ARTICLE pubs.acs.org/IECR
Experimental Study on LiquidLiquid Macromixing in a Stirred Tank Yanchun Zhao,†,‡ Xiangyang Li,† Jingcai Cheng,† Chao Yang,*,† and Zai-Sha Mao† †
Key Laboratory of Green Process and Engineering, National Engineering Laboratory for Hydrometallurgical Cleaner Production Technology, Institute of Process Engineering, Chinese Academy of Sciences, Beijing 100190, China ‡ Graduate University of Chinese Academy of Sciences, Beijing 100049, China ABSTRACT: In this paper, the experimental data on the mixing time and power consumption of two immiscible liquids in a mechanically agitated baffled tank are presented. The electric conductivity method was taken for the measurement of the mixing time and the shaft-torque method for the power consumption measurement. Tap water was used as the continuous phase and kerosene the dispersed phase. The effects of the agitation speed, type of impeller, clearance of the impeller off the tank bottom, volume fraction of the dispersed phase, physical properties of the liquids, and probe position on the macromixing of the liquidliquid system were studied. The phenomena of macromixing are largely similar to those of single-liquid and gasliquid stirred tanks. The experiment indicates that the flow field and turbulence can be dampened at high volume fraction of the dispersed phase while enhanced at low percentage. The mixing time becomes longer with increasing viscosity of the dispersed phase. The results show that the pitched blade turbine downflow is more efficient for macromixing than the others tested in this work.
1. INTRODUCTION Dispersion of two immiscible liquids in mechanically agitated vessels is a very common industrial operation widely used in the chemical, food, petroleum, pharmaceutical, and power industries. These industries have a rich faculty of knowledge at their command, but the mixing processes are far from being fully understood. The main aspects that affect liquidliquid mixing are the drop size, the drop size distribution, the minimum impeller speed for dispersion, drop breakup and coalescence, drop suspension, phase inversion, and the influence of the material on the surface behavior, etc.1 Investigations previously conducted on the liquidliquid dispersion in stirred tanks were mainly focused on the minimum agitation speed for complete liquidliquid dispersion,2,3 the relative influence of the dispersed-phase viscosity and interfacial tension on the equilibrium drop size and drop size distribution,4 correlations for the mean size and size distribution,5 the influence of the rotational speeds, different impellers with varying diameter, and off-bottom clearance on the drop size distribution,6 and the fluid dynamic characteristics of immiscible liquidliquid dispersions at low dispersed-phase holdup by the two-phase particle image velocimetry (PIV) technique.7 Modeling these systems was also reported in recent years using the lattice-Boltzmann method and computational fluid dynamics (CFD) techniques with the multiblock model.8,9 The mixing time is one of the most significant parameters in expressing the performance of stirred tanks. There have been many studies on single-liquid systems,10,11 but few on gasliquid,12 solidliquid,13 and gasliquidsolid14 dispersions. A report on the mixing time of a liquidliquid dispersion in a stirred tank has not been seen in the open literature. The purpose of this paper is to contribute to the knowledge on macromixing of liquidliquid systems. The mixing time and power consumption in both single-liquid and liquidliquid systems were obtained. The effects of various operation conditions on r 2011 American Chemical Society
macromixing of the liquid phase were analyzed, and the mixing time data were correlated.
2. EXPERIMENTAL SECTION 2.1. Experimental Setup. The device used in this experiment is shown in Figure 1. The diameter of the stirred tank is 380 mm, and the total height is 600 mm. Four baffles with a width of 38 mm were evenly equipped at the wall. The liquid height for all experiments was set at H = T. Different types of stirrers were used (Figure 2), namely, Rushton disk turbine (RDT), half-circle blade disk turbine (HCDT), 45° pitched blade turbine downflow (PBTD), and 45° pitched blade turbine upflow (PBTU). The four impellers have the same diameter (D = T/3). The impeller blade thickness of the RDT is 1 mm, and the impeller blade width of the RDT is 26 mm. 2.2. Measurement Method. The mixing time was determined by means of the electric conductivity method in this work. In most experiments 10 mL of NaCl solution (250 g/L) was injected into the system as the tracer at point P1, which was at the liquid surface near the wall. The effect of the injection and probe positions was also examined. There were two monitoring positions; one was placed at the opposite location of the injection point 60 mm below the liquid surface and the other 80 mm above the bottom. The time elapsed between the addition of tracer and the moment when the instant conductivity reached within 95% of its final value was defined as the mixing time for each probe. The sampling frequency of the electric conductometer was 100 Hz. Every experimental condition was tested repeatedly six to eight times to get an average of the mixing time. The power consumption was measured by the shaft-torque method.15 Received: November 9, 2010 Accepted: March 31, 2011 Revised: March 27, 2011 Published: March 31, 2011 5952
dx.doi.org/10.1021/ie102270p | Ind. Eng. Chem. Res. 2011, 50, 5952–5958
Industrial & Engineering Chemistry Research
ARTICLE
Table 1. Values of the Mixing Time versus Stirred Time (Rushton Impeller, C = T/3, 329 rpm) stirred time
15 min
30 min
1h
1.5 h
2h
mixing time
12.35 s
10.82 s
9.38 s
9.54 s
9.22 s
Table 2. Physical Properties of the Continuous and Dispersed Phases density (kg/m3)
liquid
Figure 1. Experimental setup: (1) computer, (2) conductometer, (3) conductivity electrode, (4) stirred tank, (5) injector, (6) speed controller, (7) motor, (8) rotary torque transducer, (9) amplifier; A (probe 1), 60 mm from liquid surface, 175 mm from shaft axis; B (probe 2), 80 mm from bottom, 175 mm from shaft axis; P1 (injector 1), at the liquid surface, 185 mm from shaft axis; P2 (injector 2), 160 mm from bottom, 185 mm from shaft axis; P3 (injector 3), 80 mm from bottom, 185 mm from shaft axis.
viscosity
interfacial
(Pa 3 s)
tension (N/m)
water
997.7
0.001
kerosene
789.5
0.002
0.02400
paraffin oil methyl silicone oil
864.5 972.4
0.052 0.565
0.00281 0.01790
Table 3. Effect of the Probe and Injector Positions (RDT, N = 329 rpm, ud = 10%) t (s) position
A
B
P1
9.38
9.20
P2
8.70
7.9
P3
8.94
8.50
different liquids (kerosene, paraffin oil, and methyl silicone oil) are used as the dispersed phase. Kerosene was used in most experiments; paraffin and methyl silicone oil were just used to test the influence of the physical properties. The liquids were chosen to cover a wide range of physical properties with Fc > Fd and ηc < ηd. The physical properties of the liquids tested at room temperature are listed in Table 2.
3. RESULTS AND DISCUSSION 3.1. Mixing Time. 3.1.1. Effect of the Probe and Injector Positions. The mixing time for a set of operating conditions has Figure 2. Impellers used in our experiments.
The minimum stirring time before measurement was also tested. In the beginning, the hydrodynamics in the liquidliquid stirred tank is not stabilized. Only after some period of time of stirring does the experiment give a consistent measurement of the mixing time. The minimum stirring time was determined by examining the change of the mixing time versus the stirring time. The mixing time determined after 15 and 30 min and 1, 1.5, and 2 h is shown in Table 1. The data suggest that 1 h of stirring at least is needed for obtaining reliable and consistent data of the mixing time in the waterkerosene system. The stirring speed tested in the experiment was in the range of 329461 rpm. It was sufficiently above the critical speed for liquid-phase dispersion (237 rpm for the RDT, 272 rpm for the HCDT, and 329 rpm for the PBTD and PBTU in the waterkerosene system). 2.3. Materials. The model system is an oil-in-water emulsion, in which the tap water is taken as a continuous phase and three
been determined by varying the probe and the tracer injection locations, as shown in Table 3. A Rushton impeller located at T/3 above the bottom with a speed of 329 rpm was used. The volume fraction of the dispersed phase was 10%. A and B are the positions of the probes, while P1, P2, and P3 are the positions of the injectors. As shown in Figure 1, by adding the tracer into the liquid in the stirred tank from the P2 point, the shortest mixing time was observed at both monitoring positions A and B. This is because P2 was located in the discharge stream. The liquid in this area with a high velocity caused the tracer to disperse quickly and led to a short mixing time. The tracer was therefore, and for convenience of operation, added at P1 in most experiments. No matter where the tracer was injected, the mixing time tested at A (upper position) was always longer than that at B (lower probe position). The result may be explained by the flow pattern and mixing mode. The monitoring positions were set at the location opposite the injection point. The tracer injected into the liquid surface on one side flowed through the blade area with the liquid and then cycled to the liquid surface on the opposite side. The length of the circulation path for A was longer than that for B, so the mixing time observed in point A was longer than that in B. To 5953
dx.doi.org/10.1021/ie102270p |Ind. Eng. Chem. Res. 2011, 50, 5952–5958
Industrial & Engineering Chemistry Research
Figure 3. Mixing time versus impeller type (N = 427 rpm, C = T/3, jd = 10%).
attain the desired level of homogeneity at most parts of the whole vessel, the mixing time was obtained by choosing the bigger value recorded among all probes in all later tests. These considerations are also conventionally observed in experiments on mixing in single-liquid and gasliquid systems. 3.1.2. Effect of the Impeller Type. A comparison of the mixing time by the RDT, HCDT, PBTU, and PBTD impellers for the dispersed-phase volume fraction jd = 10% at the agitation speed N = 427 rpm is shown in Figure 3. The impellers were positioned at C = T/3. An increase in the mixing time has been detected due to the presence of the dispersed phase, which indicates that the macromixing process was dampened by the dispersed phase in the experimental conditions. For the same impeller, the differences in the mixing time in the two-phase system with respect to the single-phase case were moderate except for the PBTU impeller, which was influenced significantly by the second liquid phase. Kerosene was lighter than water and tended to float on the surface, but the flow produced by the PBTU was upward, which had a negative impact on the dispersing oil phase. When the liquid left the impeller and flowed up to the surface, the lighter phase tended to coalesce and stayed on the surface in the coalesced state. Consequently, the poor dispersion of the oil phase caused an increasing mixing time. 3.1.3. Effect of the Impeller Clearance. The effect of the impeller clearance on the mixing time in agitated tanks was tested for each impeller type. The values of the impeller clearance from the tank bottom for the Rushton impeller that were examined are T/6, T/4, and T/3 (Figure 4a). The experimental results in the single-phase system can be seen in line with those of Raghav Rao and Joshi.16 With a second phase the mixing time decreased as the impeller clearance decreased in the range T/3 to T/6. This is consistent with the conclusion for a single-phase system. When the impeller was located at the position T/6 from the tank bottom, the lower circulation area in the double loop flow structure was restrained, so there was only a loop left and the flow structure tended to be an axial one. The two-phase liquid dispersion circulated in one loop and flowed through the impeller region with high turbulence repeatedly, which made the mixing quicker. As the clearance from the tank bottom changed from T/6 to T/3, the flow structure of double loops recovered gradually, the radial flow was strengthened, and the mixing time became longer. As seen from Figure 4b, the value of
ARTICLE
the mixing time was found to increase with the increase in the impeller clearance for the HCDT. The effect of the clearance for the HCDT was similar to that for the RDT, and the flow patterns made by them were similar. In the case of the PBTD, the mixing time decreased as the impeller clearance increased from T/6 to T/3, while it was found to increase again with C to T/2 (Figure 4c). This trend of variation is similar to that in the single-phase case, which accords with the report of Rewatkar and Joshi.17 When the impeller was located at a clearance of T/2 far from the tank bottom, the flow was not strong enough to turn around before reaching the bottom and a new loop formed under the early one. The two relatively independent flow zones retarded the transport of tracer to the area at the bottom. The lower loop was a weak circulation area, which made no advantage for oil-phase dispersion, and therefore, a longer mixing time was observed. For a clearance of T/3, axial flow was predominant with a short mixing time. When the clearance decreased from T/4 to T/6, blending was weaker in the upper tank and the mixing time was extended. It can be seen that the PBTU with C = T/4 performed the best in both single-liquid and liquid liquid systems (but the mixing time was still inferior to those of other impellers). If the C/T ratio was either decreased or increased from this value, the mixing time was found to increase (Figure 4d). 3.1.4. Effect of the Agitation Speed. The agitation speed was varied from 329 to 461 rpm to determine the effect of the agitation speed on the mixing time in the tank agitated by the RDT, HCDT, and PBTD. The impellers were all located at T/3, and the holdup of the dispersion phase was 10%. The mixing time decreased as the agitation speed increased. At a given agitation speed, the mixing time for the PBTD was the lowest and that for the HCDT was the highest, as seen in Figure 5. Although the RDT provided higher shear and turbulence levels, the main flow pattern of the PBTD was an axial one, which favored macromixing more. The flow patterns of the RDT and HCDT were all radial, but the shear produced by the RDT was higher, which caused a shorter mixing time than that of HCDT. 3.1.5. Effect of the Dispersed-Phase Volume Fraction. As shown in Figure 6, measurements at 0, 3, 5, 7, 10, 15, and 20 vol % of the dispersed phase were performed. At volume fractions of 1020% of the dispersed phase, the mixing time increased and was longer than that measured in single-phase systems, but the measurements performed at 3%, 5%, and 7% revealed a reverse trend; i.e., the values of the mixing time were shorter than that in single-phase systems. This might indicate that turbulence could be enhanced by the dispersed phase at low volume holdup and dampened at high holdup. The result was consistent with those of two previous reports. Laurenzi et al.7 used the PIV technique to investigate the fluid dynamic characteristics at 1% dispersedphase volume fraction and found that the action of the dispersed phase was generally to promote turbulence and the droplets were characterized by both radial and axial rms velocity values greater than those of the continuous phase. The presence of the dispersed phase of low volume percentage intensified the turbulence by vortex shedding, which was caused by the motion of the droplets relative to the continuous phase. Svensson and Rasmuson18 measured the velocity vector fields by the laser Doppler anemometry (LDA) technique for 0%, 10%, and 25% dispersed phase and considered that the turbulence was dampened when the amount of oil was increased. The increase of the percentage of the more viscous dispersed phase led to a higher effective viscosity. The increased drag force between the droplets 5954
dx.doi.org/10.1021/ie102270p |Ind. Eng. Chem. Res. 2011, 50, 5952–5958
Industrial & Engineering Chemistry Research
ARTICLE
Figure 4. Mixing time versus impeller clearance (N = 427 rpm, jd = 10%).
Figure 5. Impeller speed versus mixing time (C = T/3, jd = 10%).
and continuous phase because of the increase of viscosity modified the velocities of the droplets and led to a weaker turbulence finally. The mixing time increased at high holdup in the liquidliquid system due to the presence of the dispersed phase, which is similar to that reported for solidliquid19,20 and gasliquid21 systems. 3.1.6. Effect of the Physical Properties of the Dispersed Phase. Figure 7 presents a set of experiments using a Rushton impeller operated at 358 rpm with various oils. The physical properties of the dispersed phases can be seen in Table 2. It illustrates that
higher viscosity oils produced longer mixing times ranging from 7.86 s for 1.2 mPa 3 s oil to about 10.84 s for 565.3 mPa 3 s oil. The interfacial tension of the oil phase changed as the viscosity changed, but no obvious correlation with mixing time can be found. The mixing time increased with higher density and viscosity. This can be compared to the result of section 3.1.5 of the effect of the dispersed-phase volume fraction. In that section, the increase of the dispersed-phase percentage led to a lower average density and a higher apparent effective viscosity, but the mixing time also increased. The result appears to suggest that the increase in viscosity had a more significant influence on dampening the flow field and turbulence than the change in density. 3.2. Power Consumption. The power curves of the RDT and PBTD in single-phase systems are given in Figure 8a. The blade heights of the RDT and PBTD are both T/3. The power numbers for the two impellers are 5.77 and 1.75, respectively, in the completely turbulent regime. The predicted result for the RDT by the correlation NP ¼ 6:405 55:673
t D
ð1Þ
reported by Rutherford et al.22 was 5.89, which was close to the experimental measurement. Figure 8b shows the power curves for four impellers in the liquidliquid system. At a high Reynolds number, the power number tended to be constant. The RDT had the largest power number among the four impellers, while the PBTD and PBTU had similar power numbers which were the smallest. It is 5955
dx.doi.org/10.1021/ie102270p |Ind. Eng. Chem. Res. 2011, 50, 5952–5958
Industrial & Engineering Chemistry Research
ARTICLE
Figure 6. Volume fraction of the dispersed phase versus the mixing time (Rushton impeller, C = T/3, N = 329 rpm).
Figure 8. Power number versus Reynolds number.
Figure 7. Physical properties of the dispersed phases versus the mixing time (Rushton impeller, C = T/3, N = 358 rpm, jd = 10%).
concluded that the axial impeller is more energy-efficient than the radial one in liquidliquid systems. To compare the mixing time of different impellers under equal power consumption, the data of the mixing time (C = T/3, jd = 10%) are plotted versus the power in Figure 9. For equal power consumption, the mixing time for the PBTD is the shortest, and the mixing time for the HCDT is longer than that for the RDT at higher power. This indicates that the PBTD is relatively more energy efficient when compared to the HCDT and RDT. Figure 10 shows the effect of the impeller clearance on the power consumption per volume (PV) with a Reynolds number of 9.7 104 for the RDT, PBTD, and PBTU. The RDT required less energy at T/2 from the bottom than at other positions and required the most energy at T/3. The power input of the PBTD located at T/3 was the least. Figure 11 shows that the power input decreased with increasing dispersed-phase volume fraction. The density of the emulsion decreased and the effective viscosity increased as the dispersedphase percentage increased. Although the increase of the viscosity caused the value of NP to rise, the equation P = NPFN3D5 suggests that the influence of the reduced density on the power consumption was more notable than the influence of increasing NP due to higher viscosity.
Figure 9. Mixing time versus power consumption (C = T/3, jd = 10%).
3.3. Correlation of the Mixing Time. The relation of the nondimensional mixing time and the power number in liquid liquid systems was derived from the equation for single-phase systems proposed by Grenville and Nienow:23 2 1 1 D Np1=3 ¼ ð2Þ Nm 5:2 T
where the power number NP is that for the homogeneous systems. To express the influence of the second phase, a new parameter, jd, the volume fraction of the dispersed phase, was included in eq 2. The new correlation, which is suitable for volume fractions ranging between 0% and 20%, is developed as 2 1 1=3 D ¼ 0:095Np expð4:287jd 40:49jd 2 Þ ð3Þ Nm T 5956
dx.doi.org/10.1021/ie102270p |Ind. Eng. Chem. Res. 2011, 50, 5952–5958
Industrial & Engineering Chemistry Research
ARTICLE
(4) For the RDT and HCDT, the mixing time increases as the impeller clearance increases from T/6 to T/3. For the PBTD, the mixing time decreases first and then increases with a minimum value at T/3. For the PBTU, the mixing time has the lowest value at T/4 when the impeller clearance changes from T/6 to T/3. (5) The viscosity increase of the dispersed-phase liquid has a more significant influence on dampening macromixing than the increase in its density.
’ AUTHOR INFORMATION Corresponding Author
*Tel.: þ86-10-62554558. Fax: þ86-10-82544928. E-mail: chaoyang@ home.ipe.ac.cn. Figure 10. PV as a function of the impeller clearance from the bottom and impeller type (Re = 9.7 104, jd = 10%).
Figure 11. PV as a function of the dispersed-phase volume fraction (RDT, Re = 9.7 104, C = T/3).
Compared with the experimental data of the RDT (C = T/3), the maximum relative deviation of eq 3 is 9.21% while the mean relative deviation is 5.58%.
4. CONCLUSION Liquidliquid macromixing in a stirred vessel of standard geometry has been investigated with water and three other immiscible liquids. The mixing time of the continuous phase and the power consumption have been determined by means of electric conductivity and shaft-torque techniques, respectively. The following conclusions can be made: (1) Generally, the trend for the mixing time in the liquid liquid system is similar with that in a single-phase system. (2) An increase in the macromixing intensity of the continuous liquid phase has been detected at low dispersedphase holdup, while dampening of the macromixing intensity of the continuous phase has been found at high dispersed-phase volume percentage (g10%) as compared with the single-liquid-phase system. (3) The combination of mixing time and power consumption analysis shows that the order of energy efficiency is PBTD > RDT > HCDT.
’ ACKNOWLEDGMENT Financial support from the National Natural Science Foundation of China (Grants 20990224 and 20906090), National Science Fund for Distinguished Young Scholars (Grant 21025627), 973 Program (Grants 2009CB623406 and 2010CB630904), 863 project (2011AA060704), Beijing Natural Science Foundation (Grant 2112038), and Jiangsu Province Projects BY2009133 and BE2008086 is gratefully acknowledged. ’ NOTATION C = impeller clearance off the tank bottom, mm D = diameter of the impeller, mm H = height of the liquid in the tank, mm N = agitation speed, rpm NP = power number, P/FN3D5 Nm = nondimensional mixing time, Ntm P = power consumption, W PV = power consumption per volume, W/m3 Re = Reynolds number, NFD2/η T = diameter of the tank, mm t = impeller blade thickness, mm tm = mixing time, s η = dynamic viscosity, Pa 3 s Fc, Fd = densities of the continuous and dispersed phases, kg/m3 Fm = mean density, jdFd þ (1 jd)Fc, kg/m3 jd = volume fraction of the dispersed phase ’ REFERENCES (1) Leng, D. E.; Calabrese, R. V. Immiscible liquidliquid systems. In Handbook of Industrial Mixing: Science and Practice; Paul, E. L., Atiemo-Obeng, V. A., Kresta, S. M., Eds.; John Wiley & Sons: New York, 2004; Chapter 12, pp 639753. (2) Skelland, A. H. P.; Ramsey, G. G. Minimum agitator speeds for complete liquid-liquid dispersion. Ind. Eng. Chem. Res. 1987, 26 (1), 77–81. (3) Armenante, P. M.; Huang, Y. T. Experimental determination of the minimum agitation speed for complete liquidliquid dispersion in mechanically agitated vessels. Ind. Eng. Chem. Res. 1992, 31 (5), 1398–1406. (4) Wang, C. Y.; Calabrese, R. V. Drop breakup in turbulent stirredtank contactors. Part II: Relative influence of viscosity and interfacial tension. AIChE J. 1986, 32 (4), 667–676. (5) Calabrese, R. V.; Wang, C. Y.; Bryner, N. P. Drop breakup in turbulent stirred-tank contactors. Part III: Correlations for mean size distribution. AIChE J. 1986, 32 (4), 677–680. 5957
dx.doi.org/10.1021/ie102270p |Ind. Eng. Chem. Res. 2011, 50, 5952–5958
Industrial & Engineering Chemistry Research
ARTICLE
(6) Zhou, G. W.; Kresta, S. M. Evolution of drop size distribution in liquidliquid dispersions for various impellers. Chem. Eng. Sci. 1998, 53 (11), 2099–2113. (7) Laurenzi, F.; Coroneo, M.; Montante, G.; Paglianti, A.; Magelli, F. Experimental and computational analysis of immiscible liquidliquid dispersions in stirred vessels. Chem. Eng. Res. Des. 2009, 87 (4A), 507–514. (8) Derksen, J. J.; Van Den Akker, H. E. A. Multi-scale simulations of stirred liquidliquid dispersions. Chem. Eng. Res. Des. 2007, 85 (A5), 697–702. (9) Alopaeus, V.; Koskinen, J.; Keskinen, K. I.; Majander, J. Simulation of the population balances for liquidliquid systems in a nonideal stirred tank. Part 2—Parameter fitting and the use of the multiblock model for dense dispersions. Chem. Eng. Sci. 2002, 57 (10), 1815–1825. (10) Sano, Y.; Usui, H. Interrelations among mixing time, power number, and discharge flow rate number in baffled mixing vessels. J. Chem. Eng. Jpn. 1985, 18 (1), 47–52. (11) Houcine, I.; Plasari, E.; David, R. Effects of the stirred tank’s design on power consumption and mixing time in liquid phase. Chem. Eng. Technol. 2000, 23 (7), 605–613. (12) Murthy, G. G. K; Elliott, J. F. Definition and determination of mixing time in gas agitated liquid baths. ISIJ Int. 1992, 32 (2), 190–195. (13) Kraume, M. Mixing times in stirred suspensions. Chem. Eng. Technol. 1992, 15 (5), 313–318. (14) Xu S. A.; Feng L. F.; Gu X. P.; Wang K. Mixing time in stirred tank of three-phase gas-liquid-floating particle systems. J. Chem. Eng. Chin. Univ. 2000, 14 (4), 328333 (in Chinese). (15) Wang, T.; Yu, G. Z.; Yong, Y. M.; Yang, C.; Mao, Z.-S. Hydrodynamic characteristics of dual-impeller configurations in a multiplephase stirred tank. Ind. Eng. Chem. Res. 2010, 49 (3), 1001–1009. (16) Raghav Rao, K. S. M. S.; Joshi, J. B. Liquid phase mixing in mechanically agitated vessels. Chem. Eng. Commun. 1988, 74 (1), 1–25. (17) Rewatkar, V. B.; Joshi, J. B. Effect of impeller design on liquid phase mixing in mechanically agitated reactors. Chem. Eng. Commun. 1991, 102, 1–33. (18) Svensson, F. J. E.; Rasmuson, A. LDA-measurements in a stirred tank with a liquidliquid system at high volume percentage dispersed phase. Chem. Eng. Technol. 2004, 27 (3), 335–339. (19) Bujalski, W.; Takenaka, K.; Paolini, S.; Jahoda, M.; Paglianti, A.; Takhashi, K.; Nienow, A. W.; Etchells, A. W. Suspension and liquid homogenization in high solids concentration stirred chemical reactors. Chem. Eng. Res. Des. 1999, 77, 241–247. (20) Micheletti, M.; Nikiforaki, L.; Lee, K. C.; Yianneskis, M. Particle concentration and mixing characteristics of moderate-to-dense solid liquid suspensions. Ind. Eng. Chem. Res. 2003, 42 (24), 6236–6249. (21) Pinelli, D.; Bakker, A.; Myers, K. J.; Reeder, M. F.; Fasano, J.; Magelli, F. Some features of a novel gas dispersion impeller in a dualimpeller configuration. Chem. Eng. Res. Des. 2003, 81, 448–454. (22) Rutherford, K.; Mahmoudi, S. M. S.; Lee, K. C.; Yianneskis, M. The influence of Rushton impeller blade and disk thickness on the mixing characteristics of stirred vessels. Chem. Eng. Res. Des. 1996, 74 (A3), 369–378. (23) Grenville, R. K.; Nienow, A. W. Blending of miscible liquids. In Handbook of Industrial Mixing: Science and Practice; Paul, E. L., Atiemo-Obeng, V. A., Kresta, S. M., Eds.; John Wiley & Sons: New York, 2004; Chapter 9, pp 507542.
5958
dx.doi.org/10.1021/ie102270p |Ind. Eng. Chem. Res. 2011, 50, 5952–5958
