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An investigation of gas/liquid mass transfer in mechanically agitated pressure leaching systems DeGraaf, Kenneth Brant 1984

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AN INVESTIGATION OF GAS/LIQUID MASS TRANSFER IN MECHANICALLY AGITATED PRESSURE LEACHING SYSTEMS by KENNETH BRANT DEGRAAF •A.Sc. (Chemical Engineering), Queen's University, 1 A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF MASTER OF APPLIED SCIENCE in THE FACULTY OF GRADUATE STUDIES (Department of Metallurgical Engineering) We accept this thesis as conforming to the required standard THE UNIVERSITY OF BRITISH COLUMBIA September, 1984 © Kenneth Brant DeGraaf, 1 9 8 4 In p r e s e n t i n g t h i s t h e s i s i n p a r t i a l f u l f i l m e n t of the requirements f o r an advanced degree a t the U n i v e r s i t y o f B r i t i s h Columbia, I agree t h a t the L i b r a r y s h a l l make i t f r e e l y a v a i l a b l e f o r r e f e r e n c e and study. I f u r t h e r agree t h a t p e r m i s s i o n f o r e x t e n s i v e copying of t h i s t h e s i s f o r s c h o l a r l y purposes may be granted by the head o f my department o r by h i s o r her r e p r e s e n t a t i v e s . I t i s understood t h a t copying or p u b l i c a t i o n o f t h i s t h e s i s f o r f i n a n c i a l g a i n s h a l l not be allowed without my w r i t t e n p e r m i s s i o n . Department of Metallurgical Engineering The U n i v e r s i t y of B r i t i s h Columbia 1956 Main Mall Vancouver, Canada V6T 1Y3 October 1, 1984 i i ABSTRACT Oxygen pressure leaching rates have been accelerated to the point where oxygen consumption rates of 0.75 moles of 0 2 per l i t r e of leach s o l u t i o n are obtained within a residence time of the order of 1 hour. At these high oxygen consumption rates the mass transfer of dissolved oxygen at the g a s / l i q u i d i n t e r f a c e may become rate-determing. The purpose of t h i s study has been to examine g a s / l i q u i d mass transf e r rates i n mechanically agitated pressure leaching systems. Using an 0 2 -Na 2S0 3 system to measure the oxygen mass tr a n s f e r rates, the e f f e c t of a number of process v a r i a b l e s on the mass transf e r r a t e s , impellor gas pumping rates, and the volumetric power requirements has been studied. The experimental work was done using both bench-scale (2 l i t r e s and 20 l i t r e s ) and p i l o t - s c a l e (2100 l i t r e s ) equipment. I t i s demonstrated that dimensionless mixing c o r r e l a t i o n s previously applied are not u s e f u l for extrapolating or scaling-up a mechanically agitated g a s / l i q u i d mass tr a n s f e r system. It i s shown that g a s / l i q u i d systems are more appropriately described and scaled-up i n terms of the impellor t i p v e l o c i t i e s and gas pumping c h a r a c t e r i s t i c s . *The existence of a c r i t i c a l impellor v e l o c i t y , which corresponds to the p o i n t when bubbles f i r s t form at the i m p e l l o r i s confirmed experimentally and discussed t h e o r e t i c a l l y . The g a s / l i q u i d mass transfe r rate i s found to be d i r e c t l y r e l a t e d to the gas pumping capacity of the impellor. The p r a c t i c a l Implications of t h i s p r i n c i p l e i n the d e s i g n of a g i t a t i o n f o r g a s / l i q u i d systems are shown experimentally to have great p o t e n t i a l for improving g a s / l i q u i d mass transfer rates and reducing the power consumption of an a g i t a t o r . i i i TABLE OF CONTENTS Abstract Table of Contents List of Tables List of Figures Acknowledgements Forward Chapter 1. INTRODUCTION 1 1.1 Pressure Oxygen Leaching Systems 1 1.2 Reaction Kinetics in Pressure Oxygen Leaching 3 1.3 Mass Transfer in Pressure Oxygen Leaching 8 1.4 Agitation Theory 15 1.5 Summary 24 2. GAS-LIQUID MASS TRANSFER IN PRESSURE OXYGEN LEACHING 26 2.1 Agitation Theory of Gas Dispersions - 26 * 2.2 Previous Work on Gas-Liquid Mass Transfer 32 2.3 Oxygen-Sodium Sulphite System 33 2.4 Purpose and Scope of the Present Investigation 36 3. EXPERIMENTAL DETAILS 38 3.1 Materials 38 3.2 Apparatus 38 3.3 Experimental Procedures 43 i i i v v i x iv 4. RESULTS AND DISCUSSION 46 4.1 Cominco Mixing Model Experiments 46 4.1.1 Single Impellor Systems 47 4.1.2 Dual Impellor Systems 53 4.1.3 Oxygen Concentration Effects 57 4.1.4 Small Diameter-High Speed Impellor Tests 59 4.1.5 Special Sparging Mode Experiments 60 4.1.6 Comparison of Results with Theory 61 4.2 Bench Scale Experiments 64 4.2.1 Single Impellor Systems 64 4.2.2 Dual Impellor Systems 69 4.2.3 Effect of Oxygen Concentration 70 4.2.4 Effect of Solids 71 4.2.5 Comparison of Results with Theory 71 4.3 Summary 73 5. CONCLUSIONS,APPLICATIONS AND RECOMMENDATIONS 77 5.1 Conclusions 77 5.2 Applications 79 * 5.3 Recommendations for Further Work 82 6. REFERENCES 85 FIGURES 1 to 31 88 TABLES 1 to 15 119 APPENDICES A to J 133 V LIST OF TABLES Table 1. Effect of Impellor Type 119 Table 2. Effect of Depth of Impellor Immersion 120 Table 3. C r i t i c a l Tip Velocity Correlation 121 Table 4. Effect of Impellor Diameter-Mixing Model 122 Table 5. Effect of Half Baffles on Mass Transfer-Mixing Model 123 Table 6. Effect of Impellor Immersion with Half Baffles 124 Table 7. Oxygen Depletion in Gas Bubbles 125 Table 8. Effect of Oxygen Concentration 126 Table 9. Effect of Oxygen Concentration 127 Table 10. Small Diameter-High Speed Impellor Experiments 128 Table 11. Effect of Special Sparging Mode 129 Table 12. Effect of Impellor Diameter-Bench Scale 130 Table 13. Effect of Baffle Length-Bench Scale 131 Table 14. Alternate Dual Impellor Configurations-Bench Scale 132 Table 15. Effect of Solids 133 v i Figure 1. LIST OF FIGURES Zinc Pressure Leach Process at Cominco's T r a i l Operations 88 Figure 2. Anaconda Arbiter Process 89 Figure 3. Sherritt-Gordon Ammonia Oxygen Pressure Process 90 for Ni-Cu-Co Figure 4. Reaction Steps in a Typical Leach 91 Figure 5. Liquid-Side Mass Transfer Models 92 Figure 6. Liquid-Phase Concentration Profile for Mass Transfer with a Chemical Reaction 93 Figure 7. Liquid-Solid Interface 94 Figure 8. Effect of Agitation on Liquid-Solid Mass Transfer 95 Coefficient Figure 9. Reynolds Number Correlates Dimensionless Parameters 96 Figure 10. Typical Experimental Rate Curve for the Oxidation 97 of Sodium Sulphite v i i Figure 11. Autoclave Mixing Model 98 Figure 12. Effect of Impellor Type on the Oxygen Mass 99 Transfer Rates Figure 13. Impellor Positioning in Mixing Model 100 Figure 14. Effect of Impellor Immersion Depth on Oxygen Mass 101 Transfer Rate for the 4-Bladed Axial Impellor Figure 15. Effect of Impellor Immersion Depth on Oxygen Mass 102 Transfer Rate for the 4-Bladed Radial Impellor Figure 16. Effect of Impellor Immersion Depth on Oxygen Mass 103 Transfer Rate for 6-Bladed Radial Disc Impellor Figure 17. Effect of Impellor Diameter on Oxygen Transfer 104 Rate for the 4-Bladed Axial Impellor figure 18. Effect of Impellor Diameter on Oxygen Transfer 105 Rate for the 4-Bladed Radial Impellor Figure 19. Effect of Impellor Diameter on Oxygen Transfer 106 Rate for the 6-Bladed Radial Disc Impellor Figure 20. Effect of Baffle Length on Oxygen Transfer Rate 107 in the Mixing Model v i i i Figure 21. Standard Dual Impellor Configuration Used in the 108 Commercial Autoclave Figure 22. Effect of Baffle Length on Oxygen Transfer Rate 109 for the Standard Dual Impellor Configuration Figure 23. Alternate Dual Impellor Configurations-Unsparged 110 Figure 24. Alternate Dual Impellor Configurations-Sparged 111 Figure 25. Effect of Gas Plenum Oxygen Concentration on the 112 Oxygen Mass Transfer Rate Figure 26. Effect of Impellor Type on the Oxygen Transfer 113 Rate in 20-litres Vessel Figure 27. A) Effect of Agitation Rate on Impellor Gas Pumping 114 Capacity and Oxygen Mass Transfer; B) Effect of * Surface Aeration of Impellor on Power Consumption (6-Bladed Radial Disc Impellor - 58mm diameter) Figure 28. A) Effect of Agitation Rate on Impellor Gas Pumping 115 Capacity and Oxygen Mass Transfer; B) Effect of Surface Aeration of Impellor on Power Consumption (6-Bladed Radial Impellor - 58mm diameter) ix Figure 29. A) Effect of Agitation Rate on Impellor Gas Pumping 116 Capacity and Oxygen Mass Transfer; B) Effect of Surface Aeration of Impellor on Power Consumption (6-Bladed Axial Impellor - 58mm diameter) Figure 30. A) Effect of Agitation Rate on Impellor Gas Pumping 117 Capacity and Oxygen Mass Transfer; B) Effect of Surface Aeration of Impellor on Power Consumption (Standard Dual Impellor - 58mm diameter) Figure 31. Effect of High Oxygen Partial Pressures on Oxygen 118 Mass Transfer Rate (58mm diameter) X ACKNOWLEDGEMENTS I would like to express my appreciation to Professor Ernest Peters for his assistance, guidance and support throughout the course of this project. And to Dr. G.M. Swinkels, I give thanks for his interest and support in helping to initiate this work. I would also like to acknowledge the assistance and cooperation of George Parker and Alex Mackie during my stay at the Cominco Technical Research Centre in T r a i l . Thanks i s also extended to Christine Harrison, Neil Walker, Horst Tump, Ross McLeod and Ed Klassen for their kind help and cooperation with various portions of the project. The financial support of the National Sciences and Engineering Research Council of Canada, and the B.C. Science Council is gratefully Acknowledged. Gratitude i s also extended to Cominco for their industrial sponsorship of this project. FORWARD "Science without religion is lame, religion without science is blind ... A religious person is devout in the sense that he has no doubt of the significance of those super-personal objects and goals which neither require nor are capable of rational foundation ... Subtle is the Lord, but malicious he is not ... Nature hides her secret because of her essential loftiness, but not by means of ruse " Albert Einstein The Lord is my shepherd; I shall not want." Ps. 23:1 'The Science and Life of Albert Einstein", by Abraham Pais, Oxford Press 1 9 8 2 . 1 1. INTRODUCTION 1.1 Pressure Oxygen Leaching Systems One of the most important developments i n the f i e l d of hydrometal lurgy has been the app l i c a t i on of e levated pressures and temperatures to the leaching of complex metal sulphide o r e s . Processes invo l v ing ox idat i ve pressure leaching have received a great dea l of a t ten t ion i n recent years ; for example, there i s the z inc pressure leaching process which i s i n commercial operat ion at Cominco's operat ions at T r a i l , B .C. . Even though there are as yet only a few commercial p lants i n ex is tence , and convent ional smelt ing s t i l l i s the main processing method rou t ine l y prac t i sed today, the d i r e c t ox idat ive pressure leaching of sulphides remains compel l ing . With t ime, the development of f u l l y competi t ive pressure ox idat ive leaching plants for many metals i s expected as the benef i ts of these hydrometa l lurg ica l plants are recognized. The ox idat i ve pressure leaching of metal sulphides i s an anodic d i s s o l u t i o n process and can be described in general by the f o l l o w i n g 1 : MS M m + + S ° + me~ [1.1] N n + + e" - N ( n _ 1 ) + [1.2] 2 where MS = metal sulphide and N i s an unidentified oxidizing agent. I [ | Besides oxygen, f e r r i c ion (Fe ) or n i t r i c acid can be used as an oxidizing agent in the oxidative leaching of metal sulphides. In oxidative pressure leaching processes, the leaching of metal sulphides leads to the d i s s o l u t i o n of the metal ions (M m +) i n the aqueous phase and the formation of elemental sulphur accordingly: MS + 2H + + j 0 2 M 2 + + S° + H 2 0 [ 1 .3] Once formed elemental sulphur may be oxidized further to form sulphate by oxygen or other oxidants present in the leach sytems i . e . 2 [1.4] [ 1 . - 5 ] Since oxidative pressure leaching produces sulphur In the elemental form, one of the major benefits of the process is the elimination of air 'pollution due to the production of sulphur dioxide. Furthermore, sulphur in the elemental form is easier to store and handle, and i t is directly marketable. 2 S ° + 30- + 2H o0 •* 4H + + 2S07 2 2 4 S° + eFe 4"^ + 4H„0 •> eFe"^ + 8H + + SO? 3 1 . 2 Reaction Kinetics in Pressure Oxygen Leaching The reaction mechanisms found In oxidative pressure leaching are complex and varied. In some cases, the dissolved oxygen in the aqueous phase reacts directly with the metal sulphide. In other cases, the oxygen reacts homogeneously with an intermediate species ( i . e . ferrous ion, F e + + ) ; the oxidized intermediate species in turn reacts with the metal sulphides. The controlling kinetic mechanism in an oxidative pressure leaching system has a profound influence on the optimum reactor design. Depending on the percent extraction and the average rate of reaction that i s acceptable in a reactor, the controlling kinetic mechanism w i l l dictate the use of either a plug flow reactor or a backmixed reactor 3. Where the reaction driving force is controlled by concentration gradients in the solution a plug flow reactor w i l l perform better than a backmixed reactor. In plug flow both percent extraction and average reaction rate are maximized. For autocatalytic reactions, *the choice of reactor design is governed by the percent extraction. At high enough conversions the plug flow reactor gives superior reaction rates, while at low conversions the backmixed reactor w i l l give higher reaction rates. The other extreme is where the controlling kinetic mechanism is mass transfer limited - zero-order kinetics. In this case there is no difference between selecting plug flow or backmixed configurations. However, what i s important under these circumstances i s the means by which mass transfer Is promoted. 4 Examination of the reaction kinetics in a few commercial oxidative pressure leaching processes w i l l emphasize the c r i t i c a l importance of these design considerations. Three commercial oxidative pressure leaching processes that i l l u s t r a t e this are as follows (listed in chronological order of development): 1) Sherrltt Gordon Ammonia Pressure Process for Ni-Cu-Co 2) Anaconda Arbiter Process for Cu 3) Cominco's Zinc Pressure Leach Process The Sherritt Gordon ammonia pressure leach process (shown in Figure 1) treats a pentlandite-nickel concentrate in a two-stage counter -current leach in which the pregnant liquor i s produced in the f i r s t stage. In i t s simplest form the leaching action may be described as a reaction between the sulphide minerals in the concentrate, dissolved oxygen, ammonia and water, in which copper, nickel and cobalt are converted to ammines, sulphur is converted to an oxidized form and iron is converted to hydrated ferric oxide: NiS:FeS + 3FeS + 702 + 10NH3 + 4H20 -»• Ni(NH 3) 6S0 4 + 2Fe 20 3'H 20 + 2(NH 3) 2S 20 3 [1.6] 2(NH 4) 2S 20 3 + 40 2 + 4NH3 + H20 -> NH4 »S03 »NH2 + 3(NH 4) 2S0 4 [1.7] According to laboratory and pilot-plant data, under optimum operating conditions, 95 percent Ni and 60 percent S is extracted at 70-80 °C and 5 an oxygen partial pressure of 0.68 atm to give a f i n a l solution of 45 g/1 Ni and 8 g/1 Cu **. This gives an average oxygen consumption of 0.18 g 02/l*min»atm , based on an average residence time of 19 hours. The Anaconda Arbiter process for copper (shown in Figure 2) makes use of the same ammonia chemistry as the Sherritt process. The goal of the Anaconda Arbiter process was to overcome the necessity of high oxygen partial pressures by applying mixing technology, rather than relying on high pressure reactors. The process also uses commercial oxygen rather than a i r . The process treats chalcopyrite concentrates accordingly: CuFeS2 + 4jo 2 + 4NH3 + 1 ^ 0 + C^NHg)^ + FeO(OH) + 2S0^ +2H+ [1.8] Again, under optimum conditions, 95 percent Cu and 27 percent S i s extracted from the concentrate at 60-90 °C and an oxygen partial pressure of 0.34 atm to give a f i n a l solution containing 43 g/1 Cu 5. *This gives an average rate of oxygen consumption for the Anaconda process equivalent to 0.39 g 0 2 /l*min*atm ,based on the residence time of 5 hours. Agitation for both processes is stated to have a pronounced effect on the leaching rates due to the heterogeneous nature of the reactions. However, almost a l l the data available in the literature on 6 these processes is from laboratory and pilot-plant studies* From the oxygen consumption rates in both of these processes i t i s d i f f i c u l t to determine i f mass transfer would be the limiting factor. As long as gas/liquid mass transfer is rate-limiting there is no difference in the average rate between a backmixed and a plug flow reactor. Furthermore, in a reactor that is mass transfer limited there w i l l be an equal rate of heat evolution in each compartment, therefore equal cooling capacity is required for each compartment. In contrast to both of the above processes is the Cominco zinc pressure leach process (shown in Figure 3) where zinc sulphide concentrates are converted directly to zinc sulphate solutions and elemental sulphur. The process uses oxygen rather than air to treat sphalerite concentrates as follows: ZnS + H„SO. + i o . + ZnSO. + S° + H.O [1.91 2 4 2 2 4 2 1 J FeS + H 2S0 4 + |o 2 -»• FeSO^ + S° + H20 [1.10] 2FeS04 + H 2S0 4 + -> F e 2 ( S 0 4 ) 3 + H20 [1-U] F e 2 ( S 0 4 ) 3 + ZnS -• 2FeS04 + ZnS04 + S° [1.12] Operating at 145-155 °C and an oxygen partial pressure of 7.5 atm (total 7 pressure = 11.4 atm) after 1.5 hours the extraction of zinc and the yield of elemental sulphur are 98% and 96%, respectively 6. These are much higher temperatures and pressures than the previous processes. Information available on the commercial operation shows that over 80 percent of the reaction takes place in the f i r s t compartment of the autoclave in 26 minutes 7. Since most of the reaction takes place in the f i r s t compartment, the conditions here are most important. With an i n i t i a l zinc concentration of 50 g/1 and a f i n a l concentration of approximately 120 g/1 in the f i r s t compartment the average oxygen consumption is 0.14 g 0 2 /l»min«atm . At high feed rates the oxidation of ferrous to fe r r i c iron is believed to be rate limiting 6. However, i f the f e r r o u s / f e r r i c couple i s rate l i m i t i n g there are two p o s s i b i l i t i e s ; the chemical rate constant for the reaction is rate controlling or the mass transfer of oxygen for the oxidation of the ferrous ion is rate controlling. There is no evidence in the literature to support one or the other. However, the reactor design would be affected differently for each case. Although the oxygen consumption of the zinc pressure leach appears to be considerably lower than the ammonia oxidative pressure leach processes i t is important to remember that the Anaconda data is only from the pilot plant. If commercial data were available i t would probably be found that the residence times are much higher than the zinc pressure leach process; this would make the average oxygen consumption rates comparable to the zinc pressure leach rates. If they are not 8 comparable to the zinc pressure leach rate then the descrepencies may be the r e s u l t of chemical d i f f e r e n c e s between the systems. It can be seen that for oxidative pressure leach processes leaching rates have been accelerated to the point where oxygen consumption rates of 0.75 mole 0 2 per l i t r e of leach solution are obtained within a residence time of the order of one hour. At these high oxygen consumption rates the mass transfer of dissolved oxygen at the gas/liquid or liquid/solid interface may become rate-determining. Clearly, an understanding of the mass transfer processes in an oxidative pressure leach process are important. 1.3 Mass Transfer in Pressure Oxygen Leaching Before considering the mass transfer processes in oxidative pressure leaching i t is useful to be reminded of the reaction steps that take place in the leach solution (see Figure A). The f i r s t step in any oxidative leach is the dissolution of gaseous oxygen .Into the leach * solution, requiring oxygen to be transported across the gas/liquid interface. Next, the dissolved reactants must diffuse to the solid interface of the particle being leached, involving transport across the liquid/solid interfacial film. The reactants are then adsorbed on to the solid surface which is then followed by the chemical leach reaction. After the reaction takes place any soluble products are desorbed from the interface which is then followed by diffusion of the products from the interface into the bulk solution. In an oxidative leach process any 9 s i n g l e s t e p o r c o m b i n a t i o n o f s i n g l e s t e p s c o u l d b e r a t e l i m i t i n g . I n m o s t s y s t e m s r e l a t e d t o l e a c h i n g t h e a d s o r p t i o n a n d d e s o r p t i o n s t e p s a t t h e p a r t i c l e s u r f a c e a r e f a s t c o m p a r e d t o e i t h e r t h e c h e m i c a l r e a c t i o n o r t h e m a s s t r a n s f e r a c r o s s p h a s e i n t e r f a c e s . T h e r e f o r e t h e m o s t s i g n i f i c a n t r e s i s t a n c e t o m a s s t r a n s p o r t o c c u r s a t t h e i n t e r f a c e s . F i r s t c o n s i d e r t h e g a s / l i q u i d m a s s t r a n s f e r i n p r e s s u r e l e a c h i n g . T h e b a s i c r e p r e s e n t a t i o n f o r t r a n s p o r t o f a s o l u t e g a s , w h e r e b y g a s d i s s o l v e s i n t h e l i q u i d w i t h o u t r e a c t i n g , i s b a s e d u p o n t h e c o n c e p t o f a d d i t i v i t y o f a g a s - p h a s e r e s i s t a n c e a n d a l i q u i d - p h a s e r e s i s t a n c e , a s s u m i n g t h e i n t e r f a c i a l r e s i s t a n c e c a n b e n e g l e c t e d . T h e r a t e o f a b s o r p t i o n i s t h e n , r = k G a ( P - P ± ) = k ^ C * - C A Q ) [1.13] w h e r e a i s t h e i n t e r f a c i a l a r e a b e t w e e n g a s a n d l i q u i d p e r u n i t v o l u m e * o f t h e s y s t e m , r i s t h e a v e r a g e r a t e o f t r a n s f e r o f g a s , p a n d p ^ a r e t h e p a r t i a l p r e s s u r e s o f t h e s o l u b l e g a s i n t h e b u l k g a s a n d a t t h e i n t e r f a c e , i s t h e c o n c e n t r a t i o n o f d i s s o l v e d g a s c o r r e s p o n d i n g t o e q u i l i b r i u m w i t h p ^ , a n d i s t h e a v e r a g e c o n c e n t r a t i o n o f d i s s o l v e d g a s i n t h e b u l k l i q u i d ( s e e F i g u r e 5 ) . 1 0 The c o e f f i c i e n t k^ i s the gas-side mass transfer coefficient, which refers to the gas film resistance, implying there is a stagnant film across which the soluble gas is transferred by molecular diffusion alone (while the bulk of the gas has a uniform composition). It has been demonstrated by Westerterp et a l 8 that gas-side transfer resistance for oxygen transport i s n e g l i g i b l e ( i e . k g » k^), even in a i r . This does not account for gas-side resistance due to oxygen partial pressures reduced below the oxygen partial pressures found in a i r , which can become significant. The c o e f f i c i e n t k^ i s the physical liquid-side mass transfer coefficient in the absence of a chemical reaction. There are two main types of models that represent the physical significance of Ic^: the film models and the surface-renewal models. The film models assume a stagnant film at the surface of the liquid next to the gas while the rest of the liquid away from the film boundary i s kept uniform in composition by turbulent agitation 9. The film models consequently predict that k^ i s proportional to D^ , the d i f f u s i v i t y of the dissolved *gas in the li q u i d . Although this i s a simple model i t does not agree 1 / 2 well with the experimental evidence that suggests k^ varies as (D^) The surface renewal models assume the periodic replacement of elements of liquid at the interface by liquid from the interior of bulk composition, where the rate of absorption is a function of the time of exposure of the element 1 0' 1 1. The surface renewal models r e a l i s t i c a l l y 1 / 2 predict that k varies as ( D ) . However, both lead to the same 1 1 predictions concerning the effect of the driving force ( C. - C. ) on A Ao average rate of gas transfer. Except for very slow chemical reactions, the basic effect of a chemical reaction on the mass transport i s to hold C^Q equal to zero in the bulk liquid; that i s , as soon as A is dissolved in the bulk liquid i t is consumed by a chemical reaction ( see Figure 6 ) . Therefore, the process is essentially one of physical absorption followed by reaction in the bulk liquid. It was found for the three commercial processes discussed earlier that there are significant differences in the oxygen mass transfer rates. These processes each employ similar agitator designs; therefore the differences in the oxygen mass transfer rates must exist in the details of the chemical and mass transfer processes. It i s not possible to discuss these details for any of the processes because the information in the literature is incomplete. However, i t is possible to examine how the oxygen mass transfer could be affected by different chemical conditions. The oxygen consuming reaction in a pressure oxidative leach must take place in one of three possible zones1*8. One possible zone i s on a solute monolayer at the gas/liquid interface which i s a reducing agent reactive to oxygen. Secondly, the oxygen could be dissolved in the bulk liquid and after dispersion by diffusion and convection react with a reducing agent dissolved in the solution or suspended as a solid such as the mineral being leached. The third zone i s intermediate, and oxygen is consumed somewhere in the boundary layer by a reducing agent dissolved in the solution. In the f i r s t and last cases i t i s essential that there be chemical reactions that decompose the mineral forming dissolved reducing agents without simultaneous consumption of oxygen. For example in the ammine processes cuprous ammine is such a reducing agent being formed by the reaction of the mineral with cupric ammines, while in acid systems i t is ferrous ion. The zone where the consumption of oxygen takes place depends on the concentration profile of the dissolved reducing agent across the boundary layer and the chemical reaction rate. Therefore, in different chemical systems there are likely to be differences in the concentration profiles, as well as in the chemical rates, even though both the systems may be gas/liquid mass transfer limited. This point i s illustrated by comparing some preliminary experiments using a cuprous ammine oxidation against the sodium sulphite oxidation experiments. It was found that the oxygen consumption rate in the ammine system is equal to 1.02 g 02/l*min»atm as compared to 0.27 g 02/l»min»atm in the sulphite system under similar agitation conditions. In both systems gas/liquid mass transfer is the rate-limiting step, but i n the ammine system, i t is possible that the cuprous complex is oxidized at the bubble interface, and the mass transfer rate of t h i s species would then be rate-determining. In the second case, i t i s probable that most of the reacting oxygen dissolves, passes through the boundary layer, and does not react with sulphite u n t i l a catalyst ion is available to mediate the reaction. In this case i t is the mass transfer rate of dissolved oxygen that is rate-limiting. In both cases, however, the gas/liquid interface area determines the overall rate, even though the same amount of area cannot lead to exactly the same oxygen consumption rate. The other mass transfer process that takes place in oxidative pressure leaching is liquid/solid mass transfer. As with gas/liquid mass transfer, the transport of a dissolved species in the liquid phase to the surface of a solid particle i s based upon the concept of transport resistances in series, assuming that Interfacial resistance can be neglected. The rate of transfer is then: D R = k T ( C - C ) = r ^ - ( C - C ) = r [1-14] L s L v s c l J where k T i s the liquid-side mass transfer coefficient; C, C and C are s c the concentrations of the transport species in the bulk solution, at the liquid/solid interface and at the receding chemical reaction front; r is the chemical reaction rate; Dg is the effective d i f f u s i v i t y through the product layer and L is the product layer thickness (see Figure 7 ) 3. The l i q u i d - s i d e mass transfer coefficient (k^) is treated in liquid/solid mass transfer the same way as i t is in gas/liquid mass transfer. In leach reactions where a product layers forms there may be significant resistance to mass transport across the layer. In porous product layers the reactants diffuse by means of molecular diffusion through cracks and channels in the product layer. 14 Since leaching systems are generally agitated the effect of f l u i d hydrodynamics on the mass transfer coefficients is an important consideration. In the case of solids, the agitation usually does not effect the particle size dramatically. Furthermore, once particles are suspended off the bottom the mass transfer rates increase l i t t l e as the particles approach uniform suspension ( see Figure 8 ) . Once the particles are in a uniform suspension, the particles are essentially in free f a l l most of the time. Consequently, the liquid-side mass transfer coefficients can be based on correlations using the terminal s l i p velocity of the given particle sizes. In cases where internal diffusion through a product layer is mass transfer controlling the liquid hydrodynamics have no effect. Generally, in gas/liquid mass transfer the liquid-side mass transfer coefficient i s the limiting factor in oxidative pressure 'leaching. In gas/liquid mass transfer the product of the mass transfer c o e f f i c i e n t and the interfacial area per unit volume of system (a) is controlled directly by the hydrodynamics. As in liquid/solid mass transfer, the liquid-side coefficient i s relatively insensitive to the degree of agitation. However, the interfacial area i s greatly affected by the hydrodynamic conditions. Therefore, the purpose of agitation in gas/liquid dispersions i s primarily one of increasing the surface area per unit volume. 15 Clearly, the design of agitators i s far more c r i t i c a l in gas/liquid systems than in liquid/solid processes. It must be understood that the process of dispersing a gas in liquid is very different from the process of mixing a liquid. The mixing of a liquid is meant to reduce nonuniformities or gradients i n compostion, properties, or temperature of material in bulk . The function of agitation in gas dispersion is to create gas/liquid interfacial area. This distinction is important, and w i l l be emphasized in the discussion on agitation theory. 1.4 Agitation Theory In general, agitation results in fluid motion. Fluid motion in any agitated system must obey the laws of conservation of mass and momentum. For a constant-density, Newtonian liq u i d , the mass and momemtum balance in terms of local pressure and velocity i s represented by the Navier-Stokes equation : « = ~g c P + v + g [1.15] where Dv/Dt is the substantial time derivative of velocity, i s the t density, g i s the acceleration due to gravity, g c is the gravitaional conversion factor, p is the pressure, is the viscosity, i s the vector differential operator, and is the Laplacian operator. It is not intended in this thesis to give an exhaustive treatment of the details of the Navier-Stokes equation, or for that matter, agitation. The primary purpose is to give an appreciation of the fundamentals as they are applied to the engineering of agitation systems. More specifically, the limitations of the concepts behind agitation theory w i l l be addressed. Other sources are available for a more detailed „. 13,14,15 treatise As can be seen from the Equation 1.15, agitation involves a large number of independant variables which include geometrical, operational, and physical variables. It i s common practise in the engineering disciplines to simplify the Navier-Stokes equation by means of dimensional a n a l y s i s 1 6 . The method of dimensional analysis rewrites the Navier-Stokes equation in a dimensionless form by selecting characteristic quantities to represent the principal dimensions of length, time and mass. The characteristic length used in agitation i s the impellor 'diameter, D. The characteristic time is the reciprocal of the agitator rotation speed, 1/N. The characteristic mass i s the product of liquid density, p and the cube of the impellor diameter, D 3. 1 7 Accordingly, dimensionless lengths and times are defined as: x' = x/D y' = y/D z' = z/D [1.16] [1.17] [1.18] t' = t/D [1.19] The dimensionless velocity, V , i s the ratio of the actual velocity, V , to the characteristic velocity, ND: V = V/ND [1'20] The dimensionless pressure i s defined as follows, where g £ i s the gravitational conversion factor, and p Q i s a reference pressure which is selected to simplify the boundary conditions: (P " P j g c pN D If these dimensionless variables are substituted into equation 1.15 and rearranged 1 7 a dimensionless form of the Navier-Stokes equation results, D V u g = -V'p' + ( ) V * V + ( ) [1.22] Dt' pD 2N DN2 In Equation 1.22, two dimensionless groups appear as parameters. The Reynolds Number for agitation, Re = pD2N/u , which represents the ration of i n e r t i a l to viscous forces, appears in reciprocal form as the coefficient for the viscous dissipation term. Also, the Froude number for agitation, Fr = DN2/g , which represents the ratio of i n e r t i a l forces to gravitational forces, appears. From Equation 1.22 i t follows 18 that for a given set of i n i t i a l and boundary conditions that require geometric similarity, the velocity and pressure distributions can be expressed as functions of the Reynolds and Froude numbers: V ' C x'.y'.z'.t') - f(Re,Fr) [1.23] p'(x',y',z\t') = f(Re,Fr) [1.24] It i s common practice to consider gravitational effects unimportant where the liquid surface is essentially f l a t , such as in f u l l y baffled tanks, therefore the velocity and pressure distributions are determined only by the Reynolds number. Impellor power consumption i s related to the pressure distribution along the face of the impellor blade. Power, P, is the product of rotational speed and applied torque, where the applied torque is determined by integrating the pressure distribution over the surface of a flat-blade turbine. Thus, power is related to flu i d pressure at the blade surface: f P" Po }blade [1.25] It follows from Equation 1.21 that dimensionless pressure and power are related as follows: 19 From Equation 1.24 the Power number, N^, is found to be only a function of the Reynolds number when gravitational effects are not a factor. Therefore, the equation usually takes the following form for correlating agitation power data: P g c — = f(Re) [1.27] pN3D5 In most applications of an agitator the in e r t i a l forces dominate the viscous forces; these are considered turbulent conditions and are associated with large Reynolds numbers (Re>101+). Under these conditions the terms in the Navier-Stokes equation which represent the viscous and gravitational forces can be neglected, and this yields an equation for flu i d motion known as Euler's equation 1 1 +: DV* = -V'p' [1.28] Dt' * In Equation 1.28, the Reynolds number i s not included as a parameter, therefore the velocity and pressure distributions are constant for this limiting case. As shown in Figure 9, the Power number is constant at high Reynolds number therefore we find the power draw of an impellor is proportional to the agitation speed and impellor diameter as follows, for a given geometric arrangement: 20 P « pN3D5 [1-29] If a constant tip speed, ND, is maintained then according to Equation 1.29 the power consumption w i l l vary with D2. An important point in this power correlation is that i t applies only to single phase systems, therefore power consumption in gas/liquid systems cannot be described by this equation. As important as understanding power in agitated systems i s understanding the behaviour of f l u i d velocity, since by definition agitation i s the fl u i d motion produced by impellor rotation. This point is frequently missed, consequently the volumetric power dissipation in a vessel is often,used to describe the degree of agitation. However, i f we examine the dimensionless f l u i d velocity in the turbulent region (see figure 9) i t is found to be fixed just as the Power number i s , therefore the same degree of fl u i d motion can be achieved with different power levels in the same vessel by varying the impellor diameter and agitator speed. Another important feature of an agitator is the impellor pumping capacity. The conclusions related to velocity behaviour can be applied directly to pumping capacity. The relationship between the average velocity across an area and the pumping capacity through that area i s stated as follows: [1.30] where Q Q - « - [1-31] A D 2 Therefore, i f Q/D2 is substituted into the dimensionless average velocity relationship the the following results: 1 7 v (Q/D 2i Q - ^ s . = = = f(Re) [1.32] ND ND ND3 Since the behaviour of the pumping number is similar to the behaviour of velocity, the pumping number i s found to be fixed in the turbulent region as shown in Figure 9. A detailed examination of the theoretical and empirical relationships between discharge velocities and flow rates, impellor geometry and rotational speed is given by Uhl and Gray in Volume I, Chapter 4. 1 3 As was the case for power consumption, i t is not fundamentally correct to apply the correlation for pumping capacity (Equation 1.32) to gas/liquid systems. These correlations are only correct for a constant density l i q u i d . A correlation has not yet been published in the literature that successfully predicts power consumption and impellor pumping characteristics in gas/liquid.systems. The present correlations 22 do not account for the effect of sparging or the impellor pumping gas into the liquid from the gas plenum. Dimensional analysis has been applied to gas dispersion only in limited applications. Remembering that the primary effect of agitation in gas dispersion is to create interfacial area, i t is found that under some conditions properties such as bubble size can be correlated by dimensionless parameters. In gas dispersions there are two regimes; one where the effect of gas sparging predominates and the other where gas pumped into the liquid from the gas plenum exerts control. Dispersions are brought about by f l u i d dynamical forces which have to overcome the static forces of surface tension. Surface tension forces resist dispersion by attempting to retain bubble sphericity and prevent gross distortion leading to breakup. These dynamic forces which induce a shear stress to produce a dispersion may be due to induced fl u i d flow or buoyancy. * In systems where gas sparging is present, and bubbles are formed by means other than surface aeration, dimensionless analysis of bubble break-up yields a dimensionless parameter called the Weber number, We. The Weber number is a ratio of i n e r t i a l forces to the surface tension t forces, which is formulated for agitator systems as follows: We = " • p fl-331 23 In most practical systems the impellor is li k e l y to pump gas into the liquid from above, therefore this type of correlation i s of limited value. The power consumption in gas sparged systems has been predicted by correlating the power consumption under gassed conditions to the power consumption under ungassed conditions by means of the impellor pumping number20: ^ = f ( ^ ) [1-34] where P and P are the power consumptions under ungassed and gassed 8 conditions, respectively. However, this method has been shown to not correlate a l l data taken over a wide range of conditions 2 1. A correlation soundly based on physical principles has not been published. In systems where the gas pumped into the liquid by the impellor from the gas plenum predominates, dimensional analysis has been shown to seriously underestimate the interfacial area at high rates of agitation. Calderbank 2 2 demonstrated the deficiency of such an analysis where surface aeration exists at the impellor. It i s seen that dimensional analysis does not account for the formation of bubbles due to surface aeration of the impellor. As • Calderbank 2 2 has shown the interfacial areas where surface aeration i s present are much higher than where bubbles are formed only by sparging. Therefore, dimensional analysis cannot be applied to the design of 2 4 practical gas/liquid mass transfer systems where surface aeration i s usually present. 1.5 Summary Examination of the development of pressure oxidative leaching processes indicates that mass transfer aspects are becoming increasingly important. In pressure oxidative leaching gas/liquid mass transfer is particularly important. In gas/liquid mass transfer i t has been demonstrated that the l i q u i d - s i d e mass transfer, k^a , i s rate c o n t r o l l i n g , i n general. Furthermore, the l i q u i d - s i d e mass transfer coefficient, k^, has been shown to be relatively insensitive to agitation; whereas, the effect of agitation on the interfacial area, a , i s quite substantial. Therefore, the primary purpose of agitation in a gas/liquid system i s to generate interfacial area. The interfacial areas capable of being produced by surface aeration are considerably larger than those 'produced solely by gas sparging i n an agitated system. However, cla s s i c a l dimensional analysis in the theory of agitation does not account for the gas pumped into the liquid by the impellor from the gas plenum, and for this reason i t cannot be applied to the design of industrial gas/liquid operations where obtaining large interfacial areas is the aim. This leads to the conclusion, that for the successful operation of pressure oxidative leaching processes a clearer understanding of the effect of agitation on gas pumping and the generation of interfacial area in gas/liquid dispersions is necessary. This thesis is addressed to this topic in gas/liquid mass transfer as i t applies to pressure oxidative leaching. 26 re 2. Gas-Liquid Mass Transfer in Pressure Oxidative Leaching  2.1 Agitation Theory of Gas Dispersions 2.1.1 C r i t i c a l Agitation Speed As Calderbank 2 2 and others**'z3,24,25,26 n a v e demonstrated» the is a c r i t i c a l agitation speed, NQ. Below NQ, the gas hold-up and the gas/liquid interfacial area are essentially dependent on the gas sparge rate, while above N Q the gas hold-up and the interfacial area depend much more on the agitation speed. This c r i t i c a l agitation speed corresponds to the hydrodynamic conditions where bubbles form at the impellor due to surface entrainment of gas, as opposed to cavitation effects. However, the c r i t i c a l agitation speed has only been correlated by empirical equations with no discussion of the underlying fundamental theories. 8 2 3 24 2 5 Westerterp and others ' ' have used the following * relationship, or a modified version, to correlate the c r i t i c a l agitation speed to different mechanically agitated vessels: NQD T - = A + B ( ) [2.1] (og/p) 1 / l t D where T i s the vessel diameter, p is the liquid density, a i s the 27 surface tension, and g is the acceleration of gravity. A and B are constants which are a function of the agitator type. Pandit et a l 2 6 Tiave correlated the c r i t i c a l agitation speed based on liquid circulation velocities to give the following: N 0 D T T 1/3 r2.21 = 0.865 ( )( ) ' 1 J (V G/e G) D 1.0 where T i s the vessel diameter in metres, i s the superficial gas velocity of the sparged gas, and is the fractional gas hold-up. Although these relaionships have been shown to correlate well under the conditions they were applied, t h e i r u n i v e r s a l i t y i s questionable. F i r s t l y , i t i s d i f f i c u l t to believe that there is no effect of impellor depth of immersion. Neither equation includes the effect of impellor depth of immersion; surely, for similar conditions, an impellor at a depth of 1 metre is not going to give the same c r i t i c a l •agitation speed as an impellor at a depth of 10 metres. Furthermore, for Equation 2.2 the effect of superficial gas velocity and gas hold-up on the c r i t i c a l agitation speed is d i f f i c u l t to accept. In the lim i t , as Vg •* 0 (no gas sparging) equation 2.2 implies that N Q -*• 0; this i s contrary to experimental evidence. A more appropriate approach to predicting the c r i t i c a l agitation speed has been postulated by Peters 2 7 which is based on fundamental 28 considerations of momentum and viscous effects in a gas/liquid system. The viscous effects are a function of the complex shear stresses around the impellor, these shear stresses are induced by flu i d dynamical forces which are influenced by the velocity distribution around an impellor. Neglecting the viscous effects, the c r i t i c a l agitation speed can be approximated by calculating i t on the basis of momentum effects only. This assumes that the kinetic energy needed to generate a bubble at the blade tip of an impellor immersed at a depth, h, in the liquid is equivalent to the potential energy a gas bubble positioned at a height, h, below the free liquid surface loses in rising to the liquid surface in the absence of drag: mgh = 1/2 mV2 [2.3] This is also equivalent to the potential energy of a drop of liquid positioned at a height, h, above the free liquid surface. Therefore, on the basis of momentum effects only, the c r i t i c a l tip velocity i s : « * 1/2 V = (2gh) i / Z [2.4] where h i s the depth of impellor immersion, g i s the acceleration due to grav i t y , and V i s the c r i t i c a l tip velocity. It i s important to note that there is no density effect since the mass terms in Equation 2.3 cancel each other. Therefore, as i t i s written this equation would be equally applicable in liquid mercury. On this basis we expect to observe an effect of depth of immersion. Momentum effects can be complicated by the existence of standing waves due to wall effects, therefore vessel geometry can possibly affect the c r i t i c a l tip velocity. Furthermore, the viscous effects are also expected to change with depth of impellor immersion due to changes in the velocity distribution around the impellor. The dispersion of a gas by the impellor pumping gas into the l i q quid from the gas plenum above the liquid surface is considered to be a two step process. The f i r s t step is the formation of a bubble at the tip of an impellor blade; the formation of a bubble i s governed by the c r i t i c a l tip velocity as discussed above. The second process i s the break-up of the bubble that has been formed. The breakup of a bubble i s brought about by flui d dynamical forces which have to overcome the static forces of surface tension which were discussed in section 1.4; the equilibrium state between these two forces determines the maximum bubble size that can exist. The concept of a two step process in gas dispersion is important because i t demonstrates that the prediction of 'bubble size, gas hold-up or interfacial area is meaningless unless the conditions for bubble formation have been satisfied f i r s t . 2.1.2 Gas/Liquid Interfacial Area It was established earlier that the primary function of agitation in gas/liquid systems i s to create interfacial area. Above the c r i t i c a l tip velocity, the gas/liquid interfacial area, a, increases 30 linearly with impellor tip velocity and is affected less by the gas 8 24 25 sparge rate ' ' . The l i n e a r relationship between interfacial area and impellor tip velocity i s a compound effect of gas hold-up and bubble diameter as demonstrated by Westerterp 2 8. The average bubble size i s seen to decrease rapidly near the c r i t i c a l tip velocity. This i s mainly due to the fact that near the c r i t i c a l tip velocity the bubbles are not yet homogeneously dispersed In the l i q u i d . Below the impellor the bubble density i s s t i l l much lower than above the impellor. The average bubble size rapidly reaches a constant value. Only at extremely high agitation rates, and above a fractional gas hold-up of approximately 0.4, does the bubble size decrease. According to Westerterp 2 8, the gas hold-up is shown to have a non-zero value below the c r i t i c a l tip velocity, and this is due to the presence of gas sparging in his experiments. Nevertheless, Westerterp shows how the fractional gas hold-up increases linearly with the * agitation rate. Finally the fractional gas hold-up reaches an upper limit of 0.4; thereafter, i t increases very slowly with increases in the agitation rate. The most dense spherical packing of gas bubbles gives a theoretical upper limit where the fractional gas hold-up is approximately 0.6. The gas hold-up ceases to be linear with respect to agitation rates only at t i p velocities much higher than typically encountered in industrial processes. In summary, increases in the interfacial area are mainly the result of increased fractional gas hold-up, while the average bubble size i s essentially constant over the range of agitation rates typically used in industry. It is interesting to recall that the impellor pumping capacity was also linear with respect to agitation rate acording to Equation 1.32. This suggests that a relationship may exist between the gas hold-up (or interfacial area) and an impeller's capacity to pump gas. In the experimental results section evidence w i l l be presented that confirms this relationship. 2.1.3 Impellor Power Consumption As discussed previously, the prediction of power consumption in gas dispersions must account for the reduced effective density due to presence of gas in the liquid. The power consumption of an impellor is expected to rapidly decrease above the c r i t i c a l tip velocity due to unloading at the impellor caused by bubble formation. The power consumption in gassed systems has been correlated to the dimensionless pumping number (Q/ND3), where Q i s the gas sparge rate. Correlation of the power consumption has also been related to the superficial gas velocity. However, these correlations only consider the effects of gas sparging, and cannot be applied where the impellor i s pumping gas into the liquid from the gas plenum above the liquid surface. 32 In surface aerated systems the effective liquid density is going to be related to the gas hold-up. Since the gas hold-up is related to the impellor tip velocity in surface aerated systems the effective density w i l l likewise be related. In particular, the effective density i s probably closely related to the gas pumping capacity of a given impellor type as follows: Q a ND^ [2.5] g where w i s the impellor blade width, and Q i s the impellor gas pumping capacity. Therefore, any correlation for predicting the power consumption of a surface aerated impellor must account for the effect of impellor unloading due to bubble formation at the impellor. 2.2 Previous Work on Gas/Liquid Mass Transfer in Agitated Systems The following is a brief review of the literature available on *gas/liquid mass transfer i n mechanically agitated vessels. Comprehensive reviews have been published by Sideman et a l 2 9 , V a lentin 3 0, Nagata 3 1, Charpentier 3 2 and Shah 3 3. The most systematic experimental work has been carried out by Westerterp et- a l 8 , Reith 3 1*, Mehta and Sharma23, Miller 2 1* and Pandit and J o s h i 2 6 . These authors confirm the existence of a c r i t i c a l tip velocity, however, only empirical correlations are reported. There is 33 no discussion of the underlying fundamentals related to the c r i t i c a l tip velocity in the literature. The gas/liquid interfacial area has been reported to be increased by increases in ionic strength, viscosity, by the presence of solids or Immiscible liquid, or by a decrease in liquid surface tension. There have been some exhaustive studies done to examine velocities, flow rates, and flow patterns for rotating impellors by 35 36 37 Nagata et a l . ' ' However, velocity measuring tubes were used and i t i s expected that the velocity profiles would be affected by the measuring devices. Nevertheless, they show some Interesting effects of impellor depth of immersion on the velocity profiles. A i b a 3 8 has shown that tangential velocities are greatly decreased by the insertion of baffles and both axial and radial velocities are increased. Charpentier 3 2 has given an excellent review of both physical and chemical methods for measuring gas/liquid interfacial areas and mass 'transfer coefficients. 2.3 Oxygen-Sodium Sulphite System for Measuring Interfacial Area and  Gas/Liquid Mass Transfer Rates The methods used to measure the specific interfacial area and mass transfer rates in gas/liquid contacting devices can be classified into two categories: local measurement with physical techniques such as 3 4 light scattering or reflection, photography, or electric conductivity methods; and secondly, global measurements with chemical techniques. Experimental work in this project has only concerned the global measurement of oxygen mass transfer rates. To be specific, the chemical reaction of oxygen with dissolved sodium sulphite has been used to measure the oxygen mass transfer rates. We w i l l briefly examine the reaction of oxygen with sodium sulphite. The absorption of oxygen in sulphite solutions i s very often used to measure mass transfer rates and interfacial areas as reviewed by Reith and Beek. 3 9 The chemical reaction i s : CoSO^ 2Na 2S0 3 + 0 2 — ---*• 2Na2S0lt [2.6] with the rate equation represented by: r k c C 0 2 CNa 2 S0 3 ^oSO^ [ 2 ' 7 ] ^The reaction i s considered a rapid pseudo-mth-order type which i s catalyzed by the presence of cobalt ions. When the reaction between 0 2 and Na 2 S0 3 in the liquid phase is mth-order in 0 2 and nth-order in Na 2 S0 3 , under conditions where the concentration of Na 2 S0 3 i s the same everywhere in the bulk solution (k C„ " i s constant) the reaction is 3 mn Na 2S0 3 ' said to be pseudo mth-order in 0 2> 3 5 The reaction is reported to be second order with respect to oxygen (m=2) for partial, pressures of oxygen below 1 atm and f i r s t order with respect to oxygen above 1 atm.4*0 Therefore, reactions carried out with air at atmospheric pressures are second order with respect to oxygen. The reaction rate i s independent of the sodium sulphite concentration in the concentration range 0.4 - 0.8 M, this represents zero order k i n e t i c s . 4 0 Below 0.4 M, the reaction i s reported to be f i r s t order with respect to sodium sulphite concentration. The influence of catalyst concentration (Co"1-1") gives a reaction order of unity for cobalt concentrations in the range of 1 X 10 - 4 to 2 X 10"3 M. 4 0 Any cobalt concentrations higher than these w i l l form a solid precipitate with the sulphite. The oxidation of sodium sulphite is inhibited by the presence of trace amounts of copper in s o l u t i o n . 4 1 The effect of ammonium sulphate ((NH^)280^) concentration on the rate of oxidation has not been determined in the literature. It i s posssible that cobalt ion w i l l complex with the ammonium species under certain conditions. If a constant concentration of ammonium sulphate i s used throughout the presence of ammonium sulphate should not be important. The most obvious effect of ammonium sulphate i s the decrease in bubble diameter related to the increased ionic strength of the solution. The average volumetric rate of absorption of oxygen per unit of inter f a c i a l area i s independent of the gas hold-up 3 2; however, the volumetric rate varies proportionally with the interfacial a r e a 4 0 . 3 6 Therefore, the reaction of oxygen with sodium sulphite is applicable for the measurement of gas/liquid mass transfer rates and the interfacial areas in the mechanically agitated systems being studied. A typical experimental rate curve from this study demonstrates zero order kinetics indicating the reaction is in the mass transfer controlled regime 3 (see Figure 10). Furthermore, under the appropriate conditions discussed above this reaction system can be used to measure the average interfacial area. 4 + 2 2.4 Purpose and Scope of the Present Investigation This thesis i s an examination of the gas/liquid mass transfer in pressure leach processes. The emphasis i s on the creation of gas/liquid interfacial area in mechanically agitated vessels. It w i l l be demonstrated that the mechanisms by which gas dispersion take place are different from the mechanisms by which f l u i d mixing is accomplished. Furthermore, the conditions that optimize gas dispersion in gas/liquid systems w i l l be discussed. The optimum conditions for mixing are not *the same as for gas dispersion; therefore, a clear distinction between mixing a fl u i d and dispersing a gas is necessary to achieve optimum process conditions in a gas/liquid system. Furthermore, i t w i l l be demonstrated that because of the above distinction dimensionless correlations are not useful for extrapolating or scaling up a gas/liquid mass transfer system. Instead i t w i l l be 3 7 shown that such gas/liquid systems are more appropriately described and scaled-up by means of the impellor tip velocity and gas pumping characteristics. This view of gas/liquid systems i s shown to follow from fundamental considerations of the momentum transfer and the viscous effects around an impellor. The effect of a number of variables on the oxygen mass transfer rates, and volumetric power requirements has been studied. These variables include the type of impellor, the impellor positioning, sparging modes, oxygen partial pressure, the presence of solids, and the use of f u l l length and half length baffle configurations. These variables were studied using both single and dual impellor arrangements. The experimental work was done using both bench scale equipment (a 2 l i t r e autoclave and a 2 0 l i t r e vessel) and a pilot scale reactor ( 2 1 0 0 l i t r e s ) . The pilot scale experiments were performed at Cominco's Technical Research centre in T r a i l , B.C.. A novel approach to gas dispersion i s demonstrated to have potential for a much reduced power consumption over the conventional systems presently employed. 38 3. Experimental Details o 3.1 Materials Technical grade anhydrous sodium sulphite (Na 2S0 3) was used for the oxidation tests. The 100 pound bags in which the sodium sulphite was provided did not provide any vapour barrier, therefore i t i s possible that atmospheric moisture was absorbed. As a catalyst in the oxidation tests, cobalt was dissolved in aqueous solution. This was added as reagent grade cobalt chloride (CoCl 2 '6H20). To simulate the mixing environment of an electrolyte ammonium sulphate with surface active agents removed was dissolved in the aqueous solution. Reagent grade ammonium sulphate was used for the bench scale experiments. The only distinguishable difference between this and commercial grade ammonium sulphate was the amount of froth produced. 3.2 Apparatus 3.2.1 Cominco Mixing Model Tests The oxidation tests performed at Cominco's Technical Research Centre were performed in a scaled-down model of the f i r s t compartment of 39 the commercial autoclave. The model was constructed of a 1.22 metres diameter cylindrical section of lucite with one bulged end which is constructed of fiber reinforced plastic (see Figure 11). Inside the tank were two baffles 0.18 metres wide which were located equidistant from the agitator shaft at either end of the tank. Two lengths of baffles were used interchangeably: one set was 48 inches long and went from top to bottom, the other set was 0.61 metres long and was placed in the lower half of the tank. It was possible to remove impellors from the agitator shaft and use various impellor combinations. Three types of impellors were tested: i) four-bladed 45° pitched blade (axial flow) impellor i i ) four-bladed 90° f l a t blade (radial flow) impellor i i i ) six-bladed 90° flat blade (radial flow) impellor The diameter of these impellors could be varied to be either 46 cm (18 * inches) or 53 cm (21 inches)(see Appendix A for dimensional details). The agitator shaft was driven by a constant speed 3-phase 11 kW motor. By changing sheaves on the motor shaft, the jack shaft or the agitator shaft i t was possible to study an extended range of agitation rates (50-300 rpm). These agitation rates were measured by a hand held tachometer. 40 The agitator drive motor was instrumented so that power measurements could be made during operation. A voltmeter, wattmeter and primary and secondary ammeters were provided. The voltage was 5:1 and the current stepdown was 2.73:1 (ratio of primary to secondary components), therefore the actual wattmeter reading must be multiplied by 13.65 to obtain the power draw. There is some question as to how accurately these power measurements represent the actual power input because with this size motor the efficiency i s low converting energy form el e c t r i c a l to mechanical. The power draw when in air is 0.982 kW. Gas could be sparged into the tank. The sparger with an o r i f i c e diameter of 15 mm was used to sparge under the bottom impellor (see Appendix B). Another sparger was designed to sparge gas just above the bottom impellor. This sparger has two or i f i c e s , each having a diameter of 15 mm (see Appendix B). Gas flow to the sparger was measured with a Fischer and Porter rotameter using B6-35-10 tube and BSVT-64 float. A second F and P rotameter FP-1-35-G10/27 was used to measure oxygen flow rates when oxygen enriched air was sparged. Both rotameters were fitt e d with a pressure gauge so pressure corrections could be made. It was possible to introduce air or oxygen into the vapour space of the mixing model through PVC tubing; flow through the tubing was measured using the rotameters used for the sparger (see Appendix C). Measurements of dissolved oxygen were made with a YSI model 51B dissolved oxygen meter. The sensing probe was inserted through the lucite port in the bulged end of the tank. 41 For the purposes of a special experimental arrangement to test the effect of a small diameter high speed impellor a 1/2 Hp variable speed (0-1200 rpm) motor was mounted above the opening of the mixing model. A 6-inch six-bladed disc impellor was mounted on the shaft and was provided a depth of immersion up to 5.5 inches. 3.2.2 Bench-Scale Experiments The bench-scale tests at U.B.C. were performed in a 2- l i t r e pressure autoclave and a 20-litre atmospheric vessel. The 2-li t r e pressure autoclave was constructed of titanium with a zirconium lining; the vessel diameter was 10 cm. The 20-litre vessel was constructed from thick-walled pyrex glass with a 28 cm diameter c y l i n d r i c a l cross-section. Inside each reactor four b a f f l e s were placed symmetrically around the walls. The baffle widths were equivalent to one-tenth the diameter of the vessel. Two lengths of baffle were used Interchangeably: one set of baffles went from top to bottom of the vessel, the other set of baffles went from the bottom of the vessel to * halfway up the vessel wall. Both reactors were fi t t e d so that i t was posssible to sample the solution during an experiment. the pressure autoclave was fitted with a 600 psi pressure gauge. As with the pilot-scale tests, i t was possible to remove the impellors from the agitator shaft and use various impellor combinations. The impellors used in the bench-scale experiments were the same three types used in the pilot-scale tests (see Appendix D). The diameter of 42 these impellors could be varied to be either 58 mm or 40 mm. The agitator shaft for the two-litre pressure autoclave was driven by a constant speed 1-phase 44 Watt motor. By changing pulleys i t was possible to use a range of agitation rates (150-1450 rpm). The agitator shaft for the 20-litre vessel was driven by a variable speed (0-1400 rpm) 75 Watt motor. The agitation speeds were measured by a hand-heId tachometer. It was not possible to make power draw measurements on the 2-litre autoclave. However, the motor mount for the 20-litre vessel was constructed with a mechanism for measuring the torque delivered by the motor. This was done by mounting the motor vertically along the shaft axis on a thrust bearing so that the motor casing rotates freely. The motor was connected by means of a radial arm to a spring balance (500g) which measured the force exerted by the rotating motor (see Appendix E for schematic). For the 20-litre vessel, a portion of the free liquid surface area was enclosed so that gas bubbles rising out of solution could be trapped, and the volumetric flow rate of bubbles could be measured (see Appendix F). The gas flow rate was measured using Gilmont gas flowmeters ( a 2000 ml/min capacity and a 8000 ml/min capacity). 43 3.3 Experimental, Procedure 3.3.1 Cominco Mixing Model Tests The rate of oxygen mass transf e r to the s o l u t i o n was measured by sampling the s o l u t i o n at i n t e r v a l s (usually every 3 minutes) and analyzing the concentration of Na 2S0 3. Once the rate of depletion of Na 2S0 3 was esta b l i s h e d , the rate of oxygen mass transf e r was determined by stoichiometry according to the reaction: 2Na 2S0 3 + 0 2 -• 2Na2SO^ [3.1] This a n a l y s i s for Na 2S0 3 was done by t i t r a t i n g with 0.1N iodine-iodide s o l u t i o n using an a c e t i c - s t a r c h i n d i c a t o r . A c a t a l y s t concentration of 5 ppm Co was used i n the ammonium sulphate s o l u t i o n . With only a few exceptions, the test work was performed i n a s o l u t i o n of ammonium sulphate (14kg/m 3). Following d i s s o l u t i o n of the ca t a l y s t and ammonium sulphate a measured amount (1-4 kg) of Na 2S0 3 was added. The Na 2S0 3 was dissolved by s t a r t i n g and stopping the agitator repeatedly, so as not to cause any gas dis p e r s i o n , u n t i l the dissolved oxygen meter f e l l to less than 0.3 ppm. I f sparging was used, the sparger was f i r s t s t a rted, then the ag i t a t o r was s t a r t e d . Solution samples were taken at i n t e r v a l s u n t i l the dissolved oxygen meter read 44 greater than 0.3 ppm. The volume of the mixing model tank i s 2116 l i t r e s at 100% f u l l . The testwork was performed at 82% f u l l which is a liquid volume of 1735 l i t r e s . For the oxygen enriched tests no additional air was admitted to the vapour space and the open top of the tank was covered. Gas samples were taken from the headspace and analyzed for percent oxygen using an Orsat analyzer. For the purpose of mixing the solution homogeneously during the small diameter high speed impellor tests a single 46cm four-bladed axial impellor was placed a half impellor diameter above the bottom of the tank and rotated at 104 rpm. _ At this agitation rate the single impellor at the tank bottom did not cause any disturbance of the free liquid surface. Throughout the experimental study unstable mixing conditions were prevented by maintaining agitation rates that did not cause loss of solution through the top of the mixing model due to splashing. 3.3.2 Bench-Scale Experiments For the bench-scale experiments a catalyst concentration of 0.1g/l CoCl 2»6H 20 was used in a l l experiments. A l l test work was 45 performed in 14kg/m3 ammonium sulphate solutions. For experiments in the 2-litre pressure autoclave 8-15g of Na2SG"3 was added at the start of the experiment. In the 20 l i t r e vessel the amount of Na 2S0 3 dissolved in solution was i n i t i a l l y 20-60 g. The solutions were stirred gently with a stirring rod unt i l a l l components were completely dissolved. A sample of the solution was taken and analyzed before the agitator was started. After the agitator was started samples were taken at intervals (usually 3-5 minutes) and analyzed immediately; this continued un t i l the Na 2S0 3 concentration f e l l below 0.05 kg During experiments with the 2 0 - l i t r e v e s s el, the power consumption of the impellor and the gas flowrate due to rising bubbles were recorded. Tests done in the 2-litre autoclave were performed with a liquid volume of 1.5 l i t r e s , and the work in the 20-litre vessel used a liquid volume of 16 l i t r e s . For the oxygen enriched tests done in the 2-litre pressure autoclave the autoclave was f i r s t purged with pure oxygen to remove any residual a i r . After the purge was complete, the autoclave was brought to the operating pressure then the agitator was started. 46 4. Results and Discussion 4.1 Cominco Mixing Model Tests The impellor configuration that is presently used in the commercial autoclave has a four-bladed 45° pitched blade (axial flow) impellor on the upper part of the shaft, and a four-bladed 90° flat blade (radial flow) impellor on the bottom (see figure 11). This impellor configuration is used as the standard impellor configuration for comparison purposes, and curves on any graph representing this configuration are labelled 'STD'. Although the model and the commercial autoclave are geometrically similar, the agitator speed in the autoclave that gives the same process results as the model w i l l not be the same speed. the scale-up formula used for vertical cylindrical tanks to give equal mass transfer i s : 4 3 , DU2/3 [4.1] n2 = n l ( D ~ ) where n^ and n 2 are the agitator rpm for the model and autoclave, respectively, and D^  and D 2 are similarly the impellor diameters. Although this scale-up correlation applies only to a single phase system i t i s used here only to give an idea of the order of magnitude that scale-up represents. Applying this equation to the horizontal cyl i n d r i c a l vessel, the test results for the model at 198 rpm should be approximately equivalent to the results for the commercial autoclave at 100 rpm. It is not correct to apply this equation to the bench scale experiments because the systems are not geometrically similar. Nevertheless, for the purpose of rough comparisons the test results for the model at 198 rpm would be equivalent to the results for the bench scale impellors at 785 rpm. 4.1.1 Single Impellor Systems 4.1.1.1 Effect of Impellor Type A series of oxidation tests were done to determine the effect of impellor type on the rate of oxygen mass transfer (g 0 2/l*min). The results are shown in Figure 12 for a range of impellor tip speeds. Clearly, a single six-bladed flat disc impellor i s more effective than the standard dual impellor configuration used in the commercial autoclave for promoting gas/liquid mass transfer. The six-bladed disc impellor was more effective than the other two types of impellors probably for two reasons: i t provides greater shear force than the axial impellor, and due to the disc i t forces more gas to flow through the high shear zone at the tip of the impellor blades. Similarly, the six-bladed flat disc impellor made more effective use of the energy i t consumed to produce oxygen mass transfer. A 48 measure of this effect i s the ratio of the oxygen mass transfer rate (g0 2/l*min) to the volumetric power consumption of the agitator (kW/m3), and is termed the relative mass transfer efficiency (g02/l*min)/(kW/m3). A comparison of the relative mass transfer efficiency of each impellor type at a constant impellor tip velocity i s given in Table 1. The values for the power draw (kW) have had the power draw of the impellor rotating i n a i r (P = 0.98 kW) subtracted from them. Although, in a 3- X X" non-gassed liquid, a six-bladed disc impellor consumes more power than the other two impellors, i t was found that in gas/liquid systems the six-bladed disc impellor consumed no more power than the four-bladed f l a t impellor. This lower power consumption is due to the extensive gassing of the impellor. In the unsparged cases, the top impellor in a dual impellor system is mainly doing the work to create gas/liquid mass transfer. The four-bladed 45° pitched blade impellor is used as the upper impellor in the standard autoclave configuration. It i s shown here that the four-bladed 45° pitched blade impellor is much less effective than the six-bladed disc impellor; therefore, i t i s expected that the six-bladed disc impellor alone would be more effective than the standard dual impellor configuration. 4.1.1.2 Effect of Depth of Impellor Immersion The depth of the impellor immersion below the free liquid surface had a different effect for each type of impellor. Each impellor 49 was studied at three depths - 22.9 cm (0.5 impellor diameter), 34.3 cm (0.75 diameter), and 45.7 cm (1.0 diameter). A cross-sectional view of the impellor positions in the mixing model is shown in Figure 13. The results for each impellor type are shown In Figure 14, Figure 15 and Figure 16. The oxygen mass transfer rates produced by the six-bladed f l a t disc impellor were found to be the most sensitive to the depth of impellor immersion. The four-bladed flat blade impellor appeared to be sensitive only to placements of the impellor below a depth of 34 cm. While the four-bladed 45° pitched blade impellor was affected significantly only by impellor placements above a depth of 34 cm. If the concept of a c r i t i c a l tip velocity, as explained in section 2.1.1, is representative of a surface aerated impellor then an effect of depth of immersion is expected to be observed. The difference between the three types of impellors is likely due to the difference in the velocity profiles around each type of impellor (this represents a viscous effect). The sensitivity of the six-bladed disc impellor to the depth of immersion is l i k e l y due to the partitioning effect the disc has on f l u i d flow (there is no flow axially through the impellor). Thus the six-bladed disc impellor w i l l have a very different velocity profile from the other two impellors. The relative mass transfer efficiency for each impellor displays trends similar to the oxygen mass transfer rates. Table 2 shows the effect of impellor immersion depth on the impellor power draw (kW/m3). Only the power draw for the six-bladed disc impellor was greatly 50 affected by the depth of impellor immersion; the power increased rapidly with depth, likely due to reductions in the gas pumping capacity of the impellor. * As expected from Equation 2.4, the c r i t i c a l tip velocities, V , for the six-bladed impellor were found to be proportional to the square 1/2 root of the immersion depth, h , with a correlation coefficient of 0.997 (see Table 3). The effect of depth of immersion did not correlate 1/2 as well to h for the other two impellors; this i s l i k e l y due to viscous effects. 4.1.1.3 Effect of Impellor Diameter For each of the impellor types a comparison of the oxygen mass transfer rate was made for two impellor diameters - 46cm and 53 cm. The impellors were placed at the same depth of 23 cm for each experiment. The results for each impellor type over a range of agitation rates i s shown in Figure 17, Figure 18 and Figure 19. For each of the impellor types an increase in impellor diameter increases the oxygen mass transfer rate once the agitation rate is above the c r i t i c a l impellor tip velocity. The c r i t i c a l tip veloctiy corresponds to the point where the extrapolated linear portion of the curve intersects the x-axis. From the graphs i t appears that the c r i t i c a l t i p v e l o c i t y i s not significantly affected by a change in the impellor diameter. Again, considering Equation 2.4 a change in the impellor diameter should have no effect on the c r i t i c a l tip velocity as shown by the graphs. 51 Although, i n general, the oxygen mass transf e r rate increased with an increase i n the impellor diameter, the e f f e c t was only marginal fo r the six-bladed disc impellor. The increased rates are thought to be due to the increased circumferential area swept out by the t i p and the increased c i r c u l a r c r o s s - s e c t i o n a l area, which both combine to give increased gas pumping rates through the impellor. The reason the six-bladed disc impellor was only marginally a f f e c t e d by the increase i n diameter i s l i k e l y due to the presence of the d i s c which r e s u l t s i n only a small increase i n the c i r c u m f e r e n t i a l area swept out by the t i p (there i s e s s e n t i a l l y only r a d i a l f l u i d flow); there i s e f f e c t i v e l y no increase i n the cross-sectional area. Table 4 summarizes how impellor diameter affected the impellor power consumption and the r e l a t i v e mass t r a n s f e r e f f i c i e n c y . In a l l cases, the power consumption increased with the diameter of the impellor. Only i n the case of the four-bladed r a d i a l impellor did the r e l a t i v e mass transfer e f f i c i e n c y increase with a larger diameter impellor; a l l others showed a decrease i n the r e l a t i v e mass transfer e f f i c i e n c y . The r e l a t i v e mass transfer e f f i c i e n c y of the four-bladed r a d i a l impellor increased as the diameter increased probably due to the shear forces increasing more than power consumption increased. An increase i n shear forces means more surface area i s produced f o r r e a c t i o n . 52 A.1.1.4 Effect of Baffle Length Conventionally, the design of baffles in a chemical reactor has the length of the baffle run from the bottom of the tank to just above the liquid level of the solution. A series of experiments were done to examine the effect of using baffles which are shorter in length. The baffles used were 58cm long (half the mixing model diameter) and 19cm wide, and were placed at the bottom of the tank. The results of experiments using the four-bladed axial impellor show the effect of short length baffles in Figure 20. The use of short length baffles resulted in a more uniform increase of 0.0014 kg02/m3*min in the oxygen mass transfer rate. The use of short length baffles resulted in a more pronounced vortex. Improvements in the impellor power draw and the relative mass transfer efficiency were also observed when the short length baffles were used instead of the f u l l length baffles as shown in Table 5. It appears that as the agitation rate increases, the power saved using short baffles instead of long baffles becomes more substantial. However, the improvement in the relative mass transfer efficiency i s larger at the lower agitation rates than i t is at the higher agitation rates, when short baffles are used instead of long baffles. The reduced power draw when short baffles are employed is probably due to lower viscous dissipation of energy in the solution. This would also explain the more pronounced vortex. Furthermore, half length baffles become less beneficial at higher agitation rates probably because at higher 53 agitation rates the viscous dissipation of impellor power is so high that removal of baffles in the top portion of the tank has only a marginal effect. A similar diminishing improvement is observed when the effect of short length baffles is considered at different depths of impellor immersion as shown in Table 6. The benefits of using short baffles instead of long baffles at a shallow depth are about a factor of 3-3.5 greater in each case than the benefits at a deep depth of impellor immersion. 4.1.2 Dual Impellor Systems 4.1.2.1 The Standard Dual Impellor Configuration The standard dual Impellor arrangement (a four-bladed 45° pitched blade-upper, and a four-bladed 90° flat blade-lower) that is used in the commercial autoclave was studied over an extended agitation range (125-290 rpm). The results for both unsparged and sparged (0.28 normal m3/min or 10 normal ft 3/min) are shown in Figure 21. For both the sparged and the unsparged conditions there appears to be two linear regions. In the unsparged case, the slopes of the lines on the two region curve can be explained by considering the effect of each impellor separately. The slope of the line in the f i r s t region is approximately 54 equal to the slope expected for a single 4-bladed 45° pitched blade impellor. The c r i t i c a l tip velocity of the f i r s t region i s much lower than that measured for a single four-bladed 45° pitched blade impellor, and i s li k e l y due to the assistance provided by the additional momentum of the bottom impellor. The discontinuity in the curve has been attributed to effect of the lower impellor becoming dramatically important. The slope of the second linear region is approximately equal to the slope for a four-bladed 45° pitched blade impellor (at 0.5 impellor diameter immersion) plus the slope for a four-bladed flat-blade impellor (at 1.0 diameters immersion). If the second linear region i s extrapolated to the c r i t i c a l tip velocity i t i s found to be 4.5 m/s which i s approximately equal to 3.85 m/s for a four-bladed radial blade impellor immersed at a depth of 46 cm (see Figure 15). The same trends are observed for the sparged case, however, because oxygen gas is already present in the solution without any agitation due to the sparge gas, the f i r s t linear region intersects the y-axis at a slight positive oxygen mass transfer rate. 4.1.2.2. Effect of Baffle Length on Standard Configuration The results from the single impellor studies suggested that oxygen mass transfer rates could be improved by using short baffles instead of long baffles; therefore, i t was reasonable to assume that that same effect would be observed in a dual impellor system. The results of tests done using the standard dual impellor configuration are 55 shown in Figure 22. As i t was in the single impellor experiments, there was an increase in the oxygen mass transfer rate when short baffles were used instead of long baffles. However, this was true only up to an impellor tip speed of approximately 4.5 m/s. Beyond this agitation rate the oxygen mass transfer rate increases only marginally, and appears to reach a plateau value of approximately 0.011 kg0 2/m 3min. This agitation rate corresponds approximately to the c r i t i c a l tip velocity of the second linear region found in the unsparged experiments. This suggests there is some negative effect of the lower impellor, however an adequate explanation for this phenomena cannot be provided. 4.1.2.3 Alternate Dual Impellor System - Unsparged Considering the results of experiments done using single impellor systems a series of impellor configurations were tested which were believed to give higher oxygen mass transfer rates than the standard impellor configuration. The six-bladed disc impellor had been shown to be the most effective impellor, therefore i t was the f i r s t choice for the upper impellor. Unfortunately, there was only a single six-bladed disc impellor available, so only a four-bladed radial impellor or a four-bladed axial impellor could be selected for the lower impellor. For the case where no sparging was used, a comparison of different impellor arrangements against the standard configuration is shown in Figure 23. The combination of six-bladed disc impellor-upper 56 and four-bladed radial impellor-lower was found to be nearly four times more e f f e c t i v e than the standard impellor configuration. The four-bladed radial impellor as the lower impellor produced an oxygen mass transfer rate that is 27 percent higher than the four-bladed axial impellor as the lower impellor. This was expected based on the single impellor studies. Because the top impellor i s the one mainly responsible for the gas dispersion in the unsparged case there was expected to be only a small difference between the 21 inch diameter and the 18 inch diameter upper six-bladed disc impellor. This was found to be true in the dual impellor system. This follows from the single impellor studies of the oxygen mass transfer rates. A.1.2.4 Alternate Dual Impellor Systems - Sparged The same series of experiments that were performed above were run with a sparger under the lower impellor. The effect of sparging on the oxygen mass transfer rate for this series of impellor configurations is shown in Figure 24. The use of sparging as compared to no sparging showed improvements ranging from 60 to 100 percent higher oxygen mass transfer rates. Again, the impellor combination with the six-bladed disc impellor as the upper impellor, and the four-bladed f l a t blade Impellor as the lower impellor proved to have the highest oxygen mass transfer rates. It i s clear from the results that the sparged gas dispersed by the bottom impellor i s very important; the oxygen mass 57 transfer rates were almost doubled. The surprising result in this comparison is that the impellor arrangement with the upper 53 cm six-bladed disc impellor had a lower oxygen mass transfer rate than the 46 cm impellor. However, because the comparison is based on the impellor t i p velocity of the upper impellor only, and because the bottom impellor has a very significant effect, the lower oxygen mass transfer rate in the 53 cm arrangement is the result of the lower impellor in the 53 cm arrangement having a lower tip velocity than the lower impellor in the 46 cm arrangement. 4.1.3 Oxygen Concentration Effects 4.1.3.1 Oxygen Depletion in Reacted Gas Bubbles A set of experiments were done to examine the effect of oxygen concentration in the headspace on the oxygen mass transfer rate when only surface incorporation of the gas is used (i.e. no sparging). The results showed that the oxygen mass transfer rate increases linearly with oxygen concentration in the headspace, as shown in Figure 25. This linear relationship agrees with the mass transfer equation discussed in section 1.3 for conditions where C^o=0: r • v < c 2 - c A O > l 1 ' 1 3 ! The dissolved oxygen concentration in the bulk liquid is considered to 58 be zero in the oxidation tests, and in the pressure leach. Furthermore, an experiment was designed so that the gas bubbles rising from the solution could be trapped and analyzed for their oxygen concentration. This was achieved by collecting rising bubbles in inverted beakers before the gas could mix with the fresh air in the headspace. The respective oxygen concentrations in the headspace and the reacted gas bubbles are given in Table 7. The results show a significant depletion of oxygen in the bubbles. The depletion of oxygen can present serious problems for oxygen mass transfer i f the gas bubbles experience long residence times in the solution before they rise to the headspace to be replenished with oxygen. Although only air was used in the experimental work there appeared to be no effect of oxygen depletion over the course of an experiment; a l l experimental rate curves displayed zero order kinetics (i.e. linear with time). 4.1.3.2 Oxygen Enrichment for the Dual Impellor Systems The oxygen mass transfer rates were measured for the six-bladed disc-upper/ four-bladed radial-lower arrangement using a pure oxygen sparge under the bottom impellor and a pure oxygen purge in the headspace. The improved oxygen mass transfer rates are compared in Tables 8 and 9 against air sparge for the same impellor arrangement. 59 4.1.4 Small Diameter-High Speed Impellor Experiments A series of experiments were designed to demonstrate a novel approach to gas dispersion, and to illustrate that the dispersion of gas (production of interfacial areas) and the mixing of fluids each require different conditions to be optimally effective. In general, for aqueous systems very l i t t l e agitation i s necessary to provide a well-mixed, homogeneous solution; this can be achieved with low agitation rates and low impellor power consumption. For gas dispersion we have seen thus far that impellor type, impellor tip velocity and depth of immersion are important operating variables; moreover, there does not seem to be any correlation between the impellor power consumption and the oxygen mass transfer rate. With these assumptions considered, a 15 cm diameter six-bladed disc impellor was selected to serve the gas dispersion function which was run at agitation rates between 500-1000 rpm so as to provide impellor tip speeds similar to the standard configuration. The small impellor was placed at a depth of 14 cm in the solution. To provide solution mixing a 46 cm four-blade axial impellor was placed a half impellor diameter above the bottom of the tank, and agitated at 104 rpm (at this rate there is no disruption of the free liquid surface). A comparison of the experimental results against the standard impellor configuration is given in Table 10. The results indicate that with the small impellor at 1000 rpm the oxygen mass transfer rates are 39 percent higher than the standard impellor configuration. Furthermore, although the actual power draw of 60 the small impellor could not be measured, the maximum i t could be i s equivalent to the motor's power capacity, which is rated at 0.375 kW (0.5 Hp). The total volumetric power of two impellors, under this assumed maximum, is 0.359 kW/m3. This compares to a power draw for the standard impellor configuration, at an agitation rate of 169 rpm, equal to 0.832 kW/m3. The emphasis in this experimental series was to demonstrate the distinction between the mechanisms of mixing and the mechanisms of gas dispersion. Clearly, the oxygen mass transfer was provided by the small impellor and the mixing of the solution done by the large impellor at the bottom. Therefore, oxygen mass transfer can be achieved with much lower power consumption. According to theoretical calculations done using power curves, in an ungassed aqueous system the small impellor would consume 1.90 kW (2.55 Hp), and assuming that the efficiency of the agitator drive train is 90 percent a motor of at least 2.11 kW (2.83 Hp) would be required to operate the Impellor at 1000 rpm (see Appendix E for details). Nevertheless, the small impellor was able to operate with only a 0.37 kW (0.5 Hp) motor. Therefore, the power required to operate the gassed impellor is less than 20 percent of the power required to operate the ungassed impellor. However, i t would be necessary to provide a 2.11 kW (2.83 Hp) motor for start up conditions. 4.1.5 Special Sparging Mode Experiments Two experiments were done to examine the difference between 61 sparging above an axial flow impellor and sparging below an axial flow impellor (the axial flow of solution was downward). The results of experiments using a six-bladed disc impellor-upper and a four-bladed axial impellor-lower are shown in Table 11. There was only a marginal improvement in the oxygen mass transfer rate when sparging is done from above as opposed to sparging from below. There is no apparent difference in power consumption. 4.1.6 Comparison of Results with Theoretical Predictions It was proposed in section 2.1.1 that the c r i t i c a l tip velocity should be related to momentum effects, and according to the relationship developed to describe these effects (Equation 2.4) the c r i t i c a l tip v e l o c i t y , V , should be directly proportional to the square root of the 1/2 depth of impellor immersion, h . For the six-bladed disc impellor, when a linear regression analysis was performed on a straight line plot * 1/2 of V versus h a good correlation coefficient of 0.997 was obtained. This experimental evidence gives excellent support to the idea that Equation 2.4: * 1/2 V = (2gh) i /^ [2.4] is a fundamentally correct concept. The other two impellor types did not correlate as well to the above equation. This i s believed to be due to other effects such as non-radial momentum vectors through the impellor that are not accounted for by this obviously simple equation. 62 The above equation represents only the f i r s t step in what is proposed to be a two step process for the dispersion of a gas in a l i q u i d . The f i r s t step i s the formation of the bubbles, which i s governed by the c r i t i c a l t i p v e l o c i t y , V , should not be affected by changes in the impellor diameter according to Equation 2.4. For each of the impellor types the experimental curves showed essentially no change in the c r i t i c a l tip velocity as the impellor diameter was changed. The second step after the formation of the bubbles i s establishment of a stable bubble diameter which is determined by the hydrodynamics at the impellor. The bubble diameter is rapidly established and then only the gas hold-up is significant in determining the gas/liquid interfacial area. It was suggested that the gas hold-up is directly related to gas pumping capacity of the impellor. This is supported by the increase in oxygen mass transfer rates which followed increases in the impellor diameter; the gas pumping capacity of an impellor is expected to increase as impellor diameter becomes larger. The results obtained for the dual impellor systems were also well explained by the theoretical concepts put forward. Examination of the dual impellor system revealed a compound effect of the two impellors which resulted in a discontinuity in the rate of oxygen mass transfer as the impellor tip velocity increased. The discontinuity in the curve is explained in terms of the individual contribution of each impellor in the dual arrangement. The slope and c r i t i c a l tip velocity for each linear section of the curve was found to relate well to the additive 63 effect of the predictions for each single impellor. It has been stressed that f l u i d mixing and gas dispersion are optimized under different process conditions. Classical mixing practice uses baffles to improve f l u i d mixing i n a reactor, especially in cylindrical reactors. By their nature, baffles change the direction and decrease velocities in fl u i d flow which results in energy losses. In gas dispersions, however, high shear stresses in the region where bubbles are formed are desirable to minimize the size of bubbles in solution. High shear stresses are enhanced by high velocities. Baffles provide high shear only in the immediate vacinity of the baffles; therefore, the removal of baffles from the top half of the mixing model resulted in higher oxygen mass tansfer rates and lower impellor power consumption. Normally, the removal of baffles would have a detrimental effect on fl u i d mixing, but in the case of gas dispersion the optimum conditions are different, consequently the removal of baffles is found to Improve the dispersion of gas. The small diameter-high speed impellor experiments similarly discriminated between flu i d mixing and the dispersion of gas. These experiments were found to further confirm the concept of a two step process for gas dispersion. According to classical mixing theory the power draw of an impellor can be greatly reduced by decreasing the diameter of the impellor (Power <* D 5),or at constant tip velocity, ND (Power tt D 2). If oxygen mass transfer rates can be maintained while lowering the power consumption, there is potential for many benefits. 64 Bubble formation is only dependent on the c r i t i c a l tip velocity, which is mainly a function of the impellor immersion depth. If a smaller diameter Is desired then a higher agitation rate is required to maintain tip velocity; power is not as strong a function of agitation rate as i t is of impellor diameter (Power <* N 3). Furthermore, i f the impellor is immersed at a shallower depth the c r i t i c a l tip velocity reduces, and the required increase in agitation rate w i l l not be so large. The small diameter-high speed impellor experiments demonstrated well these implications of the two step process for gas dispesion. This arrangement was able to improve the oxygen mass transfer rates beyond the rates obtainable with standard dual impellor configuration, while the small diameter-high speed impellor system only consumed approximately half the power of the standard dual Impellor configuration. 4.2 Bench-Scale Experiments 4.2.1 Single Impellor Systems 4.2.1.1 Effect of Impellor Type A series of oxidation tests were done in a 20-litre cylindrical vessel to further examine the effect of impellor type on the rate of oxygen mass transfer. The results are shown in Figure 26 for a range of impellor tip speeds. In these small-scale experiments the difference between each impellor type is not as pronounced as i t was in the 65 experiments using the mixing model (2100 l i t r e s ) . However, the six-bladed disc impellor is s t i l l found to be more effective than the other two impellor types. The diminished difference between the impellor types i s l i k e l y due to two reasons: geometric similarity between mixing model and the 20-litre vessel has not been maintained, and secondly, the scale is a hundred times smaller than the mixing vessel. The former i s probably the most c r i t i c a l criterion. 4.2.1.2 Impellor Gas Pumping Rates For each impellor type a series of experiments was done to measure the gas pumping rate as a function of agitation rate. The gas pumping rate is based on a measurement of gas flow from the annular region amounting to 80 percent of the total surface area of the liquid. The measurements cannot be considered an absolute measure of the impellor gas pumping rate, therefore the volumetric flow measurements have been termed the relative gas pumping rate. The results for each type of impellor are shown in Figure 27a, Figure 28a and Figure 29a. The results clearly show that the relative impellor gas pumping rate i s a linear function of the agitation rate. As discussed earlier, the oxygen mass transfer rate i s mainly dependent on the gas hold-up, the gas hold-up has in turn been shown to be a linear function of the agitation rate. It i s believed that the gas 66 hold-up i s directly related to the gas pumping rate of the impellor, therefore the gas pumping rate i s also expected to have a linear dependence on the agitation rate. Furthermore, not only i s the impellor gas pumping rate linear with respect to the agitation rate as i s the oxygen mass transfer rate, but the gas pumping rate has a c r i t i c a l tip velocity that closely coincides with the c r i t i c a l tip velocity for the oxygen mass transfer rate. This supports the idea that the gas hold-up in a surface aerated vessel is directly related to the gas pumping rate of the impellor. 4.2.1.3 Impellor Power Consumption It was of interest to examine the power consumption of each impellor type as a function of the agitation rate. The power consumption for each impellor type was measured under two conditions using a torque measuring device: f i r s t l y , the power was measured under conditions where no gas was entrained in the solution; secondly, the power consumption for each impellor was measured under conditions where surface aeration takes place and the Impellor becomes gassed. The results for each impellor type are shown in Figure 27b, Figure 28b and Figure 29b. The ungassed curve for each impellor type displays the class i c a l relationship between power and agitation rate, where the power increases with respect to the cube of the agitation rate, N 3. However, the gassed curve has a much reduced power consumption beginning at the agitation rate that corresponds to the c r i t i c a l tip velocity. This is the point where the impellor begins to pump gas. After the gassing begins, the power curve flattens out i n i t i a l l y , then eventually increases monotonically with the same shape as the ungassed curve, but at a much reduced power consumption. The reduced power consumption is at least partially due to the much lower effective solution density and viscosity as a result of the impellor pumping gas into the solution. If Figures 27, 28 and 29 are each examined there appears to be a correlation between the oxygen mass transfer rate, the gas pumping rate and the gassed power curve. Both the gas pumping rate and the oxygen mass transfer rate are linear functions of agitation rate with approximately coincident c r i t i c a l tip velocities. Furthermore, the gassed power curve deviates from the ungassed power curve at an agitation rate which is approximately equal to the c r i t i c a l tip velocity. 4.2.1.5 Effect of Impellor Diameter The effect of changes in the impellor diameter on the oxygen mass transfer rate were done using a 40 mm and a 58 mm six-bladed radial disc impellor. The results are shown in Table 12. As i t was found in the mixing model results, the oxygen mass tranfer rates increase with an increase in impellor diameter. The increase in oxygen mass transfer rate correlated well to the increase in gas pumping rate that resulted from the increase in impellor diameter. Both the oxygen mass transfer rate and the impellor gas pumping rates double when the impellor diameter was increased from 40 mm to 58 mm. The impellor power consumption increased as expected when the impellor diameter was 68 increased. However, the power consumption increased by only 77 percent. Whereas, i f the classical relationship between power and Impellor diameter is governing (P <*• D^), then the power consumption for the 58 mm impellor should have been 2.1 times higher than the power consumption for the 40 mm impellor. This much lower increase in the impellor power consumption is li k e l y due to the increased gas pumping that results from a larger impellor diameter. The increased gas pumping rate reduces the effective solution density and in part counteracts the effect of increased power consumption that normally results from larger impellors. 4.2.1.6 Effect of Baffle Length The effect of baffle length was examined using a 58 mm diameter six-bladed radial disc impellor. The results of these experiments are shown in Table 13. The oxygen mass transfer rates were found to be lower when half length baffles are used instead of f u l l top-to-bottom length baffles. This is contrary to the effect short length baffles had in the mixing model experiments. The difference is most li k e l y due to the difference in the geometry of the two vessels, and possibly the vessel scale has an effect. The impellor gas pumping rate was also lower when half length baffles are used instead of f u l l length baffles. Furthermore, the impellor power consumption was found to be lower when half length baffles are used. This result agrees with the results obtained in the mixing model experiments. 6 9 A.2.2 Dual Impellor Systems 4.2.2.1 The Standard Dual Impellor Configuration The standard dual impellor configuration (a six-bladed 45° pitched blade-upper and a six-bladed 90° f l a t blade-lower) was examined to measure the effect of agitation on the impellor power consumption and the impellor gas pumping rate. The results are shown in Figure 30. The oxygen mass transfer rate and the impellor gas pumping rate are both linear functions of the agitation rate. The slope of oxygen mass transfer rate curve is the same as the slope for the single six-bladed axial impellor; this is expected since the upper blade of the standard dual configuration is also a six-bladed axial impellor. It should be noted that the standard dual impellor configuration in the mixing model experiments used only four-bladed impellors, whereas the standard configuration in the bench-scale experiments used six-bladed impellors. Although, there was found to be l i t t l e difference between the mass transfer rates for a four-bladed and a six-bladed radial disc impellor in the mixing model experiment, i t appears to be important for radial impellors and axial impellors. For example, the slope of oxygen mass transfer rate curve for the six-bladed bench-scale impellor i s 0.0082 g 02*s/m»1»min, while the slope the four_bladed mixing model impellors is 0.0042 g 02»s/m*l«min. As with the single impellor experiments, the power consumption for the dual impellor configuration i s lower when the impellors are 7 0 gassed due to surface aeration. However, the reduction in power consumption is not as large as that found for single impellor studies. The reason for this is because the lower impellor in the configuration is s t i l l significantly ungassed by surface aeration at the impellor tip speeds. Furthermore, there was no discontinuity observed in the bench-scale experiments because the agitation rates were not sufficiently high. 4.2.2.2 Alternate Dual Impellor Configurations The most effective dual Impellor configuration was compared against the standard dual impellor configuration. The results are shown in Table 14. The six-bladed radial disc-upper/six-bladed radial-lower configuration had an oxygen mass transfer rate slightly higher than the standard dual configuration. This difference in oxygen mass transfer rates i s not as large as was found in the case of the mixing model experiments. Furthermore, there was found to be only a small difference in the values of the relative gas pumping rate and the impellor power consumption. These differences between the results from the bench-scale experiments and the results from the mixing model experiments are attributed to changes in geometry and scale. 4.2.3 Effect of Oxygen Concentration The effect of oxygen concentration was studied over a more extended range of oxygen partial pressures than was possible with the mixing model apparatus. The range of oxygen partial pressures examined was 0.21 atma (3.1 psia) to 6.8 atma (85.3 psia). The results are plotted in Figure 31. The rate of oxygen mass transfer is found to vary linearly with respect to the oxygen partial up to 6.8 atma as would be expected according to Equation 1.13, where the dissolved oxygen concentration is essentially zero, C^ q=0, in the bulk liquid. 4.2.4 Effect of Solids Experiments were done to examine the effect on gas mass transfer rates when solids are present in the gas/liquid system. For these experiments s i l i c a particles (240 mesh) were used in a solids concentration of 46 kg/m3. The results of the experiment are shown in Table 15. The oxygen mass transfer rate and the impellor gas pumping rate are lower when solids are present, however the effect i s marginal. The solids also had only a small effect on the impellor power consumption. 4.2.5 Comparison of Results with Theoretical Predictions The gas hold-up has been reported to be a linear function of the agitation rate as discussed earlier in section 2.1.2. The gas pumping rate has been predicted to be directly related to the gas hold-up. i f the gas hold-up and the impellor gas pumping rate are directly related to each other then the impellor pumping rate should be a linear function of the agitation rate, just as the gas hold-up i s . The experimental results obtained from measurements of the impellor gas pumping rates firmly support these predictions. The impellor gas pumping rates were found to vary linearly with the agitation, as did the oxygen mass transfer rates. According to the classi c a l dimensional analysis of power consumption for an impellor the power draw of an impellor should increase proportionally with the cube of the agitation (P « N 3) for the case of ungassed impellors. The measured power curves were found to correlate well to this relationship for ungassed impellors. However, the gassed impellors were found to have a lower power consumption rate than the ungassed impellors. For impellors which were affected by surface aeration the power curve was coincidental with the curve for an ungassed impellor until the agitation rate was approximately equivalent to the impellor's c r i t i c a l tip velocity. Below the c r i t i c a l agitation rate no gas Is incorporated into the solution due to surface aeration, so the impellors are expected to behave similarly. At the c r i t i c a l agitation rate the impellor becomes gassed, and the two power curves diverge with the gassed impellor curve having a lower power consumption. This was found to be true for both dual and single impellor systems. The classical analysis of power curves relates the power consumption to the Reynolds number; these power functions do not account for the effect of gassing at the impellor and are inadequate for describing a gas dispersion system which is affected by surface aeration. 73 Examination of the effect of impellor diameter showed that as the impellor diameter is increased the oxygen mass transfer rate increases. The oxygen mass transfer rate is a function of the gas hold-up, and the gas hold-up i s a function of the impellor gas pumping rate. Based on this analysis the oxygen mass transfer rate could be expected to increase with an increase in the impellor diameter because an Increase in the impellor diameter w i l l result in a higher gas pumping rate. The classical power curves state that the power consumption of an impellor varies according to the f i f t h power of the impellor diameter (P * D 5). This was not found to be true in the case where an impellor experiences surface aeration and becomes gassed. The power consumption increased at a much lower rate. This i s due to the increase in gas pumping rate with an increase in impellor diameter, therefore the effective solution density i s reduced, thus the power consumption i s reduced. 4.3 Summary 4.3.1 Mechanisms of Gas/Liquid Dispersions It has been demonstrated that the dispersion of a gas into a liquid solution is well represented by a two step process. The f i r s t step in the process is the formation of gas bubbles at the tip of the impellor blades. The second step i s the establishment of the bubble size. Once the bubble size has been set the gas/liquid interfacial area 74 i s determined almost exclusively by the gas hold-up. The formation of bubbles depends on the c r i t i c a l tip velocity, * V , which i s a function of both momentum and viscous effects. Based only on the momentum effects, the c r i t i c a l tip velocity i s shown to vary with the square root of the impellor depth of immersion according to the fundamental equation: * 1/2 V = (2gh) i / Z [2.4] This i s a simplified representation which serves only to provide a fundamental viewpoint of a complex situation. A more comprehensive model would be expanded from this equation to include both momentum and viscous effects. Momentum effects are altered by the presence of fl u i d flow deflections from walls and baffles; these could add to or subtract from momentum vectors arising from the agitator blade, leading to geometrical distortions of the c r i t i c a l velocity criterion by the system. Viscous effects are related to the distribution of solids and gases in a solution and physical properties of the f l u i d . A l l these aspects must be incorported to improve the accuracy of the c r i t i c a l tip velocity model. The size of bubbles in a gas dispersion i s determined by an equilibrium between shear stresses in the solution and the surface tension forces present. The bubble size i s relatively insensitive to the agitation rate, except at high agitation rates which are not typical 75 in industrial applications. The gas hold-up is directly related to the interfacial area which ultimately determines the gas mass transfer rate. The gas hold-up has been shown to be linearly related to the gas pumping rate of an impellor, as long as the agitation rate i s above the c r i t i c a l tip velocity. The power curves that describe the power consumption of a surface aerated impellor are vastly different from the classi c a l dimensionless power curves. The power consumption of an impellor gassed by surface aeration is as much as 75 percent lower than the power consumption for an ungassed impellor. The power consumption i s a complex function of not only impellor diameter and agitation rate, but also a function of the impellor type, impellor positioning, f l u i d properties and vessel geometry. The dispersion of gas using dual impellor systems was demonstrated to be the compound effect of the individual impellors making up the configuration. Therefore, based on single impellor studies i t is possible to predict the composite effect of two impellors placed in a dual configuration. 4.3.2 Scale-Up Considerations The problem of scale-up has not been specifically addressed in t h i s study, n e v e r t h e l e s s some co n c l u s i o n s can be made. 76 The scale-up of gas/liquid dispersions for equal interfacial area, which corresponds in most cases to equal gas mass transfer rates, cannot be based on the c r i t e r i a of Reynolds number (or similar dimensionless correlations) or on the basis of power inputs per unit volume of liquid. When geometric similarity i s maintained in both the impellors and the vessel shape, the impellor t i p velocity remains likely as an independent parameter of scale-up c r i t e r i a . In applications to oxygen pressure leaching, i t seems compelling that agitator size, type, depth of immersion, and baffle configuration need to be selected on the basis of gas pumping and c r i t i c a l tip velocity c r i t e r i a rather than mixing theory, i f i t has been determined that gas/liquid mass transfer is a substantial resistance in the rate of the autoclave process. 77 5. Conclusions, Applications and Recommendations  5.1 Conclusions This study has been an examination of gas/liquid mass transfer in mechanically agitated oxidative pressure leaching systems with particular emphasis on the effect of agitation on gas dispersion in a liquid phase. The effect of a number of variables on the oxygen mass transfer rate and the impellor power consumption has been examined. These variables include the type of impellor, the impellor positioning, oxygen concentration, sparging modes, the presence of solids in solution and the use of f u l l length over half length baffle configurations. The following conclusions can be made: 1) It has been shown that the dispersion of a gas in a liquid can be described by a two step process. 2) The mechanisms by which a gas is dispersed in a liquid are different from the mechanisms by which f l u i d mixing is accomplished. Consequently, the optimum conditions for each process are not the same. 3) The c r i t i c a l tip velocity, which i s the agitation rate corresponding to the point when bubbles are f i r s t formed at the impellor, can be modelled by a theoretical equation based 7 8 on the fundamental momentum effects, Vmin " ( 2 « h I 2'**] The experimental results support the basic nature of this model, however i t i s necessary to further develop this model to include both viscous and momentum effects. 4) Classical dimensionless correlations are not adequate for extrapolating gas/liquid mass transfer systems where surface aeration of the impellor is present. 5) The scale-up of a gas/liquid mass transfer system cannot be done on the basis of dimensionless correlations nor on the basis of power inputs per-unit volume of liquid. 6 ) It is more appropriate to describe and scale-up by means of the impellor tip velocity and gas pumping characteristics. Some of the more practical implications of these conclusions are as follows: 1) The most effective type of impellor for the dispersion of gas was found to be the flat-bladed radial disc impellor, followed by the flat-bladed radial impellor and la s t l y the 45°-pitched flat-bladed axial impellor. At equal impellor tip speeds, increased impellor diameter was found to increase the 79 oxygen mass transfer rates. 2) Gas sparging can significantly enhance the oxygen mass transfer in dual impellor systems. 3) Under certain conditions the use of half length baffles was found to improve the gas mass transfer rates and reduce the power consumption. 4) The oxygen mass transfer rate was found to be a linear function of the oxygen partial pressure over the range studied ( 0.21 - 6.8 atma). 5) The presence of solids in a gas dispersion were found to marginally lower the oxygen mass transfer rates. 5.2 Applications The application of these concepts could enhance the operation of a variety of hydrometallurgical processes where gas/liquid mass transfer i s rate limiting. In processes where the rate limiting process i s undefined, but where gas/liquid mass transfer is potentially important, the use of these concepts can provide a guide to process changes that w i l l eliminate any potential gas/liquid mass transfer problems In an effective and economical way. The zinc pressure leach process i s an example of a process where 80 the rate limiting process i s undefined. The process is believed to be limited in the f i r s t compartment by either the gas/liquid mass transfer of oxygen to the solution, or by the reactions associated with the ferrous-ferric (Fe^/Fe 1 1 1 ) couple. There i s no evidence in the literature at present to indicate that one or the other process is rate limiting. Although the principles presented here for improving the dispersion of gas should have no effect on the ferrous/ferric couple, the rate of oxygen mass transfer to the solution could be improved by these methods until i t is clear that gas/liquid i s not rate-determining in the zinc pressure leach process. The results obtained in the mixing model indicate that i t is possible to increase the oxygen transfer rates much more than current industrially experienced rates are in the autoclave. Another oxidative pressure leaching process where gas/liquid mass transfer is likely to play a v i t a l role i s the Equity Silver Mines leach process for removing antimony and arsenic from high-silver and gold bearing copper concentrates. The leach plant, which began operating in 1981 at Houston, B.C., makes use of sodium hydroxide and sodium sulphide (fed as sodium hydrosulphide) to leach soluble antimony and arsenic from the concentrate. The leach liquors are then processed to sequentially crystallize antimony, arsenic and excess sodium sulphate. 4 5 It is in the antimony precipitation step that gas/liquid mass transfer i s likely to play a v i t a l role. In the antimony autoclaves the oxidation of S~~, SbS 3 and AsS 3 ions takes place i n so l u t i o n 4 9 : 81 Na2S + 20 2 •*• Na2S01+ Na 3AsS 3 + 0 2 + 3H20 + Na3As01+ + 3H2S01+ Na 3SbS 3 + 13/2 0 2 + 4NaOH + H20 -• NaSb(OH)6 + 3Na2S01+ [5.3] [5.1] [5.2] At the operating temperature of 150 UC, the oxidation of these species is very rapid and the reactions are probably oxygen mass transfer limited in view of the large quantities of oxygen that are consumed. Therefore, improved dispersion of the oxidizing gas in the solution would likely enhance the reaction rates. The accelerated leaching of gold and silver concentrates using microbiological systems has been successfully demonstrated in the laboratory, along with processes for the biological leaching of copper, zinc and uranium.1*1* In microbiological leaching, oxygen and carbon dioxide are essential constituents which must be dissolved in the leach solution by means of gas/liquid contacting. In one laboratory study on gold and silver recovery using biological leaching 1* 6 i t was reported * that considerably higher oxidation rates were obtained in a stirred tank as compared to a pachuca tank. The improved oxidation rate in the stirred tank is l i k e l y due to the increased oxygen mass transfer rate. Clearly, gas/liquid mass transfer, and more specifically gas dispersion, is l i k e l y to be important in the commercial use of biological leaching. wherever i t is necessary to disperse a gas in a liquid in Order to successfully operate a process the principles discussed in this paper 82 can be applied to aid in the design of the gas/liquid contacting equipment. 5.3 Recommendations for Further Work This study has emphasized the Importance of understanding the mechanisms of gas dispersion. In this regard the study cannot be considered complete, but serves to demonstrate that there is s t i l l much research yet to be done in this area. There is s t i l l no unified theory of mixing which incorporates a l l the efforts related to f l u i d mixing and gas dispersion. As an extension of this study i t i s recommended that the following be undertaken: 1) The effect of f l u i d properties should be examined to determine their effect on gas mass transfer and power consumption in gas dispersion. Of particular importance are f l u i d viscosity, f l u i d density, surface tension, ionic strength of the solution, and the effect of surface active agents. 2) The positioning of impellors in solution relative to the liquid surface should be examined more extensively. In dual impellor configurations the relative position of one impellor to the other should be investigated. 3) In dual impellor systems, the methods of sparging gas into 8 3 the solution should be studied to determine whether sparging is beneficial only at the lower impellor. 4) Although velocity profiles of impellors have been studied in the past, measurements have been limited to pitot-tube and hot-wire anemometer instruments. By measuring with these devices the velocity profiles are changed due to the physical presence of the instrument i n the system. A more accurate means of studying velocity profiles would make use of laser-Doppler velocimetry. Velocity measurements have been made in both single phase and two phase gas/liquid systems 4 7. A careful study of how the impellor diameter, depth of impellor immersion, agitation rate, and baffle length affect the velocity distribution would give a greater understanding of viscous effects. 5 ) The effect of vessel geometry should be given some attention, this is often dismissed as unimportant but there i s no evidence to support this conclusion in the case of gas/liquid systems. Finally, in a more general way, the effect of agitation on other mass transfer processes should be investigated to give a more complete picture of the mass transfer in a hydrometallurgical process, since either gas/liquid or liquid/solid transfer can be rate limiting: 8 4 1) The effect of agitaion on l i q u i d / s o l i d mass transfer should be studied. A possible system for investigating this would be the oxidation of s o l i d cuprous chloride (CuCl) to cupric chloride (CuCl 2) which i s soluble in aequeous systems. 85 REFERENCES 1. Hiskey, J.E. ; Wadsworth, M.E.; "Electrochemical Processes in the Leacing of Metal Sulphides and Oxides", Process and Fundamental  Considerations of Selected Hydrometallurgical Systems, AIME (1981); M.C. Kuhn, Ed., p303 2. McKay, D.R.; Halpern, J.; Trans AIME, vol 212; 301 (1958) 3. Levenspiel, 0.; Chemical Reaction Engineering, 2ed ; John Wiley & Sons, N.Y. (1972) 4. Forward, F.A.; Mackiw, V.N.; Trans AIME, vol 203, 457-463 (1955) 5. Kuhn, M.C.; Arbiter, N.; Kling, H.; CIM Bulletin, (Feb 1974) p62-74 6. Parker, E.G.; CIM Bulletin, (may 1981) pl45-150 7. Parker, E.G.; McKay, D.R.; Salmon-De-Friedberg, H.; Proc. of 3rd  Int. Symp. on Hydrometallurgy, AIME; Atlanta, Georgia (March 1983) p927-940 8. Westerterp, K.R.; Van Dierendouck, L.L.; de Kraa, J.A.; Chem Eng Sci (1963) vol 18 P157-196 . 9. Whitman, W.G.; Chem. Metall. Eng., 29, 147 (1983) 10. Higbie, R.; Trans. Am. Inst. Chem. Eng., 35, 365 (1935) 11. Danckwerts, P.V.; Ind. Eng. Chem.,43,1460 (1951) 12. Oldshue, J.Y.; Chem. Eng. (Junel3, 1984) p82-108 13. Uhl, V.W.; Gray, J.B.; Mixing; Theory and Practice, vol I and II, (1966), Academic Press, N.Y. * 14. Bird, R.B.; Stewart, W.E.; Lightfoot, E.N.; Tranport Phenomena, Wiley, N.Y. (1960) 15. Nagata, S.; Mixing, Principles and Applications, Halsted Press, Wiley,N.Y. (1975) 16. Langhaar, H.W.; Dimensional Analysis and Theory of Models, John Wiley & Sons, N.Y. (1951) 17. Dickey, D.S.; Fenic, J.C.; Chem Eng (January 5, 1976) pl39-145 18. Hinze, J.O.; AIChE J., 5, 289 (1955) 19. Dickey, D.S.; Hicks, R.W.; Chem Eng (February 2, 1976) p93-100 86 20. Calderbank, P.H.; Trans Inst Chem Engrs (London), 36 443 (1958) 21. Michel, B.J.; Miller, S.A.; AIChE J., 8,262 (1962) 22. Calderbank, P.H.; Trans Inst Chem Engrs (London), 37,175 (1959) 23. Mehta, V.D.; Sharm, M.M.; Chem Eng Sci, 26, 461 (1971) 24. Miller, D.N.; AIChE J., 20, 445 (1974) 25. Boerma, H.; Lankester, J.H.; Chem Eng Sci, 23, 799 (1968) 26. Pandit, A.B.; Joshi, J.B.; Chem Eng Sci, 38, 1189 (1983) 27. Peters, E.; Private Communication (1982) 28. Westerterp, K.R.; Chem Eng Sci, 18, 495 (1963) 29. Sideman, S.; Horatacsu, 0.; Fulton, J.W.; Ind Eng Chem, 58, 32(1966) 30. Valentin, F.H.H.; "Absorption in Gas-Liquid Dispersions", Spon, London (1967) 31. Nagata, S.; "Mixing", Wiley, N.Y. (1975) 32. Charpentier, J-C; "Mass Transfer Rates in Gas-Liquid Absorbers and  Reactors", Advances in Chemical Eng., vol II, Academic Press, Inc (London), (1981) 33. Shah, Y.T.; "Gas-Liquid-Solid Reactor Design", McGraw-Hill,N.Y. (1979) 34. Reith, T.; Ph.D. Thesis, Delft University (1968) 35. Nagata, S.; Yamamoto, K.; Hashimoto, K.; Naruse, Y.; Mem. Fac. Eng.  Kyoto Univ., 20, 336 (1958) * 36. Ibid; 21,260 (1959) 37. Ibid; 22,68 (1960) 38. Aiba, S.; AIChE J., 4, 435 (1958) 39. Reith, T.; Beek, W.J.; Chem Eng Sci, 28, 1331 (1973) t 40. Laurent, A.; Doctoral Thesis, University of Nancy, France (1975) 41. Wesselingh, J.A.; Van't Hoog, A.C; Trans. Inst. Chem. Engrs., 48, T69 (1970) 42. DeWaal, K.J.A.; Okeson, J.C.; Chem Eng Sci, 21, 559 (1966) 87 43. Rautzen, R.R. et a l . ; Chem Eng (October 25, 1976) p21 44. Bruynesteyn, A.; Working Paper on Microbiological Leaching as a  Means for Metal Recovery, B.C. Research, Vancouver, B.C. (1983) 45. Dayton, S.; Eng. and Mining J. (January, 1982) p78-83 46. Lawrence, R.W.; Bruynesteyn, A.; CIM Bulletin (September, 1983) pl07-110 47. Mahallngham, R.; Llmaye, R.S.; Brink, Jr., J.A.; AIChE J. (1976) volume 22, No. 6 48. Gollakota, S.V.; Guin, J.A.; Ind Eng Chem Process Pes Dev (1984) volume 23, p52-59 49. Jones, D.; Private Communication (1984) 8 8 Nickel Concentrate First Stage Leach Water 1 Gases NH 3 Scrubber Nitrogen Air-NH-Water Gases- Solution Solids Second Stage Leach Weak Liquor NH3(Aaua) NH3(Aqua) Solids Counter Current Wash Wash Water H2S-Solids Final Tailings Steam Copper Sulphide Copper-Free Solution To Nickel-Cobalt Precipitation And Ammonium Sulphate Recovery F i g u r e 1. S h e r r i t t - G o r d o n Ammonia Oxygen P r e s s u r e P r o c e s s f o r Ni-Cu-Co Ammonia Recycle Moke Up Ammonia Oxygen Concentrate Feed From Smelter Cooling Wote^ Scrubber Water I Ammonia Ammonia Ment Recovery Froctionator Lime Slurry Lime Boil Sysjem Condensate Return Steam Process, Water r Boiler Feed Water Steam Plant Raffinate Gypsum Residue Recycle I Ammonia Vent Leach Reactors CCD Thickeners Wash Water j : Dilution and Launder Spooy Water Leach Reside Flotation — C Z I Moke -UP H . s a Pregnant Liquor Fil " .Filter Itration r Cake al CO 1X> Solvent Extraction miuen I ? t Flotation Tails •Return to Smelter Make Up Water Electrowinning L Z = Cathode. Copper F i g u r e 2 . A n a c o n d a A r b i t e r P r o c e s s Fortified Spent Electrolyte Zinc Concentrate LtPh Flash Steam i i Oxygen Conditioning Tank [Molten S° .Sffjotatjon . j ~\ ! Cone i _ l Zinc Sulphate Slurry (To Zinc Recovery) p"p"ty"*-fiWfl- ~* Clean Sulpher Steam Unreacted Sulphides (To Roasters) o F i g u r e 3 . ° 2 Gas Phase 0o(ln Solution) F e 2 ° 3 X H 2 ° F i g u r e 4 . R e a c t i o n S t e p s i n a T y p i c a l L e a c h Cos Film 1 6 Liquid Film L ! 1 Liquid Layer of 1 Infinite Depth • _ . i Liquid Layers of Infinite Oeplhs ! Bulk of 1 Gos ' •I [ ! 1 p« CA 1 ! V 1 Bulk of j \ j Liquid | \« So i 1 1 t, » Film Model in the Liquid Higbie Model in the Liquid Donckwerts Model in the Liquid no F i g u r e 5 . L i q u i d - S i d e M a s s T r a n s f e r M o d e l s 92 Figure 6 . Liquid-Phase Concentration Profile for Mass Transfer with a Chemical Reaction 94 Liquid Film Diffusion Rote=klAC Liquid Solid Interface — Receding Reaction Front Reaction Rate = kC„ Unreacted / Solid F i g u r e 7 . L i q u i d - S o l i d I n t e r f a c e 5 6 9 6 it w Q. E 3 \ \ \ \ Viscous Transition , Turbulent 1 1 1 1 Reynolds Number (a) Power number Q Z £ 3 Z o> c "a. E 3 a . Reynolds Number D (b) Dimensionless velocity I Viscous I Transition I Turbulent Reynolds (c) Pumping number Number D F i g u r e 9 . R e y n o l d s N u m b e r C o r r e l a t e s D i m e n s i o n l e s s P a r a m e t e r s 97 10 15 Time (min) 25 Figure 10. Typical Experimental Rate Curve for the Oxidation of Sodium Sulphite 9 8 9 9 Figure 1 2 . E f f e c t of Impellor Type on the Oxygen Mass Transfer Rates 100 Figure 13. Impellor Positioning in Mixing Model 101 Figure 14. Ef f e c t of Impellor Immersion Depth on Oxygen Mass Transfer Rate for the 4-Bladed A x i a l Impellor 1 0 2 1 1 P„ '21 kPo I Impellor Tip Speed (m/s) Figure 15. Effect of Impellor Immersion Depth on Oxygen Mass Transfer Rate for the 4-Bladed Radial Impellor 1 0 3 1 r l4ko/rn3(NH4)2S04 Impellor Tip Speed (m/s) Figure 16. Eff e c t of Impellor Immersion Depth on Oxygen Mass Transfer Rate for 6-Bladed Radial Disc Impellor 1 0 4 Figure 17. Effect of Impellor Diameter on Oxygen Transfer Rate for the 4-Bladed A x i a l Impellor 1 0 5 Impellor Tip Speed (m/s) Figure 18. Effect of Impellor Diameter on Oxygen Transfer Rate for the 4-Bladed Radial Impellor 106 Impellor Tip Speed (m/s) Figure 19. Eff e c t of Impellor Diameter on Oxygen Transfer Rate for the 6-Bladed Radial Disc Impellor 107 8 30 x 3| cf a) 2-0 o or <*-lO c o 10 to .£ 10 c a) x O 14 kg/m* (NH^S04 P. «2l:'kPa.t °2 1 22 9 cm 4B-A-I8 STD 2-5 50 Impellor Tip Speed (m/s) 7-5 Figure 20. Effect of Ba f f l e Length on Oxygen Transfer Rate i n the Mixing Model 1 0 8 Impellor Tip Speed (m/s) Figure 21. Standard Dual Impellor Configuration Used In the Commercial Autoclave 109 l4kg/m8(Nr^S04 21 kPo 1 2 2 9 c m 4B-A-I8 4B-R-I8 2 2 - 9 c m Effect of B a f f l e Length on Oxygen Transfer Rate for the Standard Dual Impellor Configuration 1 1 0 O O 2 0 -CM O CP 0) B w c o CO CO o c cn • X O 10 Sparge Rate =0 Normal mVmin Upper Impellor Tip Velocity = 405 m/s l4kg/m3(NH4)2S04 per wer Q. O 3 cm cm to CO IT) (J to Q 15 •5 o or •o — o CD 5 3 a > a . ? a . o 3 _ l S § CO CO u u> Q * 5 "D -O cc "O <D TJ _ _ o -aci co a . 5 C L o 3 —I 6 E D o co u> o to o _ •6 .9 o x or < • a TJ = o CD co "S a TO c o CO Figure 23. Alternate Dual Impellor Configurations-Unsparged I l l Sparge Rate = 0-28 Normal mvmin Upper Impellor Tip Velocity = 405m/s 14 kg/m3 (NH4)2 S04 O o X ' £ 0 ~ 3 3t O a 2 c o o c C P X O 0 aj w Q. a. o 3 E £ u u ro Q0 i n u ST '6 "5 T5 -o or a> "O o GQ 10 w. a> a. Q. o 3 £ u cm CD co u tn Q "5 o or •a T3 O CQ CO SI S3 S i tn g _ '•o o o x or < " O a> o -CD CO ^J" O T3 C O (?) Figure 24. Alternate Dual Impellor Configurations-Sparged 112 c 1 0) o or c o 0 0 8 O06 ^ 0O4 o c > » X o 0 0 2 1 1 1 — Tip Speed =4-75 m/s I4kg/m? (NH4)2S04 i • / • 1 1 -2 2-9 cm 4B-AH8 6B-RD-I8 t \ V - i — i 2 2-9 cm / t / — s y 1 i i 20 40 60 80 Oxygen Concentration (%Volume) 100 Figure 25. E f f e c t of Gas Plenum Oxygen Concentration on Oxygen Mass Transfer Rate 1 1 3 Impellor Tip Velocity (m/£) F igure 26. E f f e c t of Impel lor Type on the Oxygen Trans fe r Rate i n 2 0 - l i t r e s Vesse l 1 1 4 2 5 •g £ o or s a. m 8 12 E or 68058 14 kg/hf (N^)2S04 Full Baffles 15 e .9 a. I .0 % o a . I 5 •8 x 1 "E HO * « o or in e .e 2 >> O 1 2 3 4 Impellor Tip Velocity (m/s) 50mm _ a 1 A Ungassed / l4KgV(NVVzS04 / Full Baffles / • A A / / - 0 - 0 -Gassed • < 0 1 2 3 4 Impellor Tip Velocity (m/s) Figure 27. A) Effect of Agitation Rate on Impellor Gas Pumping Capacity and Oxygen Mass Transfer; B) Effect of Surface Aeration of Impellor on Power Consumption (6-Bladed Radial Disc Impellor - 58mm diameter) o 5 X 1 £ o az 6 3 3 0. | 2 & e $ 11 01 Wkg/hT (N^) 2 S0 4 Full Baffles 8 x £ 1-0 * o a : c o 0-5 c 0) o> >> K O 1 2 3 4 Impellor Tip Velocity (m/s) 15 c o 3 O u tt * I 5 tt CL E I4kg/ri3 (NH4)2S04 Full Baffles 6BR58 I Ungassed A w 0 I 2 3 4 Impellor Tip Velocity (m/s) Figure 28. A) Effect of Agitation Rate on Impellor Gas Pumping Capacity and Oxygen Mass Transfer; B) Effect of Surface Aeration of Impellor on Power Consumption (6-Bladed Radial Impellor - 58mm diameter) 2 5 o or E 3 Q. in o I 2 & E « > I 1 IT I4kg/n 3 (NH 4) 2S0 4 Full Bottles 1 1 6 50mm 6BA58 X 10 o o or 01 § 0-5 1 2 3 4 Impellor Tip Speed (m/s) * 10 o. E 3 in c o O v I 0) Q . E Wkg/ta' (N^jSC, Full Baffles 50mm 68A58 Ungassed A 1 2 3 4 Impellor Tip Speed (m/s) in in a 2 e a> a> >» O Figure 29. A) Effect of Agitation Rate on Impellor Gas Pumping Capacity and Oxygen Mass Transfer; B) Effect of Surface Aeration of Impellor on Power Consumption (6-Bladed Axial Impellor - 58mm diameter) g £ 5 1 1 7 o or ? "5. £ 5 4-3-& 2 -2 « or -68A58« 6BR58c T 50 mm 60 mm l4kg/m s(N^LS04 Full Baffles -lA-O o & E o IO o. 4) O or tn § 0-5 »i in o c a> l~ 2 3 4 Impellor Tip Velocity (m/s) 15 e .2 Q. E 3 in e o O a> s. a a . E 10 Kkg/m3 {NH4)2S04 Full Baffles 50mm 60mm 6BA58 6BR58 Ungassed / 0 I 2 3 4 Impellor Tip Velocity (m/s) A) Effect of Agitation Rate on Impellor Gas Pumping Capacity and Oxygen Mass Transfer; B) Effect of Surface Aeration of Impellor on Power Consumption (Standard Dual Impellor - 58mm diameter) 1 1 8 Figure 31. Effect of High Oxygen Partial Pressures on Oxygen Mass Transfer Rate (58mm diameter) 119 TABLE 1 EFFECT OF IMPELLOR TYPE IMPELLOR TYPE AGITATION RATE VOLUMETRIC POWER RELATIVE MASS TRANSFER EFFICIENCY m/s kW/m3 (kg0 2/m 3»min)/(kW/m 3) Single 6-Bladed Radial Disc - 46 cm 4.05 0.558 0.0341 Single 4-Bladed Radial - 46 cm 4.05 0.558 0.0054 Single 4-Bladed 45°-Pitched Blade Axial - 46 cm 4.05 0.224 0.0027 Standard Autoclave Configuration - 46 cm 4.05 0.832 0.0059 VOLUMETRIC POWER: The power draw of the impellor in air has been subtracted = 0.566 kW/m3 SINGLE IMPELLOR IMMERSION DEPTH = 23.9 cm = 9.0 inches 120 TABLE 2 EFFECT OF IMPELLOR IMMERSION IMPELLOR TYPE DEPTH OF IMMERSION 23cm 34 cm 46cm kW/m3 RMTE kW/m3 RMTE kW/m3 RMTE 6-Bladed Radial Disc - 46 cm 0.558 0.0341 0.929 0.0078 1.067 0.0007 Single 4-Bladed Radial - 46 cm 0.558 0.0054 0.477 0.0050 0.495 0.0012 Single 4-Bladed 45°-Pitched Blade Axial - 46 cm 0.224 0.0027 0.268 0.0015 0.260 0.0008 STD Autoclave Configuration - 46 cm 0.832 0.0059 R^MTE: Relative Mass Transfer Efficiency (kg02/m3•min)/(kW/m3) AGITATION RATE = 4.05 m/s = 169 rpm 121 TABLE 3 CRITICAL TIP VELOCITY CORRELATION IMPELLOR DEPTH (m) EXPERIMENTAL1 VALUE (m/s) V c r i t i c a l PREDICTED2 VALUE (m^) V c r i t i c a l 0.229 2.42 2.12 0.343 2.93 2.59 0.457 3.27 2.99 1. Refer to Figure 16. 2. Equation 2.4 122 TABLE 4 EFFECT OF IMPELLOR DIAMETER IMPELLOR TYPE VOL. POWER (kW/m3) RMTE kg02/mJ 'min ( kW/m3 ' AGITATION (m/s) 6-Bladed Disc 1.200 0.0188 4.05 - 53 cm 6-Bladed Disc 0.558 0.0341 4.05 - 46 cm 4-Bladed Radial 0.811 0.0107 4.05 - 53 cm 4-Bladed Radial 0.558 0.0054 4.05 - 46 cm 4-Bladed 45°- 0.403 0.0012 4.05 Pitched Blade -53cm 4-Bladed 45°- 0.224 0.0027 4.05 Pitched Blade -46cm STD Autoclave Configuration 0.832 0.0059 4.05 * RMTE: Relative Mass Transfer Efficiency . 123 TABLE 5 EFFECT OF HALF BAFFLES ON MASS TRANSFER BAFFLE LENGTH AGITATION (m/s) POWER (kW) VOL. POWER (kW/m3) RMTE kg02/m3 •min ( — > kW/m3 HALF 4.05 0.388 0.224 0.0080 FULL 4.05 0.388 0.224 0.0027 HALF 6.30 1.268 0.731 0.0115 FULL 6.30 1.748 1.008 0.0071 46 cm-FOUR-BLADED IMPELLOR 45°-PITCHED IMPELLOR RMTE: Relative Mass Transfer Efficiency 124 TABLE 6 EFFECT OF IMPELLOR IMMERSION ON HALF BAFFLES BAFFLE LENGTH DEPTH OF IMMERSION 23cm 34cm 46cm kW/m3 RMTE kW/m3 RMTE kW/m3 RMTE HALF 0.224 0.0080 0.221 0.0050 0.260 0.0027 FULL 0.224 0.0027 0.268 0.0015 0.260 0.0008 RMTE: Relative Mass Transfer Efficiency (kg02/m3•min)/(kW/m3) AGITATION RATE = 4.05 m/s = 169 rpm 46 cm-Four-Bladed 45°-Pitched Blade Impellor 125 TABLE 7 OXYGEN DEPLETION IN GAS BUBBLES OXYGEN CONCENTRATION (% VOLUME) HEADSPACE GAS REACTED BUBBLES 20.6 13.0 53cm- 6-BLADED RADIAL DISC IMPELLOR-UPPER 46cm- 4-BLADED RADIAL IMPELLOR-LOWER IMPELLOR TIP SPEED =2.90 m/s - UPPER 2.49 m/s - LOWER OXYGEN MASS TRANSFER RATE = 0.0061 kg0 2/m 3»min 126 TABLE 8 EFFECT OF OXYGEN CONCENTRATION OXYGEN MASS TRANSFER RATE (kgO 2/m 3 •min) AIR SPARGE PURE OXYGEN SPARGE 0.0327 0.1137 46cm- 6-BLADED RADIAL DISC IMPELLOR-UPPER 46cm- 4-BLADED RADIAL IMPELLOR-LOWER IMPELLOR TIP SPEED =4.05 m/s TABLE 9 EFFECT OF OXYGEN CONCENTRATION OXYGEN MASS TRANSFER RATE(kg02/m3min) AIR SPARGE PURE OXYGEN SPARGE 0.0409 0.1277 53cm- 6-BLADED RADIAL DISC IMPELLOR-UPPER 46cm- 4-BLADED RADIAL IMPELLOR-LOWER IMPELLOR TIP SPEED = 4.72, m/s - UPPER 4.05 m/s - LOWER 1 2 8 TABLE 10. SMALL DIAMETER-HIGH SPEED IMPELLOR EXPERIMENTS Agitator Speed (m/s) Oxygen Mass Transfer Rate (g/1 02/min) Small High Speed Impellor 3-93 7-98 00029 0-0068 Standard 4 0 5 0 0 0 4 9 1 22-9 cm e -6B-RD-6 • 4B-A-I8 _] 140 cm ~~f T 4B-A-I8: Tip Speed = 2 4 9 m/s 6B-RD-6 6Bladed radial disc impellor 15-2 cm diameter 4B-A-I8 4Bladed axial impellor 458 cm diameter 129 TABLE 11 EFFECT OF SPECIAL SPARGER MODE SPARGER MODE kW/m3 RMTE SPARGER ABOVE BOTTOM IMPELLOR 0.850 0.929 0.0369 0.0290 AVERAGE 0.890 0.0330 SPARGER BELOW BOTTOM IMPELLOR 0.890 0.0327 RMTE: Relative Mass Transfer Efficiency (kgO 2/m 3'min)/(kW/m 3) 130 T A B L E 1 2 E F F E C T O F I M P E L L O R D I A M E T E R D I A M E T E R ( m m ) O X Y G E N M A S S T R A N S F E R R A T E ( k g 0 2 / m 3 m i n ) R E L A T I V E G A S P U M P I N G R A T E ( m 3 / m i n ) P O W E R C O N S U M P T I O N ( W ) 4 0 0 . 0 0 4 2 0 . 0 0 1 1 2 . 5 5 5 8 0 . 0 0 7 9 0 . 0 0 2 2 4 . 5 0 6 - B L A D E R A D I A L D I S C I M P E L L O R , F U L L B A F F L E S T I P S P E E D = 2 . 8 2 m / s , D E P T H O F I M M E R S I O N = 5 0 mm 1 3 1 TABLE 13 EFFECT OF BAFFLE LENGTH BAFFLE LENGTH OXYGEN MASS TRANSFER RATE (kg02/m3 min) RELATIVE GAS PUMPING RATE (m3/min) POWER CONSUMPTION (W) HALF 0.0100 0.0026 2.75 FULL 0.0134 0.00531 4.60 6-BLADE RADIAL DISC IMPELLOR, 58 nm DIAMETER TIP SPEED = 3.26 m/s , DEPTH OF IMMERSION = 50 mm 1. Extrapolated Value 132 TABLE 14 ALTERNATE DUAL IMPELLOR CONFIGURATION IMPELLOR OXYGEN MASS RELATIVE GAS POWER CONFIGURATION TRANSFER RATE PUMPING RATE CONSUMPTION (kg02/m3 min) (m3/min) (W) Standard 0.0118 0.0040 10.4 Configuration 6-Bladed Radial * Disc-Upper/ 0.0123 0.0032 10.3 6-Bladed Radial -Lower FULL BAFFLES , 58 mm DIAMETER TIP SPEED *> 3.28 m/s , DEPTH OF IMMERSION = 50 mm -UPPER = 110 mm -LOWER * Extrapolated Value 1 3 3 TABLE 15 EFFECT OF SOLIDS SOLID CONCENTRATION (kg/m3) OXYGEN MASS TRANSFER RATE (kg02/m3 •min) RELATIVE GAS PUMPING RATE (m3/min) POWER CONSUMPTION (W) 0 0.0113 0.0038 9.50 46 0.0104 0.0036 9.33 STANDARD DUAL IMPELLOR CONFIGURATION, 58 mm DIAMETER, FULL BAFFLES TIP SPEED - 3.23 m/s , DEPTH OF IMMERSION - 50 mm -UPPER = 110 mm -LOWER 1 3 4 APPENDIX A MIXING MODEL EXPERIMENTS Below a r e shown t h e t h r e e t y p e s o f i m p e l l o r s used d u r i n g t h e p i l o t - s c a l e work a t Cominco*s T e c h n i c a l R e s e a r c h c e n t r e ( t h e d i a m e t e r s v a r i e d between 46cm and 53cm): 4-Bloded Axial Impellor 4-Bladed Radiol Impellor 6- Bladed Radiol Disc Impellor m -23cm-T J 75em - 2 6 5 c m — •19cm-15 I cm cm-J - 2 3 c —26-5cm— 1 3 5 APPENDIX B SPARGING MODES F o r c e r t a i n e x p e r i m e n t s i n t h e m i x i n g model t h e r e was s p a r g i n g u s e d . One s p a r g e r was d e s i g n e d t o r e s t on t h e bottom o f t h e v e s s e l and s p a r g e from u n d e r n e a t h t h e bottom i m p e l l o r . The o t h e r s p a r g e r was made so t h a t gas c o u l d be sp a r g e d from j u s t above t h e lower i m p e l l o r . Mode-. Below Lower Impellor Mode: Above Lower Impellor 136 APPENDIX C AIR AND OXYGEN FLOW DIAGRAM D i f f e r e n t m i x t u r e s o f oxygen and n i t r o g e n were used i n t h e m i x i n g model e x p e r i m e n t s t o v a r y t h e oxygen c o n c e n t r a t i o n . Relow i s a p i p i n g s c h e m a t i c showing t h e placement o f v a l v e s and gauges. Pressure Gauge Valve Xo Rotameters • (X) Valve o Pressure Gouge ® Valve 6 Oxygen Cylinder Compressed Air or Oxygen Cylinder To Headspace Purge To Sparger 1 3 7 APPENDIX D BENCH-SCALE EXPERIMENTS Below a r e shown t h e t h r e e t y p e s o f i m p e l l o r s used d u r i n g t h e e x p e r i m e n t s a t U.B.C. ( t h e d i a m e t e r s v a r i e d between 40mm and 58mm): 6-Bloded .Axial Impellor 6-Bloded Radial Impellor 'J2 7-5 Tim —20mm— • 29mm — 6-Bladed Radial Oisc Impellor 138 APPENDIX E POWER MEASUREMENTS In t h e b e n c h - s c a l e e x p e r i m e n t s power measurements were made from t h e t o r q u e m e a s u r i n g d e v i c e shown s c h e m a t i c a l l y below. Knowing t h e r o t a t i o n a l speed o f t h e i m p e l l o r and t h e a p p l i e d t o r q u e t h e power was c a l c u l a t e d as f o l l o w s : s p r i n g b a l a n c e r e a d i n g : lOOg o r 0.1kg arm l e n g t h : 0.1m r o t a t i o n a l s p e e d : 900 rpm Power: 0.1kg x 0.1m x 9.81m/s x 2 x 900rpm x 1/60 9.25 Watts 139 APPENDIX F VESSEL DESIGN FOR MEASURING IMPELLOR GAS PUMPING RATES Below i s shown a s e c t i o n a l d r a w i n g o f t h e 2 0 l i t r e v e s s e l used t o measure t h e r e l a t i v e i m p e l l o r gas pumping r a t e i n t h e bench s c a l e e x p e r i m e n t s . A s e a l e d gas plenum e x i s t e d above t h e l i q u i d , e x c e p t f o r t h e gas s a m p l i n g p o r t which c o u l d be used t o draw gas samples from t h e plenum. By drawing t h e gas out o f t h e plenum t h r o u g h a r o t a m e t e r a t a r a t e t h a t k e pt t h e l i q u i d l e v e l c o n s t a n t t h e s t e a d y s t a t e i m p e l l o r gas pumping r a t e c o u l d be re a d from t h e r o t a m e t e r : Plenum Gas Sampling Port "150mm" 100 mm O 01 00 u c 8 e l i •130 mm" 'Liquid Sampling Port Gas Seal Around Perimeter •o w u a a . cn 300 mm •f Each Set of Four Baffles is Removable 140 APPENDIX 6 POWER DRAW CALCULATION FOR SMALL DIAMETER HIGH SPEED IMPELLOR P Newton's law conversion Power = P Impellor Diameter = D Power number = N„ Impellor r o t a t i o n a l speed = N Impellor blade width = w Density =^ f a c t o r = g P = 4.63 x 10 -16 N PC For a small i m p e l l o r (15.2 cm) 6-bladed d i s c i m p e l l o r (w/D=l/5) P = 4.63 x 1 0 " 1 6 (5.0)(1.0)(1000) 3(15.2) 5 = 1.90 kW = 2.55 Hp Assuming the motor has a 90 % e f f i c i e n c y , a 2.11 kW motor i s re q u i r e d . N = 5.0 P N = lOOOrpm f 1.0 D 15.2 cm (6 inches) 141 APPENDIX H C a l c u l a t i o n of Oxygen Consumption Rate i n Anaconda A r b i t e r Process Reference: Kuhn, M.C., et a l ; CIM B u l l e t i n ( F e b . 1974), p62-74 Temperature = 60-90 °C Residence Time = 5 hours 0 2 P a r t i a l Pressure = 34kPa = 0.34 atm Weed Concentrate Composition: wt % Cu Fe S 26.6 . 21.4 33.2 % E x t r a c t i o n : 95.0 25.0 26.7 The i r o n i s e x t r a c t e d as Fe(0H)g. 3 The f i n a l s o l u t i o n has a copper concentration of 45.25 kg/m . n 3 0.95 x 0.266 kg concentrate/m 3 = 45.25 kg/m ^ ^ ^ 3 S o l u t i o n Composition (kg/m ): Cu FeCFeCOH)^) S 45.25 9.58 15.87 Oxygen Demand: Cu 45.25 x 16/63.5 = 11.40 Fe 9.58 x 24/55.8 = 4.12 S 15.87 x 48/32.0 = 23.81 3 Total Oxygen Demand = 39.33 kg/m 3 Total Oxygen Demand (kg O^ /m min atm) = 0.39 142 APPENDIX I Calculation of Oxygen Consumption Rate in Sherritt Gordon Ammonia Leach Reference: Forward, F.A., et a l ; Trans AIME, vol 203 (1955) p457 Nashner, S.; CIM Bulletin, vol 58, (1955) p212 Temperature = 70-80°C Residence Time = 19.2 hours 0 2 Partial Pressure = 68kPa = 0.68 atm Concentrate Composition: wt % Ni Fe S 12.0 30.0 28.0 Extraction % 95.0 60.0 60.0 The iron is assumed to extract as FeCOrl)^  and the S as S0^~. 3 The final solution has a nickel concentration of 45.0 kg/m . 3 45.0 kg/m3 3 kg concentrate/m = = 395 kg/m 0.95 x 0.12 Solution Composition: (kg/m3) Ni Fe(Fe(0H)3) S(S04 ) 45.0 71.1 66.4 Oxygen Demand: Ni 45.0 x 16/58.7 = 12.3 Fe 71.1 x 24/55.8 = 30.6 S 66.4 x 48/32.0 = 99.6 3 Total Oxygen Demand = 142.5 kg/m 3 Total Oxygen Demand (kg O^ /m min atm) = 0.18 143 APPENDIX J Calculation of Oxygen Mass Transfer Rate in Cominco Zinc Pressure Leach References: Parker, E.G.; CIM B u l l e t i n (May 1981), p 145 Parker, E.G., et a l ; Proc 3rd Inter. Symp. on Hydrometallurgy, AIME (March 1983), p927 Temperature = 145-155°C Residence Time (1st Compartment) = 26min O2 P a r t i a l Pressure 750kPa =7.5 atm Concentrate Composition: wt % Fe S Pb Zn 11.5 30.5 6.5 49.0 % Extraction 82 ! o(Fe 3 +) 8o!3(S°) 82.0 82!o It i s assumed that 2.4% of the sulphur reacts to form S0^~. 3 The f i n a l solution has a zinc concentration of 120 kg/m in the 1st compartment and an i n i t i a l concentration of 50 kg/m in the feed. 3 1 2 0 " 5 0 3 kg concentrate/m - = = 174 kg/m 0.82 x 0.49 Oxygen Demand: Zn (120-50) x 16/65.4 = 17.13 Fe (0.115x174x0.82) x 24/55.8 = 7.06 Pb: {0.065x174x0.820 x 16/207 = 0.72 S(S0 4 ) (2.4/100x174x0.305) x 64/32.0 = 2.55 Total Oxygen Demand = 27.46 kg/m 3 Total Oxygen Demand (kg 02/m min atm) = 0.14 3 

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