The Canadian Journal of Chemical Engineering

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1 The Canadian Journal of Chemical Engineering A U G U S T V O L U M E 8 3 N U M B E R 4 Message from the Incoming Editor K. Nandakumar Articles A Review of Recent Technological Advances in the Brightening of High-Yield Pulps Yonghao Ni Review on Mixing Characteristics in Solid-Liquid and Solid-Liquid-Gas Reactor Vessels Gopal R. Kasat and Aniruddha B. Pandit Hydrodynamic Characteristics of a Powder-Particle Spouted Bed with Powder Entrained in Spouting Gas Qunyi Zhu, C. Jim Lim, Norman Epstein and Hsiaotao T. Bi Écoulement de Taylor-Couette en géométrie finie et à surface libre Mahamdia Ammar, Bouabdallah Ahcène et Skali Salah Eddine The Impact of Local Phenomena on Mass Transfer in Gas-Liquid Systems Olaf Bork, Michael Schlueter and Norbert Raebiger On Criteria for Occurrence of Azeotropes in Isothermal and Isobaric Binary Systems Ronald W. Missen Numerical Simulation of Pollutant Formation in Precalciner Lai Huang, Jidong Lu, Shijie Wang and Zhijuan Hu Dibenzothiophene HDS Over Sulphided CoMo on High-Silica USY Zeolites Gerardo Lara, José Escobar, José A. De Los Reyes, María C. Barrera, José A. Colín and Florentino R. Murrieta Heat Removal from Reverse Flow Reactors Used in Methane Combustion Sukumar Balaji and S. Lakshminarayanan Simulation of Radiation Field in Multilamp Photoreactor Using the LSDE Model Quan Yang, S. O. Pehkonen and Madhumita B. Ray A Comparison of PID Controller Tuning Methods Michael W. Foley, Rhonda H. Julien and Brian R. Copeland Attainment of PI Achievable Performance for Linear SISO Processes with Deadtime by Iterative Tuning D. B. Goradia, S. Lakshminarayanan and G. P. Rangaiah Recyclage d un déchet, une boue rouge, comme catalyseur pour l élimination des composés organiques volatils J. F. Lamonier, F. Wyrwalski, G. Leclercq et A. Aboukaïs Méthodologie d étude et de modélisation de la déphosphatation d effluents aqueux L. Montastruc, S. Domenech, L. Pibouleau, C. Azzaro-Pantel The Effect of Particle Size Distribution on Pressure Drop through Packed Beds of Cooked Wood Chips Quak Foo Lee and Chad P. J. Bennington R&D Notes A Study on Hydrodynamics and Heat Transfer in a Bubble Column Reactor with Yeast and Bacterial Cell Suspensions Nigar Kantarci, Kutlu O. Ulgen, Fahir Borak Décantation assistée par ultrasons des eaux résiduaires Nicolas Augis, Philippe Hus, Pierre-Xavier Thivel et Jean-Yves Viau Dispersion and Mass Transfer Effects on the Performance of an Immobilized Lipase Packed Bed Reactor During the Hydrolysis of Rice Bran Oil V. Ramachandra Murty, Jayadev Bhat, P. K. A. Muniswaran and S. Sivasankaran Challenging the Traditional Counter-Current Water Circulation System Case: Washing of Soap in a Kraft Pulp Fibreline Kari Ala-Kaila, Outi Poukka, Pekka Tervola 2005 Published by the Canadian Society for Chemical Engineering Publié par la Société canadienne de génie chimique CJChE A7(6) ISSN

2 The Canadian Journal of Chemical Engineering A U G U S T V O L U M E 8 3 N U M B E R 4 D e v o t e d t o t h e p u b l i c a t i o n o f c h e m i c a l e n g i n e e r i n g s c i e n c e, i n d u s t r i a l p r a c t i c e a n d a p p l i e d c h e m i s t r y Editor K. Nandakumar, Department of Chemical and Materials Engineering, University of Alberta, Edmonton, AB Associate Editors R. E. Hayes, Department of Chemical and Materials Engineering, University of Alberta, Edmonton, AB B. Huang, Department of Chemical and Materials Engineering, University of Alberta, Edmonton, AB R. S. Sanders, Syncrude Canada Ltd., Edmonton, AB Editorial Assistants S. Peake Editors Emeriti L. W. Shemilt, N. Epstein, C. W. Robinson and P. Carreau Editorial Board John J. Carroll, Gas Liquids Engineering, Calgary, AB (2001) Michel Perrier, École Polytechnique, Montréal, QC (2001) Ron H. Crotogino, PAPRICAN, Pointe-Claire, QC (2001) Thomas-J. Harris, Queen s University, Kingston, ON (2001) Alexander Penlidis, University of Waterloo, Waterloo, ON (2001) Ex-Officio Milena Sejnoha, NRCan, Ottawa, ON (2004) Souheil Afara, University of Western Ontario, London, ON International Advisory Board G. Chen, Zhejiang University, Hangzhou, China J. B. Joshi, University of Bombay, Bombay, India F. Coeuret, École Louis de Broglie, Bruz, France A. W. Nienow, The University of Birmingham, Birmingham, U.K. D. De Kee, Tulane University, New Orleans, LA, U.S. N. Wakao, Yokohama National University, Yokohama, Japan D. Dochain, CESAME, Université Catholique de Louvain, Louvain-la-Neuve, Belgique A. Tamir, Ben Gurion University of the Negev, Beer Sheva, Israel A. A Adesina, University of New South Wales, Sydney, NSW, Australia The Canadian Journal of Chemical Engineering is published every two months by the Canadian Society for Chemical Engineering, a constituent society of The Chemical Institute of Canada, 130 Slater Street, Suite 550, Ottawa, ON, Canada K1P 6E2. Formerly published as The Canadian Journal of Technology (Vol. 1 35). Unless it is specifically stated to the contrary, the Society assumes no responsibility for the statements and opinions expressed herein. Subscriptions are payable in advance print subscription rates: Canada Individual CAN$285, Institution CAN$335; U.S. and Foreign Individual US$285, Institution US$340. Members $75. For manuscript submission instructions, see the inside back cover. We can print images in colour for a fee of $2,000, payable in advance. Authors of papers are assessed a page charge of $30 per printed page to help cover costs of publication. While it is not mandatory for authors to pay this charge, the publisher believes that granting or sponsoring organizations of an author s work accept this view that some portion of the cost of preparing the paper for public distribution is a fair charge against a research budget. Authorization to photocopy items for internal or personal use, or the internal or personal use of specific clients, is granted by the Canadian Society for Chemical Engineering for libraries and other users registered with the Copyright Clearance Center (CCC) Transactional Reporting Service, provided that the base fee of $1.00 per copy, plus.25 per page is paid directly to CCC, 21 Congress Street, Salem, MA 01970, USA /83 $ Advise Production Office in Ottawa, in advance of change of address, giving old as well as new address, and subscriber I.D. no. Enclose address label if possible. Claims for missing issues will not be allowed if received more than 60 days from date of mailing, plus time normally required for postal delivery of journal and claim. No claims allowed because of failure to notify the circulation department of a change of address, or because copy is missing from files. Indexed by Engineering Index Inc. Postage paid in cash in Ottawa, ON. Publications Mail Registration Number The Canadian Journal of Chemical Engineering is published six times per year. Printed in Canada by Gilmore Printing Services Inc. Production Office 130 Slater Street, Suite 550, Ottawa, ON, Canada K1P 6E2 Publishing Editor: M. Piquette Production Coordinator/Graphic Designer: R. Lalonde Translator: B. Sagnier We acknowledge the financial assistance of the Government of Canada, through the Publications Assistance Program (PAP), toward our mailing costs.

3 MESSAGE FROM THE INCOMING EDITOR This issue marks a new beginning for The Canadian Journal of Chemical Engineering, as my role as the new Editor is reflected in this issue s cover. My new Associate Editors and I, along with our Editorial Assistant, are excited to have the responsibility and opportunity to guide the Journal through its next phase of evolution. We are cognizant of the revolutionary changes that are taking place in the publication industry. Examples include the prevalence of on-line manuscript submission and review systems and the ever-growing importance of electronic publications. The recent and upcoming changes in the Journal signify our commitment to position it at the forefront of chemical engineering research publications. Now, more than ever, we need your support as authors, as readers and as reviewers. At this juncture I wish to thank Pierre Carreau for more than eight years of dedicated service to the Journal. I am inspired by his contributions and by those of his predecessors, Cam Robinson and Norman Epstein. Each of these individuals has left his mark on the Journal; and each has spent valuable time with me and given invaluable advice as I prepared to take this job. For these contributions, I am most grateful. Over the past year we have made significant progress in two areas: we have cleared a substantial backlog of manuscripts and have rolled out an on-line manuscript submission and tracking system. While the on-line system helps to reduce the time between submitting a manuscript and receiving a decision, the success of this endeavour ultimately depends upon the cooperation of researchers like you acting as reviewers for the Journal. During the past seven months, we have made decisions in as little as 17 days; yet for some manuscripts, it has taken more than six months to reach a decision. If you are called upon to review an article, please accept the call and be co-operative in following the suggested time line. Without your help and timely input as reviewers, it will not be possible to improve the quality of the Journal and provide timely service to the authors and readers. Although we have cleared the backlog of manuscripts to be published, we expect that the typesetting and publication steps will still take six to eight weeks for each issue. The time delay associated with these steps is of particular concern to us. We are presently exploring the possibility of collaborating with established, professional publishing organizations to expedite this part of the process. No decisions have been made and I welcome your input on the direction that the Journal should take in the coming years. For those of you, who are unfamiliar with our on-line publication, please note that it is available on our Web site: If your library does not have a subscription to the on-line version of the Journal, please encourage them to get one. We are also exploring the possibility of making all past issues available on-line. Another area of renewal for the Journal is its International Advisory Board. I thank the members of our International Advisory Board for their service to the Journal. Some have served for a very long time. During the coming year many outstanding, internationally recognized researchers will replace our long-serving members. I promise that its membership will reflect the usual high standards expected of the Journal. In addition, the new members varied areas of expertise will embody the fundamental changes taking place in chemical engineering research around the world. On a similar theme, I have begun to invite contributions from leading chemical engineering researchers to a new series entitled, Frontiers of Chemical Engineering Research. I expect these contributions to focus on the interdisciplinary nature of chemical engineering research and to offer perspectives on particular research areas. These articles will describe the current state-of-the-art in a given field, where research is heading in the near future and its anticipated major impacts. The authors will be leading researchers in their respective fields and the topics will cover such diverse areas as microfluidics, molecular dynamics, nanomaterials, polymer composites, fuel cells, tissue engineering and regenerative medicine. The common theme among the articles in the series is the role that chemical engineering research plays in shaping these new frontiers. If you have an idea for such a topic or you are at the frontier yourself, please contact me directly. I look forward to assisting you in disseminating your research to the global chemical engineering community. In that process, we will together take The Canadian Journal of Chemical Engineering to new heights. K. Nandakumar Editor VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 609

4 REVIEW BASED ON 2004 SYNCRUDE INNOVATION AWARD LECTURE A Review of Recent Technological Advances in the Brightening of High-Yield Pulps Yonghao Ni* Limerick Pulp and Paper Centre and Department of Chemical Engineering, University of New Brunswick, Fredericton, NB, Canada E3B 5A3 Peroxide brightening/bleaching is an important unit operation for pulp and paper mills producing value-added paper grades containing bleached mechanical pulps. In this paper, a number of process innovations have been discussed. They include: hydrosulphite assisted chelation process, the so-called Q y process; DTPA spray; P N2 process; P M process; Mg(OH) 2 -based peroxide process. It will be shown that these processes can improve the peroxide brightening/bleaching processes in terms of (1) decreasing the production cost, (2) improving the product quality, and (3) decreasing the environmental impact. Le blanchiment ou lustrage au peroxyde est une opération unitaire importante pour les usines de pâtes et papiers produisant des grades de papier à valeur ajoutée à base de pâtes mécaniques blanchies. On examine dans cet article un certain nombre d innovations de procédés. Cellesci incluent : le procédé de chélation assisté à l hydrosulfite, appelé procédé Q y ; la vaporisation de DTPA; le procédé P N2 ; le procédé P M le procédé au peroxyde basé sur Mg(OH) 2. On montre que ces procédés peuvent améliorer les procédés de blanchiment ou de lustrage en ce qui a trait à (i) la diminution du coût de production, (ii) l amélioration de la qualité des produits et (iii) la diminution de l impact environnemental. Keywords: high-yield pulp, mechanical pulp, peroxide bleaching brightness The Canadian pulp and paper industry is the largest manufacturing industry in the country. Pulp fibres are currently obtained from lignocellulosic plants, such as wood, through either chemical pulping or mechanical pulping processes. In chemical pulping, the resulting pulp is called chemical pulp, and the kraft process is dominant. Sodium hydroxide and sodium sulphide are used in removing lignin (the inter fibre bonding material) so that pulp fibres can be separated from each other. In mechanical pulping, the resulting pulp is called mechanical pulp (also known as high-yield pulp), and fibre separation is achieved via mechanical procedures, such as grinding or refining. Usually, the chemical pulps are much stronger than mechanical pulps, but at the cost of much lower pulping yield. High-yield pulps (HYP), such as thermomechanical pulp (TMP), play an important role in the Canadian pulp and paper industry. This is largely due to the following two factors: (1) the electricity cost in Canada is significantly less than that in most other countries; and (2) Canada has very good wood species, such as spruce, which can produce high quality highyield pulps. * Author to whom correspondence may be addressed. address: yonghao@unb.ca 610 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

5 Table 1. Comparison of the residual Mn contents of the Q y and Q treated pulps (Ni et al., 1998) (spruce TMP, PH 5.8, 0.5% DTPA, 3% pulp consistency, 70ºC, 30 min) Sample ID Description Manganese contents (ppm) Unchelated pulp 144 Q 0.5 (washed) 0.5% DTPA, washed thoroughly after chelation 11 Q 0.5 (pressed) 0.5% DTPA, pressed to 30% pulp consistency after chelation 20 Q 0.2 (washed) 0.2% DTPA, washed thoroughly after chelation 20 Q 0.2 (pressed) 0.2% DTPA, pressed to 30% pulp consistency after chelation 35 Q y 0.5 (washed) 0.1% hydrosulphite, 0.5% DTPA, washed thoroughly after chelation 3 Q y 0.5 (pressed) 0.1% hydrosulphite, 0.5% DTPA, pressed to 30% pulp consistency after chelation 15 Q y 0.2 (washed) 0.1% hydrosulphite, 0.2% DTPA, washed thoroughly after chelation 8 Q y 0.2 (pressed) 0.1% hydrosulphite, 0.2% DTPA, pressed to 30% pulp consistency after chelation 17 In order to find application in a wider range of paper grades, the quality of HYP needs to be further improved in particular the brightness. Peroxide brightening/bleaching is the dominant process for HYP. Such processes are usually carried out by mixing an alkaline peroxide bleaching solution (known as bleach liquor) with the pulp. The bleach liquor may consist of hydrogen peroxide, caustic soda, sodium silicate, magnesium sulfate, and chelants (such as diethylene triamine pentaacetic acid, DTPA, ethylene diamine tetraacetic acid, EDTA). The mixture of the pulp and bleach liquor is then held in a tower for a required residence time at a desirable temperature. Much progress has been made recently in terms of both the fundamentals and the technological aspects of the process, (Colodette et al., 1989; Xu, 2000, 2002; Liu, 2003; Froass et al., 1996; Johnson et al., 2002; Leduc et al., 1997; Leduc et al., 2001). This paper will summarize the progress made on this subject recently at the Limerick Pulp and Paper Centre of the University of New Brunswick. THE HYDROSULPHITE ASSISTED CHELATION (THE Q Y PROCESS) Wood fibres have very strong ion exchange properties with metal ions and most of them bonded with the pulp fibres tightly. Some of the transition metals, such as manganese, have negative effects during various pulp and paper processes. For example, manganese can catalyze the hydrogen peroxide decomposition during the alkaline peroxide bleaching (Chirat and Lachenal, 1994; Lapierre et al., 1995; Wekesa and Ni, 2003). Therefore, it is desirable to maximize the removal of transition metal ions from pulp fibres prior to a peroxide stage. Conventionally, transition metal ions are removed from pulp fibres by chelating with sequesters, such as DTPA or EDTA. More advanced techniques have been developed to enhance the removal of harmful transition metal ions. For example, a Swedish patent (Lindahl, 1981) describes a process whereby transition metal ions in kraft pulps are effectively removed by the combined action of sodium sulphite and DTPA. It was assumed that the transition metal ions present in the original wood chips are converted to insoluble sulphides and/ or hydroxides during the course of kraft cooking and that they are thus strongly retained in the fibres. By the action of bisulphite ions, the metal sulphides (hydroxides) are reduced and dissolved, and as a result, the DTPA to metal ion complexes are formed upon the addition of DTPA. These can subsequently be removed from kraft pulps by pressing and/or washing. Gellerstedt and Pettersson, 1982 also reported that a treatment with chelant in combination with sodium bisulphite is more powerful in removing transition metals from kraft pulps than the chelant alone. Nye, 1996 introduced sodium hydrosulphite to the bleaching of chemical pulps in a DTPA chelation stage and found that the subsequent peroxide stage exhibits enhanced performance as a portion of DTPA is replaced with sodium hydrosulphite. Transition metal ions present in pulp fibres may be mostly in their high oxidation states. Unfortunately, transition metal ions with a high valency may form highly stable complexes with lignin structures and/or other ligands in pulp fibres. This may hinder their removal during chelation with sequester agents, such as DTPA. However the complexes formed between the same ligands and the same transition metal ion but at a low oxidation state are much less stable. Therefore, it is possible to improve the removal of transition metal ions from wood pulps by reducing transition metal ions from their higher oxidation states to the lower oxidation states simultaneously during chelation, or prior to chelation. For the Q y process, our objective was to enhance the removal of transition metal ions from mechanical pulps so that the subsequent peroxide bleaching can be improved by applying the combined actions of reduction and chelation (Ni et al., 1998, 1999). Since sodium hydrosulphite is a common chemical for bleaching of mechanical pulps, it was chosen as the reducing agent for this purpose, and the DTPA chelation and reduction of transition metal ions can be performed in a single stage since the process condition of a sodium hydrosulphite treatment is compatible to that of a DTPA chelation. Typical results for the comparison between the Q y process and the conventional chelation process (Q) are given in Table 1 (Ni et al., 1998). The initial manganese content of the Eastern Canadian TMP was 144 ppm. The chelation conditions were 3% pulp consistency, 70ºC, 30 min. At the completion of chelation, the pulp slurry was transferred to a buchner funnel for filtration. Subsequently, the pulp cake was either pressed to 30% pulp consistency or washed thoroughly with deionized water. Based on the results in Table 1, VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 611

6 Table 2. Bleaching Responses of Different Chelated Pulps (Ni et al., 1998) (Spruce TMP chelation conditions, same as those in Table 1, P Stage conditions: 12% pulp consistency, 3% Na 2 SiO 3, o11% MgSO 4, 70ºC, 2 hrs) H 2 O 2 charge Sample NaOH charge H 2 O 2 consumption Brightness (%) (%) (%) (% ISO) 1 Q 0.2 (pressed) Q y 0.2 (pressed) Q 0.5 (washed) Q y 0.5 (washed) Q 0.2 (pressed) Q y 0.2 (pressed) Q 0.5 (washed) Q y 0.5 (washed) Q 0.2 (pressed) Q y 0.2 (pressed) Q 0.5 (washed) Q y 0.5 (washed) Q 0.2 (pressed) Q y 0.2 (pressed) Q 0.5 (washed) Q y 0.5 (washed) one can conclude that the Q y process led to enhanced transition metal removal compared to the conventional chelation. Subsequently, four of the chelated pulps, in accordance with the above techniques namely Q 0.2 (pressed), Q y 0.2 (pressed), Q 0.5 (washed) and Q y 0.5 (washed) were subjected to alkaline peroxide bleaching with various peroxide charges. The number after Q or Q y is the DTPA charge, % on pulp. The results were presented in Table 2. Other conditions for the peroxide stage were: 12% pulp consistency, 3% Na 2 SiO 3, 0.1% MgSO 4, 70ºC, 120 min. Table 2 shows that the Q y chelated pulps had a better response in the subsequent peroxide bleaching, producing bleached pulps with a higher brightness compared to the Q chelated pulps at the same hydrogen peroxide charge. Furthermore, the hydrogen peroxide consumption was less. In conclusion, the hydrosulphite assisted DTPA chelation (Q y ) is an effective approach to improve the removal of manganese from mechanical pulps. The sodium hydrosulphite charge as low as 0.1% is effective to achieve the observed improvement in manganese removal. The brightness gain during the subsequent peroxide bleaching is improved for the Q y chelated pulps than for the Q chelated pulps. The sodium hydrosulphite assisted DTPA chelation (Q y ) has been successfully implemented in an Eastern Canadian mill (Ni et al., 1999). DTPA SPRAY It has been generally accepted that the presence of some metal ions in pulp and paper products negatively affects their brightness. The above is especially true for bleached mechanical pulps (Polcin and Rapson, 1972) and may be partly attributed to the Figure 1. Effect of water quality on the brightness of bleached groundwood pulp (Ni et al., 1997) (Chelation: 60ºC, 0.5% DTPA, ph 5.8, 2% pulp consistency, 30 min; Peroxide bleaching: 60ºC, 10% pulp cons., 2 hrs, 3% Na 2 SiO 3, NaOH/H 2 O 2, mass ratio of 0.8) 612 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

7 Table 3. Spraying DTPA solution to handsheets made with tap water to recover the brightness loss DTPA spraying room temperature (Ni et al., 1997) (Spruce SGW, 0.5% DTPA charge ph 6) Sample ID Brightness (% ISO) Sheets made with Sheets made with Sheets made with deionized water tap water tap water and DTPA spraying formation of highly coloured lignin-metal ion complexes (Polcin and Rapson, 1972; Meshitsuka and Nakano, 1973; Gupta, 1970; Ghosh and Ni, 1998; Ni et al., 1999, 1998, Ni and Mosher, 1999: Peart and Ni, 2001). It appears that iron, both ferrous and ferric, has a strong negative effect on the brightness. Copper has less effect than iron on pulp brightness due to its low concentration in the pulp. With the ever increasing environmental regulations, the trend is clearly towards minimum effluent discharge achieved by increased filtrate recycle. As a result, the process water may be expected to contain more metal ions (Bryant et al., 1993). Consequently, the brightness decrease due to the formation of the coloured lignin-metal ion complexes may attain alarming proportions. The brightness loss of bleached mechanical pulps due to the quality of process water is illustrated in Figure 1 (Ni et al., 1997). A spruce groundwood pulp was bleached in a single stage hydrogen peroxide stage at various charges. The bleached pulp samples were then made into handsheets with either deionized water or tap water. Figure 1 shows that the brightness of the bleached handsheets made with deionized water is consistently 2 to 3 units higher than that of the corresponding handsheets made with tap water, while the difference in the brightness of unbleached pulp sheets made with either deionized water or tap water is only 1.4% ISO. The stronger negative effect of water quality on the brightness of bleached groundwood pulp compared to that of the unbleached pulp can be explained by the fact that new lignin functional groups are generated during hydrogen peroxide bleaching. These new functional groups can form coloured complexes with metal ions. Most likely, they are phenolic hydroxyl and carboxyl groups, since it was reported (Joachimides, 1989) that the content of both these groups increases during an alkaline hydrogen peroxide treatment of mechanical pulps. Because transition metal ions form much more stable complexes with sequesters, e.g. DTPA, EDTA, than those with pulp fibres, theaddition of DTPA or EDTA to a system consisting of metal ions and pulp fibres can lead to the formation of metal ion to chelant complexes, thus releasing metal ions from pulp fibres. Therefore, one would expect that the impregnation of pulp and paper products by a metal sequestering solution would reduce the negative impact of these metal ions on brightness. However, Janson and Forsskåhl, 1989 found that the brightness loss due to these metal ions may be recovered only by washing out the chelated metals. We found that in-situ metal chelation by spraying metal containing handsheets with a DTPA solution resulted in increased brightness, as shown in Table 3. The brightness of the handsheets made with tap water (Column 3) is always 3 to 4 units lower than those made with deionized water from the same pulp (Column 2). However, by spraying a DTPA solution onto the sheets, the brightness loss due to the water quality can be completely recovered, as evidenced by the similar brightness in Columns 2 and 4, respectively. The brightness loss of bleached mechanical pulps during the papermaking process due to the presence of metal ions can be significant (2 to 3 brightness units). However, the problem can be corrected by the addition of a DTPA solution to the wet web. The following options may be considered: impregnate either the top or the bottom felt with a DTPA containing solution, and allow the transfer of the DTPA solution from the felt to the wet paper web to take place in the first press; if available, apply the DTPA solution to the wet web using a smoothing press, either by a one-sided or two-sided treatment. THE P N2 PROCESS During bleaching of mechanical pulps with an alkaline hydrogen peroxide, in addition to chromophore removal (bleaching) reactions, which are responsible for the observed Figure 2. The brightness development during the process and the conventional peroxide stage (Ni et al., 2000) (H 2 O 2 charges: 3.5%, 5%, 6.5% and 8%; NaOH charges: 2%, 2.5%, 3.5%, 4%; 80ºC; 12% pulp consistency; 3 hrs) VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 613

8 brightness increase, other reactions that generate more chromophores, such as thermal yellowing, alkali darkening, are also taking place. Alkaline darkening suppresses the development of brightness during peroxide bleaching (Kutney and Evans, 1985; Leary and Giampaolo, 1998; Giust et al., 1991; He et al., 2004), and it is caused by the formation of ortho-quinones and coniferaldehydes (Giust et al., 1991). The presence of oxygen in the system further enhances the intensity of alkaline darkening (He et al., 2004). On the other hand, if the bleaching of mechanical pulps with hydrogen peroxide is carried out in the presence of an inert gas such as nitrogen (the so-called P N2 process (Ni et al., 2000)) some of the counter-productive reactions, i.e. the formation of chromophores, may be suppressed and, as a result, a higher brightness can be reached. Figure 2 compares the brightness development between the P N2 process and the conventional peroxide process (P). It is evident that the P N2 process leads to a higher brightness than the P process under otherwise the same conditions. These results can be explained by the suppressed chromophores formation during the course of reactions. Implied here is that any effort that would minimize the presence of oxygen/air in the process can improve the brightness gain during peroxide bleaching. THE P M PROCESS Sodium silicate and magnesium sulfate ( Epsom salt ) are commonly known as stabilizers that reduce the transition metal ion-induced peroxide decomposition. They are part of the bleach liquor for peroxide bleaching of mechanical pulps. The cascade make-up system (Presley and Hill, 1996) is usually used to prepare the bleach liquor. In this system, magnesium sulfate and sodium silicate are added to water, followed by the addition of caustic soda and finally by the addition of hydrogen peroxide. The resulting liquor is prepared for utilization for bleaching by mixing it with the pulp slurry. Here, hydrogen peroxide and others required for peroxide bleaching, such as caustic soda and the stabilizers, are mixed to the pulp simultaneously as part of the mixture. Also, the in-line bleach liquor make-up system is practised in industry (Presley and Hill, 1996). Again, the stabilizers (caustic soda and hydrogen peroxide) are mixed first before they are added to the pulp fibres. The difference between the in-line and cascade make-up systems is the lack of cascade tanks in the former. The advantages of the in-line over the cascade liquor make-up system are (Presley and Hill, 1996): the system is more flexible during grade change; there is less chance of scaling. We found that the transition metal ion-induced peroxide decomposition is not minimized during the conventional peroxide bleaching process and developed the so-called P M process, the modified peroxide process (Ni et al., In press)], in Table 4. Comparison of the P M process, and the conventional peroxide process (Ni et al., In press) (Mill chelated TMP, 4% H 2 O 2, 2.0% NaOH, 3% Na 2 SiO 3 (as solution), 0.05% MgSO 4, 10% pulp consistency, 60 C, 120 min) Process Residual Peroxide Brightness (% on pulp) (% ISO) Conventional process The P M process which sodium silicate and/or other peroxide stabilizers, such as DTPA, EDTA or magnesium sulfate, are added to the pulp first along with part or all of the sodium hydroxide required, followed by the addition of hydrogen peroxide. The comparison of the P M process with the conventional peroxide process (P) is made in Table 4 (Ni et al., In press). The results show that the P M process produced a pulp with a much higher brightness than the P process. The peroxide consumption during the course of reaction is significantly lower in the P M process than the P process. These results support the notion that the P M process is very efficient in bleaching mechanical pulps, leading to the production of bleached TMP with much higher brightness compared with the conventional P process at the same hydrogen peroxide charge. The following are proposed to account for the improvement of the P M process over the conventional P process: transition metal ions, such as manganese, are better stabilized in the P M process so that the peroxide decomposition is suppressed; in the P M process, acetyl groups in pulps (associated with hemicellulose) are hydrolyzed prior to the addition of hydrogen peroxide so that peracetic acid is not formed during peroxide bleaching. Since peracetic acid is much less stable than hydrogen peroxide under the peroxide bleaching condition, avoiding the formation of peracetic acid decreases the amount of peroxide consumed in wasteful reactions. Recently, it was proposed that the manganese-induced peroxide decomposition occurs via an efficient redox cycle (Wekesa and Ni, 2003). Oxidation reactions: 4 Mn 2+ + O H 2 O 4 Mn OH! 2 Mn 2+ + H 2 O 2 2 Mn OH! Reduction reaction: 2 Mn 3+ + H 2 O OH 2 Mn 2+ + O H 2 O Since the oxidation and reduction reactions are all extremely fast under the bleaching conditions, this explains why a small amount of manganese is so efficient in decomposing an alkaline peroxide solution. In the P M process, stabilizers, such as silicates, magnesium sulfate and chelants, are added prior to the addition of hydrogen peroxide. Manganese present in the system is stabilized so that the redox cycle is disrupted. Lidén and Öhman, 1997 proposed a solid solution concept to account for the redox stabilization of manganese by magnesium precipitates (hydroxides and carbonates). Mn 2+ isomorphically replaces magnesium in the carbonate or hydroxide precipitates formed during alkaline peroxide bleaching. It is expected that Mn 2+ and Mg 2+ will form similar precipitates with silicates. Also, Mn 2+ not Mn 3+ will be possible to form these solid solutions because the ionic radius of the former is comparable to Mg 2+. This explains the experimental finding that stabilizers added to Mn 2+ not Mn 3+ containing system is much more efficient in decreasing the Mn-induced peroxide decomposition (Lidén and Öhman, 1997; Ni et al., 2000; Qiu and Ni, In press). Wood contains acetyl groups and usually hardwood species have more than softwood. The acetyl groups are mainly bonded with hemicellulose through ester linkages at C2 and C3 positions and they are readily hydrolyzed under alkaline conditions. Thus, kraft and soda pulps contain negligible amounts of acetyl groups. However, the acetyl groups largely survive during the typical mechanical pulping process. 614 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

9 Table 5. Comparison of the NaOH-based and the Mg(OH) 2 -based peroxide processes (Li et al., In press) (mill chelated spruce TMP, 2% H 2 O 2, 5 hrs, 10% pulp consistency, 70ºC) Process NaOH -Based Process Mg(OH) 2 -based process (3% Na 2 SiO 3 ) (0.1% DTPA) P P M NaOH or Mg(OH) 2 (%) End ph Residual peroxide (%) Brightness (% ISO) An earlier study (Kang and Ni, 2001) showed that peracetic acid is formed from the reaction of hydrogen peroxide with the acetyl groups of hemicellulose during the conventional peroxide bleaching of TMP. This reaction is very fast, and under peroxide bleaching conditions, it can be completed within the first 2 minutes. The amount of peracetic acid formed depends on the acetyl group content in the pulp, and bleaching conditions (particularly the hydrogen peroxide charge). Under a typical peroxide stage about 20% hydrogen peroxide can be converted to peracetic acid. Since peracetic acid is less stable than hydrogen peroxide, the formation of peracetic acid represents the consumption of hydrogen peroxide in wasteful reactions. On the other hand, in the P M process, the alkaline solution (silicates and caustic soda) is mixed with pulp first so that the acetyl groups are hydrolyzed, prior to the addition of hydrogen peroxide. This prevents the formation of peracetic acid, resulting in improved peroxide stability, and thus an increased bleaching efficiency. THE Mg(OH) 2 -BASED PEROXIDE PROCESS The conventional peroxide bleaching process uses sodium hydroxide (NaOH) as the alkali source. However, as a strong base, sodium hydroxide can cause the dissolution of carbohydrates, particularly hemi-cellulose, into the bleach effluent and is responsible for the increased COD and BOD load to the secondary treatment system. Also, the high alkalinity of the process leads to the formation of a significant amount of anionic trash, which is carried onto the wet-end of the paper machine. This can have negative effects on papermaking operations, such as increased polymer/additive cost, reduced drainage and decreased product quality. In recognition of the drawbacks associated with the NaOHbased peroxide process, peroxide bleaching using weak alkali sources such as magnesium hydroxide (Mg(OH) 2 ), has received much attention recently. Griffiths and Abbot, 1994 studied the use of magnesium oxide as the alkali source in a peroxide bleaching of pine TMP, and found that brightness gain is inferior to that of NaOH based peroxide process (about 2 units lower at the same peroxide charge). Soteland et al., 1988 observed that the size and shape of magnesium oxide particles have a significant impact on the bleaching results and under optimum condition, the magnesium oxide-based peroxide process produced a bleached pulp with about 1.5 units lower in brightness than the NaOH-based process under otherwise identical conditions. Suess et al., 2001 reported that the application of magnesium hydroxide as an alkali source to bleach mechanical pulps is very attractive. The COD is decreased by 30 to 40% and the sodium silicate charge can be decreased significantly. Johnson et al., 2002 recently concluded many advantages associated with using Mg(OH) 2 as the alkali for bleaching mechanical pulps. A comparison of NaOH-based and Mg(OH) 2 -based peroxide processes is shown in Table 5 (Li et al., In press). For the Mg(OH) 2 system, three Mg(OH) 2 dosages were tested and the P M process (Ni et al., In press) was investigated. The results show that at the optimum Mg(OH) 2 dosage of 1% under the conventional mode (the P process), the Mg(OH) 2 -based peroxide process gives a slightly lower brightness than the NaOH-based process. The P M process (DTPA and Mg(OH) 2 added to pulp slurry first, and 1 min later, H 2 O 2 added), results in a higher brightness and a higher H 2 O 2 residue. This confirms that the P M process is applicable to the Mg(OH) 2 -based process, and improves the performance of the peroxide bleaching process. It is noted that in Table 5, 3% silicate was used as the stabilizer for the NaOH-based process while 0.1% DTPA was used for the Mg(OH) 2 -based process. This is because the alkalinity is very mild for the Mg(OH) 2 -based process and the metal ion-induced peroxide decomposition is much less so that a small amount of DTPA is effective to stabilize peroxide. Due to a low alkalinity associated with the Mg(OH) 2 -based peroxide process, pulp fibres may not swell to the same extent, which may affect fibre to fibre bonding. This could result in decreased strength properties of the bleached pulp. A comparison of strength properties of the two bleached pulps is shown in Table 6. The difference in tensile, tear and burst are very small. There is also a small increase in bulk (from 2.52 to 2.59 cm 3 /g). This is quite different from the results obtained from other pulps, such as maple CTMP (Zhang et al., 2004). A possible explanation is that the freeness of the TMP pulp (spruce with a small amount of fir) is about 70 ml, significantly lower than the Table 6. Pulp strength properties from NaOH-based and Mg(OH) 2 - based processes Li et al., In press (spruce TMP, about 70% ISO brightness, NaOH process conditions: 1.6% NaOH, 3% Na 2 SiO 3, 2% H 2 O 2, 5 hrs. 10% pulp consistency, 70ºC; Mg(OH) 2 process conditions: 1% Mg (OH) 2, 0.1% DTPA 2% H 2 O 2, 5 hrs, 10% pulp consistency, 70ºC) Parameter NaOH process (P) Mg(OH) 2 process (P M ) Tensile index (N m/g) Tear (mn m 2 /g) Burst (kpa m 2 /g) Bulk (cm 3 /g) VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 615

10 BCTMP maple grade (about 350 ml). The pulp fibres of the TMP are well developed, and the change in alkalinity from the NaOHbased process to the Mg(OH) 2 - based process would have a negligible effect on fibre development for the TMP. Other benefits of the Mg(OH) 2 -based peroxide process include: decreased bleaching cost (Li et al., In press); decreased anionic trash formation (Li et al., In press; He et al., submitted); Elimination of oxalate-related scaling (Yu et al., 2004). CONCLUSION The hydrosulphite assisted DTPA chelation, or the Q y process, is an effective approach to improve the removal of harmful transition metal ions from mechanical pulps. The results showed that the residual manganese content in the pulp after the Q y treatment is substantially lower than that after the conventional chelation with DTPA only (Q). Furthermore, the brightness gain during the subsequent peroxide bleaching is better for the Q y chelated pulp than for the Q chelated pulps. The presence of metal ions in the process water has a strongly adversed effect on the brightness of bleached mechanical pulps, which can be explained by the formation of coloured complexes between lignin functional groups and metal ions. Impregnation/ spray of these metal ion containing paper sheets with a DTPA solution can minimize the brightness loss caused by the impurities in the process water. The P N2 process produces bleached CTMP maple pulps with a higher brightness at a same hydrogen peroxide charge/consumption compared to the conventional peroxide bleaching. The better bleaching results of the P N2 process is attributed to the fact that the presence of nitrogen suppresses some of the darkening reaction during the alkaline peroxide bleaching. The P M process is an effective way to improve the performance of alkaline peroxide stage to bleach mechanical pulps. At the same peroxide charge, an improvement in brightness of about 2 units was observed for the P M process in comparison with the control. The peroxide consumption is much less. The P M process has shown a significant decrease in the manganeseinduced peroxide decomposition. In addition, the elimination of the peracetic acid formation may also be responsible for the improved performance of the P M process over the control. The Mg(OH) 2 -based peroxide process decreases the bleaching cost while producing bleached TMP with similar optical and strength properties. The process has been successfully installed commercially. Benefits include: a lower bleaching cost; decreased COD in the bleaching effluent; decreased anionic trash. ACKNOWLEDGEMENTS I would like to thank those who have been involved in these projects, in particular: Students: Z. Li, Z. He, A. Ghosh, M. Wekesa, Y. Ju, Z. Qiu, C. Peart, Q. Jiang, M. Rae, L. Yu, S. Krishnan Colleagues: A. R. P. van Heiningen, G. J. Kang, M. M. Sain, H. Xiao and K. Li Government funding agencies: NCE Wood Pulps Network, NSERC, Canada Research Chairs Program, Canadian Foundation for Innovation Industrial collaborators: Holmen Paper, Irving Paper, Martin Marietta, Neill and Gunter, Tembec REFERENCES Bryant, P. S., K. Robarge and L. L. Edwards, Transition Metal Profile Monitoring and Control in Closed Kraft Mill Fibrelines, Proceedings, Tappi Environmental Conf., 617, Boston, MA (1993). Chirat, C. and D. Lachenal, Beneficial and Adverse Effects of Metal Ions in ZP Bleaching Sequence, Proceedings, Tappi Pulping Conf., 1239, Atlanta, GA (1994). Colodette, J. L., S. Rothenberg and C. W. Dence, Factors Affecting Hydrogen Peroxide Stability in the Brightening of Mechanical and Chemimechanical Pulps. Part I: Hydrogen Peroxide Stability in the Absence of Stabilizing Systems, J. Pulp Paper Science 14(6), J126 (1988); Part II: Hydrogen Peroxide Stability in the Presence of Sodium Silicate, J. Pulp Paper Science 15(1), J3 (1989); Part III: Hydrogen Peroxide Stability in the Presence of Magnesium and Combination of Stabilizers, J. Pulp Paper Science 15(2), J45 (1989). Froass, W., R. Francis, C. Dence and G. Lefevre, The Interactions of Calcium, Magnesium and Silicate Under Peroxide Bleaching Conditions, Proceedings, 82th Annual Meeting, CPPA.TS, A228, Montréal, QC (1996). Gellerstedt, G. and I. Pettersson, Chemical Aspects of Hydrogen Peroxide Bleaching, J. of Wood Chem. Technol. 2(3), 231 (1982). Ghosh, A. and Y. Ni, Lignin Complexes and their Relationship to the Brightness of Bleached Mechanical Pulps, J. Pulp Paper Science 24(1), 26 (1998). Giust, W., F. McLellan and P. Whiting, Alkaline Darkening and its Similarities to Thermal Reversion, Pulp Paper Sci. 17(3), J73 (1991). Griffiths, P. and J. Abbot, Magnesium Oxide as a Base for Peroxide Bleaching of Radiata Pine TMP, Appita J. 47(1), 50 (1994). Gupta, V. N., Effect of Metal Ions on Brightness, Bleachability and Colour Reversion of Groundwood, Pulp and Paper Mag. Can. 71(9), T391 (1970). He, Z., Y. Ni and E. Zhang, Alkaline Darkening and its Relationship to Peroxide Bleaching of Mechanical Pulp, J. Wood Chem. and Techn. 24(1), 1 (2004). He, Z., M. Wekesa and Y. Ni, Formation of Anionic Trash from the Mg(OH) 2 -Based Peroxide Bleaching and its Impact on Filler Retention, Pulp Paper Can. In Press. Janson, J. and I. Forsskåhl, Color Changes in Lignin-Rich Pulps on Irradiation by Light, Nordic Pulp Paper Res. J. 3(4), 197 (1989). Joachimides, T., High Brightness Mechanical Pulps, Proceedings, Tappi Pulping Conf., 131, Seattle, WA (1989). Johnson, D., S. Park, J. Genco, A. Gibson, M. Wajer and B. Branch, Hydrogen Peroxide Bleaching of TMP Pulps using Mg(OH) 2, Proceedings, Tappi Fall Conference and Trade Fair (2002). Kang, G. J. and Y. Ni, The Formation of Peracetic Acid and its Impact on Peroxide Bleaching of Mechanical Pulps, Proceedings, 11th International Symposium on Wood and Pulping Chemistry, Nice, France (2001). Kutney, G. W. and T. D. Evans, Peroxide Bleaching of Mechanical Pulps. Part 1 Alkali Darkening The Effect of Caustic Soda, Svensk Papperstidn. 88(6), 78 (1985). Kutney, G. W. and T. D. Evans, Peroxide Bleaching of Mechanical Pulps. Part 2 Alkali Darkening Hydrogen Peroxide Decomposition, Svensk Papperstidn. 88(9), 84 (1985). 616 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

11 Lapierre, L., J. Bouchard, R. M. Berry and B. van Lierop, Chelation Prior to Hydrogen Peroxide Bleaching of Kraft: An Overview, J. Pulp Paper Science 21, J268 (1995). Leary, G. and D. Giampaolo, The Darkening Reactions of TMP and BTMP During Alkaline Peroxide Bleaching, Preprints, 84th Annual Meeting, CPPA.TS, A31, Montréal, QC (1998). Leduc, C., M. M. Sain, C. Daneault, R. Lanouette and J. L. Valade, Use of New Oxidizing Agents (Peroxide Perborate, Activated Peroxide) in the Refiner A Route to High Brightness Pulp, Proceedings, International Mechanical Pulping Conf. 63 (1997). Leduc, C., S. Rouaix and C. Daneault, Use of Zeolites for the Bleaching of Mechanical Pulp, Proceedings, Tappi Pulping Conf. 435 (2001). Li, Z., G. Court, R. Belliveau, M. Crowell, R. Murphy, A. Gibson, M. Wajer, B. Branch and Y. Ni, Using Magnesium Hydroxide (Mg(OH) 2 as the Alkali Source in Peroxide Bleaching at Irving Paper, Pulp and Paper Can. 106(6), 24 (2005). Lidén, J. and L. O. Öhman, Redox Stabilization of Iron and Manganese in the +II Oxidation State by Magnesium Precipitates and Some Anionic Polymers, J. Pulp Paper Sci. 23(5), J193 (1997). Lindahl, J. A. I., Swedish Patent, 416, 481 (1981). Liu, S., Chemical Kinetics of Alkaline Peroxide Brightening of Mechanical Pulps, Chemical Engineering Sci. 58(11), 2229 (2003). Meshitsuka, G. and J. Nakano, Effect of Metal Ion on Color of Lignosulfonate and Thiolignin, Tappi 56(7), 105 (1973). Ni. Y., A. Ghosh, Z. Li, C. Heitner and P. McGarry, Photostabilization of Bleached Mechanical Pulps with DTPA Treatment, J. Pulp Paper Science 24(8), 259 (1998). Ni, Y., A. Ng and M. Mosher, The Formation of Coloured Metallic Extractive Complexes and its Effect on the Brightness of TMP Pulp. Part I. Model Compound Study, J. Wood Chem. and Techn. 19(3), 213 (1999). Ni, Y., G. Court, Z. Li, M. Mosher, M. Tudor and M. Burtt, Improving Peroxide Bleaching of Mechanical Pulps by an Enhanced Chelation Process, Pulp and Paper Mag. Can. 100(10), 51 (1999). Ni, Y., J. Yu and H. Ohi, Further Understanding on the Manganese-Induced Peroxide Decomposition, J. Pulp Paper Sci. 26(3), 90 (2000). Ni, Y., Q. Jiang, Z. Li, G. Court and M. Burtt, Improved Transition Metal Removal in a Reducing Agent-Assisted Chelation Stage: A laboratory Study, Pulp and Paper Mag. Can. 99(8), 77 (1998). Ni, Y., Q. Jiang and Z. Li, The P N2 Process to Bleached CTMP Maple Pulp, Appita J. 53(5), 404 (2000). Ni, Y., Z. Li, G. Court, R. Belliveau and M. Crowell, Improving Peroxide Bleaching of Mechanical Pulps by the P M Process, Pulp and Paper Canada, 104(12), 78 (2003). Ni, Y., Z. Li and A. R. P. van Heiningen, Minimization of the Brightness Loss Due to Metal Ions in Process Water for Bleached Mechanical Pulps, Proceedings, 83rd Annual Meeting, CPPA.TS, 43, Montréal, QC, (1997). Nye, J., Reductive Bleaching of Chemical Pulps, Proceedings, International Pulp Bleaching Conf., 485, Tappi Press, Atlanta (1996). Peart, C. and Y. Ni, UV-Vis Spectra of Lignin Model Compounds in the Presence of Transition Metal Ions and Chelants, J. Wood Chem. and Techn. 21(2), 113 (2001). Polcin, J. and W. H. Rapson, Sapwood and Heartwood Groundwood of Western Hemlock and Jack Pine. Part III. Influence of Solvent Extraction of the Bleaching of Pulps, Pulp and Paper Mag. Can. 73(1), 86 (1972). Presley, J. R. and R. T. Hill, Pulp Bleaching Principles and Practice, C. W. Dence and D. W. Reeve, Eds., 480, Tappi Press (1996). Qiu, Z. and Y. Ni, Methods to Decrease the Mn-Induced Peroxide Decomposition, Appita J., 56(5), 355 (2003). Soteland, N., F. A. Aberdie Maumert and T. A. Arnerik, Use of MgO and CaO as the Only Source in Peroxide Bleaching of High Yield Pulps, Proceedings, International Pulp Bleaching Conf., 231 (1988). Suess, H. U., M. Del Grasso, K. Schmidt and B. Hopt, Options for Bleaching Mechanical Pulp with Lower COD load, Proceedings, Appita Conf., 419 (2001). Wekesa, M. and Y. Ni, Further Understanding of the Chemistry of Manganese-Induced Peroxide Decomposition, Can. J. Chem. Eng. 81(10), 1 (2003). Xu, E. C., H 2 O 2 Bleaching of Mechanical Pulps. Part I: Kinetics and Mechanism, J. Pulp Paper Science 26(10), 367 (2000); Part II: ph and Temperature, J. Pulp Paper Science, 26(11), 407 (2000); Part III: Thermodynamics, J. Pulp Paper Science 28(1), 26 (2002); Part IV: H 2 O 2 Consumption, J. Pulp Paper Science, 28(11), 379 (2002). Yu, L., M. Rae and Y. Ni, Formation of Oxalate from the Mg(OH) 2 -Based Peroxide Bleaching of Mechanical Pulps, Proceedings, Paptac Annual Meeting, Montréal, QC, A-211 (2004), J. Wood Chemistry and Technology 24(3), 341 (2004). Zhang, X. Z., Y. Ni, Y. Zhou and D. Joliette, Mg(OH) 2 -Based Peroxide Process for CTMP Hardwood Pulp, Proceedings, Paptac Annual Meeting, Montréal, QC, B-133 (2004). Manuscript received November 12, 2004; revised manuscript received March 10, 2005; accepted for publication May 24, VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 617

12 REVIEW Review on Mixing Characteristics in Solid- Liquid and Solid-Liquid-Gas Reactor Vessels Gopal R. Kasat and Aniruddha B. Pandit* Chemical Engineering Division, University Institute of Chemical Technology (UICT), University of Mumbai, Matunga, Mumbai , India Mechanically agitated reactors with single and multiple impeller systems are used in the industry for the various three-phase mixing processes such as crystallization, fermentation, and hydrogenation, etc. The paper reviews the experimental work reported in the literature along with different techniques used for the measurement of the specific quantities such as minimum or critical impeller speed for solid suspension. The work critically surveys the literature and makes specific recommendations for the use of appropriate correlations and conditions to be used for the success of such equipment. This assessment will put all the relevant literature on a common footing and will help to validate work reported earlier. Les réacteurs agités mécaniquement munis d une seule turbine et de turbines multiples sont utilisés dans l industrie pour divers procédés de mélange triphasiques, tels que la cristallisation, la fermentation, l hydrogénation, etc. On examine dans cet article les travaux expérimentaux présentés dans la littérature scientifique ainsi que les différentes techniques utilisées pour la mesure de quantités spécifiques, telle la vitesse de turbine minimale ou critique pour la suspension de solides. On effectue une étude critique de la littérature scientifique et on propose des recommandations pour le choix de corrélations et conditions appropriées pour une bonne utilisation de cet équipement. Cette évaluation mettra toutes les publications pertinentes à un même niveau et aidera à valider le travail présenté antérieurement. Keywords: three-phase reactor, solid suspension, minimum impeller speed, multiple impeller In carrying out the physical mass transfer processes and chemical reactions in the gas-liquid-solid systems, in which liquid is the continuous phase and where the solid and gas phase are the discontinuous phases, there is often a need to suspend the solid particles in the presence of gas. Stirred tanks are often employed for this purpose as slurry reactors, especially for discontinuous/batch operations and for smaller production levels. The main task of the stirrer in an aerated suspension system is to suspend the solids particles completely so as to expose their total surface area for the reaction and/or for mass transfer, to disperse the gas uniformly throughout the vessel and to provide liquid mixing patterns for an intimate contact between the phases. Thus, mechanically agitated reactors are extensively used in the industry to handle the three-phase systems and find application in a wide range of processes such as catalyzed oxidation and hydrogenation reactions, polymerization, evaporative crystallization, aerobic fermentation, froth floatation, microbial coal desulphurization, gold leaching, and bacterial sulphide oxidations. In all these applications, simultaneous dispersion of gas and the suspension of the solids by the impellers are of vital importance. For a reliable design of a three-phase stirred tank reactor, it is important to know the complex hydrodynamic problems (i.e. phase-phase interactions) associated with the demands of simultaneous gas dispersion and solids suspension with a single or multiple impellers. One of the most important parameters for the assessment of the performance of the impellers for solid suspension (both in the absence and in the presence of gas), is the minimum impeller speed required for the complete offbottom suspension of the solids in the vessel. It is denoted by N js, for solid suspension in the absence of gas and by N jsg, for solid suspension in the presence of gas. In this paper, these notations have been used whenever reference has been made to this speed. This speed is also called as the critical impeller speed for solids suspension. Importance of N js and N jsg in the design of two- or three-phase stirred tank reactors has often been discussed in the literature (Zwietering, 1958; Wiedmann et al., 1980; Chapman et al., 1981, 1983; Frijlink et al., 1990; Rewatkar et al., 1991a; * Author to whom correspondence may be addressed. address: abp@udct.org 618 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

13 Dylag and Talaga, 1994). According to the literature, in general, N jsg is always higher than N js. For a given application (i.e. for a given three-phase system) N jsg has been shown to depend on the superficial gas velocity, impeller diameter and design, off-bottom clearance of the impeller, sparger design and location, and vessel geometry. By choosing a right combination of the above parameters, enormous improvement in the solid suspension and gas dispersion characteristics of the impeller can be made thereby resulting in substantial energy savings. In the present paper, the main emphasis will be on reviewing the performance characteristics (in terms of the N js, N jsg ) of the single and multiple impeller system for suspending the solids in the presence of gas and for validating the reactor performance over the range of operating parameters likely to be used in the industry. This will be based on the various earlier studies reported in the area of solid suspension by impellers in both, i.e. the presence and the absence of gas. Final design recommendation will be based on the data available and analyzed from the existing literature. The deficiencies in the present literature will be pointed out that can be taken up for future research work. STATES OF SUSPENSION For measuring the performance of solid suspension equipment, it is necessary to know the different states of the suspension occurring in a stirred vessel with respect to the increasing impeller speed. Kraume (1992), Bujlaski et al. (1999) have observed the following states of suspension occurring in the solidliquid stirred vessel with respect to the stirrer speed (Figure 1): 1. for low stirrer speed, all the solids rest on the bottom of the vessel (Figure 1a); 2. with increasing stirrer speed, the solids get lifted by the circulating liquid flow and become suspended up to a certain height; 3. with further increase in the stirrer speed, all the solids are lifted from the bottom of the tank. At this stage, the applied power input is sufficient to prevent the settling of the solids to the vessel bottom for more than 1 or 2s. It can be said that complete off-bottom suspension is achieved and the corresponding impeller speed is termed as N js. At this stage an interface appears between the suspension and clear liquid layer in the upper part of the vessel (Figure 1c); 4. an increase in the impeller speed beyond N js results in an increase in the height of the interface of the solid liquid suspension from the vessel bottom (i.e. height of the suspension); 5. when the slurry height amounts to 90% of the total liquid height, the slurry height criterion for determination of N js (as defined by Bujlaski et al., 1999) has been said to be fulfilled (Figure 1d); Figure 1. States of suspension in a stirred vessel (Kraume, 1992) Finally, a still further increase in stirrer speed distributes the solids throughout the liquid volume confirming the complete homogeneous distribution of solids throughout the tank volume (Figure 1e). In solid-liquid/gas-solid-liquid stirred tanks, two main suspension states can be defined, namely complete off-bottom suspension and homogeneous suspension. Complete off-bottom suspension is said to be achieved, when all the solids are lifted of the vessel bottom and no solids remain on the vessel bottom for more than a short period, e.g. 1 2 s (Zwietering, 1958). Under this condition, all the surface of the solids is exposed to the fluid, thereby ensuring that maximum or all the solid surface area is available for chemical reactions (occurring at the solid surface) and the transport processes. At this condition, the distribution of solids over the vessel may be extremely non-uniform, especially in high aspect ratio tanks. This condition is sufficient in the case of a process involving dissolution of solids in liquid or a chemical reaction with solids as the catalyst, because in these processes the exposure of the total solid surface area and not the axial solids distribution is important. The minimum impeller speed at which the solids are completely suspended off the vessel bottom is called as N js or N jsg in the absence and the presence of the gas respectively. A homogeneous suspension is said to be achieved, when the solids concentration as well as the size distribution (for ranges of particle size) is constant throughout the tank. This condition is particularly desirable when a continuous and representative flow of solids from the system is required. The impeller speed at which the homogeneous suspension of the solids is achieved is called as critical impeller speed for homogeneous suspension or critical impeller speed for ultimate suspension. Beyond this speed, any increase in the impeller speed does not improve the quality of suspension (Dohi et al., 2001). These speeds are denoted as N us and N usg for solid suspension in the absence and in the presence of gas, respectively. Throughout the discussion, these notations will be used whenever reference is made to these speeds. According to Bujalski et al. (1999), complete off-bottom suspension is achieved at stage (iii) (Figure 1c), while homogeneous/ultimate suspension is attained at stage (vi) (Figure 1e), i.e. at an impeller speed substantially higher than that required for complete off-bottom suspension. In many cases, homogeneous or ultimate suspension is not desirable or necessary and would be very difficult to realize with rapidly settling particles. In such cases, attainment of ultimate suspension may require high impeller speeds resulting into large amount of energy dissipation. Hence, the criteria for the design of an agitated vessel for solid suspension should be based on N js or N jsg rather than on N us or N usg. Also, the rate of various transport processes from the solid surface (e.g. solid-liquid mass transfer) increased significantly with an increase in the impeller speed up to N js and N jsg. Beyond this speed, there is only a slight increase in the rate of the transfer processes with an increase in the impeller speed (Harriot, 1962). However, in some cases (e.g. bioreactors) the axial distribution of solids is also important. With an increase in the impeller speed above N js and N jsg the solids distribution becomes more uniform throughout the vessel up to N us and N usg and after that, there is no improvement in the homogeneity of the suspension (Dohi et al., 2001). Hence, N us and N usg are also important from the design point of view. In general, the accessibility of maximum or all the solid surface area and the elimination of high-localized solids concentration zones with minimum energy requirement have made N js and N jsg the most useful parameters for characterizing the solid suspension operation. VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 619

14 EXPERIMENTAL MEASUREMENT OF N JS AND N JSG The performance of the solids suspension equipment is often measured in terms of the fraction of the solids resting at the vessel bottom for a specified interval of time (Baldi et al., 1978). Usually this criterion is related to N js and N jsg. Both experimental investigation and the theoretical (empirical) modelling of the solids suspension process in an agitated liquid continuum have received much attention in the past. Inevitably, the experimental determination of N js or N jsg is somewhat subjective depending on the accuracy of the measuring technique used and definition of the suspension criteria. Several experimental methods, visual and otherwise have been devised and have claimed to allow the determination of critical impeller speed required for solid suspension (Bourne and Sharma, 1974; Buurman et al., 1986; Raghava Rao et al., 1988; Rieger and Ditl, 1994), but the most simple and widely adopted method is the visual method introduced by Zwietering (1958). Visual Methods Zwietering (1958) introduced the visual observation method to determine the N js. The motion of the solids on the tank bottom was visually observed through the transparent tank wall and tank bottom with an aid of a mirror placed directly underneath it. The bottom of the tank was illuminated by photoflood light to aid the observation. N js was measured as the speed at which no solids are visually observed to remain at rest on the tank bottom for more than 1 or 2 s. This method is very subjective and measures the N js with an accuracy of ±5% for a given observer. Even when the gas was dispersed throughout the tank, the motion of the solids on the tank bottom can be clearly observed to determine N jsg. It was found that it sometimes took 2 to 3 min to attain steady state with regards to particle motion, especially at high gas rates. This method requires complete suspension of all the solids off the vessel bottom. It was also observed that a small amount of solids settle in relatively stagnant regions (e.g. around the periphery of the vessel bottom near the baffles or at the centre of the vessel bottom, where the liquid circulation is weak as compared to the bulk of the vessel) to form fillets. These fillets are generally insignificant from the practical point of view. It is necessary to increase the impeller speed substantially beyond N js (up to 20 50% more of N js ) to lift solids from these insignificant fillets resulting in more energy consumptions. It was observed that any extra input of energy in terms of higher impeller speed with the intent of raking particles out of the fillets tended to homogenize the solids concentration in the bulk of the tank before the elimination of fillets. Caution should, therefore, be exercised in applying the criterion of particles spending not more than a second or two at the tank base, for complete offbottom suspension. The loss of small amount of active solids in these fillets is often insignificant compared with the energy savings obtained by ignoring these fillets. Hence, wherever necessary, while determining N js by this method, a small amount of solid particles is allowed to settle in the relatively stagnant regions on the vessel bottom to form fillets (Nienow, 1992), i.e. the criterion for the determination on N js should be applied to the bulk of the vessel bottom excluding the fillet formation cites. Evidently, higher N js values would be observed if the view of observation includes the fillet cites also. Einenkel and Mersmann (1977) have proposed another method based on the visual observation of the height of the interface between the slurry (solid-liquid suspension) and the clear liquid. They defined N js as the speed at which the height of the interface (from the vessel bottom) between the slurry and the clear liquid is approximately 90% of the total liquid height (Figure 1d). This method does not appear to be a superior suspension criterion because the small particles suspended (for a wide particle size distribution), come to the top of the tank before the last particle lifts off the tank bottom. According to Kraume (1992), this speed was approximately 20 25% higher than that predicted by Zwietering s (1958) method. It is worth mentioning here that results of Kraume (1992) differ significantly from the findings of others for particle sizes less than 300µm. Hicks et al. (1997) measured the settled bed height at different impeller speed to determine the N js. They defined the N js as the speed at which the height of the settled bed is zero and a further reduction in the impeller speed will cause the solids to settle on the vessel bottom. One of the disadvantages of this method is that, due to the complicated flow pattern in the stirred tank, the bed surface is usually uneven, making it difficult to measure the accurate bed height. Hence, the bed height measured at a plane just above the tank bottom and at a point half way between the two consecutive baffles was used as the representative bed height. The main advantage of these visual methods is their intrinsic simplicity. It must be pointed out that these methods are affected by subjective evaluation, and only with a careful and skilled observation is it possible to get about ±5% reproducibility at best in the case of dilute suspensions. However, no data can be reliably reproduced for solids concentration above 8% w/w (Oldshue and Sharma, 1992). In addition, for all the visual method discussed above, it is necessary to build transparent vessels in order to observe the movement of the solids at the vessel bottom; this is possible for laboratory scale equipments, but it is impracticable for large-scale installations. Other Methods In order to measure the N js and N jsg for large-scale installations and opaque vessels, where the visual observation method cannot be applied, some additional methods have been reported in the literature. The first method is based on the measurement of variation in impeller power consumption with respect to the impeller speed (Figure 2) (Rewatkar et al., 1991a). Rewatkar et Figure 2. Typical impeller power number behaviour for the twoand three-phase system (Rewatkar et al., 1991a) 620 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

15 al. (1991a) have used this method to determine N js and N jsg for reactor size ranging from 0.57m to 1.5m. Figure 2 is a graphical representation of the power number behaviour with impeller speed for the two- and three-phase system. The exact values for the power number with impeller speed for a range of vessel and impeller geometry are reported in Rewatkar (1990). The power consumption behaviour for a solid-liquid system (for, X ~ %) is shown in Figure 2 (Rewatkar et al., 1991a) by the curve EFGH (dotted line). At low impeller speeds, the power number decreased with an increase in the impeller speed (along the line EF) due to the formation of fillets at the centre and along the periphery of the vessel bottom, which made the apparent bottom streamlined. These fillets reduce the impeller pumping action (false bottom effect) and promote a streamlined flow. At a speed corresponding to the point F, the fillets start breaking up leading to the suspension of small amount of solids. Along the line FG, more and more solids are suspended, thereby increasing the energy dissipation at the solid-liquid interface. Also, due to more number of particles in suspension, the average density of the suspension increases. Because of these two reasons, the power number increases with an increase in N along the line FG. Beyond point G, any further increase in the impeller speed was not found to change the power number and the power number remains constant along the line GH. Also at the same impeller speed, visual observations have indicated that the suspension of solid is complete (as per the 1 2s criterion), i.e. N = N js at point G. Thus, for a solidliquid system, N js can be taken as the speed beyond which power number remains constant with an increase in the impeller speed. The power consumption behaviour in a three-phase system is shown by the curves EFGG H and EFGG H (solid lines in Figure 2) for low (<6.67%w/w) and high (>6.67%w/w) solid loadings, respectively. In the presence of gas, the hydrodynamics of a three-phase system is more complex as compared to a solidliquid system. The formation of gas-filled cavities behind the impeller blade reduces the form drag behind the impeller blade, and the power consumption decreases. Beyond point F, break-up of these gas cavities start (i.e. shedding of bubbles), but the solid suspension begins only after the complete dispersion of the gas. Along the line FG, more and more gas bubbles are formed, and an increasing amount of solids are suspended. The increasing energy dissipation at the gas-liquid and the solid-liquid interfaces and the cavity break-up phenomenon results in an increase in the power number along the line FG. At lower solid loading (<6.67%w/w), the recirculation of the three-phase mixture becomes predominant before the critical impeller speed for solid suspension (N jsg ) is reached and N P starts decreasing. For higher solid loadings (X>6.67%w/w), the rise in the power number during the process of cavity break up and particle suspension follows the path FG. At higher solid loadings, the reduction in the impeller power due to the re-circulation of gas-liquid-solid mixture is compensated by the increase in the power due to more solids in the suspension and hence, the power number increased continuously though re-circulation of gas-liquid-solid mixture started much before N jsg is reached. In this case, the visually observed value of N jsg corresponds to the speed at which the power number has the maximum value, i.e. at point G. Beyond G, the power number continuously decreased with an increase in the impeller speed. The values of N js and N jsg measured by power consumption method were found to be within 5% of those measured by visual observation (Rewatkar et al. 1991a). Figure 3. Typical liquid phase mixing time behaviour in single, two- (gas-liquid and solid-liquid) and three-phase stirred tank reactors (Rewatkar et al., 1991a) The second method is based on the measurement of mixing time (θ mix ) with respect to the impeller speed (N) in liquid alone, solid-liquid, gas-liquid and gas-liquid-solid systems keeping all other geometrical and operating parameter constant (Raghava Rao and Joshi, 1988; Rewatkar et al., 1991a). In this method, in order to measure N js, the θ mix -N curve for a solid-liquid system is compared with the single-phase (only liquid) system, whereas to measure N jsg, the θ mix -N curve for a three-phase system is compared with that of gas-liquid system (Figure 3). Figure 3 is a representation of the mixing time behaviour with respect to the impeller speed and the actual values are reported in Rewatkar (1990) for the range of variables covered in their study. For single-phase system (in the absence of solids and gas sparging), the mixing time continuously decreased with an increase in the impeller speed as shown by line LM in Figure 3 (Rewatkar et al., 1991a). In the presence of solids, the values of mixing time were found to be higher than that of single-phase system. The liquid phase in the vicinity of settled solids is relatively stationary as compared to the bulk liquid, resulting in higher values of mixing time at a speed in the vicinity of point P. Above point P, the higher values of mixing time in the presence of solids may be attributed to the energy dissipation at the solid-liquid interface, which reduces the energy dissipation for the liquid circulation (reduction in the average circulation velocity) thereby increasing the mixing time values. At point Q, the difference between the mixing time values for single-phase and solid-liquid system is maximum. Also, the visual observations have indicated that, solids are completely suspended at an impeller speed corresponding to point Q. Thus in the case of a solid-liquid system, N js can be taken as the speed at which the difference between the mixing time in the presence and absence of solids is maximum. In the case, of a three-phase system (line PP Q QR in Figure 3), the mixing time behaviour is different as compared to a solidliquid system. The difference lies in the range P Q, where mixing time increased with an increase in the impeller speed due to the formation of more and more gas bubbles, similar to a gas-liquid system (line PP Q R). The increase in the energy VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 621

16 Figure 4. Typical count rate behaviour for the determination of N js (Rewatkar et al., 1991a) Figure 5. Typical solid concentration behaviour obtained by sampling near the vessel bottom (Chapman et al., 1983a) dissipation at the gas-liquid interface makes less energy available for liquid-phase circulation and this result in an increase in the mixing time. A maxima occurs in the mixing time curve at point Q, and the impeller speed at this point corresponds with the critical impeller speed for gas dispersion (N CD ). With further increase in the impeller speed, beyond point Q, the liquid-phase mixing time starts decreasing again. In this speed range, solids start suspending from the bottom of the vessel. As more and more particles get into suspension, the mixing time values start deviating (increasing) from that of gas-liquid system (similar to that of the deviation of the mixing time for solid-liquid system from the liquid alone), and the maximum deviation occurs at point Q, which matches with the N jsg observed visually. The values of N js and N jsg obtained with this method were found to be within 5% of those measured by the visual observations or by the power consumption method. One of the advantages of these two methods is that they can be used for opaque vessels and for large-scale installations. Rewatkar et al. (1991a) have also described the radioactive tracer technique for the measurement of N js and N jsg. This technique uses the signals (count rate) from radioactive tracer particles (solid particles labelled with radioactive material) to trace the motion of the solid particles within the stirred tank. The variation in the count rate with impeller speed agrees very well with the visual observation of solid concentration and helps in the prediction of N js or N jsg. For this purpose, the labelled solid tracer particles were added to the vessel along with the bulk of solid. The physical properties of the tracer particles were identical with the bulk of solids in the vessel. The digital signals (count rate) given by the tracer particles were continuously monitored with respect to impeller speed. The variation of scintillation count rate with impeller speed depends upon the detector position. Rewatkar et al. (1991a) have found that for a disc turbine, the optimal detector position was below the bottom and at the centre of the vessel, whereas for PTD the optimal position was below the bottom and at the periphery of the vessel. It was found that, with an increase in the impeller speed, the count rate continuously decreased along the line AB (Figure 4) as more and more particles are suspended. At point B, the count rate suddenly falls and remains practically constant beyond it. This shows that at point B the solids are completely suspended. The impeller speed at point B was found to agree well with the visually observed values of N js within 5%. Though simple, none of the above method provides information about the solids concentration profile (axial or radial) or the solids concentration distribution existing in the stirred vessel and hence, it is not possible to determine the values of N us or N usg with the help of these methods. Concentration Measurement Method (Local Values of Solids Concentration) Bourne and Sharma (1974) and Musil (1976) have proposed a more objective definition for N js, which not only gives the N js values, but also gives information about the solid concentration distribution prevailing in the vessel. According to them, N js can be defined as the impeller speed at which the particle concentration just above the vessel bottom shows a maximum or a discontinuity with respect to impeller speed. This method is also called the concentration peak method. Bourne and Sharma (1974) have used the sampling technique for the measurement of solids concentration. They showed that experimentally, for a wide range of geometries and particulate solids, a peak occurs when a graph of solids concentration in the withdrawn sample is plotted against the impeller speed (Figure 5), provided the sample be withdrawn from a point just above the vessel bottom. The peak in the concentration was found to coincide with the visually observed N js value. At low impeller speeds, majority of the particles rest on the vessel bottom resulting into low solids concentration in the sample withdrawn. With an increase in the speed, the particles gradually start suspending and hence the solids concentration just above the vessel bottom increases. Eventually, a speed is reached when the source of particle on the base is practically exhausted and hence, further increase in the impeller speed had a tendency to disperse the pre-suspended solids particles through out the vessel. Thus, the local solids concentration just above the vessel bottom reduces after giving a maximum (peak). The impeller speed corresponding to this peak is taken as the N js as shown in Figure 5. However, this 622 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

17 Figure 6. Typical solid concentration behaviour at different axial location (Dohi et. al., 2001) rather idealized description does not always apply. Musil (1976) noted that occasionally a kink or discontinuity in the form of sharp decrease in the gradient rather than a peak would occur. In order to obtain information about the axial solids concentration profile, the solids concentration should be measured at various points along the height of the vessel rather than measuring it only just above the vessel bottom. The sampling should be done isokinetically to ensure representative sample. This is not always easy, as the changing impeller speed will change the sampling conditions and this should be accounted for. Unlike visual methods or the mixing time and power consumption measurement methods, the method of measurement of the solids concentration along the height of the vessel gives information about ultimate suspension speed, N us and N usg. Dohi et al. (2001) have studied the determination of N usg with the help of measurement of solids concentration. According to them, when the solids concentrations measured at three different horizontal planes along the vessel height, (one near the top, one in the middle of the vessel and one near the vessel bottom) are plotted against the impeller speed, a point is reached where the slope of the concentration vs. impeller speed curve showed a sharp change. The impeller speed corresponding to this point is the N usg (Figure 6). From Figure 6, it can be seen that the concentration of solids withdrawn from three different locations along the vessel axis coincide with each other at N usg. Beyond it all, the three curves shows a constant value implying that the solid concentration is constant throughout the vessel and beyond N usg, there is no change in the degree of homogeneity of the suspension. Various methods have been reported in the literature for the measurement of solids concentration in a solid-liquid stirred vessel. These include sampling technique (Bourne and Sharma, 1974; Musil, 1976; Dohi et al., 2001), conductivity method (Musil and Vkl, 1978), optical method (Fajner et al., 1985), etc. The sampling technique is the most widely used method for the measurement of the solids concentration in the vessel although it is associated with certain drawbacks. Mac Taggart et al. (1993) have shown that the sample tube design (shape, diameter and tip angle) can significantly affect the solids concentration and the particle size distribution in the sample withdrawn. They concluded that sampling at isokinetic velocity (sampling velocity = liquid circulation velocity) with a tapered sampling tube is the ideal mode of sampling. Barresi et al. (1994) have also reported similar findings. In the sampling technique, there is no guarantee that the sample withdrawn will have a representative concentration due to the turbulent nature the fluids. Also, two or more samples need to be collected to eliminate the statistical variation in the reading. One of the main problems associated with sampling technique is that the introduction of the mechanical sampler (sampling tubes need to be inserted inside the vessel in order to collect the samples in the radial direction) may disturb the fluid velocity patterns particularly in an un-baffled small vessels. In order to overcome these problems related with the sampling technique, the methods such as conductivity measurement method (Musil and Vkl, 1978) and optical methods (Fajner et al., 1985) have been introduced. The conductivity method is based on the measurement of the change in the conductivity of a salt solution due to suspension of the non-conducting particles present between two electrodes (Musil and Vkl, 1978). The conductivity values are then calibrated to give the solids concentration. This method also requires the insertion of the conductivity probe inside the vessel, which may result in the disturbance of the flow pattern. Fajner et al. (1985) have used the optical method to measure the solids concentration inside a stirred vessel. The measurement system consists of a Light Emitting Diode (LED) as the light source and a miniature silicon-diode as the receiver (both source and receiver located outside the vessel). Any disturbance in the source light due to the solid particles is recorded on the receiver and the data from the receiver is transferred to a computer. Solids concentration has been measured from the calibration curves. It may be noted that as both the source and the receiver are located outside the vessel, there is no chance of any disturbance in the fluid velocity pattern. But this method needs either a transparent vessel or transparent windows on the opposite side of the vessel wall. Pressure Measurement Method (Overall Values of Solids Concentration) The techniques described above give the values of the solids concentration at particular points, i.e. local values of solids concentration. Brucato et al. (1997) have modified the pressure measurement technique of Biddulph (1990) in order to get the overall values of the solids concentration in the vessel rather than the local values. The variation in the pressure at the vessel bottom due to the suspension of particle was measured with the help of pressure probe. The pressure readings are calibrated to get the value of the solids concentration in the vessel. In order to understand how pressure measurement can give information on the amount of suspended solids, one may consider that when solid particles lie on the vessel bottom their apparent weight is borne by the bottom itself by direct mechanical interaction with the solids. On the other hand, when particles are lifted into suspension, their apparent weight is borne by the liquid, which eventually discharges it on the vessel bottom as a pressure increase with respect to the case where no particles are suspended. Also, one may consider that when particles are suspended, the apparent density of the agitated liquid phase increases, resulting in a greater hydrostatic head on the vessel bottom. Brucato et al. (1997) have proposed a relationship between the mass of suspended solids (M S ) and the observed pressure increase ( P S ) as follows: VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 623

18 ( ) ( ) VS ρs ρ1 g 1 ρl / ρs Ps = = A 2 b πt / 4 Ms (1) It may be noted that as the no-agitation (zero impeller speed) condition ensures that all solids are at rest on the vessel bottom, the pressure reading at any point on the bottom in this condition, corresponds to the hydrostatic head of the liquid inside the vessel. This pressure is considered as the reference pressure and it is subtracted from the readings obtained at higher agitation speeds, when some or all of the solids are suspended. The value of pressure increase obtained in this way would then be immediately converted by means of Equation 1 into the corresponding mass of suspended solids. The dynamic effects due to the liquid motion inside the vessel under agitated condition also give rise to a contribution to the measured pressure. It is therefore necessary to account for the dynamic head effects in order to extract the pressure increase solely due to particle suspension from the observed total pressure increase. For this, Brucato et al. (1997) and Micale et al. (2000) have considered that above the complete suspension the fluid-dynamic properties of the twophase mixture become independent of agitation speed. The compensation for the dynamic effects can be done as follows: fitting a suitable parabola to carefully chosen points in the high agitation speed region (solid parabola in Figure 7a); subtracting the estimated dynamic effects from the total pressure measured at the vessel bottom and translating it by means of Equation 1, this result into a fractional solids suspension data (empty squares in Figure 7b). Figure 7 shows negative values of the pressure (Figure 7a) and the corresponding solid concentration (Figure 7b) at low agitation speeds. According to Micale et al. (2000), these negative values are probably due to the fact that at low agitation speeds, the vessel bottom is reshaped by the settled solids (false bottom effect). Hence, the liquid flow field near the vessel bottom is strongly affected, resulting in a local pressure decrease at the particular location where the measurements were taken. This observation was supported by the consideration that no negative pressure values were observed at the same location when no solids were present in the system. These authors have also observed the same negative pressure effect for solid-liquid system. But the extent of this negative pressure was smaller than that observed with threephase systems. The reason for this was unclear. Brucato et al. (1997) have shown that when the fraction of the suspended solids was plotted versus agitation speed, an S-shaped curve was obtained that can be fitted by Weibull function as follows: 0 For N<N min 2 XS = N N 1 - exp min For N N N min span One of the advantages of using Weibull function in this case is that the values of the two parameters (N min and N span ) have an immediate physical meaning. N min is the value of N at which the suspension phenomena starts, while N span is such that twice its value gives the range of N in which most of the suspension takes place. At N-N min = 2 N span the value of X is 0.98, indicating that at this agitator speed less than 2% of the solids are still settled on the tank bottom. It is worth noting that, the lack of participation of these 2% solids (may be present as fillets) in the processes could hardly be considered significant. Based on the above (2) Figure 7. Typical pressure and solid concentration behaviour with correction for the dynamic head effect (Micale et al., 2000) considerations, Brucato et al. (1997) have defined the criteria for attaining the sufficient suspension as: N ss = N min + 2 N span (3) where, N ss minimum impeller speed required for sufficient suspension. It is clear that the N ss defined here is somewhat similar to the N js defined by Zwietering and very close to the 98% suspension criteria introduced by Chaudak (1982). Brucato et al. (1997) have compared the values of N ss calculated experimentally by pressure measurement method (using Equation 3) with the N js values estimated by Zwietering s (1958) correlation. They showed that the adoption of the visual criterion (N js ) instead of the quantitative criterion (N ss ) results in an unjustified over deign of the reactor. This method can also be used to measure N jsg (Takenaka et al., 2001) since the gas hold up does not contribute significantly to the static head as compared to the static head due to liquid and solid. Comments: In general, the visual observation of the suspension to determine N js or N jsg, though simple and widely used, is concerned only with the behaviour of solids in the proximity of the vessel bottom and does not give any information about the distribution of the solids throughout the vessel. Also, it is applicable for transparent vessels with subjective evaluation and no data can be really reproduced for solid concentration above 8% by weight (Oldshue and Sharma, 1992). In order to get the data on the distribution of solids throughout the vessel one has 624 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

19 Figure 8. The process of gas-liquid-solids dispersion in a threephase mechanically agitated reactor (Pantula and Ahmed, 1998) to use the methods based on measurement of solid concentration in the vessel, which gives information not only about the just suspension and the ultimate suspension speed, but also gives an idea of the distribution of the solids throughout the vessel. For concentration measurement, sampling method is simple and less expensive but requires the insertion of sampling tube inside the vessel, which may disturb the fluid velocity pattern inside the vessels. The non-intrusive optical methods, thus, seems to be more effective to measure the solids concentrations inside the stirred vessel, without disturbing the fluid velocity pattern. This is a more suitable method for large-scale installations and opaque vessels. The experimental concentration profiles obtained by this method were also found to be in good comparison with those obtained by CFD models (Montante et al., 2001). The only drawback associated with the concentration measurement method is that it gives local values of the solids concentration in the vessel rather than the overall values. In order to get the later, measurement of static head, due to liquid and suspended solids, at the vessel base with the help of pressure probes may be used. Adoption of the sufficient suspension criterion, based on the pressure measurement technique, in place of conventional visual criterion for complete off-bottom suspension is shown to result into important savings in agitation costs. MINIMUM IMPELLER SPEED FOR SOLID SUSPENSION IN PRESENCE OF GAS (N JSG ) Three-phase systems involve the simultaneous dispersion of gas and the suspension of solids in the liquid, which is the continuous medium. The energy required is supplied by mechanical agitation. As discussed earlier, the main parameter to characterize the impeller hydrodynamics in a three-phase system is N jsg. This parameter strongly depends on the gassing rate, impeller design, location of the impeller from the vessel bottom, type of sparger and location of sparger, vessel geometry (especially the design of the vessel bottom), physical properties of the liquid and solid, etc. Thus, while designing a three-phase mechanically agitated reactor, it is of utmost importance to consider the effect of the operating and geometric parameters on N jsg. Figure 9. The generation of hold-up in a mechanically agitated G-L-S reactor. The correspondence to these regimes identified in Figure 8 is shown (Pantula and Ahmed, 1998) Effect of Gas Sparging Several research reports (Zlokarnik and Judat, 1969; Queneau et al., 1975; Arbiter et al., 1976; Subbarao and Taneja, 1979; Wiedmann et al., 1980; Chapman et al., 1983c; Wong et al., 1987; Rewatkar et al., 1991a) have shown that the solid suspension mechanism in a three-phase system is mainly determined by the gas-liquid hydrodynamics of the impeller. In a stirred gasliquid or gas-liquid-solid system the gas flow rate, distributable by the stirrer, is limited and dependent on the stirrer speed. According to Dylag and Talaga (1994) at flooding condition, large gas bubbles were also observed detaching themselves from the gas trails (cavities) formed behind the impeller blades. These gas bubbles flow along the impeller shaft towards the free liquid surface. An impeller under flooding condition does not generate any significant liquid phase circulation in the vessel. The energy conveyed by the gas makes up the major part of the energy supplied to the system. In the three-phase system, the flooding of the stirrer reduces the impeller power draw, thereby reducing the liquid-phase circulation resulting into the sedimentation of the suspended particles. An increase in the impeller speed (at constant gassing rate) or decrease in the gas flow rate (at constant impeller speed) may lead to the re-suspension of the solid particles. The dispersion of solid particles in a three-phase system only takes place when the impeller speed has increased beyond the critical impeller speed for gas dispersion (N CD ). The sequence of events in terms of gas dispersion and solids suspension in a three-phase reactor is depicted in Figure 8, while the typical gas hold-up behaviour as a function of the power input shown in Figure 9, for Rushton turbine and Pitched blade turbine (Pantula and Ahmed, 1998). At low impeller speeds, gas is dispersed around the impeller shaft only. The solids are settled at the vessel bottom and the vessel behaves like a partial bubble column with single point sparger (Figure 8a). This condition of the impeller is termed as flooding condition. For a mechanically agitated threephase system; this regime is characterized by gas-liquid mixing with very few solids in the suspension. The exception would be the cases where the density difference (between solids and liquid) or where the particle size of the solids is low, leading to VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 625

20 more suspension of the solids. On further increase of the impeller speed, the gas is dispersed throughout the region above the impeller, thus resembling a bubble column with multi-point sparger as in Figure 8b. At this stage, the concentration of particles suspended is still not sufficient to influence the gasliquid hydrodynamics of the impeller. For regimes given by Figure 8a and b, the gas hold-up increases linearly with the increase in impeller power input (Figure 9). This region is shown as Region A in Figure 9. Figure 8b shows the gas dispersion region, where sufficient power is supplied by the impeller to distribute the gas throughout vessel. The corresponding impeller speed would be the critical impeller speed for complete gas dispersion (Nienow et al., 1978). Very few solid particles are in suspension even at this impeller speed. However, on further increase in the power input, a sharp change in the rate of increase in the gas hold-up is observed (a marked change in slope) as indicated by Figure 9. This region is shown as Region B in Figure 9. With further increase in power input, more and more solids are suspended. Figure 8c shows the just suspension point, as described by Zwietering (1958). At this point, another transition in the gas hold-up characteristics is observed. The solids are not uniformly distributed throughout the vessel at this stage. There may be a clear liquid layer at the top. The solids reach a certain height, which depends on the vessel geometry, particle properties, and the gassing rate. On further increase in the power input, uniform suspension is achieved as shown in Figure 8d and gas is completely dispersed with some secondary circulation above the impeller. In this region however, the increase in impeller speed tends to produce diminishing returns in terms of hold-up. This region is shown as Region C in Figure 9. Arbiter et al. (1976) were the first to deal with the three-phase systems. They have observed drastic sedimentation of the suspended particles when the gas flow number, Fl G (= Q G /ND 3 ) exceeds a critical value. This critical value of gas flow number was found to be a function of particle size and coincided with the point, on the P G /P-Fl G curve, where the power drawn by the agitator decreased suddenly with a small increase in the gassing rate. According to Nienow et al. (1978), the minimum in the P G /P-Fl G curve also corresponds with the flooding point. This indicates that above critical gas flow number (corresponding to the flooding point), the impeller is flooded and hence the capacity of the impeller to suspended solids decreases, resulting into sedimentation of the particle. Whereas, the sudden drop in the P G /P-Fl G curve for the geometry used by Arbiter et al. (1976) was explained as being due to particles blocking the suction port of the impeller (shrouded impeller), which is unlikely to occur in the un-shrouded system. Queneau et al. (1975) and Wiedmann et al. (1980) have also observed a quantitatively similar phenomenon of the sedimentation of suspended solids on the introduction of gas. Zlokarnik and Judat (1969) have found that an approximately 30% higher impeller speed over N js is required to ensure the re-suspension of the solids on the introduction of gas (i.e. N jsg =1.3N js ). All the above studies have reported the sedimentation of the solids on the introduction of the gas, but have not explained why and under what condition this phenomenon occurs. According to Subbarao and Taneja (1979), on aeration an entrained liquid flow, proportional to the gassing rate, would be set-up in a counter-current direction to the propeller induced liquid flow and hence hinder the suspension of the solids. They summarized the effect of gassing on solids suspension in terms of a critical gas flow number below, which good dispersion of gas and suspension of particles is possible, defined as: ( ) = FlG crit QG (4) N N jsg D ( ) This critical gas flow number agreed closely with the critical gas flow number proposed by Arbiter et al. (1976). However, this was probably a coincidence since the blocking mechanism, which causes solids sedimentation in the close impeller geometry proposed by Arbiter et al. (1976) for the sudden sedimentation of the suspended solids on the introduction of the gas, was not possible in the open impeller geometry used by Subbarao and Taneja (1979) for which the sedimentation was not sudden but rather occurs over a range of gassing rate. Chapman et al. (1983c) explained the sedimentation phenomenon on the basis of the decreased liquid pumping capacity and the power input of the impeller on introduction of the gas. According to them, any decrease in the liquid pumping capacity and power input reduces all the parameters responsible for solid suspension. They did not notice the dramatic collapse of suspension as reported by Arbiter et al. (1976). At a constant gassing rate, an increase in the impeller speed from N js to N jsg ensures the complete re-suspension of the solids. Under a gassed condition, such that N<N jsg, a reduction in Q G ensures the complete re-suspension of the solids. In fact, they observed a unique combination of gassing rate and impeller speed at which the just suspended condition is achieved for a particular system property. As a result, a net increase in the power input from the agitator is required to cause suspension under aerated condition suggesting that the gas presence has an additional effect other than a reduction of liquid flow from the impeller region, in terms of damping of the local turbulence and liquid velocities near the vessel base. Warmoeskerken et al. (1984) have explained the sedimentation phenomena on the basis of decrease in the impeller power consumption due to the formation of gas-filled cavities behind the impeller blades. The cavities reduce the pumping capacity of the impeller to a point where the liquid flow generated by the impeller is no longer sufficient to keep the particles suspended. The solid suspension process in a three-phase system is greatly affected by the hydrodynamics of the impeller in the gas-liquid system that in turn, is significantly influenced by the gas-filled cavities formed behind the impeller blades on introduction of gas. Three characteristic cavity forms, which develop successively with an increasing gassing rate at a constant stirrer speed (for a Rushton turbine), are described as vortex, clinging, and large cavities. Warmoeskerken et al. (1984) have observed that, in a three-phase system with Rushton turbine, suspension is at first (at a low gassing rate) more easily maintained than in a liquid-solids system. They have observed an improvement in the suspension conditions, on introduction of gas, over the range of gassing rate up to Fl G = This was more remarkable since the impeller power demand falls to about half its un-gassed value in this region. Above Fl G = 0.03 sedimentation of suspension was observed under otherwise constant conditions with an increase in Fl G. It was only beyond this point that higher stirrer speeds were required to keep the particles in suspension. Warmoeskerken et al. (1984) concluded that suspension, as such, is not determined solely by the specific impeller power input. The local airlift pumping effect associated with the introduction of gas from the sparger may be helping the suspension process, though the power input from this source is only a few percent of that from the agitator. Also, it may be possible that flow irregularities resulting from the somewhat indeterminate conditions around vortex and clinging cavities contribute to 626 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING VOLUME 83, AUGUST 2005

21 the improved suspension performance at low gassing rate (Fl G 0.03). These effects are counterbalanced at higher gassing rate by the steady development of large cavities with the associated loss of impeller power and pumping capacity. Studies by Warmoeskerken et al. (1984) have confirmed that the scale of cavities determine the impeller power under aeration, and the major decrease in the impeller power draw is observed only after the formation of large and clinging cavities. Also, the introduction of gas changes the liquid flow pattern in the vessel. Chapman et al. (1983b) have observed following mixing stages with disc turbine when the impeller speed is increased for a particular gas rate: negligible dispersion; upper part of vessel acting as bubble column; gas circulation in the upper part of the vessel with occasional movement in the lower region (onset of flooding); gas circulation throughout the whole vessel (solid suspension starts only after this stage); secondary loops form and gross re-circulation. Frijlink et al., (1990) have reported similar findings for disc turbine and inclined blade impellers. However, they observed that the impellers such as the disc turbine are less sensitive to the gassing rate as compared to the pitched blade turbine. Thus, it can be said that the effect of introduction of the gas on the suspension process also depends on the type and geometry of the impellers. According to Rewatkar et al. (1991a), reduction in the impeller power in a three-phase system is due to two reasons viz. i) formation of solid fillet at the centre and along the periphery of vessel bottom (i.e. at the junction of vessel wall and vessel bottom). These fillets reduce the baffle action and promote a streamlined flow, which results in decrease in power consumption; ii) formation of gas filled cavities behind the impeller blades. The cavities reduce the form drag behind the impeller blade resulting into reduction of the impeller power draw. The reduction in the power consumption due to cavity formation is of greater magnitude than the reduction due to fillet formation. This also confirms that, in a three-phase system, the capacity of impeller to suspend solids is mainly determined by the impeller hydrodynamic in gas-liquid system rather than that in a solidliquid system. Warmoeskerken et al. (1984) have also reported that the solids have no measurable influence on the gas-liquid hydrodynamics of the impeller. Also, it should be noted that with a three-phase system, the entire gas becomes dispersed before the suspension of solids. Therefore, the statement can be postulated that whether or not a disperse gas-liquid-solid system is formed is determined by the conditions under which the suspension is generated (Dylag and Talaga, 1994). Thus, if impeller is flooded with gas then suspension of the solids is not possible. Effect of Type of Impeller The main source of energy in a stirred contactor is the power dissipation by the impeller rotation. In a three-phase system, the main task of an impeller is to suspend solids as well as to disperse gas simultaneously and effectively. The gas-liquid hydrodynamics of three-phase stirred contactor depend on the type of the impeller used (Frijlink et al., 1990). Different impeller generates different liquid-phase flow pattern and hence, has different hydrodynamic properties in a three-phase system. Thus, the choice of a proper impeller to satisfy the requirement of simultaneous solid suspension as well as gas dispersion with minimum power requirement is the key for success and the economy of the process. Figure 10. Schematic of the various impellers and their typical flow patterns Researchers have studied a variety of conventional impellers, like the Rushton turbine (radial flow), angled blade impeller, propeller (axial flow) and pitched blade turbine, in both uppumping and down-pumping mode (mixed flow impeller) for the solid suspension in a three-phase system. Recently, studies on the solid suspension in a three-phase stirred contactor with some modern impellers, like Scaba 6SRTG, Chemineer HE3 and Lightnin A315 etc., have also been reported (Wong et al., 1987; Neale and Pinches, 1994; Lehn et al., 1999). The schematic of the various impellers studied is shown in Figure 10. According to Zwietering s (1958) 1 2 s criteria, the N js or N jsg depend on the position on the vessel bottom from which the last particle is suspended. This position depends on the impeller type and impeller geometry. The radial flow impellers suspended particles last from an annulus around the centre of the vessel bottom, particularly at the four points in line with the baffles. The axial and mixed flow impellers tended to suspend the last particles from around the periphery of the vessel bottom, i.e. at the junction of the vessel wall with the vessel bottom. This is because of the different flow pattern generated by these two different types of the impellers. The flow pattern of an axial flow impeller favours easier suspension in comparison to the flow pattern produced by a radial flow impeller. The suspension efficiency of the mixed flow impeller remains in between. It VOLUME 83, AUGUST 2005 THE CANADIAN JOURNAL OF CHEMICAL ENGINEERING 627