
8.148
8.604
8.58
8.48
8.875
8.536
8.362
HI
--
8.785
8.756
8.856
8.587
8.246
8.582
8.84
R
0.095
0.637
0.645
0.821
0.183
0.629
0.229
0.571
I + II + HI =25.471
PREPARATION OF "ULTRAFINE" WHITE CARBON BLACK
277
Taking into account the fact that the average size of the product should be as small as possible and that the operation must be stable, the optimal operating conditions determined are A~, B2, C~, D~, E~, F3, G~, and H~, i.e., the optimized conditions are: reaction temperature of 38 °C, rotary speed of the propellers 900 rpm, concentration of -3 sulfuric acid 8 kmol.m -, position tbr feeding at the inlet of the drawing tube, feeding time 30 s, reaction time after feeding 900 s, amount of the dispersion agent zero, and ageing time after reaction 1800 s. Under these conditions the SCISR operated in semibatch mode can produce a product of white carbon black with particles sizes from 0.5 to 2.0 gm, the average size ranging from 1.1 to 1.6 gm. The results relating to the influence of the feeding position indicate that the best position is at the inlet of the drawing tube. According to the principles of impinging streams, the essential condition t0r the enhancement of transfer and/or mixing is the impingement between the opposing streams at a certain impinging velocity. The material, a solution or suspension, fed either at the outlet of the drawing tube or at the center of the reactor cannot be accelerated effectively, so that it cannot mix well with other stream(s), the poorest position for material feeding leading to the poorest mixing status is at the center of the reactor. Poor mixing must result in a slow reaction and thus is unable to create a high and uniform supersaturation for precipitation. So, the results described above indicate that the only option for the feeding position is at the inlet of the drawing tube(s). The results relating to the influence of Factor C, i.e., the concentration of sulfuric acid, show that the particle size of the product tends to increase with the H2SO4 concentration increasing. The increase in H2SO4 concentration favors increasing the supersaturation caused by reaction and thus promotes nucleation, on one hand, while the change in H2SO4 concentration also affects the electrical field environment around the newly formed particles, favoring coalescence of particles, on the other. It seems that, for achieving high supersaturation, the lowest concentration of H2SO4 originally selected is high enough and any further increase would be harmful. The influence of Factor B, the rotary speed of the propellers N, actually reflects the effect of the impinging velocity, u0. For the same SCISR u0 is a monodrome function of N. It was somewhat unfortunate that during this investigation the measurement of u0 at various N could not then be carried out and so the rotary speed was used as the influencing factor; now, for the approximate dependence of u0 on N one may refer to Fig. 11.2 in Chapter 11. The results on the mean diameter of the particles versus N exhibit a turning influence, which is similar to that obtained by Chen et al. [165]. In principle, from the flow configuration in the SCISR, the increase in N enhances micromixing in the impingement zone and thus should favor nucleation in quantity, while on the contrary, at too high rotary speed the mean size of the particles increases, as indicated by the data in Table 13.3. The following three facts may account for the phenomena' (1) Too strong micromixing may lead to excessive nucleation, leading to an enhanced coalescence tendency (2) Higher rotary speed increases the collision probability between the fine particles newly formed also leading to an enhanced coalescence tendency; and (3) As the rotary speed increases, the flow rate transported by the propellers increases, suggesting that the amount (volume) taking part in the
278
IMPINGING STREAMS
reaction increases, leading to, under the same other conditions, decreased supersaturation in the reaction region. These three items imply that there exist opposite influences on the mean diameter of the particles when increasing the rotary speed of the propellers, resulting in the overall turning influence. Unfortunately, with existing technical tools it is difficult to determine exactly the reasons for the phenomena described above and further investigations are necessary. The purpose of setting further reaction time after the dripping of H2804 is finished is to exhaust the reactants and to precipitate the hydrated silica fully. During the reaction, the following three actions may occur simultaneously: continuous nucleation, coalescence of particles and deactivation of the particle surface. From the experimental data it follows that Factor F, the reaction time, also exhibits a turning influence on the mean size of the particles and a maximum value appears at a time of about 600 s. It is possible that, at that time, infirm coalescence of the particles occurred and some of the coalescent particles can be broken into individual particles under the conditions of stirring-mixing. Therefore, in order to produce a finer product, it is necessary to arrange a period of time for further reaction after the dripping of H2SO4 is finished under continuous stirring conditions. The ageing time, Factor H, exhibits a monotonous increasing influence on the mean size of the particles, suggesting that the particles of hydrated silica have a significant tendency to coalesce under stationary conditions. Therefore, the ageing time before further processing of the precipitate should not be too long. Both the Factors E, the dripping rate of H2SO4, and G, the amount of dispersion agent, exhibit monotonously increasing influences on the mean diameter of particles; but both their effects are insignificant. Globally, the extreme differences, R in Table 13.3, for various factors are not large. This is because the result of a number of investigations on this topic have been referred to so that all the conditions selected for the present study are essentially in the operational ranges.
13.4.2 Continuous operation of the SCISR To determine the optimal conditions, experiments of continuous operation are carried out at a reaction temperature of 25°C, while solutions of Na2SiO3 and H2804 are fed at the inlets on the drawing tubes on the two sides, respectively; the mean residence time of the reaction mixture in the SCISR is 900 s. At a steady state of operation, the overflow suspension is collected for ageing, and then the precipitate is separated from the liquor, the size distribution of particles in the precipitate are sampled and measured. In order to examine the possibility of increasing the concentration of the slurry for a larger capacity reactor, experiments with various concentrations of Na2SiO3 solutions are carried out; all the other conditions are the same as the optimized ones determined in the last section. The results are illustrated in Table 13.4. The measurement under the condition of the Na2SiO3 concentration being 0.84 kmol.m -3 is repeated many times, while the variation of the mean diameter of particles in the precipitate is small and is
PREPARATION OF "ULTRAFINE" WHITE CARBON BLACK
279
within the error range for the measuring instrument, suggesting the operation is very stable. The data listed in the table are the averages. From the data listed in Table 13.4 it follows that an increase in the concentration of the reaction slurry is possible, but the magnitude of the permitted increase is not large, otherwise gelation would occur, resulting in operation break-up. The reason is clear: the increase in the concentration leads to increased viscosity of the slurry, negatively affecting mixing. The most significant conclusion obtained by these experiments is that the SCISR can be operated continuously in the preparation of ultrafine white carbon black by the common (one-step) precipitation process, and the average size of the particles in the product is essentially the same as those obtained by semi-batch operation. To date, all the commercial reactors for preparation of white carbon black by precipitation processes are STRs operated in semi-batch mode. Globally the flows in these devices have a perfect-mixing feature. It is likely that the special flow configuration of the perfect mixing-plug flow in series in the SCISR is the major reason that it can be operated continuously. It is clear that continuous operation is normally superior to batch or semi-batch, especially for production on a large scale. Table 13.4
Results for different concentrations of Na2SiO3 solution Concentration of Si02, kmol.m-3
Particle size, gm d32
d,.lo
dvso
dvgo
dvmin
d .....
0.66
1.245
1.118
1.268
1.373
0.667
1.556
0.80
1.331
1.004
1.431
1.793
0.222
2.556
0.84
1.381
1.106
1.459
1.571
0.370
2.334
0.90
Gelation
13.4.3 Comparative experiments in semi-batch operation Comparative experiments are carried out between the SCISR and the STR for further verifying the good performance of the SCISR. Both the reactors are operated in semibatch mode and under the same optimized conditions as before. The structure and dimensions of the experimental reactors have been described in Section 13.2. The size distributions of the particles in the precipitates from the two reactors are illustrated in Fig. 13.2. Obviously, the product from the SCISR is finer with a narrower size distribution, i.e., more uniform in size. It should be noted that the effective volume of the experimental SCISR is six times that of the STR, suggesting the scales favor the
280
IMPINGING STREAMS
STR. Thus, these results indicate an obvious difference between the performances of the two kinds of reactor. 40
30 0 0
~
20
r~
10
_1_Jl._1
1.0
i ~
2.0
....
3.0
Particle size, ~tm Figure 13.2 Comparison between particle size distributions of white carbon black prepared in SCISR and STR, respectively. A--0.6xl0 -3 m 3STR .--3.6x10 -3 m 3 SCISR.
13.4.4 Study of the final treatment of the reaction product As mentioned above, the major goal of the present investigation is to produce white carbon black product as fine as possible, reaction-precipitation being the key operation for its production. In order to focus attention on the major problems, all the measurements of size and size distributions made above are with the reacted wet precipitates. To examine the size stability of the product during the final treatment, experiments on spray drying of the reacted wet precipitate are carried out. The reaction-precipitation takes place continuously under the optimal conditions determined in Section 13.3.1; the washed cake separated from the liquor and the washing water is made into slurry again by stirring and is then spray-dried in a tower of 500 mm in diameter with hot airflow to yield the dry product. Since there is a great difference between the capacities of the reactor and the spray dryer, the reactionprecipitation must be operated for a long time, over 48 hours, until the amount of wet precipitate collected is large enough for the spray dryer operation for, at least, 2 hours; and then the dryer can be operated. Therefore the mean time for maturation of the precipitate is very long, over 24 h and the longest can be 48 h. All the dried products collected at the bottom of the dryer, from the cyclone and the bag fitter, respectively, are put together and mixed fully, and then sampled and measured for size and size distribution. Data for the characteristic parameters of sizes are listed in Table 13.5.
PREPARATION OF "ULTRAFINE" WHITE CARBON BI,ACK
1281
From the data listed in Table 13.5 it can be seen that the Sauter mean diameter of the dried product, d32, is larger than that of the wet precipitate obtained under the same reaction conditions by about 10%, or by 0.15 gm. An obvious fact is that no matter whether at the bottom of the dryer or in the cyclone or in the bag filter, the recovery of the finer particles must be lower than that of the larger particles. These differences between the recoveries of particles of different sizes must lead to an increased mean diameter of the product. If this fact is taken into account, the sizes of the particles can be considered to be stable enough during the final treatment of the precipitate, without coalescence of particles occurring. Table 13.5 Characteristic sizes of the spray-dried product from SCISR operated continuously Characteristic size lam
d-,_~
d, 1!~
d, 50
dvg0
dvmin
d,m~,x
1.491
i.294
1.530
1.702
0.556
2.0
13.5 CONCLUSIONS Utilizing its features of efficient micromixing and very strong pressure fluctuation and the special flow configuration ot perfect mixing-plug flow in series, the submerged circulative impinging stream reactor (SCISR) is used for the preparation of "ultrafine'" white carbon black by the common (one-step) precipitation process; comparative experiments are also made between the SCISR and the traditional stirred tank reactor (STR). The following main results are obtained: ~,1) In the SCISR of 3.6x10 ~ rn ~ m effective volume, the common (one-step) precipitation process operated Jn semi-batch yields ~ultrafine" white carbon black consisting of particles sized () 5 to 2.0 gm, the average sizes ranging from 1.1 to 1~6 btm. (2) The main factors affecting the dzc and ;i~:.-, distribution of the particles in the product of white carbon black are detcrmined experimentally, for which the optimal conditions are: the reaction temperature is the common (room) temperature; the concentration of H-,SO4 solution 8 kmol.m-3; the feeding position at the inlet of the drawing tube(s): the reaction time after feeding of all the reactants 900 s; and the ageing time 1800 s. (3) The SCISR can also be operated continuously, and the product so prepared ha,, essentially the same size and size distribution a~ that obtained by operation in semi-batch mode under the s~me condition~. (4) The results of the comparative exp~:rimer.t~ operated in semi-batch m~de indicate that the product prepared with the SCISR is finer and with a narrower distribution than that from the STR of 0.,q×10 --~ m 3 in effective volume, suggesting that the performance of the SCISR i,: ,,uperior to t>,at or" the STR.
282
IMPINGING STREAMS
(5) The results of the spray drying experiment of the wet precipitate show that the particle sizes of particles in the product produced in the present investigation are stable, and no coalescence of particles during the final treatment of the reacted wet-precipitate is observed.
-14PREPARATION OF ULTRAFINE POWDERS BY REACTION-PRECIPITATION IN IMPINGING STREAMS I1: NANO COPPER AND ITS SURFACE IMPROVEMENT
14.1 INTRODUCTION As is well known, the nanometer, nm, is a measurement of materials and represents a length of 10-9 m, which is equal to the scale of about 10 atoms. The term "nano material" indicates those materials consisting of particles sized less than 100 nm in every dimension, i.e., solid materials consisting of ultrafine particles sized from 1 to 100 nm. Since the dimensions of the composition phase or the crystalline particles are near molecular size, nano materials have a number of excellent characteristics which normal materials do not possess and so can be widely used in the fields of electricity, magnetism, optics, superconductors, intelligent materials, hydrogen-storage, biomedicine, nano-medicine, functional eyes, functional ceramics, functional fibers etc., and have very high application value. As mentioned in the previous chapter, many countries, including the industrially most developed ones, have adopted nanotechnology as one of the key technologies for development in the 21st century. It is no exaggeration to say that developments in nanotechnology will yield significant and farreaching influences on science and technology, economics, military affairs, and daily life etc. in the coming few decades. Put simply, nanotechnology includes the aspects of preparation, property characterization, surface improvement, and application of nano materials. It involves many disciplines and its progress needs the cooperation of scientists and engineers from various disciplines. Obviously, among the aspects mentioned above, preparation of nano materials is the basis. If there was nano material preparation any other work related to nano materials would be meaningless, like a tree without roots. Because of their important application values, ultrafine powders have been the subject of a number of investigations and developments in the last two or three decades, and many kinds of preparation methods have been proposed. By examining the status of the research and development and the various methods proposed, it is not 283
284
IMPINGING STREAMS
difficult to see that chemical engineering, as a traditional and old discipline, has played and will continue to play a very important role in research, development and application of the technologies for the preparation of ultrafine powders. With the discovery that the features of efficient micromixing and strong pressure fluctuation existing in the submerged circulative impinging stream reactor (SCISR) favor the preparation of ultrafine particles by reaction- precipitation, in addition to that of the "ultrafine" white carbon black, the preparation technologies of several nano particles were investigated experimentally with the SCISR as the reaction-precipitation equipment, and all the investigations yielded satisfactory results. In all the processes the method of reaction-precipitation were employed because only in such processes can the SCISR exhibit its superior performance. This chapter introduces investigations into the preparation of nano copper and its surface improvement together with the major results; the preparation of the other two nano materials, Titania and hydroxyapatite, will be discussed in later chapters.
14.2 PROPERTIES AND MAIN USES OF NANO COPPER Nano copper powder has the size- effects, the quantum-tunnel effect, the surface-effect and the volumetric effect similar to other nano metal powders and thus exhibits many special properties quite different from those of normally assembled copper [185-187]. Common metallic copper is purple in color, and the melting point is 1084°C, boiling point 2582°C, and density 8920 kg.m -3. While the melting point of the nano copper with an average size of 40 nm drops to 750°C and that with an average size of 20 nm drops even more sharply to the level of 39°C, the specific surface area of nano copper increases rapidly and, consequently, the surface energy increases sharply as the particle size reduces. For example, when the average size reduces from 100 nm to 10 and 1 nm, the specific surface area increases rapidly from 6.6 mZ.g-~ to 66 and 660 mZ.g-~, respectively; the surface energy increases from 590 J-mo1-1 to 5900 and 59000 J.mol -~, respectively; and, correspondingly, its reactive and catalytic activities are greatly increased. In addition, the specific heat capacity of nano copper is twice as large as common copper and the self-diffusivity of nano copper crystalline is 1016 to 1019 times that of common copper crystalline, e.g. the diffusivity of the nano copper sized 8 nm is of the value of 2.6x10 -2° m2.s-~. Nano crystalline copper has a greater strength than that of the common copper, and exhibits plastic ductility. Its coefficient of elongation is over 5100% and the phenomenon of hardening will not appear during handling. Its hardness and yield-strength are higher than those of common copper by 50 and 12 times, respectively. Nano metallic copper tends strongly to electric neutrality and exhibits almost non electric conduction. Its electrical resistivity increases as the particle size reduces, while the thermo-coefficient of electrical resistivity decreases, and can even be a negative value, as the particle size reduces. Because of the extremely high activity of its surface, nano copper very easily adsorbs oxygen in the surrounding air and, simultaneously, it is oxidized. Also, nano copper has a very great ability for light-absorption, while its ability for light-reflex is very weak, normally lower than 1%. Nano copper powder has important applications in the following fields:
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
285
(1) Catalysts: Nano copper can be used as a catalyst for deep splitting of long-chain hydrocarbons in the petrochemical industry. For the hydration of acrylonitrile nano copper exhibits very high catalytic activity and selectivity [188]. It is sldo highly active in both the reactions of ethyne polymerization and the catalytic oxidation of CO. The results of the investigation on the catalytic activity of nano copper particles in the polymerization of ethyne carried out by Wang et al. [189] showed that the size of the nano copper particles has an important influence on the catalytic activity: the smaller the particle size, the higher is the yield of the product. The results of the comparative experiments to assess the catalytic activities of nano Cr, Mn, Ni, Fe and y-A1203 catalysts [190] showed that for a conversion of 100%, the reaction temperature required by nano copper is the lowest, suggesting it has the highest catalytic activity. (2) Electro-conductive rubber material: Because of its great strength and much lower price than such expensive metals as silver and palladium, nano copper or coppersilver double metal powder can be used in the electronics industry to take the place of those expensive metals for the preparation of electro-conductive rubbers [191], electro-conductive slurry, and electrode materials etc.; in addition, the copper-silver double metal powder has the characteristic of antibiosis. For such use nano copper or copper-silver powder should be needle-like crystalline; the nano copper powders of sphere-like crystalline has very low electro-conductivity. (3) Additives in greases of high grade: This is one of the fields where nano copper is most successfully applied. Copper is a soft metal. The addition of oil-soluble nano copper in grease can increase the wear-resistance performance by a wide margin and can form highly efficient arc-eliminating electro-conductive grease. Also, the addition of nano copper of sphere form gives the shaft-bearing system a selfrenovating function. In an experimental study Dou et al. [192] considered that the fine particles of metallic copper exert a bearing effect, polishing effect, metallurgical effect and strengthening effect on the interface. The fine copper particles can penetrate through the interface and thus improve the surface. Under the action of mechanical motion forces, the fine particles are pressed and inlayed on the friction surface in the form of atoms to form complex metallic structures, resulting in a greatly reduced friction drag force. Xia et al. [ 193] proposed that nano copper powder takes part in lubrication in the forms of pads and balls. The polishing scratch increases very little as the loading increases, the surface is kept smooth, and the concaves on the slipping surface are filled and leveled up, yielding improved lubricating performance; the addition of 5% triethanolamine yields a composite effect [194] that greatly increases the wear and tear of the lubricating grease. Xu et al. [195] formulated several lubricants with nano copper particles of various sizes in the range 4 to 50 nm, which were prepared by gasstream grinding in fluidized bed, and tested their performances. The experimental results showed that with the nano copper particles of 4 to 15 nm the increase in lubricating efficiency was very significant. Later, Xu et al. used the nano copper powder sized about 10 nm prepared by reduction-precipitation in the experiments carried out by Chen [196] to take the place of that prepared by gas stream grinding, and the formulated lubricant exhibits very good performances: all the indexes achieve or even exceed those of all the existing lubricants.
286
IMPINGING STREAMS
14.3 PRINCIPLES AND EXPERIMENTAL METHOD 14.3.1 Chemical reactions in preparation of nano copper by reductionprecipitation Various methods for the preparation of nano copper have been proposed. Because of its feasibilities in both technology and economics and to take advantage of the strong points of liquid-continuous impinging streams, in the investigation carried out by Chen [196] the reduction-precipitation process in liquid phase was employed, i.e., a certain reducing agent is added to a solution to react with the soluble copper salt in the liquid phase, and the copper particles formed as the product of the reaction precipitate out. Several reducing agents have been used for this purpose, such as ascorbic acid, formaldehyde, hydrazine hydrate, sodium hypophosphite, potassium borohydride, sodium hyposulfate etc; while in the preparation of various nano metal powders potassium borohydride exhibited perfect performances, although its price is somewhat higher. In the investigation by Chen, a CuCI2 solution was used as the copper source and potassium borohydride, KBH4, as the reducing agent. The major chemical reaction occurring in the liquid phase can be represented by 4CUC12 -+-KBH4 + 8 K O H = 4Cu $ + 8 K C I + K B O 2 + 6 H 2 0
(14.1)
Because of the multi-valence nature of Element Cu, the simple substance copper formed may react with Cu 2+ in the solution to form the monovalence copper ion Cu +, and the latter may also react with C1- to produce precipitate: Cu @- C u 2+ -+2Cu +
(14.2)
Cu + + C1----, CuC1 $
(14.3)
2CuC1 + H20--~ 2HC1 + Cu20 ,L
(14.4)
Therefore, to increase the yield of copper, the formation of the Cu ÷ ion in liquid phase must be restrained. For this purpose a certain amount of aqua ammonia is added, as the coordinative agent, to form the coordinate complex ion Cu(NH3)42+ for restraining the reaction producing Cu +, Eq. (14.2), and also for dissolving the precipitates of CuC1 and Cu20 to keep Cu 2+ from precipitation in the form of hydroxide yielding a product with low purity under the condition of basicity. It is known by an analysis of the equilibrium constants that the concentration of Cu 2+ ion in the solution with aqua ammonia added is very low and is independent of CuCI2, but is relative to the concentration of the dissociated ammonia. On the other hand, under the condition of KBH4 excess the reaction is carried out very completely; while the existence of the coordinative agent ammonia decreases the reducibility of KBH4. Therefore both pH and the concentration of KBH4 must be enhanced.
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
287
14.3.2 Experimental equipment and procedure The experimental equipment system for the preparation of nano copper powder by reduction-precipitation is shown in Fig. 14.1, where the submerged circulative impinging stream reactor (SCISR) has the same structure as was used in the investigations described in the previous chapters in Part II of this book, with the same effective volume of 3.6×10 -3 m~; it is also operated without the top cover; but is made of titanium for anti-corrosion of C1-~. The formulated KBH4 solution is first fed into the SCISR; then the driving motors on the propellers are turned to push the liquid so that it circulates inside the reactor and impinges against the opposite stream, while the rotary speed of the propellers is manipulated so as to achieve the required impinging velocity, u0. Water at a given temperature passes through the heat-exchanging jacket outside the reaction vessel to keep the temperature of the process material inside the reactor constant during the reaction. When both the flow and the temperature in the reactor achieve a steady state, the liquor reactants are fed into the reactor in the modes selected, which will be described later. The reactions take place for a certain interval of time after feeding the reagents, then the reaction mixture is discharged from the reactor and separated with a high-speed centrifuge; the cake is washed with water to remove the residual Cu 2+ (check with potassium ferrohexacyanide K2[Fe(CN)6] for without Cu2+), washed with acetone, and then dried at 30 °C under vacuum to yield the final product of nano copper. The make-up of the reactant solutions is an important link in the preparation of nano copper by reduction-precipitation affecting the efficiency of the process. The procedures for the two solutions are as follows: .....
t
t
Figure 14.1 Scheme of the experimental system. 1, 2-solution tank; 3-propeller; 4-drawing tube; 5-discharging port; 6-jacket for heat exchange.
288
IMPINGING STREAMS
(1) KBH4 solution: 2000 mL potassium borohydride solution is prepared for each run, its concentration being determined according to the concentration of CuCl2 to be used in the run; a little strong aqua ammonia and an appropriate amount of KOH solution with the concentration of 4 mol.L -~ are added to adjust the pH, and 10 to 40 mL of the dispersing agent is also added to the solution. (2) CuC12solution: 1500 mL of copper chloride solution is prepared for each run. The solution is made-up of solid CuClz.2H20, strong aqua ammonia, and the surfactant, 2% PVP solution. The amount of CuClz.2H20 to be added is determined from the required concentration; the amount of aqua ammonia is the stoichiometric amount for producing Cu(NH3)42+ ions; the amount of 4 mol.L -~ KOH solution > 75 mL; and the amount of the dispersing agent, 2% P V P , is in the range of 10 to 40 mL. The experiments are carried out in two stages. The major goal of the work in the first stage is to make nano copper powder with a relatively uniform size distribution; the work also aims to determine the optimal or, at least, most feasible mole ratio of CuCl2 to KBH4. In the second stage the influences of various operating parameters on the mean size and the appearance of the produced nano copper aree investigated to yield the optimal conditions. The effecting factors examined are the concentration of copper salt, the amount of PVP, the pH of the reaction mixture, the reaction temperature, and the impinging velocity (satisfied by controlling the rotary speed of the propellers). Since the global goal is to prepare finer nano copper with a narrower size distribution, the mean size of the product is always taken as one of the major criteria for analysis and comparison of the results.
14.4 RESULTS AND DISCUSSIONS ON THE PREPARATION OF NANO COPPER POWDER 14.4.1 Major results obtained in the first stage As mentioned above the major goal of this stage is to produce a qualified product. Three sets of experiments are carried out with various CuC12 to KBH4 mole ratio, CuClz:KBH4 = 1:1, 1:2 and 1:3, respectively; while the other conditions are fixed at the concentration of CuCI2 C cuc12 = 0.1 kmol.m -3, the reaction temperature T = 20 °C, the rotary speed of the propellers N = 1000 rpm, the pH of the reaction mixture pH 14, the reaction time tr = 10 rain, and the reactants are fed quickly. During each run the resulting phenomena are observed carefully; the sizes and the appearance of the product are characterized with the transmission electron mirror microscope (TEM). The major results obtained are: (1) At the CuCl2 to KBH4 mole ratio of 1:1, the product has a larger size and clumps together significantly, and a modicum of long fiber-like product is formed. The major observations are: in the earlier stage of the reaction brown particles are precipitated, then the precipitate forms gradually wadding;
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
289
the existence of Cu ~+ is detected out. (2) With the mole ratio of 1:2, the product is an average size of 16 to 20 nm, slightly clumped together; the particles are approximately of sphere form without fiber- or needle-like appearance. The observations are: during the reaction the solution is colored black, no significant precipitation within one hour; no Cu 2+ in the precipitate is detected. (3) With the mole ratio of 1:3, the product gathers seriously; there is a needle-like appearance on the surface of the clumped particles and the sizes of the spherical particles are larger. The observations are: the reaction solution is colored brown; no Cu 2+ ion is detected in it; the precipitate appears up to two days later and oxide of a red-brown color forms on the surface of the precipitate. Most of the TEM photos taken at this stage are not given here due to restrictions of space, but that of the product obtained at CuC12:KBH4 = 1:2 is shown in Fig. 14.2. It can be seen from this figure that the particles in the product are sized approximately from 16 to 20 rim. The following results obtained in the first stage of investigation are of significance:
(1) P h e n o m e n o n observations: If the reaction solution or the precipitate e.xhibits a non-black color, there must be un-reacted Cu 2+ or the product of the re-reaction of Cu, Cu +, and thus the purity of the product must not be high; normally, the sizes and appearance of the product would not be ideal, i.e., the particles are largersized, non-sphere crystalline, and significantly clumped together. If the precipitate appears too early, the sizes of the particles in the product are usually large. These observations are helpful in preliminary judgments of whether the reaction conditions employed are good or not. (2) Optimal or feasible mole ratio" According to the mean sizes and the appearances of the products obtained it is considered that, among the three mole ratios tested, CuC12:KBH4- 1"2 is the best. The feasibility of this mole ratio has been verified by the results of repeated experiments.
50 nm
t................................ !
Figure 14.2 TEM photo of the product obtained under the condition of CuCI::KBH4 = 1:2.
290
IMPINGING STREAMS
It should be noted that, from the reaction represented by Eq. (14-1), the stoichiometric ratio of CuClz:KBH4 should be 4:1. That is, with any one among the three mole ratios KBH4 is greatly excessive. This is for the complete reaction of Cu 2+ to increase the yield of the simple substance copper. From the experimental results it follows that with the mole ratio of 1:1, i.e. the excess of KBH4 is smaller, the reaction of Cu 2+ is incomplete; while for the mole ratio 1:3, i.e. the excess of KBH4 is very large, Cu 2+ can be reacted completely, although the side-reaction leads to the formation of copper oxide and, consequently, a decreased yield of simple substance copper. So, the mole ratio of CuC12:KBH4 is an important operating condition that needs to be optimized.
14.4.2 Results on the influences of various factors The influences of various factors, rather than the mole ratio of CuC12: KBH4, are studied in the second stage of the investigation and the experimental conditions are listed in Table 14.1, where the symbols in the second column for the feeding modes denote the following: I - CuC12 solution is added rapidly to the KBH4 solution; IICuCI2 solution is added slowly to the KBH4 solution; III- KBH4 is added slowly to the CuCI2 solution; and I V - I n making up the solutions, the CuCI2 solution is without the addition of KOH, but 160 mL of 4 mol.L -~ KOH is added to the KBH4 solution, while for the reaction KBH4 solution is added slowly to the CuCl2 solution. Among the influencing variables listed in Table 14.1, the rotary speed of the propellers, N, actually reflects the influence of the impinging velocity, u0; while for convenience of operation N is taken as the operation variable. For a given SCISR u0 is a monodrome function of N, and for the reactor used in the present investigation the curve shown in Fig. 10.8 in Chapter 10 is essentially applicable for the relationship between u0 and N. The experimental results obtained in this stage are described below.
Table 14.1
Experimental conditions for study on the influences of various effects Influencing variable Level
Feeding mode*
Concentration of
CuC12, mol-L-1
Temperature °C
Rotary speed rpm
Amount of PVP, mL
pH
I
o. 1
10
600
0
11
II
0.2
20
800
20
12
III
0.3
30
1000
40
13
IV
0.4
40
1200
60
14
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
291
14.4.2.1 Feeding mode The other operating conditions for investigation the influence of the feeding mode are: CcL,cl2- 0.1 kmol.m -3, T = 20°C, N = 1000 rpm, pH - 14, and CuC12 to KBH4 mole ratio 1:2. The major results are as follows" In Mode I (feeding time 1 rain), the results are similar to those obtained in the repeated verifying experiments mentioned in the previous section. The TEM photo of the product shows particle-clumping and the results of X-ray diffraction (XRD) analysis indicate that the product consists mainly of copper powder with a little Cu20. The X-ray spectrum of the product obtained in this feeding mode is shown in Fig. 14.3, which indicates the major part being Cu, with little Cu20 and no CuO. In Mode II (feeding time 10 min), the TEM photo of the product shows clumped particles with a little needle-like crystalline, and the data of XRD indicates that the major product is Cu20, with a little Cu only and no CuO. In Mode III (feeding time 8 rain), the TEM photo of the product shows the loose needle-like crystalline, with a modicum of clumped particles. In Mode VI, the TEM photo of the product shows particles clumping with a modicum of needle-like crystalline cluster sized about 20 nm. After analysis and comparison, feeding mode ! is considered to be the best. On the other hand, from the point of view of rapid reaction yielding very high supersaturation, Mode ! should be the best. Therefore feeding mode I is confirmed. It should be noted that, in Mode III the product essentially of needle-like crystalline can be produced under well controlled conditions, as illustrated in Fig. 14.4. In some cases, e.g. for the production of electro-conductive rubber etc, this kind of product may be required. Therefore the feeding mode III is also of practical significance.
i ~
ii ~i!~i~
i~ ~
i ~~
i
i ii~;~!
~ii'~!;
:~!i~
~iii~
Figure 14.3 X-ray spectrum of the product obtained in Feeding mode I.
~i;iiiii!¸
292
IMPINGING STREAMS
~e......
!:.~
i
Figure 14.4 Nano copper of needle-like form obtained in the feeding mode III.
14.4.2.2 Influence of CuCI2 concentration The other operating conditions for investigating the influence of CuCI2 concentration are: the reaction temperature T = 24°C, the feeding mode I, the rotary speed of the propellers N = 1000 rpm, pH = 14, CuC12 to KBH4 mole ratio 1:2, and the reaction time after feeding tr = 10 min. The major results are:
(1) Observations: At the concentration of
C C u C 1 2 -- 0.1 kmol.m -3 the reaction solution is colored brown, while for other concentrations the color of the solution is brown at the beginning and then turns to black. The higher the CuCI2 concentration, the quicker the color changes, suggesting qualitatively that, from the point of view of increasing the yield of the simple substance copper, the condition of Ccoc~2 = 0.1 kmol.m -3 would not be desirable. (2) Appearance of product from TEM: With Ccucl2- 0.1 kmol.m -3, the product is loosely clumped particles and needle-like appearance; at Ccuc~2 = 0.2 the product consists essentially of sphere particles without needle-like appearance; at Ccuc~2 = 0.3 the appearance of the product is essentially the same as for Ccuc~2 = 0.2; while for Ccuc~2 = 0.4 the product is loose particles and needle-like cluster of long length. (3) Mean sizes: The variation of the mean size with CuC12 concentration exhibits a minimum culminating point, as indicated in Fig. 14.5.
From both the mean sizes and the appearance characteristics it seems that at lower concentrations of CuCI2 solution, normally spherical particles of uniform sizes are produced; while for Ccuc~2 - 0.4 kmol.m -3 needle-like crystalline clusters possibly up to several micrometers long are formed, which consists of the growing particles in the
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
293
system [197]. For crystallization of copper, the thermo coefficient of the supersaturation, d(AC)/dT, approaches zero and the metastable region must be extremely narrow. In the ranges of the operating conditions tested, it is very easy for the reaction solution to be in the unstable region, leading to nucleation in quantity. On the other hand, the surfaces of the nano copper particles contain unsaturated valence bounds and thus have remainder force field. At higher concentrations, interparticle adsorption would occur, resulting in assembly of particles, i.e., the condensation at particle level. These opposite effects may account for the variation in the mean size of the products shown in Fig. 14.5. According to the experimental results, the concentration of Cc,,c~2 = 0.2 kmol.m -3 is the optimal or is, at least, feasible.
/
13 //
/
E ._~ 12 ~D
//
/
/
/
/
/
/
/
<
/ ",..... --. .
.
.
.
.
.
/
/
10 0.1
0.2
0.3
0.4
Concentration of CuCl2, kmol.m -3
Figure 14.5 Influence of CuC12concentration on mean size of the product. 14.4.2.3 Influence of reaction temperature The other operating conditions for the investigating the influence of the reaction temperature are: Concentration of CuCI2 solution Ccuc~2- 0.2 kmol.m -~" Rotary speed of the propellers N - 1000 rpm; Feeding mode I; pH=14; Cue12 to KBH4 mole ratio 1:2. The following major results are obtained"
(1) Observations: The color of the reaction solution is initially green, turns to brown, and then to black in 5 min; at 10°C, the time needed for the color to change to black is longer (about 11 rain). (2) Appearance of product from TEM: The product obtained at the reaction temperature of 10°C consists of particles and needle-like crystalline; that obtained at 20°C is roughly spherical particles without needle-like crystalline and slightly clumped; that obtained at 30°C consists of spherical particles without needle-like crystalline, but significantly clumped together; while the product obtained at 40°C consists of completely spherical particles without needle-like crystalline, but seriously clumped together. (3) Mean sizes: The mean size of the product varies with the reaction temperature in a saddle-like form, as shown in Fig. 14.6.
294
IMPINGING STREAMS
20 E d N •~ > <
j
©
35
40
18 ©
16 14
©
12 10 10
15
20
25
30
T,°C Figure 14.6 Influence of temperature on the mean size. The influence of the reaction temperature on the mean size of the particles and the appearance may be explained as follows: At a lower temperature both the nucleation and the crystal-growth are slow; the crystals grow sequentially and anisotropically, yielding needle- or edge-like particles. At a slightly enhanced temperature of 20°C, the nucleation occurs in quantity and the growth is restrained; the particles grow is©tropically in a mess, and the precipitate tends to be spherical particles. When the reaction temperature is further enhanced to a level of 40°C, the reaction rate is very high and the nucleation rate is greatly enhanced, resulting in the formation of too many nuclei. On the other hand, the enhanced temperature decreases the viscosity of liquid significantly, leading to violent motion of molecules and greatly increased collision probability. The two factors may result in condensation at the particle level; while the is©tropic growth of crystals makes the particles retain their essentially spherical form. Both the theoretical inference and the experimental results indicate that the temperature of 20°C is the optimal or is, at least, feasible. It should be mentioned that in the experiments described above the fluctuation of the reaction temperature is somewhat larger due to non-powerful control, although the deviation of the mean temperature is not large. It was found in the experiments carried out later that, under the condition of rigorously controlled temperature (with fluctuation < +0.5 °C), operation at 20°C can constantly produce a product consisting of particles with an average-size of 5-10 nm, which is superior to the best one shown in Fig. 14.6.
14.4.2.4 Influence of pH in the reaction solution The operating conditions for the investigation of the influence of pH are: the concentration of CuC12 solution Ccuc~2- 0.2 kmol.m -3, the reaction temperature T 20°C, the rotary speed of the propellers N - 1200 rpm, and the feeding mode I. The
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
295
experiments are carried out under the three conditions of pH = 12, 13 and 14, respectively; lower pH cannot be tested because of the necessary addition of KOH to the KBH4 solution. The results for pH = 12 and 13 are essentially the same and are as follows: the reaction is rapid; a large amount of gas bubbles form; the reaction produces a waddinglike black precipitate that is easily settled out. The situation for pH = 14 is quite different from the others: the reaction is slower; no gas-bubbles escape from the solution; the solution is initially brown and becomes black only after as long as 30 min; the dispersity of the particles in the product is good. It is considered that, to an extent, the reduction of Cu 2+ is restrained at high pH, favoring the action of PVP coating the surface of the particles to form small-sized particle-clusters.
14.4.2.5 Influence of the impinging velocity With the other conditions optimized, in the experiments carried out in the range of 600 to 1200 rpm the rotary speed of the propellers shows no obvious effect on either the mean size or the appearance of the products; all the products from these operations have good particle-dispersity and an average size of 5 to 10 nm. Most possibly, the impinging velocity, u0, even at the lowest rotary speed in the normal operation range of the SCISR is high enough for its effect to be covered by other more active factors. These results further indicate that the SCISR has a very high flow efficiency.
14.4.2.5 Influence of the amount of the surfactant PVP The experiments in which 0, 20, 40 and 60 mL of 2% PVP solution, respectively, were added to the solution show the following: (1) in the appropriate range the addition of PVP has a significant effect on the dispersity of the particles, while (2) too much PVP, e.g. 60 mL, may make the surface of the particles have a larger adsorption energy, leading to serious clumping.
14.4.2.6 Optimized conditions By summarizing the results described above, the optimal conditions listed in Table 14.2 can be determined. Table 14.2
Optimized operation conditions Variable
Concentration of CuC12 kmol-m -3
Reaction temperature °C
Feeding mode
pH
Amount of surfactant 2%PVP mL
Rotary speed rpm
Optimal
0.2
20
I
14
40
1200
296
IMPINGING STREAMS
14.4.3 Preparation experiments under optimal conditions In order to verify the feasibility of operation and to provide certain amounts of sample product for further application testing, experiments were carried out repeatedly under the optimized conditions listed in Table 14.2. The following are observed in the experiments: the reaction solution is initially colored green and turns to brown after 3 min, and then to brown-black after 20 min; the reaction is stopped at 40 rain when the solution is colored black. During the reaction a little gaseous ammonia escapes from the solution. The results detected by TEM for a typical product show the precipitate being black clumped particles of good dispersity and an average size of 5.1 nm with a very narrow size distribution. The TEM photo of the product is shown in Fig. 14.7.
.
,::~:~ ....................~.ii~~......;i...............~:: . ... .
~<:~'~i :::.::'?..~::::ili:iiii:~:: .
...... : .....
:~i ~' "~"~!~.~,:
.
.
.
.
.
.
.
.
.
.
.
.
,~:.
:::!~'~:.,~'::i:i::k i~,7::Z;q? ::~.~E':~;il~:~.~~;';i:~:~
.
,~i:~i::~:~::.:i~:::
Figure 14.7 TEM photo of nano copper obtained under the optimized conditions. Twenty experimental runs in total were carried out under the same conditions for testing the sample product; the average size of the mixed sample is measured by TEM to be about 10 nm, suggesting the experimental operations for the preparation of the sample product have very good stability. Furthermore, part of the sample product was used by Xu et al. [195] to formulate the nano lubricating grease, which has exhibited very good performances.
14.4.4 Comparison of the results of preparations with various technologies and devices Comparison is always of interest. Table 14.3 gives a comparison between the mean sizes of nano copper powders prepared with various technologies, different reaction agents and in various reactors. It appears that the SCISR has significant superiority for the preparation of nano materials by liquid reaction- precipitation.
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
297
Table 14.3
A comparison of average sizes of nano copper products obtained by various technologies and with various reactors Reducing agent
Temperature °C
Reactor
Mean size of product, nm
Reference
CuSO4
HCHO
70
STR
100
[ 198]
CuSO4
Ascorbic acid
85
STR
>500
[ 199]
CuSO4
N2H4-H20
60
STR
50-- 500
[200]
CuSO4.5H20
NaH:PO2
55-65
STR
50
[201]
CuSO4
KBH4
20
STR
40-100
[202]
CuCL
KB H4
20
SCISR
- 10
[ 196]
CuC12
Na~S~O4
60-70
STR
20
[203]
Copper salt
14.5 SURFACE IMPROVEMENT OF NANO COPPER" PREPARATION OF CU-AG DOUBLE METAL POWDER Surface improvement is a very important link between preparation and application of nano materials; the specific method employed for the surface improvement is closely related to the occasion where the nano material is to be used. For nano copper powder the main problem is its easy oxidation, i.e., it has no anti-oxidation nature at room temperature. The partial replacement of Cu by Ag to transform the nano copper powder into the Cu-Ag double metal powder is one of the most feasible methods for the surface improvement of nano copper, because the nano Cu-Ag double metal can meet the requirement of anti-oxidation at room temperature; also the latter can be used in the preparation of electro-conductive materials and electro-conductive rubbers etc. to replace some expensive metals such as silver and palladium. There are two methods for the preparation of Cu-Ag double metal powder by replacement reaction. One is to react nano copper with silver nitrate solution directly in the existence of polymer protecting agent to produce Cu-Ag double metal powder, which, by TEM determination, appears usually in the form of twigs. In the other method, an Ag(NHs)2 + solution is used to replace the AgNOs solution to take part in the replacement reaction, and with the other conditions being the same as those in the former method, sphere particles can normally be obtained. The present study is aimed at the preparation of nano particles of sphere form, and so the second method is employed. Both the ions of Ag + and Cu'-' are easily complexed by ammonia (amine) and the corresponding complexes are very stable [204]. In the system of silver-ammonia complex ions the oxidation-reduction standard electrode potential of silver is expressed by
298
IMPINGING STREAMS IAg(NH3)21+ + e = Ag + 2NH3, E ° =+0.373 V
(14.5)
while that of copper is presented by Cu + 4NH3- 2e = ICu(NH3)4I2+, E ° =--0.05 V
(14.6)
Therefore the replacement of Cu by Ag can occur, and after the reaction the dissociated silver deposits onto the surface of the copper particles. It is necessary for the preparation of Cu-Ag double metal powder by the replacement reaction to add a suitable kind of polymer protecting agent [205]. Otherwise serious assembling would occur. Various protecting agents have somewhat different effects on the size and the appearance of the product. With polyvinylpyrrolidone (PVP) as the protecting agent, the product is of a twig- like form, the particles have smaller sizes, and the assembling condition is obviously improved. With gelatin, the product is granular powder with good dispersity and the particles are essentially the same size as the copper powder, while with soluble starch as the protecting agent, the product is usually of a twig-like form and assembling of the particles is relatively serious. In the study worked by Chen[ 196] gelatin is selected as the protecting agent. The experiments for preparation of Cu-Ag double metal powder are carried out in a three-neck flask with magnetic agitator. A given amount of prepared nano copper and 0.5-1.0 g gelatin are added and dispersed into the deionized water contained in the flask under stirring to form a copper particles-in-water suspension; then 100 mL of the aqua AgNO3-ammonia solution prepared with the required amount of AgNO3 plus an appropriate amount of aqua ammonia is dripped slowly from a hopper into the flask to react with the copper particles for 30 min at a constant temperature. The precipitate of grey color produced by the reaction is separated from the liquor by filter and washed first with deionized water and then with acetone, and then dried at 50°C to yield the final product of Cu-Ag double metal powder. The TEM measurement shows that the product prepared with the procedure described above consists of particles sized 5 to 30 nm and their clumps. The TEM photo is shown in Fig. 14.6. In order to understand the stability of the experimental product, the appearance of the product is observed visually for one month in storage, and no change in its color, a sensitive parameter for oxidation, has been found, suggesting that the product has antioxidation capability at room temperature. Table 14.4 gives a comparison between the properties of the Cu-Ag double metal powders prepared with various technologies and/or devices. It is clear from Table 14.4 that in both the mean size and the anti-oxidation nature, among the five Cu-Ag double metal powders listed in this table, the one prepared with the nano copper powder produced in the SCISR is the best.
PREPARATION OF NANO COPPER AND ITS SURFACE IMPROVEMENT
299
t
lOOm h ::.. ................:.:::.............I:
L
Figure 14.6 TEM photo of Cu-Ag double metal powder.
Table 14.4 Comparison between Cu-Ag double metal powders prepared with various technologies and/or devices Reducing agent and/or equipments
Average size of nano Cu
Ag amount Average size of covered
for nano copper
Cu-Ag particles
%mass
Anti-oxidation
Ref.
nature at room temperature
NaH2PO2
~ 50 nm
22-85
> 100 nm
yes
[206]
N2H4.H20
50-- 500 nm
--
> 300 nm
yes
[200]
_>500 nm
- 38.84
> 500nm
yes
[199]
STR
~ 45 ~tm
- 15.22
- 45 ~tm
yes
[207]
SCISR
-- 10 nm
- 30
5-30 nm
yes
[ 196]
Ascorbic acid
KBH4
14.6 CONCLUSIONS T h e p r e p a r a t i o n of nano c o p p e r p o w d e r by reduction-precipitation in the s u b m e r g e d circulative i m p i n g i n g stream reactor ( S C I S R ) with KBH4 as the r e d u c i n g agent, strong aqua a m m o n i a as the c o m p l e x a n t , and P V P as the dispersing agent, and the surface
300
IMPINGING STREAMS
improvement of the nano copper powder so prepared for anti-oxidation were investigated experimentally. The following can be concluded: (1) The optimal conditions determined for the preparation of nano copper by reduction-precipitation are: the concentration of CuCI2 solution Ccuc~2 = 0.2 kmol-m-3; the rotary speed of the propellers N ~ 800 rpm; the reaction temperature T = 20°C; the pH of the reaction solution pH = 14; the amount of 2% surfactant PVP addition 40 mL; the mole ratio CuClz:KBH4 = 1:2; and the feeding mode I, i.e., CuCI2 solution is added quickly to KBH4. (2) The repeated experimental operation under the optimal conditions described in Item (1) can yield a stable nano copper product average-sized 5 to 10 nm with a narrow size distribution and a high content of the simple substance copper. (3) The pH of the reaction solution has an important influence on the properties and appearance of the nano copper product; the escape of ammonia in the form of bubbles during and after the reaction favors the dispersity of the particles; the rotary speed of the propellers has no significant influence on the particle size in the product in the range of speed tested. (4) With the nano copper produced in the SCISR the Cu-Ag double metal powder is prepared by the replacement reaction with AgNO3 as the reactant. The TEMdetermined results show that the product of Cu-Ag double metal powder prepared consists of particles sized 5 to 30 nm and their clumps; after storage for one month at room temperature no change in the color of the Cu-Ag double metal powder has been found, suggesting the product has anti-oxidation nature at room temperature. (5) In comparison with the nano copper powders prepared with various processes and/or devices, the products prepared in the present investigation are finer with a narrower size distribution, showing that the SCISR with the features of efficient micromixing and very strong pressure fluctuation and the special flow configuration of perfect mixing flow-plug flow in series is very suitable for the preparation of ultrafine powders by reaction-precipitation and, in fact, exhibits good performance in the present investigation.
-15PREPARATION OF ULTRAFINE POWDERS BY REACTION-PRECIPITATION IN IMPINGING STREAMS II1: NANO TITANIA
15.1 PROPERTIES OF NANO TITANIA AND CHEMICAL REACTIONS IN ITS PREPARATION Titanium dioxide (Titania) is a white powder which is inert, insoluble in water, organic and weak inorganic acids, while being slightly soluble in alkali and soluble in saturated potassium acid carbonate. It can be completely dissolved in strong sulfuric acid and hydrofluoric acid after boiling for a long time. It is thermo-stable and melts gradually at temperatures over 1800°C [208]. Titania has two main stable crystalline forms: titanic and anatase. Titanic is the most stable one, and anatase converts into titanic form at high temperatures over 915°C. From the thermodynamic data, the formation heat of anatase is 8-12 kJ-mol -~, and so it should be more stable than titanic [209]. On the other hand, the TiO2 of anatase has lower hardness, and is more suitable for use in cosmetics. Nano Titania is one of the earliest nano materials to be applied commercially. It has a number of superior properties, such as super strong scattering and anti-ultraviolet capabilities, special electromagnetism and catalysis characteristics, especially the photocatalysis ability decomposing microbes, and also extremely high surface activity [210]" and can be used as active ingredients in high-grade coatings, anti-ultraviolet cosmetics, hygiene ceramics, self-cleaning glasses, composite polymer materials, photoelectric cells, electronic ceramics, semiconductors, catalysts, etc. The application of nano Titania is still in the stage of initial development and a number of possible applications of great potential have not yet been put into practice, e.g., its use as a catalysis-active component, etc. As for other inorganic nano materials, a number of preparation methods have been proposed for nano Titania. Because of its economic advantages and its simplicity, Li [211] employed the process of TIC14 hydrolyzation-precipitation in impinging streams. In practice, the chemical process whereby TIC14 hydrolyzation yields TiO2 precipitate is very complex [212]. Since it is not the major topic of this book, the chemical reaction mechanism will not be discussed in detail here. Put simply, under
301
302
IMPINGING STREAMS
controlled conditions TIC14 is converted by hydrolyzation-ionization-hydrolyzation into TiO2 precipitate [213], as follows.
Hydrolyzation of TIC14: TIC14 + H20 ~ TiOH 3÷+ H + + 4C1-
(15.1)
Ionization of the middle product: TiOH3+___~ TiO 2++ H +
(15.2)
Hydrolyzation of the secondary ion: TiO 2++ H20 ---* TiO2 + 2H +
(15.3)
where Reaction (15.1) is a rapid reaction. When a higher concentration of TIC14 solution is added, the hydrogen ions from Reaction (15.1) may restrain the desired reactions, Reactions (15.2) and (15.3), to yield a limpid solution containing TiOH 3+. With the addition of an ammonium sulfate solution containing hydrochloric acid, at higher concentration the sulfate ion can combine with the TiO 2+ ion to produce TiOSO4 precipitate, promoting Reaction (15.2). At higher temperature, the solubility of TiOSO4 increases, and so the titanium is mainly in the form of TiO 2+' while when the temperature reaches 95°C, the precipitation, Reaction (15.2), occurs. It is clear that all three reactions, Reactions (15.1) to (15.3), produce hydrogen ions. To avoid pH variation too quickly resulting in a non uniform composition of the precipitate, it is necessary to add a certain amount of aqua ammonia to form a buffer solution so that the rate of titanium oxide formation is kept suitable for the precipitation of TiO2. In addition, to set a period of time for keeping warm after the reaction operation promotes the formation and growth of the TiO2 nucleus. Actually, during the reactions, the meta-titanic acid HzTiO3 is precipitated first, which adsorbs considerably large amounts of water and soluble impurities such as ammonium sulfate, ammonium chloride etc, and also contains some insoluble impurity particles such as oxides of iron, strontium, and calcium etc. Therefore, it is essential to fully wash the precipitate to obtain a product of high purity. The insoluble and part of the soluble impurities can be removed by multiple washings; ammonium chloride and ammonium sulfate are decomposed and released as the precipitate is later heated to 300°C and 700°C, respectively, during the calcination. The meta-titanic acid obtained directly by hydrolyzation is amorphous, with extremely unstable surface properties. For convenience of storage and application, it is usually calcined to convert it into anatase or titanic powder. During calcination, the primary particles are formed gradually as dewatering. The growth of TiO2 particles during calcination are usually in the two schemes: isothermal growth and agglomerative growth, the latter yielding larger particles outside the scope of nano materials. For the preparation of nano titanium oxide, the fully washed precipitate is usually calcined at a constant temperature, and the temperature of calcination must be controlled to avoid agglomeration, or, at least, to reduce agglomeration as much as possible. The calcination temperature for anatase is round 600°C, while for titanic crystalline is
PREPARATION OF NANO TITANIA
303
800°C. During calcination the particles grow by themselves, but the growth rate is smoothed after achieving certain sizes; and the crystalline form may also be partially changed. For example, some of the titanic crystalline may also appear in the product calcined at 600 °C for a long time.
15.2 EXPERIMENTAL EQUIPMENT AND PROCEDURE The experimental equipment system for the preparation of nano titanium oxide by TIC14 hydrolyzation-precipitation is the same as that shown in Fig. 14.2 in Chapter 14, and the submerged circulative impinging stream reactor (SCISR) is also made of titanium material for anti corrosion of C12 and CI-, with an effective volume of 3.6x 10 -~ m ~. As with the preparation of nano copper powder, the make-up of the solutions of the reactants TIC14 and (NH4)2SO4 is an important link in the preparation of nano titanium oxide. Naturally, TC14 is very easily hydrolyzed and so cannot be fed directly into the reactor. Otherwise, TIC14 would be quickly hydrolyzed and produce a large amount of precipitate and no nano product would be obtained. To solve this problem, TIC14 is first dissolved in an aqua solution of hydrochloric acid to obtain the aqua TiC14hydrochloric acid solution in order to keep the concentration of Ti 4+ high enough at the beginning of reaction. The make-up of the solution is carried out in a flask at low temperature. The flask is put into an ice-water bath, and is filled with a certain amount of deionized water and strong hydrochloric acid, and then, under stirring, a certain amount of TIC14 is slowly added from a funnel into the flask to yield a transparent TiC14-hydrochloric acid solution slight yellow in color. Another process solution is that of ammonium sulfate. The dissolution of (NH4)2SO4 is endothermal, and its solubility and dissolving rate are greatly affected by temperature, so that the use of hot water is needed. First put a certain amount of (NH4)2SO4, and then hot water into an ordinary beaker, quickly agitate the mixture to speed up dissolution of (NH4)2SO4, and a limpid, transparent solution is obtained. The experimental procedure is as the following: Put a certain amount of deionized water into the SCISR; turn on the propellers to push towards the water circulation and to make the opposing streams impinge against each other inside the SCISR; adjust the rotary speed of the propellers to control the impinging velocity; let the cold water passing through the heat exchanging jacket of the SCISR cool the contents in the reactor to keep the temperature inside the reactor constant at 2 to 3°(7. When both the flow status and the temperature in the reactor are stable, slowly add the titanium chloride solution into the reactor and then drip the aqua ammonium sulfate solution into the reactor. During mixing between the reactants the temperature is controlled to be higher than 15 °C" the initial concentration calculated after mixing is 1.0 to 1.2x 10-~ mol.m -~" while the mole ratios Tia+/H+- 15 and Ti4+/SO42- - 2/1. When full mixing is achieved, let the heating medium pass through the jacket of the SCISR to enhance the temperature inside the reactor to a certain level and keep it for a certain time for the
304
IMPINGING STREAMS
hydrolyzation-precipitation reaction; add strong aqua ammonia to adjust the pH to the given value. Then cool the reacted suspension to room temperature by cold water circulating through the jacket; leave for 12 h for ageing; separate the precipitate from the liquor by filter; wash the cake with deionized water to remove C1- ions (check with 0.1M AgNO3 solution), wash with alcohol another three times and then dry the washed precipitate at room temperature to yield the reaction product. Finally, calcine the dried reaction product at the given temperature for 3 h to get the nano Titania product. In the experimental investigation the influences of the following factors were examined: (1) (2) (3) (4) (5) (6)
The The The The The The
concentration of TIC14 solution and pH of the reaction mixture; impinging velocity (by adjusting the rotary speed of the propellers); reaction or mixing time and temperature; feeding time; feeding position; and keeping warm (ageing) time after reaction.
The experiments for the preparation of nano TiO2 were carried out in three stages as described below. In the first stage the major factors affecting the size of product were examined to determine the primary optimal operation conditions. In the second stage, on the basis of the determination of the primary optimal conditions, the influences of certain factors on both the particle size and the yield of titanium were examined for further optimization of the conditions. For arrangement of the experiments carried out in both the first and the second stages the normal design technique was employed. In the third stage the experiments were carried out for mass-preparation of nano Titania with commercial TIC14 as the raw material, instead of chemical reagents, and under the conditions optimized in the last stage. Also, some supplemental studies were made at this stage, including examination of the influence of the neutralization time (rate) with ammonia on particle size; while part of the experiments organized with the uniform design technique were for further optimizing the conditions and for examining the influence of calcination temperature on the particle size. Finally, the preparation experiments were also carried out with the traditional stirred tank reactor (STR) as the hydrolyzation- precipitation equipment for comparison.
15.3 RESULTS AND DISCUSSIONS 15.3.1 Major results obtained in the first stage The first stage of the investigation is aimed at the preparation of the product of smallest-sized particles, and the normal design technique is employed for arranging the experiments. The normal-designed operation conditions are listed in Table 15.1.
PREP/XRATION OF NANO TITANIA
305 Table 15.1
Normal-designed experimental conditions in the first stage Effecting factor*
Level
A
B
C**
D
1 1.0 600 I 60 2 1.1 900 II 120 3 1.2 1200 III 180 * A--concentration of TIC14, kmol.m -3" B--rotary speed of propellers, rpm; C--feeding point; D--feeding time, s ** I--inlet of drawing tube; II--outlet of drawing tube; III--center of reactor
Table 15.2
Results of normal-designed experiments in the first stage A
B
C
D
1
2
3
4
1
1
1
1
1
15.67
2
1
2
2
2
16.09
3
1
3
3
3
18.02
4
2
1
2
3
17.28
5
2
2
3
1
17.78
6
2
3
1
2
15.44
7
3
1
3
2
13.70
8
3
2
1
3
10.16
9
3
3
2
1
12.32
Ki
49.78
46.65
41.27
45.77
K2
50.5
44.03
45.69
45.23
K3
36.18
45.78
49.50
45.46
Ki
16.59
15.55
13.76
15.26
K~
16.83
14.68
15.23
15.08
K~
12.06
15.26
16.50
15.15
R
4.77
0.87
2.74
0.18
Run No
_
Average size, nm
T=KI+K2+ K3 =136.46
T/9= 15.16
306
IMPINGING STREAMS
The experiments are carried out following the procedure described in the previous section and the washed and dried reaction products are calcined at 600 °C for 3 h. The results are listed in Table 15.2, where the data for the average sizes listed in the sixth column are calculated from the X-ray spectrum; the X-ray spectrum shows that the products are of the crystalline form of anatase. The following can be seen from the data listed in Table 15.2: (1) the influencing significances of various factors tested on the average size of product are in the order of A > C > B > D; (2) the set of operation conditions yielding the smallest average size is A3B2C1D3, i.e., the initial concentration of TIC14 is 1.2 kmol.m -3, the rotary speed of the propellers N = 900 rpm, the feeding position at the inlet of the drawing tube, and the feeding time 180 s. The product prepared under these conditions is sized 10.16 nm. For more detailed results one may refer to Ref. [211 ]. The fact that the initial concentration of TIC14 most significantly affects the average size is easily understood: the higher concentration enhances the reaction rate to produce higher supersaturation resulting in nucleation in large amounts to yield a finer-sized product. However, the initial concentration of 1.2 kmol.m -3 cannot be confirmed to be optimal as it is the limit of the range tested, i.e., the highest concentration. The rotary speed of the propellers exhibits a turning influence on the average size of the product; this is similar to the results obtained by Chen et al. [165] in their investigation on an analogous problem in a stirred tank reactor and also similar to the results on the preparation of nano copper described in the previous chapter. It results from the mutual effect between macro- and micro-mixing, as mentioned before. The experimental results on the influence of feeding position are consistent with those obtained in the investigation on the preparation of "ultrafine" white carbon black, i.e., the inlet of the drawing tube is the best position and this can be considered as a general regularity for the SCISR being used for the preparation of ultrafine particles by reaction-precipitation and is consistent with the initial design idea. The fresh feed passes first through the drawing tube where essentially no mixing and, consequently, no reaction take place so that it keeps its initial composition of high reactant concentration. Such a feed stream enters the impingement zone and suddenly undergoes strong micromixing and thus a very rapid reaction, favoring the creation of a high and uniform supersaturation, thus promoting nucleation in huge quantity. The feeding time has no significant influence on the main size of the product. The major reaction producing meta-titanic acid precipitate occurs in the heated reaction mixture while the TIC14 solution is fed completely before heating of the mixture. These facts may account for the phenomena described above.
15.3.2 Experiments and major results in the second stage On the basis of primary determination of the optimal conditions made in the first stage, the experiments are carried out for further optimization of the operation conditions in the second stage of the investigation, and also the influences of the reaction
PREPARATION OF NANO TITANIA
307
temperature, the reaction (with keeping warm) time and the pH of the reaction mixture on both the average size and the yield of titanium are examined at this stage. The yield of titanium is calculated as the ratio of the amount of titanium in the product to that in the fed solution. The other fixed conditions are: the rotary speed of the propellers N 900 rpm, the initial concentration of TIC14 1.2 kmol.m -3. The normal design technique is still employed, and the experimental conditions so designed are listed in Table 15.3. T a b l e 15.3
Normal designed experimental conditions in the second stage Effecting factor* Level A
B
C
1
75
1.0
5.0
2
85
1.5
6.0
3
95
2.0
7.0
* A--temperature, °C; B - reaction (heating) time, s C--pH of the reaction mixture The procedure for the treatment of reaction product is the same as for the first stage. The results obtained are given in Table 15.4 (over page). The data listed in Table 15.4 illustrate the following: (1) the influencing significances of various factors tested on the average size of product are in the order of C > B > A; (2) the set of operation conditions yielding the smallest average size is A~B~C~, i.e., the reaction temperature 75°C, the reaction (keeping warm) time 1 h, and the pH of the reaction mixture 5.0. The operation under these conditions yields a product consisting of particles average-sized 7.33 nm, and the yield of titanium is as high as 98.3%. Obviously, both results are very good. According to the general principles of dissolution-crystallization, the temperature (Factor A) has contradictory influences on the mean size of the product. At higher temperature, the reaction goes taster to produce larger amounts of the substance to be precipitated, favoring enhanced supersaturation, on one hand; while, at the same time, the solubility of the substance under consideration may increase, yielding a negative influence on the supersaturation, on the other. The overall tendency of the mean size variation depends on the balance between the two contradictory factors. From the experimental results, the reaction at 75°C produces the product with the smallest-sized particles and also with a higher yield of Ti. However, because 75°C is the lowest of the temperatures tested, it cannot be verified whether an even lower temperature would be better and so further study may be needed. According to the results obtained in similar investigations with other reactors, the operation temperature determined for the hydrolyzation-precipitation was mostly higher than 75°C. Therefore this temperature can be considered to be, at least, reasonable and feasible.
308
IMPINGING STREAMS Table 15.4
Results of normal-designed experiments in the second stage A
B
C
1
2
3
1
1
1
2
1
3
Run No
Yield of Ti, %
Average size, nm
1
98.3
7.33
2
2
90.1
13.94
1
3
3
75.4
8.5
4
2
1
2
87.0
11.33
5
2
2
3
89.4
11.05
6
2
3
1
93.6
14.85
7
3
1
3
88.2
15.22
8
3
2
1
92.6
12.1
9
3
3
2
92.9
10.67
Size nm
Yield %
K1
9.92
11.56
11.43
K2
12.41
12.36
11.98
12.66
11.34
11.59
Rs
2.74
1.02
0.55
Kl
87.93
91.17
94.93
K2
90.01
90.70
90.00
K3
91.23
87.30
84.30
3.30
3.87
10.63
Rv
Tsize =
104.9
Tsize/9 = 11.7
Tyield -807.5 Tyield/9 = 89.72
The reaction (keeping warm) time (Factor B) exhibits little influence on the mean size of the product, but significantly affects the yield of titanium. However the direction of the influence is somewhat unexpected: the longer the reaction time, the lower is the yield of Ti. This indicates that re-dissolution of the precipitate obviously occurred, but, as yet, the phenomenon is difficult to explain exactly or reasonably and further investigation is needed. After all, the reaction mixture is very complex. The pH of the reaction mixture (Factor C) has the most significant effect on the yield of titanium, while exhibiting abnormal behavior: pH the yield of Ti decreases as pH increasing. Wu et al. [214] also observed similar phenomena in their investigation.
PREPARATION OF NANO TITANIA
309
The most possible reason is that the molecules of ammonia have contradictory influences on the nucleation. Wu et al. considered that, to an extent, the increase in the concentration of ammonia can speed up the peptizing; however, at excessively high concentration the molecules of ammonia may combine with the Ti 4+ ions to form a complex, causing H2TiO3 to be hydrolyzed resulting in a low yield of Ti. Also, pH has a certain influence on the mean size of the product: the higher the pH value, the greater is the mean size of the product. This is not difficult to understand: if the ammonia leads to hydrolyzation of H2TiO3, the finer particles must be hydrolyzed first, leaving the larger ones. Summarizing the results obtained in the first and second stages, the optimal operation conditions can be preliminarily determined as follows: the initial concentration of TIC14 1.2 kmol.m -~, the feeding position at the inlet of the drawing tube, the reaction temperature after feeding 75°C, the pH of the reaction mixture pH = 5, the rotary speed of the propellers N = 900 rpm, the reaction (keeping warm) time 1 h.
15.3.3 Experiments of mass preparation and the results In order to examine the operational conditions determined and to make the results more feasible for industrial application, and also to provide sample product for further application testing, experiments for mass-preparation are carried out under the optimal conditions determined above, following the procedure described in Section 15.2 and with commercial TIC14 as the raw material, the composition of which is indicated in Table 15.5. Table 15.5
Composition of the commercial TIC14 used
Component Content
TIC14 %
Fe ppm
Sr ppm
Ca ppm
Na ppm
K ppm
Mg ppm
96-99.9
6.76
6.61
4.73
<1.0
The products from the multiply repeated operations were fully mixed; the major part of the mixed product was calcined at 600 °C for 6 hours to yield the sample product for further application testing, while the surplus was later used for investigating the influence of the calcination temperature. The results on the particle sizes in the product calcined at 600°C measured by TEM are listed in Table 15.6, while those calcined at various temperatures are given in Table 15.7. The data listed in Table 15.6 show that the nano Titania product prepared by TIC14 hydrolyzation-precipitation in the SCISR had very small sizes and a very narrow size distribution, the weighted mean size of which was calculated to be 9.64 nm.
310
IMPINGING STREAMS Table 15.6
Sizes of the sample product calcined at 600°C Size, nm
Percentage, %
--5
10
--10
88
~17
2 Average size: 9.64 nm
Table 15.7
Average sizes of nano Titania products calcined at different temperatures
Calcination temperature, °C
400
600
800
Average size, nm
5.47
8.84
26.84
The results for the crystalline forms of the products from the analysis of X-ray spectrum (which is not given here) indicate that the products calcined at 600 and 800°C for two hours are of the anatase and the titanic crystalline forms, respectively. The data given in Table 15.7 indicate that, as the calcination temperature increases, the variation of crystalline form is accompanied by crystal lattice growth. In particular, the size of the product calcined at 800°C is significantly enlarged: the average size is larger than that of the product calcined at 600°C by about 20 nm. The TEM photos of the products calcined at 600 and 800°C are shown in Figs. 15.1 and 15.2. As can be seen from the discussions above, the sizes of the nano Titania products are affected most significantly by the pH of the reaction mixture. On the other hand, from the point of view of chemistry, the shift of titanium from the form of ion to complex-ion to precipitate is related closely to the pH. In order to examine the possibility of controlling the size of the product by NH3- neutralization operation and to obtain some information for understanding the reaction mechanism, the influence of the neutralization rate on the mean size and the regularity of pH variation during neutralization are studied experimentally. The results of the influence of the neutralization rate on the mean size of the product are shown in Fig. 15.3, while the variation of pH with an amount of ammonia added-in is given in Fig. 15.4.
PREPARATION OF NANO TITANIA
311
Figure 15.1 TEM photo of the nano TiO2 calcined at 600°C.
100 nm I
I
Figure 15.2 TEM photo of the nano TiO2 calcined at 800°C.
12
©
E 11 ~
10
>
9
©
<
200
©
400 600 800 Time for NH addition, s
Figure 15.3 Relationship between average size of product and NH3 addition time.
312
IMPINGING STREAMS
15.3.4 Experiments of neutralization with aqua ammonia As can be seen in Fig. 15.3, the influence of the neutralization rate on the mean size of the product is not routine. There may be two reasons accounting for the phenomena: One is that the addition rate of aqua ammonia may reduce the complicated effects on the mixing condition. As described in Chapter 10, the impinging streams promote micromixing very efficiently, but the occurrence of micromixing is limited to the impingement zone only, while in the major part, about 80%, of the space in the SCISR essentially no mixing occurs. Therefore it is indeed possible that the addition rate of aqua ammonia affects the variation of the mixed composition with time in the impingement zone, resulting in changed conditions for the precipitation. Another possible reason is that the behavior of the titanic ions, the complex-ions and the metatitanic acid are related to the pH of the reaction mixture, while the concentrations of the three substances have contradictory influences on the mean size, as mentioned above. Therefore the increase in the addition rate of aqua ammonia leads to the saddle-shaped variation of the mean size. The results shown in Fig. 15.4 are very interesting: like the titration curve in the volumetric analysis, there is a mutation point on the pH curve of the reaction mixture versus the amount of aqua ammonia added. This is opposite to the property of a buffer solution. Although it is difficult to precisely explain the phenomena illustrated in Figs. 15.3 and 15.4 at present, the information contained in these figures is helpful for further analysis and investigation on the mechanism and regularities of the reactionprecipitation process, and for searching for a possible method for size-control of the product.
[] [] I
[]
[] I
[] I [] [] [] []
:~ 3
/
[]
0
[]
200
400
/
F-1o[]
[]
,, [ ]
600
800
1000
Amount of NH 3 added, mL Figure 15.4 Variation of pH of the reaction mixture during neutralization.
PREPARATION OF NANO TITANIA
313
15.3.5 Experiments for final optimization of conditions and the results As mentioned above, some of the primary optimal conditions determined in the first and second stages cannot yet be confirmed as really optimal, because they are the boundary values tested. In order to optimize these conditions further, some supplementary experiments were carried out in wider ranges to examine the influences of the pH, the reaction temperature, and the concentration of TIC14 solution on the mean size of the product, for which the uniform design technique was employed. The experimental conditions and the results are listed in Table 15.8, where the values for the average sizes of particles are calculated from the X-ray spectrums. Table 15.8
Conditions and results of experiments uniform-designed
Level
pH
Temperature °C
1
4
60
2
4
3
5
4
5
70
Concentration of TIC14, kmol-m-~
Average size, nm
Yield of Ti, %
1.5
7.34
96.15
80
1.4
5.07
95.12
50
1.3
8.41
93.57
1.2
8.84
95.12
A non-linear regression analysis of the data listed in the table is made by Li [211]. By comparison between the regression coefficients the following are concluded: (1) The pH has the most significant effect on the mean size, the influence of the TIC14 concentration takes second place, while the temperature has little effect; and (2) Similarly, the pH and the TIC14 concentration exhibit the most significant influences on the yield of titanium. From the point of view of the influences on both the mean size and the yield of titanium, it seems that the pH could be lower; however, excessively low pH may aggravate corrosion of the equipment, while an excessively high concentration of TIC14 would result in difficulties of operation. Combining these considerations, the optimal values finally determined for these three parameters are as follows: pH of the reaction mixture: pH - 4; Concentration of TIC14 solution" 1.4 kmol.m-S; and Reaction temperature: 80°C.
15.3.6 Comparative experiments between SCISR and STR and the results In order to further verify the superiority of the SCISR in the preparation of ultrafine powders by reaction-precipitation, a set of comparative experiments were carried out
314
IMPINGING STREAMS
with the SCISR and STR as the reaction-precipitation devices. The SCISR is that used in the experiments above; and the STR has the same structure and the same effective volume as that used in the investigation described in Chapter 13 (with the effective volume of 0.6x10 -3 m3). All the operation conditions for both the SCISR and STR are the same, i.e., the optimal ones determined in the previous section; also the specific effective power inputs to the two reactors are the same. The results of the comparative experiments are indicated in Table 15.9. It is clear that the product from the SCISR has finer sizes with a narrower size distribution than that from the STR, and the yield of Titanium obtained in the SCISR is also higher than that in the STR. Table 15.9
Results of comparative experiments Size of TiO2 particles, nm Reactor type
Yield of Ti Maximum
Minimum
Average
SCISR
0.962
16.0
2.0
5.68
STR
0.927
20.85
3.34
11.28
15.4 CONCLUSIONS An investigation on the preparation of nano Titania by TiC14-hydrorizationprecipitation was carried out with the submerged circulative impinging stream reactor (SCISR) as the reaction device and the results compared with those obtained with the traditional stirred tank reactor (STR). The following can be concluded: (1) With the goal of obtaining product with the smallest-sized particles, the following optimal operation conditions are determined: the concentration of TIC14 solution 1.4 kmol.m -3, the rotary speed of the propellers 600 to 1200 rpm (corresponding to the impinging velocity u0 - 0.24 to 0.27 m.s-~), the reaction temperature 80°C, the reaction (keeping warm) time after reaction 1 hour, the terminal point of neutralization with aqua ammonia pH = 4. (2) Under the conditions of the TIC14 concentration 1.2 kmol.m-3, pH - 5, the reaction temperature 75°C, and the optimized other conditions, the sample product calcined at 600°C of over 2 kg is prepared with commercial TIC14 as the raw material by multiply repeated operations. The results measured by TEM show the product being average-sized 9.64 nm with considerably narrower size distribution.
PREPARATION OF NANO TITANIA
315
(3) The data on the neutralization with aqua ammonia show that the influences of the addition rate of ammonia on the mean size is non humdrum (see Fig.15.3); and there is a mutation point on the curve of pH of the reaction mixture versus the amount of aqua ammonia added (Fig. 15.4). (4) The results of investigation on the influence of calcination temperature indicate that, as the calcination temperature increases, the variation of the crystalline form of TiO2 is accompanied by significant growth of particles in the product. The product calcined at 600°C is of the anatase form; while that calcined at 800°C is of the titanic crystalline form, with a grown average size greater than that calcined at 600°C by about 20 nm. The results of the comparative experiments show that, under the same operating conditions and the same specific effective power input, the product prepared with the SCISR is finer with a narrower size distribution than that prepared with the STR.
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-16PREPARATION OF ULTRAFINE POWDERS BY REACTION-PRECIPITATION IN IMPINGING STREAMS IV: NANO HYDROXYAPATITE
16.1 INTRODUCTION The replacement of some substances in the human body by manually synthesized materials is an important field of investigation in both medical and material sciences. Hydroxyapatite (HAP), the molecular formula of which is Cas(PO4)3(OH) or Cal0(PO4)dOH)2, is the major inorganic constituent in bone, teeth, etc. in the human body. HAP has essentially the same chemical composition and crystalline structure as those of human bone and so has good bio-compatibility. For a long time, it has been widely used as a sclerotin material in setting broken bone, filling teeth, etc. [215]. In addition, HAP can also be used as a food additive and moisture-sensitive element, etc. Because of its important application value, much research and development has been devoted to the processes for synthesis of hydroxyapatite; since the 1950s not only HAP monocrystals of high purity [216] but also carbonate-containing HAP [217], which has a composition extremely close to that of human bone, and nitrogencontaining HAPs of high stability have been synthesized. Furthermore, the polycrystal ceramics HAP with a strength and tenacity very close to those of teeth has also been prepared by compact cementation technology [218]. Because of their excellent biocompatibility with the human body, these kinds of hydroxyapatites have been widely used as synthesized bone and organization materials [219, 220]. A number of methods have been proposed for the synthesis of HAP, which can be classified into two types [221 ]: solid phase processes and aqua phase ones. One of the solid phase processes is dry synthesis, i.e., by direct calcination of mixed Ca3(P04)2 and Ca(OH)2: 3Ca3(PO4)2 + Ca(OH)2 = Cal0(OH)2(PO4)6
(16.1)
This reaction requires very high temperature and the conversion that can be achieved is relatively low. The improved solid phase process is hydrothermal synthesis, in which the mixture of Ca3(PO4)2 and Ca(OH)2 or CaCO3 is calcined under the condition of steam passing through the material bed [215,222], and the following reactions occur:
317
318
IMPINGING STREAMS CaCO3 (or Ca(OH)2) = CaO
(or H20) (g)
(16.2)
CaO +3Ca3(PO4)2 +H20 (g) = Cal0(OH)2(PO4)6
(16.3)
+ CO 2
The chemical reactions are carried out in the solid phase, although diffusion from gas to solid and that in the opposite direction are involved. The conversion of Reaction (16.3) can be 0.8 to 0.85 and so a purification procedure is needed for the reacted material in order to achieve a product of high purity. There are various processes for synthesizing hydroxyapatite in aqua phase, such as hydrolyzation [223], acid-alkali reaction [224], hydrothermal process [225], coprecipitation [226], sol-gel synthesis [227], aerosol pyrolysis [228], and micro emulsion synthesis [229], etc. As mentioned in Chapter 13, ultrafine materials have a number of excellent characteristics; therefore it is of interest to prepare HAP sized at the nano level. Some of the methods mentioned above can yield ultrafine particles of HAP, and progress has been made in the control of particle size and coalescence of the products; however, problems relating to particle appearance, dispersion and size distribution etc. have not yet been completely solved. Employing the submerged circulative impinging stream reactor (SCISR) Yuan et al. [230] preliminarily investigated the synthesis of nano hydroxyapatite, and prepared a very good product. Similarly, in order to make use of the superiorities of the SCISR's efficient micromixing and strong pressure fluctuation, an aqua phase process, the double decomposition-precipitation, was used for the synthesis, and (NH4)zHPO4 and Ca(NO3)2 used as the reactants and the sources of the calcium and the phosphate. The overall chemical reaction involved in the double decomposition precipitation process may be presented as follows: 6( NH4)2HPO 4 +10Ca(NO3) 2 +8NH 3 + 2 H 2 0 (16.4) = Cal0(OH)2(PO4) 3 ,~ +20NH4NO 3 Actually, the mechanism for the formation of HAP crystalline during the process under consideration is very complex and some suggestions proposed for the mechanism cannot be supported by experimental evidence so details of the chemical aspect will not be discussed here. In any case, precipitation, i.e., crystallization from a solution, must be involved in the process so the fact that liquid-continuous impinging streams cause strong micromixing and pressure fluctuation would favor the synthesis of nano HAP.
16.2 EXPERIMENTAL EQUIPMENT AND PROCEDURE 16.2.1 Equipment The structure and dimensions of the SCISR used for the synthesis of nano hydroxyapatite by double decomposition-precipitation are the same as those for the
PREPARATION OF NANO HYDROXYAPATITE
319
preparation of nano Titania and nano copper powder, while it is made of stainless steel 316L. For details one may refer to Fig. 10.2 in Chapter 10.
16.2.2 Procedure of the experimental operation The experimental procedure is as follows: the aqua solution of calcium nitrate with a concentration of 0.28 mol.L -~ is first fed into the SCISR, then, under stirring, the aqua solution of di-ammonium phosphate of 1.07 mol.L -~ is dripped into the reactor at room temperature. The amounts of Ca(NO~)2 and (NH4)2HPO4 are essentially in the stoichiometric ratio. The pH of the mixed solution is adjusted to 10.5-12.5 with aqua ammonia. Once the dripping of the (NH~)2HPO4 solution is accomplished, keep the propellers on for 10 min for the reaction to continue, then transfer the whole mixture into a three-neck flask with agitator: heat the mixture under the conditions of back-flow and stir for 3 to 5 hours; leave the mixture for ageing for 24 hours; then separate the precipitate from the liquor by a high-speed cooling centrifuge, wash first with water and then with acetone three times, respectively. Finally dry the washed precipitate under vacuum at 30°C to yield the final product. The products obtained are determined by the energy spectrum for the compositions, mainly for the Ca/P mole ratio, and characterized by infrared spectroscopy with the Fourier transformation intra-red spectrophotometer (FTIR) of Type Nicolet 510P made by Nicolet Co., thermal analysis on a thermo- gravimetric/differential thermal analyzer (TG/DTA) of Type ZRY-2P, X-ray diffraction (XRD) analysis with the X-ray diffractometer of Type XD-5 made by Shimadzu Co., scanning electron microscopy (SEM), and transmission electron microscopy (TEM) with the transmission electron mirror microscope of Type JEM-100SX type made by JEOL Co. In the preliminary investigation eight sample products of nano HAP were synthesized under various conditions, as listed in Table 1. Table 16.1
Experimental conditions for synthesis of the sample products Sample No
pH
Dripping rate of (NH4)2HPO4, mL.min_ 1
Back-flow time, h
HAP 1
12.5
54
0
HAP2
12.5
54
3
HAP3
12.5
54
4
HAP 4
12.5
54
5
HAP 5
12.5
96
3
HAP 6
12.5
24
3
HAP 7
11.5
54
3
HAP 8
10.5
54
3
320
IMPINGING STREAMS
16.3 RESULTS AND DISCUSSIONS 16.3.1 Influences of some factors For nano HAP product, its size and distribution and appearance of the particles are the major quantity indexes. In addition, chemically the mole ratio of Ca/P essentially characterizes the purity of the product. All these are taken as the criteria for comparison and analysis of the experimental results.
16.3.1.1 Influence of the back-flow time The back-flow time is set for ageing of the precipitate under the conditions of stirring and heating or keeping warm. The Ca/P mole ratios in HAP products obtained at various back-flow times are listed in Table 16.2. It is clear that the Ca/P mole ratio in the HAP product simply increases as the back-flow time increases in the range of the experimental operation conditions tested, suggesting that a certain ageing time is necessary for the shift from calcium phosphate to HAP; while too long a back-flow time, e.g., 5 hours, results in decreased purity of the product, Ca/P ratio up to 1.783, suggesting a certain amount of Ca(OH)2 is precipitated into the solid phase. The most likely reason for the variation of Ca/P ratio during back-flow from 4 to 5 hours may be that part of the ammonia escaped unavoidably from the reaction mixture, resulting in decreased pH of the mixture. From the data listed in Table 16.2, the conditions yielding the sample products HAP 1 and HAP4 can be considered as unfeasible from the point of view of product purity. Table 16.2
Influences on Ca/P ratios at different overflowing time Sample No Back-flow time, h Ca/P mole ratio
HAP 1
HAP2
HAP3
HAP4
0
3
4
5
1.56
1.601
1.681
1.783
The SEM images of the same sample products involved in Table 16.2 are shown in Fig. 16.2. The crystals in the sample product HAP1 seem no good and are seriously clumped together, as shown in Fig. 16.2(a), while the interface of the sample product HAP4 shown in Fig. 16.2(d) is vague. So, the images in Fig. 16.2 also indicate that the conditions yielding the sample products HAP1 and HAP4 are unfeasible. The sample products HAP2 and HAP3 do not exhibit significant difference, as can be seen from their images in Fig. 16.2(b) and (c), implying that the crystalline form of HAP essentially undergoes no change during back-flow in the range of 3 to 4 hours.
PREPARATION OF NANO HYDROXYAPATITE
(a) Back-flow time 0 h (HAP1)
(c) Back-flow time 4 h (HAP3)
321
(b) Back-flow time 3 h (HAP2)
(d) Back-flow time 5 h (HAP4)
Figure 16.2 SEM images of nano hydroxylapatite products. The results given in both Table 16.2 and Fig. 16.2 suggest that a certain back-flow (keeping warm) time after reaction-precipitation inside the SCISR is necessary for full double decomposition and/or, probably, surface deactivation of the newly formed particles, while too long a back-flow time leads to the product becoming worse in both purity and appearance. From the results given in Table 16.2 and Fig, 16.2, three to four hours can be considered to be the optimal condition for the back-flow (ageing) time.
16.3. 1.2 Influence of pH The results of the influence of the pH of the reaction mixture on the Ca/P mole ratio of the product are illustrated in Table 16.3. As can be seen in the table, the mole ratio of Ca/P increases as the pH decreases so it can be concluded that pH is an important factor affecting the composition or purity of the product. As is well known, in addition to HAP, there are three solid products with different Ca/P ratios which can possibly be combined in the product under various pH conditions yielding variation of the Ca/P mole ratio in the product during precipitation: Ca(OH)2 of Ca/P = oo, Ca~(PO4)2 of Ca/P = 1.5, and CaHPO4 of Ca/P = 1.0.
322
IMPINGING STREAMS Table 16.3 Ca/P mole ratios of the products obtained at different pH
Sample No pH of reaction mixture Ca/P mole ratio of product
HAP2
HAP7
HAP8
12.5
11.5
10.5
1.637
1.568
Chemically, the data listed in Table 16.3 should lead to the following inference in the range of pH tested: Higher pH favors the reaction below to occur, resulting in a high Ca/P ratio: Ca(NO3)2 + 2NH4OH = Ca(OH)2 ~ + 2NH4NO3 Lower pH promotes the reaction yielding a lower Ca/P ratio: 3Ca(NO3)2 +2(NH4)3PO4 = Ca3(PO4)2 ~ + 6NH4NO3 while even lower pH causes the reaction below to occur, leading to an even lower Ca/P ratio of the product: Ca(NO3)2 + (NH4)2HPO4 = CaHPO4 ~ + 2NH4NO3 However, it is somewhat unfortunate that such an inference cannot yet be verified with experimental evidence because of the complexity of the reaction system involved. According to the data listed in Table 16.3, the optimal or feasible pH should be between 12.5 and 11.5, from the point of view of product purity. On the other hand, comparing the SEM images of the sample products in Table 16.3 shown in Fig. 16.2(b), Fig. 16.3(a) and (b), respectively, the distinctness of the interface of HAP2 seems a little better than the others, so the pH value of 12.5 is determined as the optimal. ~ / : ~
~'i!!!iv :,:;i" .
~:
~
.
~,
•. . . .
~........... .............
(a) pH = 11.5 (HAP7)
(b) pH = 10.5 (HAP8)
Figure 16.3 SEM photos of the sample products HAP7 and HAP8.
PREPARATION OF NANO HYDROXYAPATITE
323
16.3. 1.3 Influence of dripping rate of di-ammonium phosphate solution The Ca/l:' mole ratios of the products obtained at various dripping rates of the (NH4)2HPO4 solution are listed in Table 16.4. It should be noted that the variation of the dripping rate implies a change in the concentration of phosphate and thus in the reaction rate of the double decomposition. The results in Table 16.4 illustrate the lack of purity of the sample product HAP5, while HAP2 and HAP6 have essentially the same Ca/l:' ratio. On the other hand, the SEM image of HAP6 shown in Fig. 4(b) indicates the bad appearance of the product. Therefore the optimal or, at least, most feasible dripping rate of the di-ammonia phosphate solution is determined to be 54 mL.L -~. An interesting phenomenon is that the sample product HAP5 has terrible purity, while exhibiting a good appearance of particles, as shown in Fig. 16.4(a). This is still difficult to explain exactly. The results given in Table 16.4 and Fig. 16.4 also indicate that the faster reaction rate inside the SCISR resulting from the appropriately shorter dripping time is necessary for synthesizing nano HAP in the process under consideration.
Table 16.4
Ca/P mole ratios of the products obtained at various dropping rates of (NH4)2PO4 solution Sample No Dripping rate, mL.min -~ Ca/P mole ratio
(a) HAP5
HAP2
HAP5
HAP6
54
96
24
1.910
1.624
(b) HAP6
Figure 16.4 SEM photos of the sample products obtained at various dripping rates of (NH4)2PO4 solution.
324
IMPINGING STREAMS
16.3.2 Optimal conditions for synthesis of nano HAP Summarizing the results described in the previous section, the optimal major conditions for the synthesis of nano hydroxyapatite by double decomposition-precipitation in the submerged circulative impinging stream reactor with calcium nitrate and di-ammonium phosphate as the reagents are determined to be as follows: the dripping rate of the (NH4)2PO4 solution 54 mL.L -~, pH of the reaction mixture 12.5, and back-flow time 3-4 hours; the amounts of Ca(NO3)2 and (NH4)2HPO4 are in the stoichiometric ratio.
16.3.3 Characterization of nano HAP product
16.3.3.1 TEM determination Figure 16.5 shows the TEM photo of a typical nano hydroxylapatite product synthesized in the SCISR following the operation procedure described in Section 16.2. As can be seen, the particles are about 15 nm in diameter and 50-70 nm long, well dispersed, and have a very regular shape and appearance.
~i ¸¸
.........
.............
:~!~i~ii ~!~ i
N:~:::2 !
,~i;:ii
i:?ii!:':
Figure 16.5 TEM photo of the nano hydroxylapatite product.
16.3.3.2 FTIR analysis The FTIR spectrum of the nano hydroxylapatite product is shown in Fig. 16.6. On the spectrum, 3570 cm -1 and 632 cm -~ are the absorption peaks reflecting the expandingcontracting and winding vibrations of OH-, respectively, suggesting that the oxhydryls do exist in the synthesized molecules; 566 cm -~, 604 cm -~, 959 cm -~, 1030 cm -1 and 1094 cm -~ are the typical absorption peaks of phosphoric acid radical. Somewhat unexpected peaks are those at 1420 cm -~ and 1470 cm -~, which are the typical
PREPARATION OF NANO HYDROXYAPATITE
325
absorption peaks of carbonic acid radical, implying the existence of CO~- in the synthesized product. The most possible source is the aqua ammonia, which absorbs CO2 from the surrounding air. The result suggests that the carbonyl easily enters the crystal lattice of apatite, which may be useful information.
"t
,
/
•,
I "~ i
,,I'
"-'-'~~,
l~,~.
I'''1
,4%
¢
,7
I
r5 :~: e~
2 Figure 16.6 FITR spectrum of the nano hydroxylapatite product.
16.3.3.3 XRD analysis The XRD spectrum of the nano hydroxylapatite product is given in Fig. 16.7 and, for understanding in detail, the corresponding diffraction data are listed in Table 16.5, where the data of the referential sample [231] are also given for comparison. From the data listed in Table 16.5 it is verified that, chemically, the product synthesized is hydroxylapatite.
I
20
!
30
I
40
I
50
I
60
20 Figure 16.7 XRD spectrum of the nano hydroxylapatite product.
326
IMPINGING STREAMS Table 16.5
Data of XRD spectrums of the synthesized nano hydroxylapatite product and the referential sample Synthesized product
Referential sample [231]
Crystal face 20
d ~t~
Ill 1
20
dA
I/Ii
200
21.494
4.1307
13
21.780
4.0772
7.0
111
22.579
3.9346
12
22.866
3.8859
2.6
002
25.725
3.4601
46
25.867
3.4415
37.4
102
28.003
3.1835
16
28.120
3.1707
10.5
210
28.763
3.1012
18
28.946
3.0821
14.3
211
31.692
2.2809
100
31.785
2.8130
100.0
300
32.777
2.7300
57
32.924
2.7182
63.0
202
33.97
2.6367
27
34.063
2.6299
24.5
130
39.721
2.2673
17
39.825
2.2617
20.1
401
46.556
1.9491
35
46.416
1.9547
1.3
132
47.966
1.8950
17
48.101
1.8900
15.7
213
49.268
1.8479
45
49.485
1.8404
33.6
231
50.353
1.8106
19
50.514
1.8053
20.2
410
51.112
1.7855
16
51.300
1.7795
14.0
402
51.872
1.7611
14
52.101
1.7540
17.1
004
53.065
1.7244
23
53.185
1.7208
18.4
16.4 CONCLUDING REMARKS The preliminary investigation on the synthesis of nano hydroxylapatite by double decomposition-precipitation is made with (NH4)zHPO4 and Ca(NO3)2 as the reactants and the submerged circulative impinging stream reactor (SCISR) as the reactionprecipitation equipment. The following are preliminarily concluded: (1)
The three operation variables, i.e., the back-flow (ageing) time after reactionprecipitation inside the SCISR, the pH of the reaction mixture, and the dripping rate of the di-ammonium phosphate solution into the reaction mixture exhibit influences on both the Ca/P mole ratio and the appearance of particles in the product.
PREPARATION OF NANO HYDROXYAPATITE (2)
327
The optimal conditions determined according to the preliminary investigation are: back-flow (ageing) time 3-4 hours, pH = 12.5, and dripping time of the diammonium phosphate solution 54 mL-min -~. (3) The results of the characterization of the product synthesized with Fourier transformation infra-red spectrophotometer (FTIR), X-ray diffractometer (XRD), scanning electron mirror microscope (SEM) and the transmission electron mirror microscope (TEM) illustrate that the product synthesized by the process of double decomposition-precipitation with calcium nitrate and di-ammonium phosphate as the reactants in the SCISR consists of well dispersed particles of about 15 nm in diameter and 50-70 nm long, having a very regular shape and appearance; and is confirmed to be hydroxylapatite. (4) Also, the results above verify again that the submerged circulative impinging stream reactor is especially suitable for the preparation of ultrafine particles by reaction-precipitation and exhibits very good performance.
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-17RESEARCH AND DEVELOPMENT OF LIQUIDCONTINUOUS IMPINGING STREAM DEVICES AND APPLICATION FORECASTING
As can be seen from the discussions in the previous chapters, due to the strong interaction between opposing streams impinging against each other, the liquidcontinuous impinging streams (LIS) method has the features of very efficient micromixing and very strong pressure fluctuation, and so has good application prospects. As an engineering discipline, Chemical Engineering should take industrial application as the starting-point and the end-result of all its investigations because converting the achievements of chemical engineering R&D into the related equipment to push improvement and innovation of technical devices is the most direct and efficient way of converting science and technology into productivity. On the basis of the investigations supported by the National Natural Science Foundation of China, the author of the present book has developed several technical devices for various unit operation processes based on the use of impinging streams, some of which have been discussed in the relevant chapters, while some LIS equipment is introduced below. The development of application is as important as the development of the device itself, the two things complementing each other, especially for new technologies and new devices. As the intention behind application development is to discover and expand new areas of application, the last section of this chapter will, taking the features of LIS into account, forecast some possible new applications and their prospects.
17.1 THE VERTICAL CIRCULATIVE IMPINGING STREAM REACTOR In the previous chapters much space has been devoted to describing the features and application cases of the submerged circulative impinging stream reactor (SCISR). Indeed, in all these applications the SCISR exhibited excellent performances, and thus suggested a possible way of simplifying the preparation technologies for nano materials by improving the technical equipment for the reactions, while the processes and/or formula to prepare nano product of high quality have also been improved. On the other hand, the SCISR also has some disadvantages. The first is that the operational impinging velocity is limited to an extent. An excessively high impinging
329
330
IMPINGING STREAMS
velocity may result in too tempestuous a liquid surface and droplet spattering and thus unstable operation. Secondly, parts of the walls of the reaction vessel are plates so that, mechanically, there is a little difficulty in constructing equipment to be operated under pressure or vacuum. In order to overcome these disadvantages and extend the application area, the vertical circulative impinging stream reactor (VCISR) was designed and patented [232]; its structure is shown in Fig. 17.1.
11 9
i ~
/12
1-vessel 2-drawing tube 3, 4-propellers 5-discharging port 6-feeding tube 7-overflow outlet
IZ
8, 9-outlet or inlet for heat exchanging medium 10-jacket for heating or cooling 11-outlet for exhaust gas 12-cover IZ-impingement zone
i
~5
Figure 17.1 The vertical circulative impinging stream reactor.
R&D OF LIS DEVICES AND APPLICATION FORECASTING
331
The major difference between the VCISR and the SCISR is the former's vertical structure, while the SCISR has a horizontal structure. This changes the flow configuration from horizontal impinging streams in the SCISR to vertical in the VCISR, and also the container form of the VCISR is changed to a cylinder. The upper and lower drawing tubes (2) are mounted vertically and co-axially in the container of the reactor; and two corresponding propellers (3 and 4) are installed coaxially inside the drawing tubes. The spiral blades of the two propellers have opposite screw directions -- left and right -- both being driven by the same motor. The vessel is filled with the process medium, liquid or suspension, and its level is over the inlet of the upper drawing tube; the space under the level makes up the effective region for the reaction or other processing, while the space over the level is kept for gaseous phase to meet the requirements of most reaction systems. In order to enhance the flow efficiency and to avoid any dead corners appearing, the bottom of the reactor is made to have a half sphere form. The two propellers (3) and (4) push the streams (process liquid or suspension) in opposite directions to flow through the drawing tubes towards the center of the reactor and impinge against each other, forming the impingement zone IZ around the impingement plane, where the violent relative motion of the fluid elements in different directions results in effective mixing and contact between fluid elements and/or between solid and liquid phases, promoting reaction or other processes occurring in the region. After impingement, the steams pass through the annular chamber between the upper and lower drawing tubes and the side wall of the reactor to return to the upper and lower sides of the effective vessel, and then are pushed by the propellers (3) and (4) downward and upward, respectively, to flow through the drawing tubes towards the center of the reactor and to impinge against each other, and so on. The VCISR can be operated either continuously or in batch mode. In batch mode, an appropriate amount of the reaction material is fed into the container all at once; when the reaction is carried out after a certain period, the reacted mixture is discharged from the reactor through the discharging tube (5) at the bottom of the reactor. In continuous operation, the fresh liquid or suspension is fed into the reactor through the upper feeding tube (6); or, if necessary, two streams can be fed separately; furthermore, another feeding tube can be set with its outlet at the inlet of the lower drawing tube, and thus two streams are fed through the upper and lower drawing tubes simultaneously. The reacted mixture overflows from the reactor through the overflow outlet (7). The mean residence time in the reactor can be arbitrarily set by suitable selection of the input flow rate(s). In addition, the VCISR can also be operated in semibatch mode (or semi-continuously), i.e., one or more reactant stream(s) is (are) first fed all at once into the reactor, while other(s) is (are) fed continuously through the feeding tube (6); when the reaction is finished, all the reacted mixture is discharged through the discharging tube (5) at the bottom of the reactor. Any gas or vapor generated during the reaction is expelled through the outlet (1 1) at the top of the reactor. There is a jacket outside the reaction vessel for heat exchanging, which enables one to cool or heat the process mixture inside the reactor. When cooling the mixture, cold water or another cooling medium enters the jacket through the lower inlet tube (8) at the bottom and flows out through the upper outlet tube (9), while for heating the
332
IMPINGING STREAMS
mixture the pipeline is connected in the opposite way. If necessary, a coiled heat exchanging tube can also be arranged outside the drawing tubes (2) (not shown in Fig. 16.1). The VCISR can have a cover (12). By suitably designing the mechanical strength and the connecting mode, the reactor can be operated under pressure or vacuum, while for open operation, the cover can be removed. It can be seen from the description above that the VCISR works according to the same principles as the SCISR and so has the inherent features of a circulative impinging stream reactor, i.e., it utilizes all the excellent features of liquid-continuous impinging streams, has the special flow configuration of circulative perfect mixingplug flow in series and mean residence time arbitrarily set and has the identical characteristics of high energy efficiency and low power consumption. Also, the VCISR has all the basic function of the traditional stirred tank reactor (STR), such as sealing ability, capability of being operated under pressure or vacuum, or continuously or in batch mode, and has heat exchanging capability, etc. The change of the vessel's form to a cylinder creates the moderately forced condition of the reactor so that it becomes much easier to design and manufacture the VCISR to be operated under pressure or vacuum, while this is somewhat difficult with the SCISR. In addition, the use of only one driving motor makes the mechanical structure simpler, although it gives rise to the problem of a relatively long shaft. The trial operation shows that the design objectives for the VCISR have been essentially realized and the disadvantages that the SCISR has are overcome, giving a wider area of application of the flow configuration of the circulative impinging streams. However, if the VCISR is operated at a very high impinging velocity, e.g., 2 m.s -1 or even higher, so that a very high rotary speed of the propellers must be employed, the propellers may lead to a strongly rotating movement of the process mixture inside the container, producing great centrifugal force. The latter causes the mixture to have an uneven level with large height difference: much higher near the wall while very low near the upper drawing tube, sometimes even lower than the inlet of the upper drawing tube, resulting in unstable operation of the reactor. The vertical-circulative impinging stream reactor of Type II was designed [233] to solve this problem and its structure is shown in Fig. 17.2. In comparison with the VCISR, the most important change with Type II is the addition of a ring-like back plate (13) mounted just under the overflow outlet port (7), which is for restraining too much uneven level due to centrifugal force. The other differences of the Type II from the original VCISR are the use of the tapered bottom (14) and the addition of the lower feeding tube (15). These have the advantages of easier manufacture and of more easily ensuring the homocentric degree of the related parts, including the propellers, the drawing tubes and the cylindrical vessel, as well as a more flexible arrangement for material(s) feeding, so that it can be applied in a wider range for various process systems. There are also other possible solutions to overcome the problem of uneven level due to centrifugal force.
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Crystallization is an important unit operation process encountered widely in the chemical, medical and light industries. The goal of the operation is usually to produce a product of uniformly large-sized crystalline. Such crystalline product not only has a significantly improved appearance and the properties of moisture absorption and blocking but also favors enhancing the technical and economic indexes in the separation operations after crystallization, e.g., by vacuum filter or centrifuge etc. According to the modes of creating supersaturation, crystallization from solution can be classified into the three modes: cooling crystallization, evaporative crystallization,
334
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and reactive crystallization, among which cooling crystallization is most widely applied, while the use of reactive crystallization has increased rapidly in recent years. Regardless of the type of crystallizer, in order to obtain uniformly large- sized crystalline, the supersaturation must be suitably controlled. At very low supersaturation, the growth rate would be small, yielding low efficiency; it is even possible that crystals cannot be separated out without the addition of crystal seeds, while at too high supersaturation, fine crystals would be obtained finally because of nucleation in great quantity. In addition, the supersaturation must be uniform; otherwise the local concentration differences may also yield fine crystals so that the uniformity of crystal sizes becomes poor. Mean supersaturation is always easily controlled. For example, for the case of cooling crystallization, the only requirement is to control the temperature according to the concentration of the solution in which the crystallization occurs; for reactive crystallization mean supersaturation at the given value can be ensured by appropriately manipulating the feeding conditions, including concentration and flow rate, under the condition of temperature meeting the requirement of the system. In order to create a uniform supersaturation environment, good micromixing conditions are needed, because, after all, both the nucleation and the crystal-growth occur at the molecular scale. A number of crystallizers have been applied industrially for many years; but the micromixing conditions they provided are not ideal. The crystallizers with drawing tubes developed in the period from the 1950s to the 1970s, such as DTB [234], DP [235], Standard-Messo [236] crystallizers, are in the same category, and, to date, are still the most advanced of the existing continuous crystallizers. The common features of the crystallizers in this category are: providing the measures of vacuum-cooling, evaporation, or direct contacting cooling with an easily-volatile liquid etc to produce supersaturation, and creating uniform supersaturation by circulating the suspension at large flow rate. At the same time, they have the following disadvantages: The circulation of large amount of suspension leads to higher power consumption; (2) The micromixing condition is not so good so that for some substances with poor crystallization properties it is rather difficult to obtain uniformly large-sized crystalline; and (3) Generally they are not applicable for reactive crystallization. As described in previous chapters in Part II, liquid-continuous impinging streams (LIS) has the feature of very efficient micromixing and can provide a uniform and controllable supersaturation environment for the crystallization process, favoring the production of uniformly large-sized crystalline; also, it has been proved experimentally that, to an extent, LIS can enhance crystal-growth rate. For industrial application, Wu [237] designed and patented the impinging stream crystallizer, the structure of which is shown in Fig. 17.3. The essential parts included in the impinging stream crystallizer are: the cylindrical vessel (1); the upper and lower drawing tubes (2) co-axially fixed in the vessel (1); and the upper and lower propellers (3) and (4) set up inside the upper and lower drawing tubes, respectively, which are driven by the same motor and have the spiral blades with opposite screw directions, i.e., one is the left screw and the other is the right. The vessel 1 contains the process medium, crystals-in-solution suspension, with the level over the
R&D OF LIS DEVICES AND APPLICATION FORECASTING
335
upper drawing tube. The space inside the vessel and under the level constitutes the effective volume for crystallization; while that over the level is kept for gaseous phase. In order to enhance the flow efficiency and avoid any dead corners appearing, the bottom of the crystallizer is made to have a pan-like form; or, alternatively, a tapered bottom can also be used.
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Figure 17.3 The impinging stream crystallizer, l-vessel; 2-drawing tube; 3,4-propellers; 5,6feeding tube; 7-overflow cofferdam; 8-mother liquor tube; 9-classifying leg; 10-discharge tube; 11-jacket for heating or cooling; 12,13- inlet and outlet for cooling medium; 14-cover; 15-outlet for gas/vapor; A-pump for mother liquor; B-cycling pipe; IZ-impingement zone The two propellers (3) and (4) push the streams (process suspension) in opposite directions to flow through the upper and lower drawing tubes, respectively, towards the center of the effective volume and impinge against each other, causing the impingement zone IZ around the impingement plane, where the violent relative motion of the fluid elements in different directions effectively promotes micromixing, creating a uniform supersaturation environment. After impingement, the suspension streams pass through the annular chamber between the upper and lower drawing tubes and the side wall of the vessel to return to the upper and lower sides of the effective volume,
336
IMPINGING STREAMS
and then are pushed by the propellers (3) and (4) downward and upwards, respectively, to flow through the drawing tubes towards the center of the crystallizer and impinge against each other, and so on. The crystallizer can be equipped with a jacket (11) outside the vessel for heat exchanging, which is used when the heat needing to be removed is not large. In this case the cooling medium is introduced into the jacket through its inlet (12) or (13) to remove the heat released during crystallization by indirect heat exchange, and then discharged through the outlet (13) or (12). The heat exchanging jacket can also be used in reactive crystallization to keep the temperature of the suspension inside the vessel at a given level. In addition, the outlet tube (15) on the top cover (14) can be connected to the vacuum system for vacuum-evaporation cooling. It is also possible to employ a combination of the two cooling modes just mentioned. The efficient micromixing in the impingement zone IZ can ensure that the temperature and the supersaturation is stable and uniform. The crystallizer has two feeding tubes (5) and (6), so that it is applicable for all three categories of crystallization operation: cooling, evaporative and reactive crystallization. For cooling or evaporative crystallization only one feeding tube is needed, and so the other, feeding tube (5) or (6), can be ignored. Reactive crystallization normally requires at least two feeding ports. In this case, if necessary, more than two feeding tubes can be used; but the outlet of any extra feeding tube must be fixed at the inlet of either the upper or the lower drawing tube. In addition to continuous operation, the impinging stream crystallizer can also be operated in batch mode to suit the crystallization of substances with very low crystalgrowth rates, e.g., some of the organic compounds. In continuous operation, the fresh solution is fed into the crystallizer through the feeding tube described above. A certain mean residence time in the crystallizer, i.e., the crystal-growth time, can be ensured by a proper arrangement for the feeding flow rate. Because of the density difference between solid and liquid, when the direction of the flow turns, part of the crystals in the stream flowing downward out of the annular chamber leave the suspension and drop down into the classifying leg under the bottom of the crystallization vessel by gravity; Pump A then draws the liquid, clean solution without crystal, from the overflow cofferdam (7) at the upper part of the crystallizer, through the mother liquor pipe (8), and delivers it to the classifying leg for hydraulic classification of the crystals. The thick crystalline particles, with part of the mother liquor are taken out through the discharge tube (I0) and conveyed to the device for solid-liquid separation, while the small particles are carried by the classifying liquid to return to the crystallization vessel for further growth. In Fig. 17.3, Pump A and the cycling pipeline B represented by the dashed line are the complementary elements of the crystallizer. While in batch operation, the fresh solution is fed into the crystallizer all at once; the crystallization process carries on for a certain period, then the suspension is discharged from the crystallizer and is conveyed to the device for solid-liquid separation. For such operation the classifying leg (9), the overflow cofferdam (7), and the pump A are not needed.
R&D OF LIS DEVICES AND APPLICATION FORECASTING
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In comparison with existing crystallizers, including DTB, DP and Standard-Messo crystallizers, the impinging stream crystallizer has the following main advantages: (1) By utilizing the feature of LIS promoting micromixing efficiently, it can provide a very uniform supersaturation environment for crystallization; (2) Because LIS itself has the nature of efficient micromixing, the circulation of the suspension inside the crystallizer in large amounts is not necessary, but in a small amount, i.e., the cycling ratio in the impinging stream crystallizer is very small (< 2), yielding lower power consumption; (3) The special flow configuration of the perfect mixing-plug flow in series in the impinging stream crystallizer is more suitable to crystallization process and so it is much easier to prepare uniformly large-sized crystalline particles; and (4) It is multifunctional and can also be used for reactive crystallization and has a much wider range of application..
17.3 PROSPECTS FOR THE APPLICATION OF LIQUID-CONTINUOUS IMPINGING STREAMS As mentioned in the Introduction, the concept of Impinging Streams (IS) was presented as early as the beginning of the in 1960s, although the industrial application of IS evolved very slowly over a considerably long period. Since the 1990s, the development of IS application has progressed significantly and some of the achievements, such the impinging stream gas-liquid reactor and the submerged circulative impinging stream reactor etc., are pushing towards industrial application. The major reason tor the great progress lies in the fact that, after many years of investigation, the dependence of IS application on its properties has been well understood, while one of the hallmarks of the progress is that more and more scientists and engineers working on IS have used this understanding as the guiding principle in their research & development, leading to enhanced technical developments. Propelling tbrward the industrial application of a new technology, or, more generally, transfbrming science and technology into productive forces, is a problem of complicated system engineering. In addition to the advanced performances, the reliatzility and economic feasibility of the new technology itself, it is necessary to get a common view from the users -- the decision-makers of the enterprises. Without a common view, it is impossible to apply any new technology successfully, no matter what kind and how advanced it is. Making the effort to create such a common view is one of the unavoidable tasks of inventor of the new technology. There may, of course still be other factors hindering the industrial application of new technologies, Although the concept of IS was first suggested for gas-solid systems and investigations into gas-solid IS were the forerunner in the IS field, the discovery in the 1990s of the excellent features of liquid-continuous impinging streams (LIS) such as its efficient micromixing and the existence of very strong pressure fluctuation etc. changed the status of IS investigations: the emphasis was diverted to those with liquid
338
IMPINGING STREAMS
as the continuous phase and, furthermore, it changed the status of IS application. It is expected that in the coming years more applied technologies will emerge for LIS than for GIS in various processing industries. In addition to the areas of LIS application introduced above, such as chemical reactions, crystallization, and preparation of ultrafine powders, the following may be fields where LIS could be successfully applied and so are worthy of note:
(1) Crushing of cells: This is a hot topic in biotechnology. From primary analysis of the flow field it can be reasonably concluded that in the impingement zone of LIS equipment there must be a very strong shearing force field due to the strong interaction of the opposing streams, and such shearing force can be used for crushing certain kinds of cells, although it cannot of course be applied for every type of cell. In fact, a primary test using the vertical circulative impinging stream reactor (VCISR) for crushing algae cells have shown excellent results. It is notable that such a methodof crushing cells, i.e., with hydraulic energy, has a number of advantages in technical economics and environmental protection and is therefore of great significance. (2) Emulsification: The key for carrying out emulsification is dispersion, i.e., to disperse an aqua liquid uniformly in an oil liquid, or, on the other hand, to disperse an oil liquid into an aqua liquid. Obviously, the strong interaction of the opposing streams in LIS, including collision, pressing and shearing between fluid elements etc, promotes the dispersion process. Since LIS has high energy efficiency, it can be expected that, in comparison with the existing ones, LIS can yield the same emulsifying capability at a lower level of energy consumption. (3) Solvent extraction: To ensure dispersity of two insoluble liquids in each other is also an essential problem to be solved for an effective solvent extraction process. Utilizing the interaction between the opposing streams in liquid-continuous impinging streams, it is possible to achieve higher dispersity and, consequently, a larger interface area or larger volumetric mass transfer coefficient at lower energy consumption, so that the extraction process can be carried out with higher efficiency. Possibly, there are other fields where it is hoped that LIS technology can be successfully applied. As mentioned in the section "Illustration of the Translation" in the Chinese translation of Ref. [5], Impinging Streams can never be a universal tool. Furthermore, in the field of chemical engineering, many common regularities have been found and probably there will be more, but no universal tool or method has been discovered. Even for the same unit operation, no technical method can be applied to all the substance systems. What scientists and engineers can do is to use their knowledge of the specific systems and the specific technical methods to create a feasible solution to counter the specific process and the specific system; of course the impinging streams could be one feasible solution. In this way, we may discover more and more target processes and target systems for which IS can be successfully applied.
Postscript In 1995, when I contacted an Editor from The Chemical Industry Press of China in connection with the publication of the translation of "Impinging Stream Reactors: Fundamentals & Applications" by A. Tamir, the Editor asked: "Could we publish a book written by yourself? .... Ten years later." I replied. Although the answer escaped my lips immediately, it expressed exactly my basic thought: the writing of such a book must be based on thorough and in-depth investigations. However, the Editor's question planted in me the wish to write a book especially discussing Impinging Streams. After ten years of eftort, I am able to relax as the Chinese edition of this book was finally finished in 2(i)05. However, I still feel some regrets: a number of problems on the fundamentals, such as the mechanism and theoretical model of liquid-continuous impinging streams (LIS) promoting micromixing, the relationship between pressure fluctuation and micromixing, the quantitative description of promoting process kinetics, and the shearing force field in LIS etc, have not yet been solved in my investigations; some ideas, which to my mind are would be good, on the application developments of LIS have not been reinforced and, even more, no results related to these ideas can be seen. In order to complete this phase of the work, I have no alternative but to finish the writing. Alter some consideration, I feel that it is normal to have some regrets. Philosophically, understanding the absolute truth is an infinite process and something that can never be arrived at. Science and technology are always growing and advancing, never remaining at one level and, of course, impinging streams (IS) is no exception. Understanding the regularities and developments of IS cannot be achieved instantly and the problems cannot be solved by a few scientists and engineers within a finite period. It is believed that IS is a technical method with superior performances, and so has a huge sphere of application. This belief is not just based on hope but on an in-depth understanding acquired during more than twelve years of study investigations. In fact, a number of technologies employing IS have appeared, or are starting to appear, in industrial applications and are becoming advanced processes, and I expect that more and more technologies employing IS will appear in various processing industries, growing, and becoming stronger. In the last ten years and more, investigations into IS have increased, the areas in which these investigations are being carried out is constantly expanding, and the number of scientists and engineers working in this area is also increasing. All this indicates that rapid progress in both fundamental investigations and technical developments of IS will be made. Without doubt, my expectations must be realized in the near future.
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Nomenclature
A,B
component variable, dimensionless area, m
A,B
convenient constant defined by Eq. (10.12) specific surface area, me.m ~ Proportional factor (in Eq. 4.9), dimensionless
a, b, c
exponent (in Eq. 2.32), dimensionless concentration, kmol.m -~ or kg.m -~ mass fraction, kg.kg-
Ci
adjustable parameter ( i - 1, 2,...), dimensionless
Cs
solubility, kmol.m -~ or kg.m -~
AC
supersaturation, kmol.m ~or kg.m -~
CD
drag coefficient, dimensionless
C
specific heat capacity at constant pressure, kJ.kg-~.K -~
P
D
diffusivity, m-.s diameter, m or lam diameter, m
E
residence time distribution probability density function, s-~ evaporation intensity, kg.m-~.s -~ enhancement lector due to impingement between opposed streams active energy, kJ.mol -~ residence time distribution function, dimensionless frequency, Hz
f f,1
acceleration of particles in various flow regimes ( i - 1, 2, 3, 4, 5, 6), m.s -e volumetric fraction of plug flow region, dimensionless volumetric fraction of low pressure fluctuation intensity region, dimensionless dimensionless stress intensity acceleration of gravity, - 9.81 m.s -2
355
356
IMPINGING STREAMS -9
gc
conversion factor (= 1), kg.m.N-~.s -
H
height, m Fourier transform of transient pressure height of particle dropped down, m lower spacing in circulative impinging stream dryer (in Chapter 6), mm heat transfer film coefficient, W.m-Z.K -l
MV
latent heat of evaporation, kJ.kg -t
AHc
combustion heat, kJ.kg -~
I
integrated value of concentration (Eq. 3.41) integrated intensity defined by Eq. (11.7), kPa.m 3
K
overall mass transfer coefficient, m.s -~ adjustable parameter, dimensionless mass transfer film coefficient, m.s -~ constant (in Eq. 2.32), dimensionless adiabatic coefficient of air, dimensionless reaction rate constant
K~
overall crystal-growth rate constant, m.s -1 nucleation rate constant, s-~
/¢
observed frequency factor, m-s -1
ks
rate constant of crystallizing reaction on the surface, m.s -~ turbulent dynamic energy, N.m.kg -~ frequency factor, m3.kmol-~.s-~ (for second order reaction only) length or impinging distance, m length or length variable, m
l0
integral turbulent degree, m
M
constant in Eq. (1.15), dimensionless momentum, kg.m.s -~ molecular weight, dimensionless absorption rate of the absorber, kmol.s -~ mass flux, kg.m-2.s -~
m
N
mass or mass flow rate, kg or mass fraction, dimensionless sampling frequency for transient pressure, Hz number of particles/droplets rotary speed, rpm absorption flux, kmol.m-2.s -~
NOMENCLATURE
357
nucleation rate, s-~ N/
absorption rate per unit volume, kmol-m-~-s-~
N
number of sampling times at a measuring point
H 111
average number of particles per unit volume, s
Nu
Nusselt number, dimensionless
0
amount of oxygen theoretically needed for unit fuel burning, kg.kg -~
-I
pressure or total pressure, Pa probability of flow without collision defined in Eq. (2.41), dimensionless power inputted, W PSD
mean square difference of transient pressure, Pa
P
pressure or partial pressure, Pa
@
pressure drop, Pa
O
amount of heat current, kJ heat flux, kJ.m-2.s -~ -I
O
heat transfer rate, kJ.s
R
radius of accelerating tube, m universal gas c o n s t a n t , - 8.314 Pa.m~.mol-J-K -~ cycling ratio, dimensionless
RMS
mean square root
Rsl"
surface resistance for reaction defined by Eq. (8.21)
Rdi ff
resistance to diffusion defined by Eq. (8.21)
R
radial distance or coordinate variable, m radius of conduit or particle/droplet, m
Re
Reynolds number, dimensionless impinging distance, m relative supersaturation. = C/Cs
SD
standard deviation
Sh
Sherwood number, dimensionless temperature, K or °C
L
length of window function, s time, s characteristic time constant for chemical reaction, ms
IM
characteristic time constant for micromixing, ms
358
IMPINGING STREAMS
/in
(macro) mixing time, s
U
overall heat transfer coefficient, kW.m -2 velocity vector, m.s -~ velocity, m.s -1
/1/c
local sonic velocity, m.s -~
b/0
impinging velocity, m.s -1
V
volume, m 3 volumetric flow rate, m3.s-~
W
power, W or kW
W
window function defined by Eq. (11-5), dimensionless
X
selectivity or segregation index, dimensionless moisture content, kg.kg -~
x,y,z
coordinate variables or distance, m
Z
number of particles particles-to-gas volumetric ratio, dimensionless included angle between motion and vertical down directions, o Delta function width of nozzle in Eq. (1.1), m thickness of liquid film, m energy dispersion rate per unit mass, W.kg - l roughness, mm Kolmogoroff micro length, m friction resistance coefficient, dimensionless thermal conductivity, kJ.m-l.s -j viscosity, Pa.s dynamic viscosity, m2.s -j velocity potential function, dimensionless number frequency function of droplets, gm -~
rls
sulfur-removal efficiency, dimensionless shape coefficient, dimensionless axis of rotating flow mean residence time, s
P
density, kg.m -3
NOMENCLATURE
"~111
average interval between two times of collisions, s effective cross section area of collision, m 2 standard deviation interface tension, J . n f -~ dimensionless lower spacing in circulative impinging stream dryer, = h/d stream function, dimensionless
(
dimensionless axial distance non-dimensional impinging distance, = S/d local resistance coefficient, dimensionless
Superscripts: average value (1)
Steam 1 or Feeder I
(2)
Steam 2 or Feeder 2
Subscripts: A
air
A, B
component variable
AM
arithmetic-mean value
a
absorption, absorber
a, g
air, gas
ac
in accelerating stage
at
for atomization
b
bulk burning
BP
at boiling point
c
in constant rate period critical combustion
cal
calculated
cs
in collision-sleeping region
D
drying chamber
d
in decelerating stage drawing tube
ds
due to design factors
eft
effective
ex
experimental, experimentally measured
359
360
IMPINGING STREAMS
FB
in fluidized bed
f
final
fal
in falling down region
fl
flame
fv
fuel vapor
G
of gas, in gas film, in gas phase
g
of gas
h
at height h
i, in
input, inputted
IS
of impinging streams
im
in impingement zone, at impingement plane
L
of liquid or absorbent
lag
transfer lag or pure lag
lm
logarithmic mean value
m
mean in complete mixing region
max
maximum
0
at outlet
OX
of oxygen
P
of particles or droplets in plug flow region
R
of reactor
R,S
component variables
r
relative radial of radial coordinate
ST
in stirred tank reactor
S
at surface, of surface area
t
terminal total
V
volumetric, based on volume
W
of wet bubble
X, y , Z
coordinates
Z
in vertical direction
o0
infinitely far, outside burning region
0
initial
Subject Index of CISD, 147 Applied technologies of IS, research and development of, 119 Arbitrary function of time, 83 Arithmetic mean diameter, 112 Arrhenius relationship, 266 generalized, 263 Assembled and primary particles, 272 Atomization conditions of, 178 of liquid, atomizers for, 156 Atomizing agent, 168 Atomizing gas, adiabatic coefficient of, 116 Atomizing nozzles, structure of, 164 Atomizing pressure, influence of, 121 Average rate of process, 90 Axial-symmetric impinging streams, 28
Absorbent for FGD, hydrated lime, 162 with various Ca(OH)2 concentrations, spray droplets o£ 177 Absorber, structure of, co-axial cylinders. 164 Absorption chamber, 158 chemical, by impinging streams. 153 enhancement of, 156 of CO2, acetone and ammonia into water, CO~ into NaOH solution, _
160
Absorption efficiency or conversion, 155 dependence of on Ca/S, 167 Absorption enhancement, models for, 155 Absorption equipment for FGD, packed column, spray tower, swirling plate column. 162 structure of, 157 Absorption rate, 160 Accelerating tube, 42, 44, 69, 98, 136 Accelerating tube effective length of, t..... 47, 48 flow through, 92 length of, 54, 70, 93, 104 Acceleration, 49 and deceleration, stages of, 46 of droplets, 172 of particles, 4, 91, 106 motion, of first stage, 50 Active energy, 266 Active region, 61 for transfer, 67 Additional consumption of Ca, 179 Ageing of precipitate, 273 Air left reactor, 159 Airflow velocity, 121 at exits of nozzles, 112 Aluminum sulfate solution, initial concentration of, 121 Annular chamber, 136
Back mixing, 75, 90 in impingement zone, 77 Boltzmann equation, 66 Breaking-up of particles, 65 Buoyancy force, 43 Burning droplet, final radius of, temperature at the surface of, 193 Burning particle, surface of, burning rate of, 195 Ca(OH)2 concentration of, 178 concentration gradient of inside droplet, 163 Ca(OH)2 suspension, 173 atomization of, 169 Ca/S mole ratio, influence o£ 179 Ca/S ratio, influence, critical value o£ 166 Caking and cleaning, 169 of material on walls, 122 Carbon, combustion of, 195 Centrifugal separator, 129 Characteristic reaction time constant, t~, 215 361
362 Circulative impinging stream dryer (CISD) analysis for upper discharge port of, 148 comparative experiments of, results of, moisture content profile along path, 141 equipment and system design of, parameters of, 138 experimental model equipment, system scheme of, 137 experimental procedure of, overall performance of, 139 quasi industrial test of, main parameters for design of, 150 structure of, 120, 135 working principles of, 136 Circulative IS drying, 134 Circulative ratio, R, 219 Classifier in mill, 202 Coalescence, 41 Coal-gasification, Koppers-Totzek process for, 201 Co-axial gas-solid suspension impinging streams, hydrodynamics of, 32 Co-axial horizontal gas-solid impinging stream contactor (ISC), 68 Co-axial horizontal impinging stream equipment, 50 Co-axial horizontal IS drying, smaller and larger dryers, 123 Co-axial impinging streams, 25 Co-axial two impinging stream device, for measurement of droplet size distribution, 109 Co-axial two impinging streams, 11 Co-axial vertical impinging streams, 56 drying, 124 Colliding-slipping region, 88 Collision between droplets, 5, 181 influence of, 108 Collision between particles, 61 frequency of, 63 influence of, 65 similarity of, 66 Collision frequency, 64 number of, 64 probability of, 62, 89 Collisionless flow, probability for, P(t), 64 Collisions
IMPINGING STREAMS bouncing and shattering, 114 of particles on the wall and between particles, 92, 94 Collision-slipping region, 74 Combined Multifunctional Impinging Stream Gas-Liquid Reactor drawings of, 188 with replacement of foam remover, 190 for a huge amount of gas, 187 instruction of, 188 Combustion intensification of, 194 intensification of due to IS, 196 Competitive-consecutive reaction scheme, rate constants of, 226 Complete mixing, assumption of, 253 Components, intensity of, 37 Concentration of particles, 4, 62, 69 Concentration, gradient of, 2 Condensed condition, of liquid, 205 Constraining grids, 131 Continuity, 29, 30 Continuous equation, 192 Continuous phase flow of, 19 liquid as, 7 Continuous phases, differences between properties of liquids and gases, 207 Convolution integral, 75 Criterion for decision of diameters of IS device, 36 Critical nozzles, 109 Crystal seeds, 256 mean diameter of, 261 Crystal-growth, overall rate of, rate coefficient of, 255 Crystal-growth rate coefficient, data interpretation of, 256 Crystal-growth rate, comparison between in ISC and FBC, 264 Crystallization from solution, 270 Crystallizing reaction, 255 Cu-Ag double metal powder preparation chemical reactions in; procedure for, 298 comparison between various technologies and devices for, data table of, 299
SUBJECT INDEX Damped oscillations, 53 Data interpretation, 115 of FGD, 175 mathematical relationship(s) for, 78 relationship(s) for, 81 Data pre-treatment, in pressure fluctuation measurement, rotor point wiping, blank experiments, 242 Deceleration movement, of second stage, 51 Deceleration, 42 effect on, 56 of particles, 4 Deflection, 21 direction of, 20 Deformation of fluid elements, 24 6-tunction, 70 Desulfurization of flue gas, 15 Determination of micromixing, reaction system for determination, 236 Device design tk~l wet FGD, influence of, reasonableness of, 181 Difference between TIJ mixer and SCISR, 235 Diffusion resistance, 196 Diffusivity, D, 214 Dimensionless impinging distance, Shl~, 181, 182 Dimensionless mean velocity, comparison of, 37 Direction turning of flow, 95 Di-sodium phosphate, crystal-growth kinetics of, 254 Dispersal, degree of, 269 Dispersed phase, RTD of, 67 Dispersity of liquid, 65 variation of, 108 influence of IS on, 107 Dissolution of urea and sodium chloride, 209 Distance between nozzles, influence, 161 Double decomposition reaction tor preparation of white carbon black, 273 Drag coefficient, 45, 48, 57 Drag coefficient, (7~>47.49 Drag force, 44, 56, 58, 148 coefficient, CI~, 69, 93 due to friction, 92 Dried product, moisture content of. out rate of, 148
363 Driving force, 67 mean, 90 Droplets coalescence and/or re-atomization of, mean diameter of, 108 coalescence of, 61 deformation of, 198 oscillation movement of, reatomization-coalescence of, 156 size distribution, method for measurement of, 110uniformity of sizes, 112, 114 Droplets-in-gas suspensions, 107 Dry PVC powder, properties of, 139 Dry-based capacity, 129 Drying chambers, primary and secondary, 121 Dynamic energy, 19, 114 Dynamic method, input-response technique, 77 Dynamic response, 75 Eddy diffusion, 214 Eddy pressure nozzles, 170 Effective cross section for collision, 64 Elastic and non-elastic collisions, 74 Electro-conductivity, 218 Energy, conversion of, distribution of over molecules, 254 Energy balance, 93 Energy consumption, 186 Energy dispersion, 214 Enhancement parameters, expressions for, 156 Environment protection, 162 Equilibrium, limitation of, 155 Equipment design and scale-up, 91, 92 consideration of, 169 Equipment system for preparation of nano copper, drawing of, 287 for RTD measurement, 81 Ethyl acetate saponification kinetics of, experiments of, reaction rate constant of, 265 rate constant of, comparative data of in SCISR and STR, table of, 266 Experimental equipment for FGD, 165
364 for hydraulic resistance measurement, 96 for micromixing measurement, 222 for pressure fluctuation measurement, drawing of, 240 for reaction-precipitation, 273 Experimental procedure for hydraulic resistance measurement, 97 for nano copper preparation by reduction-precipitation, 288 for preparation of white carbon black, 274 of micromixing measurement, 226 External circulation, 125 Falling down region, 71 Feed flow rate ratio, influence of, plot of, 227 Feed rate of material, influence of, 145 Feed rate, influence of, 146, 147 Feeders, of particles, 96 Fick's law, 192 Field force, 43 Fine particles, surface deactivation of, 236 First deceleration stage, 48, 49 Flash-impinging stream drying, 128 Flow behavior, of free impinging, 23 Flow configurations, 91 extension in, 9 influence of, 161 of IS, complexity of, 90 Flow field, 20 Flow of continuous phase, inside device, 10 Flow passage, sudden contraction of, 95 Flow rates, variations of, 82 Flow regimes, 94 Flow regions without mixing, in SCISR, 219 Flow rotating chamber, in nozzle, 170 Flow spaces, 69 Flow stability, 79 Flow swirling plates, 123 Flows at high velocity, 91 Flows, laminar, turbulent, 20 Fluctuation, axial symmetry of, volumetric intensity distribution of, data plot of, 246 Fluctuation intensity
IMPINGING STREAMS amplitude of, frequency of, statistical analysis of, 238 profile of on various planes, data plots of, 243 profile of on various planes vertical to x - z plane, data plots of, 244 volumetric distribution of, 245 • - , Flue gas desulfurization (FGD), 162 Flue gas from coal-burning, 162 Fluidized bed crystallizer (FBC), 259 drawing of, 261 Fluidized-bed jet mill, drawing of, 202 Flying trajectories in CISD, of particles with various diameters and moisture contents, 149 Fourier Transformation, 239 Frequency of fluctuation characteristics of, 250 variation tendency of, 249 Frequency spectrum, 239 Friction, 66 Friction coefficient, 92, 100 Friction force, 56 between gas flow and droplets, 108 of fluid, 43 Gas and liquid, differences between properties, 208 Gas flow, effect of, 104 Gas flow velocity, operational range of, 59 Gas-continuous impinging streams (GIS), 67, 89, 151,154, 201 characteristics of, 17 Gas-film mass transfer coefficient, kc, 183 Gas-liquid reaction systems, natures of, 153 Gas-liquid two-phase streams, 160 GIS, 117 GIS device, hydraulic resistance of, 92, 105 GIS gas-liquid reactor major features of, 172 structure and major dimensions of, 171 GIS with liquid as dispersed phase, application of, 107 Governing variable, 216 impinging velocity u0, measurement of, 224 Granular materials, impinging stream drying of, 123
SUBJECT INDEX Gravity, 148 action of, 43 effect of, 52, 56 influence of, 44, 58 Grinding and milling, 61, 65 Haining window thnction, length of, 239 Heat and mass transfer, 151 Heat and mass transfer coefficients of; flux of, 2 driving forces for, 145 effect on, 41 enhancing, 137 interlace area lot, 157 rate of, enhancement of, driving force of, interlace tot, specific resistance tot, 1 Heat transfer coefficient, 6 measured for, 126 High turbulence, 4 Highly chlorinated PV crystalline, 150 Hold-up of dryer, 126, 127 Hopper, glide of, 78 Horizontal two impinging stream contactor, 91,97 Horizontal two impinging streams, flow configuration of, 96 Horizontal-coaxial tour impinging streams, multiple groups of, flow configuration of, 187 Hydraulic resistance, 48, 125, 134, 182 of FGD device, 174 Hydrodynamic interaction, 33 Identical velocity stage, 48 Imaging analysis, I 10 Impacting drying, 124 Impingement, 6 between opposing streams, 4, 92 intensity of, 113 Impingement chambers, 131 Impingement plane, 4, 17, 22, 32, 42-44, 50, 54, 58, 61, 65, 67, 88, 89, 95, 114, 143, 144, 200 Impingement zone major active region, 70 new definition of, volumetric fraction of, 246 pseudo-boundary of, 71 volumetric fraction of, f,,, 231
365 Impinging distance, 32, 48, 97 of CISD, influence of, 144 dimensionless, 104 influence on micromixing, 234 influence of, 99, 143 Impinging distance, S, 50, 96, 114, 117 effect of, 113 Impinging jet, IJ, 3 Impinging stream absorber operated in bobble mode, 159 co-axial, with and without partition, concentric and eccentric, 158 Impinging stream absorption, 14 Impinging stream combustion, 13, 191 Impinging stream contactor feeding or accelerating tube of, 79 response of, 86 Impinging stream crystallizer (ISC), 259 Impinging stream crystallizer drawing of, 335 experimental system scheme of, drawing of, 260 instruction of, 334, 336 introduction to, 333 Impinging stream device, 60, 68 flow configuration of, 89 hydraulic resistance of, 91 response of, 83 Impinging stream drying, 14, 119 Impinging stream gas-liquid reactor, 169 Impinging stream grinding, machines for, 201 Impinging stream loop reactor, 159 Impinging stream milling, 14 Impinging stream reactor without circulation, micromixing in, 233 Impinging stream spray drying, 121 Impinging stream steam boiler, of B KZ320-140GM type, 199 Impinging streams, IS, 3 adaptability of, 153 application status of, 12 classification of, 207 dispersed phase of; periormance of, 107 horizontal, 20 intensifying combustion, mechanisms of, experimental evidence, 197 outstanding advantages and intrinsic faults, 119
366 properties of, 13, 151 supplementary classification of, 211 three dimensional, 30 various flow configurations, 9 Impinging velocity, 21, 22, 126, 129, 186 influence of, 99, 145, 146, 250 influence on critical frequency of fluctuation, 249 influence on hydraulic resistance, 105 influence on micromixing time, 228 influence on volumetric mass transfer coefficient kGa, 184 in SCISR, measurement and control of, 241 Impinging velocity u0, 92, 102, 161,238, 262 influence of, 183 influence on integral intensity, data plot of, 248 influence on intensity at the most intensive points, data plot of, 247 measurement, data plot of, 225 related to rotary speed of propellers N, data plot of, 240 Impulse and step change, 77 Inclined U-shape tube, for hydraulic resistance measurement, 97 Incomplete combustion, 199 Incomplete mixing, influence of, 215 Inertia, 41 Initial condition, 85 Inlet concentration, of particles, 62 Input signal, 77, 78, 83, 90 Input signals to impinging stream device, 80 Input-response technique, 218 Instantaneous and irreversible reaction, 163 Intensive fluctuation region, 242 form of, 244 profile in the space, drawing of, 245 Interaction, 45 Interaction between particles and gas, 66 Interaction between particles, 60 Interaction between streams, of collision, shearing and pressing, 24 Interaction between gas and liquid, 185 Interaction between gas flow and particles, between gas flow and surrounding gas, 52
IMPINGING STREAMS Interaction between opposing streams, 12, 20, 209, 232 Interface area, 61,107, 155, 178 Internal circulation, 160 effect of, 140 Internal mixing nozzles, 118 Internal parts, 168 Interparticle collisions, 131 Irrotational flow, 29, 43 IS application, feasibility of, target system for, 106 IS combustors, 198 IS contactor, comparison of with pneumatic flash dryer, 106 IS contactor, overall resistance of, 96 IS dryers, 120 IS drying combinations, 128 IS drying, of special materials, of crystalline lysine, 132 IS drying-milling, 131 IS enhancing absorption, experimental evidence for, 160 IS enhancing transfer, effectiveness of, 124 IS gas-liquid reactor, resistance of, 185 IS mill dryer, 130 Isobars, 33 Jet impingement, structure of, free and submerged, 233 Kinetic energy, 42, 44, 51 of gas streams, 130 Kinetics, molecular collision theory of, 253 Known arbitrary function of time, 90 Known function of time, 78, 79 Kolmogoroff micro scale, ,~, 24, 214, 254 Koppers-Totzek gasifier, 3, 199 drawing of, 200 Laminar impinging streams, 25 Laplace transformation and inversion, 76 of PTR in SCISR, 221 Laplace transformation, 81 Large gas flow rate Laser-Doppler velocity meter, 37 Lengthening residence time, 133 Liquid and gas flow rates, influence of on absorption rate, 161 Liquid droplets, re-atomization and/or coalescence of, 65 Liquid film, thickness of, 6L, 154
SUBJECT INDEX Liquid flow rate, effect of on pressure drop, 185 Liquid fuel, 192 boiling point of, atomization of, 194 Liquid to gas mass flow rate ratio, influence of, 117 Liquid/gas flow rate ratio, influence of on sulfur-removal efficiency, 178 Liquid-continuous impinging streams (LIS), 205 prospects for application of, 337, 338 adaptability of for preparation of ultrafine powders, 270 features of, 271 major features of, influence of on process kinetics, 253 progress of investigations on, 207 Liquids assembly condition and intermolecule force of, 107 insoluble, miscible, 213 properties of, 205 Local concentration difference, 271 Local resistance coefficient, ~ 102, 103 Local resistance coefficient, combined, 95 Lower distance of CISD, influence of, 145 Lower spacing of CISD, influence of. 144 Macro- and micro-mixing, relationship between, 232 Macromixing time, measurement of, 218 Magnus effect, 61 Majac jet pulverizer, drawing of, 203 Mass and heat transfer, 119 Mass transfer coefficient, 107, 108, 174, 186 in LIS, 209 interpretation of, 182 volumetric, experimental data of, 161 Mass transfer, enhancement of, 6 Mass transfer model, solution of, 175 Material in dispersed phase, mean residence tome of, mixing of, 89 Mathematical theorem of proportion by addition, 82 Maximum depth of penetration, x ....... 5(i) Maximum distance of penetration, x ....... 42, 51,53 into opposing stream, 52 Maximum pressure, 33
367 Maximum radial velocity, variation of, 35 Mean diameter of droplets, Sauter, expression for, 115 Mean number of particles per unit volume, 64 Mean relative velocity of particle, 64 Mean residence time, 70, 74, 75, 134 in active region, 77 in four sub-spaces, 76 in impingement zone, 84 in SCISR, 217 in various sub-spaces, total, 88 Mean time between successive collisions, 64 Metastable region, 255 of Na2HPO4 solution, measurement of, 257 Method for measurement of RTD, 77 Micro- and macro-mixing, dependence between, distinguishing of, 213 Micro photographing, 110 Micro pressure probes, Model ACQ-062, 240 Micromixing, 210 influence on kinetics, 266 influence of, 267 in IS, 12 in LIS, comparison of investigations, dependence on impinging velocity, 235 in SCISR, 222 in TIJ mixer, data interpretation of, relationship of, 234 major results of, 226 measurement of in SCISR, system scheme drawing, 224 mixing on molecular scale, 253 molecular scale, 7 performances, comparison of between SCISR and STR, 229 Micromixing time, tM, 216 calculated and measured, data table of, 231 comparison of between measured and predicted, 230 determination of, 228 deviation of predicted from measured, 232 measurement of, chemical method for, 215
368 relationship of versus impinging velocity u0, data plot of, 229 Millets and rapeseeds, 97 properties of, 80 Milling, Trust Jet Mill, 8 Minimum gas velocity, 59 Mirror Image model, 19, 251, 32 Mixing, 6 between two streams, in impingement, 23 in active region, 154 methods for investigation of, 214 scale of, model of, 213 Mixing chamber, in nozzle, 157 Model for IS enhancing combustion, 197 Modeling-simulation, in impingement zone, 68 Moisture content profile, in annular chamber of CISD, 142 Moisture in pores of porous material, 140 Moisture removal efficiency, 124 Mole ratio Ca/S, critical value of, 173 Molecular collision, 254 Molecular collision theory, 264 Molecular diffusion, 24 Molecular motion, 66 Momentum transfer, 24 between gas and particles, 93 between particles, 43 Momentum transfer, intensity of, 210 Monte-Carlo simulation, 66 Motion equation, 57 for single particle, 44 initial conditions for, 93 Motion of a single particle, 41 Motion time, 49, 51 Movement equations of particles in CISD, 148 Movement of particles, 69 Moving direction, 56 Multifunction combination, 130 Multilayer structure of IS dryer, 128 Multiphase, 7 Multiple accelerations-decelerations motion, 51 Multistage countercurrent system, 89 Multistage drying system, 134 Multistage impinging streams, 91 Na2HPO4, crystals of, 256
IMPINGING STREAMS Na2HPO4 solution, metastable region of, data plots of, 259 Na2HPO4.2H20 crystalline and mother liquor, compositions of, 257 Na2HPO4.2H20 crystalline, crystal-growth rate measurement of, 259 Na2HPO4.2H20, crystal-growth rate coefficient of, comparative data in ISC and ISC, table of, 263 Na2HPO4.2H20, experimental procedure for crystal-growth rate measurement of, 260 Na2HPO4.2H20, overall crystal-growth rate coefficient of in ISC, data table of, 261 Nano copper chemical reactions in preparation of by reduction-precipitation, 286 major uses of, 285 of needle form, TEM photo of, 292 properties of, 284 surface improvement of--preparation of Cu-Ag double metal powder, 297 TEM photo of, 289 X-ray spectrum of, 291 Nano copper preparation comparison between various technologies and devices for, 296 data table, 297 influence of CuC12 concentration in, 292 influence of CuCl 2 concentration in, data plot of; influence of reaction temperature in, 293 influence of feeding mode in, 291 influence of impinging velocity in; influence of amount of surfactant in; optimal conditions for, data table of, 295 influence of reaction temperature in, data plot of; influence of pH in, 294 influences of various factors on, conditions for the factors, 290 phenomena observations in; optimal mole ratio for, 289 primary results of, 288 reducing agent for, 286 repeated experiments under optimal conditions of, TEM photo of, 296
SUBJECT INDEX Nano hydroxyapatite (HAP) composition, properties and uses of: methods for preparation of, 317 FITR and XRD spectrums of, 325 influences of some factors in, 320 optimal conditions for synthesis of: characteristics of, TEM photo of, 324 preparation, experimental procedure and conditions for, data table of, 319 preparation of by double decomposition-precipitation: equipment for, 318 TEM photos of products obtained at different pH, 322 TEM photos of products of, influence of dripping rate in preparation of, 323 TEM photos of products of: influence of pH in preparation of, 321 XRD spectrum data list of: concluding remarks for preparation of by double decomposition-precipitation, 326 Nano materials introduction to properties and applications of, 283 preparation of, 8 preparation of by reactionprecipitation, 208 Nano TiO2 preparation conditions for and results of second stage experiments of, 307 conditions for first stage study of. data table of, 305 effecting factors examined in; major results of first stage study of, 304 experimental equipment and procedure for, 303 influence of neutralization in, data plot of, 312 major reactions in, 302 optimization of conditions lbr, 313 results of first stage of: data table od, 306 results of second stage of, data table of, 308 results under optimized conditions of: conclusions of, 314
369 Nano TiO2 mass preparation, results of, data tables of 309, 310 Nano TiO2, TEM photos of, 311 Nano Titania TiO2, properties of, uses of, 301 Navier-Stokes equation, 31 Newton's law, 210 Newton's motion equation, 69 Non-elasticity collision, 66 Non-equivalence, of time, 67 Nozzles, 202 position, influence of, 182 Nucleation, 254 rate equation of, rate constant, 270 Nucleus, 271 Number fiequency function,,/~, 111 Number of particles, reduction of, 63 Numerical characteristics of distribution, 67 Observed active energy, comparative data in ISC and ISC, table of, 263 One-stage tangential horizontal tour IS dryer, 127 Operating conditions, contradictory effects of, 88 Operation and structural parameters of FGD device, influences of, 167 Operation elasticity, 186 Operation mode, 10 Optimal conditions, for FGD, 168 Oscillation, 41 of damped magnitude, 43 period of, 20 Oscillation movement, 44, 62 of penetration, to and fro between the opposing streams, 23 with damped amplitude, 51 Oscillation of particles, 6 Oscillation times, number of, 52 Overall pressure drop comparison between measured and calculated, 104 model for, 103 Overall residence time distribution, 84 model for, 75 Oxidant, diffusion equation for, 195 Parallel heat and mass transfer, 120 Parclose-like layer of particles, 144
370 Particle acceleration, 45 Particle concentration distribution of, influence of, local, 61 in feed streams, influence of, 63 Particle crowds, 61, 67 behavior of, 60 Particle diameter, influence of, 58 Particle motion equation for, 46 relationship for, 43 Particle velocities, variations of, 54 Particle velocity, 45 Particle velocity radial and axial profiles of, 61 Particles A and B, individual flow rates of, 82 Particles acceleration of, 99 average velocity of, mean diameter of, Reynolds number of, Rep, 93 effect of, density and mean diameter of, 104 recovery efficiency of, residence time of, 129 Particle-to-gas mass flow rate ratio, 97 influence of, 101 Particle-to-gas volumetric ratio, 61 Partition, 6 Passage area, 71 Penetration depth of, 43 of particles, 4, 5, 61 relative depth of, 53, 54 to and fro between opposing streams, 89 Penetration-oscillation, 42 Perfect mixing, 70 with a certain lag time, feature of, 87 Perfect mixing-plug flow, characteristics of, 77 Performances of impinging streams, influence on, 11 Phase condition, extension in, 11 Pillar coordinate system, for fluctuation measurement and analysis, 241 Planar two dimensional impinging streams, 26, 36 Plug flow, 70 Pneumatic atomizers, of external and internal types, 157
IMPINGING STREAMS Pneumatic nozzles, 168 Porous particles, drying of, 89 Powdery/granular solids drying of, 134 Power consumption, 91, 106 evaluation of, 105 for SCISR operation, 226 Power spectrum, analysis model for, 239 Preparation of "Ultrafine" white carbon black, normal-designed experiments for, semi-batch operation for, 275 Pressure across impingement zone, 103 Pressure distribution, 27 Pressure drop, 91, 97, 126 due to acceleration and collisions of particles, 100 due to acceleration of particles, 93 resulting from accelerating particles; caused by impingement between opposing streams, 94 resulting from impingement, caused by structural factors through accelerating tube, 98 through accelerating tube, 101 Pressure drop behavior, due to acceleration and collision of particles, 102 Pressure drop distributions, characteristics of, 98 Pressure fluctuation in SCISR, meaning of, investigation method for, 237 Pressure fluctuation, 7, 24, 61,210 arrangement of measuring points and sampling frequency for, 241 arrangement of measuring points for, drawing of, 242 existence of due to impingement between opposing streams, effect on micromixing condition, 251 in IS, 12 influence on kinetics, 266 influence of, 267 investigation method for, 238 major energy of, 249 power spectrum of at different points and different u0, data plots of, 249 Pressure profile, in impinging streams, 32 Pressure, critical value of, 194 Pressure, variation of, 34 Primary atomization, 107 Process particles B, flow rate of, properties
SUBJECT INDEX of, 79 Processes for FGD, dry, semidry and wet, 162 Processing degree, 120 Product discharge position of CISD, arrangement of, 147 Product fineness, 204 Propellers, 217 rotary speed of, 22 Property differences, between liquid and gas, 11 Proportion factor, a, 94 Pseudo flue gas, 179 PTR function, in SCISR, theoretical model for, 222 PTR in SCISR, results measured and calculated, plots of, 223 PVC drying, in CISD, experimental data of, 140 PVC temperature, variation of, 143 Radial concentration gradient, 61 Radial flow outwards, velocity of, maximum value of, 71 Radial gas flow, 147 Radial gas velocity and radial distance, relationship, between, 71 Radial velocity, 43, 51 distribution of, 33 distribution of, maximum, 34 Radiation-labeled particle, [3-active isotope TI e~14,61 Rapeseeds, drying of, 133 Rates, reaction of, diffusion of, 154 Raw gases, composition of, 200 Reaction conversion, increment of, 67 Reaction resistance, 196 Reaction scheme in wet FGD, 162 Reaction-precipitation, 8, 269 Reactions in liquid phase, features of, 154 Reactions, last and in'eversible, restricted by equilibrium, 90 Reactions of dissolved SO~ with Ca(OH)e, 163 Reactor and system design for wet FGD, conditions tot, 170 Reactor for wet FGD, overall performance of, 177 Reactor, with impinging streams, 8 Real kinetics data, 267
371 Re-atomization, 41 Regions, contributions of to absorption, 166 Regression of data, linearized, 100 Regressive equation, 116 Relationship between gas-film transfer coefficient and impinging velocity in FGD, 183 Relationship for fitting data, 84 Relative or slip velocity, 44 Relative velocity, 2, 17, 42, 46, 59, 198 between particle and gas flow, u,., 58 influence of, 156 Reproducibility of data, 80 in measurement of super solubility, 258 Residence time, 43, 54, 125, 127 of materials, 120 of particles, 5, 11,51, 52, 54, 61,65 Residence time and distribution, characteristics of, 77 Residence time distribution (RTD), 68, 74, 75 definition of, 73 of particles, 67 nature of, 76 Residence time distribution density probability (PTR), function of for SCISR, 220 Residence time distribution function, F(t), 85 Residence time distribution, model of for SCISR, 219 Residence time distribution probability function, E(t), 70, 75 Residue, processing of, 129 Resistance due to structure of device, 102 due to structure, 95 of accelerating tube, to pure airflow, 99 of reaction kinetics and diffusion, 195 majority of, 98 to gas flow, 92 variation of with impinging velocity, 185 Response of screw feeder to step change, 80 Reynolds number, 117
372 inapplicability of for scaling-up, 224 Reynolds number, Re, 31, 92, 235 Reynolds number, Rep, 45 influence of, 100 Reynolds stress equation, 37 Rigid wall, jet impinging on, 23 Rotary speed of screw, 80 Rotating packed bed (RPB), 3, 184 Rotation IS, flow configuration of, 160 Roughness, 92 RTD constituents of, contributions of various space to, 69 function, definition of, 82 measured and simulated, comparison, 87 measurement of, 82 of solid particles, 90 Sampling frequency in pressure fluctuation measurement, 239 Sauter mean diameter, 118 comparison between measured and calculated, 116 Scaling-up, criterion for, 224 Scheme of overflow, 147 SCISR comparison of with stirred tank reactor (STR), plot of data, 218 design of, drawing of, 217 dimensions of drawing, 223 flow configuration of, 219 simplified flow model of, 220 Screw feeder, 78 response of, 79, 86 Segregation index, Xs, 216 comparison of between SCISR and STR, data plots, 230 variation of with u0, 227 Segregation scale, 24, 214, 254 intensity of, 215 Semi-batch IS drying, 132 Sewage with high moisture content, drying of, 129 Shearing force, 160 Single droplet, evaporation-burning equations for, behavior of, 191 Single particle, burning equations for, 194 Single spherical droplet, symmetrical burning of, diameter of, 192
IMPINGING STREAMS Single-phase impinging streams, 21 Size distribution, 66 of droplets, measurement of, 174 of droplets, before and after impingement, 111 of particles, 69 Slide-sampling, arrangement of, 110 SO2 absorption, with limestone and NaOH enhanced limestone suspensions, 168 concentration, influence of on gasfilm transfer coefficient, 181 concentration, influences of, 180 concentration, variation of, 176 removal efficiency, variation of, asymptotical tendency of, 166 removal efficiency, r/s, 173, 174 Space inert for transfer, 75 Specific effective power, influence of on xs, 230 Spouted bed with central conduit, combination of with IS, 137 Spray droplets Sauter mean diameters of, 177 sizes of, 176 Spray dryer, with rotating impinging streams, 122 Sprays of fine droplets, 157 Staged integration, 46, 49 Stagnation jet mixer, 7 Stagnation plane, 37 Standard deviation, ~ 112, 114, 116 Static pressure, axial profile of, 99 Step change, 78 Stirred tank reactor (STR), for comparative experiments, drawing of, 274 Stokes regime, 60 Stokes, transient and turbulent regimes, 45-47, 49, 50, 53, 57, 58 Stream function, 29, 30 Streamlines, in two dimensional impinging streams, 27 Strongest fluctuation points, profile of along y-axis, data plot of, 245 Structural and operating parameters of CISD, influences of, 142 Structural factors, 92 Submerged circulative impinging stream reactor (see also SCISR), 21, 24, 216
SUBJECT INDEX Sub-pressure drops, 92 Super solubility, 255 equipment for measurement of, drawing of, 258 Supersaturation, 236, 255,270 Surface tension, 110 Suspensions solid-in-gas, 4 solid-in-liquid, 5 solid-in-gas and liquid-in-gas, thin, 41 System, complexity of, 122 System engineering, ideas of, 120 Systems of water-air and water-CO~, 115 Tangential horizontal flow multistage dryer, 128 Tangential horizontal IS dryer, 125 Tangential nozzles, 123 Target material of IS application, selection of, 151 Target system for IS application, selection of, 154, 155 Technological-economic indexes, 120 Temperature, gradient of, 2 Terminal velocity, 58-60 Thin dilution, 60 Ti, A1 and polystyrene particles, motion of, 54 TIJ mixer, micromixing determination of, major results, 234 Time domain of t _>0, 82 Time for complete burning, tb, 193, 196 Time-scattered measurement, 83 Total mean residence time, lengthening of, 137 Tracer A, step change of, 86 Tracer concentration, 80 measurement of, 86 Tracer particles A, flow rate of, concentration of, properties of, 79 Tracer particles, concentration of, 78 Trajectories of particles, 53, 57 of a single particle, 56 particles moving along, 72 Transfer, enhancement of, 209 Transfer between phases, 89 between gas and solid, 7 enhancement of, 17, 106
373 Transfer coefficients comparison between with and without IS, 124 in LIS, 208 Transfer lag or pure lag time, 75 Trial-and-error, 86 Trost jet mill, 202 drawing of, 203 Turbulence, 17, 39, 213 Turbulent degree, measure of, 36 Turbulent impinging streams, 36 Turbulent intensity, measure of, 36 Turbulent pulsation, 54 Turbulent regime, 69 Turning tube, 95 Two co-axial-cylindrical jets, impingement of, 36 Two impinging jets (TIJ) mixer, drawing of, 233 Two impinging streams, open device of, 118 Two IS dryer, with two pair of air-feeding tubes, 127 Two single-phase flows, impingement of, 19 Two-stage drying mechanism, 143 Ultrafine powder, preparation of, 13,217 Ultrafine products, preparation of, 235 "Ultrafine" white carbon black preparation Chinese national standard of, preparation methods of, common (one-step) precipitation process for preparation of, 272 comparison between size distributions of products from SCISR and STR in, data plots of, 28O final treatment of reaction product in, 280 in impinging streams, 269 primary determination of optimal conditions for; experiments in continuous operation of, 278 results description of, influences of effecting factors in, 277 results of normal-designed experiments, data table, 276 results of spray drying in, data table of; conclusions for, 281
374 results on influence of Na2SiO3 concentration in; comparative experiments in SCISR and STR in semi-batch operation in, 279 Uniformization of droplet sizes, 118 Velocities, local and average, 54 Velocity of airflow in accelerating tube, 97 of gas flow, 50, 52 of gas flow in accelerating tube, influence of, 100 of gas flow, influence of, 58 of particle, 51, 52, 56 of particles at outlet accelerating tube, influence of, 101 Velocity equation, 45 Velocity field, mean, time-averaged, 39 Velocity potential, 27 Velocity, relative, 5 terminal, fluidizing, liquid-flooding, 3 Vertical circulative impinging stream reactor drawing of, 330 instruction of, 331,332 introduction to, 329 Type II, drawing of, 333 Vertical co-axial turbulent impinging streams, 37
IMPINGING STREAMS Vertical gas-solid impinging streams, 59 Vertical impinging streams, 57 Vertical two impinging streams, 135 Vertical velocity, 72 Vibrations of fluid elements, 254 Viscosity, influence of, 31 Viscous impinging streams, 31 Volumetric evaporation intensity, Ev, 142, 144 Volumetric flow rate, 91 Volumetric mass transfer coefficient, in FGD, 176 Volumetric mass transfer coefficients, comparison between in IS gas-liquid reactor and in rotating packed bed, 184 Vortex-type burners, 199 Vortices, formation of, 21
Water, surface or flee, in pores, crystallized or bounded, 134 Wet dust-removal, function of, 172 Wet FGD experimental scheme and procedure of, 172 experimental scheme of, drawing, 173 investigations in China, 169 investigations in Israel, 164 White carbon black, properties of, 271
