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Gas-Liquid Dispersion with Buoyant Particles in a Hot-Sparged Stirred Tank Yuyun Bao, Zhengming Gao,* Xiaohua Huang, Litian Shi, John M. Smith,† and Rex B. Thorpe† School of Chemical Engineering, Beijing UniVersity of Chemical Technology, Beijing 100029, China, and Fluids and Systems Research Centre, School of Engineering (D2), UniVersity of Surrey, Guildford GU2 7XH, U.K.
Many industries use gas-liquid stirred-tank reactors with a third phase of buoyant particles. Few studies have looked at the effects of buoyant particle characteristics on solid-liquid suspension and gas-liquid dispersion, and even less have considered operation at elevated temperatures. This paper reports power consumption and gas holdup measurements made in a hot-sparged three-phase system. The vessel was a dished-base reactor of diameter 0.476 m (T) holding 0.145 m3 of liquid, agitated by a multi-impeller agitator (a hollow half-elliptical blade dispersing turbine below two up-pumping wide-blade hydrofoils, identified as HEDT+2WHU). This configuration has been recommended in previous work. Air, deionized water, and polypropylene (PP) particles were used in this work. The temperature of the water was in the range 80.582.5 °C. The results show that, as in a cold system, the effect of the solid concentration on aerated agitator power demand is negligible. There is little difference between the agitator power demand in cold- and hotsparged systems, except that the relative (gassed-to-ungassed ratio) power demand, RPD, is a little higher in hot-sparged conditions. The influence of power consumption and gas rate on gas holdup is slightly less when hot than when cold. Gas holdup in the hot-sparged system, which, with similar total gas rates and power input, is only about one-half or two-thirds of that at room temperature, is slightly increased at higher solids concentrations. The reasons for the differences in gas dispersion between cold- and hot-sparged systems are analyzed and discussed in this paper, with further experimental and CFD studies planned for future work. 1. Introduction In many industrial multiphase processes, e.g., fermentation, minerals processing, sewage treatment, and polymerization, buoyant particles are suspended in aerated stirred tanks. In a sparged system, the vapor pressure of the liquid phase may have a significant effect on the process. Water treatment and fermentation processes, and the great majority of academic research investigations, involve ambient (“cold”) conditions when this vapor pressure can be assumed to be insignificant in relation to the operating pressure. However, for many exothermic processes, e.g., oxidation, halogenation, and polymerization (i.e., “hot” systems), this assumption is likely to be invalid. For example, the stripping of the solvent and the recovering of monomer after solution polymerization of SBS or BR (cis-1,3butadiene rubber) involve vapor, water, and buoyant rubber blocks in a multi-impeller, hot-sparged three-phase system. The drawdown of the buoyant particles and the vapor dispersion are important in stripping the solvent during monomer recovery. In such hot-sparged systems, the vapor pressure should be taken into account. As adopted in previous work,1-6 in this paper, the total gas phase is considered to include the vapor along with the inlet gas. The approach is analyzed and calculated in detail in Section 2.2. Extensive studies of cold-sparged three-phase systems, usually involving settling particles, have been reported in the last 30 years.7-12 During the past decade, studies of hot gas-liquid systems have shown that hot-sparged conditions are significantly different from those in cold-gassed systems.2-6 In comparing hot- and cold-gassed operation, agitator power draw is greater at the higher temperatures; bulk and micromixing times are essentially unaltered, while retained gas fractions are substantially reduced. * Corresponding author. E-mail:
[email protected]. Tel.: +86-10-6441-8267. Fax: +86-10-6444-9862. † University of Surrey.
Recently, there has been increasing interest in the performance of multiple impeller agitators combining a hollow-blade disk turbine (HEDT) as the lowest impeller surmounted by two uppumping wide-blade hydrofoils (WHU).1,4-6,8,9,13 Bao et al.1 recommended this combination, shown in Figures 1and 2, for three-phase systems in which gas dispersion is combined with the suspension of buoyant particles. The radial bottom impeller disperses gas efficiently and operates at a higher RPD (relative power demand, the ratio of gassed to ungassed power demand, RPD ) PG/PU). The up-pumping (as opposed to down-pumping) hydrofoils can draw down buoyant particles at lower agitation speeds and power consumption in both ungassed and gassed conditions. The same gas, liquid, and buoyant solids have been used as in our previous work.1 There are some published data relating to multi-impeller agitator combinations, but there is very little for their operation in hot-sparged three-phase systems. A single example is from Dohi et al.,14 who report an experimental investigation of a boiling three-phase system with three fourpitched-blade disk turbines in a 0.20 m i.d., dished-base stirred tank. The present work with a three-stage stirred tank focuses on the effects of the concentration of solids and the gas inlet flow rate on the gas holdup and power consumption in a three-phase system with buoyant particles in the temperature range 80.582.5 °C, as shown in Table 2, when the water vapor pressure is ∼50 kPa. Correlations of power consumption and gas holdup in both cold- and hot-sparged systems are also presented here as a reference for industrial design. 2. Experimental Section 2.1. Experimental System. All experiments were carried out in a stainless steel dished-bottom cylindrical tank with a flat lid and glass windows on both sides and bottom, as sketched in Figure 3. The tank has an internal diameter T ) 0.476 m and a filled aspect ratio H/T ) 1.8. The volume of liquid is
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Ind. Eng. Chem. Res., Vol. 46, No. 20, 2007
Figure 1. Four-wide-blade hydrofoil impeller (WH).
Figure 3. Schematic of the experimental setup.
Figure 2. Six-half-elliptical-blade disk turbine (HEDT). Table 1. “Equilibrium Temperature” for Different Gas Rates and Agitation Speed at Same Heating Power, 9 kW T* (°C) for the following Qg × 103 (m3‚s-1) N
(s-1)
1.67 2.50 3.33 4.17 5.00 5.83 6.67 7.50
0
0.69
1.39
2.78
4.17
99.2 99.4 99.4 99.3 99.2 99.1 99.1 99.1
95.1 95.1 95.2 95.2 95.2 95.1 95.2 95.3
91.1 91.3 91.4 91.4 91.4 91.4 91.5 91.6
86.4 86.3 86.4 86.6 86.5 86.4 86.4 86.5
81.6 81.7 81.6 81.6 81.4 81.5 81.4 81.5
Table 2. Equilibrium Temperature for Hot-Sparged Experiment Qg × 103 (m3‚s-1)
4.17
5.0
heating power q (kW) 3 6 6 6 9 9 T* (°C) 81.0 82.0 82.5 81.3 82.0 80.5 Pv (kPa) 49.3 51.3 52.4 49.9 51.3 48.1 Qg+v × 103 (m3‚s-1) 1.56 3.25 5.30 6.30 8.11 9.17
0.69
1.39
2.22
2.78
3.47
12 82.0 51.3 11.7
∼0.145 m3. As a standard configuration, four 0.045 m wide baffles were symmetrically mounted with a 0.005 m gap between the baffle and the tank wall. Four vertical cylindrical heaters of diameter 0.034 m and length ∼0.25 m are mounted symmetrically in the bottom of the tank at a diameter of ∼0.36 m. In order to get rid of any influence from the geometry, the experiments for both room temperature and elevated temperature were all carried out with the heaters present. The electrical heat input could be adjusted from 3 to 12 kW in increments of 3 kW. During the experiments, the mixture of air and vapor passes to the shell side of a stainless steel condenser. Most of the water vapor leaving with the gas condenses and is returned to the stirred tank. The heat transfer area of the condenser tubes is 13.3 m2. The fall in liquid level during gas holdup measurements was 20%. With a viscosity change from 38 to 1 mPa‚ s, the gas holdup was reduced to about one-half. It seems clear that a decrease in liquid viscosity leads to a reduction in gas holdup. Paglianti et al.18 reviewed the influence of the liquid physical properties on the gas holdup and pointed out that the effect of viscosity has to be anayzed carefully. The gas holdup can decrease or increase according to the absolute value of the viscosity if the liquid viscosity decreases. When viscosity is reduced from 4 to 1 mPa‚s, the gas holdup decreases. However, the gas holdup increases a little when the viscosity changes from 100 to 1 mPa‚s. One explanation is that low viscosity will lead to more chances of bubble coalescence as a result of low gas holdup, but it seems more convincing to totally consider the competitive effects of the viscosity on the bubble generation, coalescence, breakup, and bubble transportation (i.e., circulation and buoyancy), which is still under investigation. An increase in interfacial tension might also cause a decrease in gas holdup, though the change in this variable over this temperature range is modest. In this work, the modest reduction in water viscosity from ∼1 to 0.36 mPa‚s lies in the low-viscosity range, consistent with the decreasing gas holdup. Previous research4 has shown that there are vortices behind each blade of turbine and hydrofoil impellers. At higher operating temperatures, with the same total gas flow rate, there will be additional evaporation into the low-pressure region in the vortices, increasing the size of the cavities and affecting the average bubble size in the stirred tank. Visual observation certainly gives the impression that the bubbles are significantly larger in a hot system than in a cold one. Moreover, the diameter of the bubbles changed in a different way for different temperature while rising from the bottom to the surface of the stirred tank. The conditions in the reactor are different when it is operating at 20 °C with a water vapor pressure of 2.3 kPa or at 82 °C with a vapor pressure of 52.4 kPa. With a pressure at the free surface of 101.3 kPa and at the bottom of the reactor of 109.7 kPa by adding the static pressure of 0.857 m water level, in cold operation the partial vapor pressure of the air in the rising bubble decreases from 107.3 to 99 kPa, while in hot conditions the reduction is from 57.2 to 48.9 kPa. The implication of this is that the expansion of bubble volume is only about one-half as much in cold conditions as it is in hot operation. That is to say, the bubbles generated at the bottom of a hot-sparged tank will grow larger than those in a cold one while rising from the impeller discharge region to the free surface. It is possible that all the above results and analysis might be very different in liquids of higher vapor pressure or even with water at other temperatures or operating pressures because of associated changes in the likelihood of cavitation at given levels of local superheat and impeller submergence. Experiments with similar systems at different temperatures are being carried out and will form the basis of future publications. These mechanisms can partly explain why the gas holdup in a hotsparged system is smaller than that in cold operation.
R2 0.979 0.986 0.981 0.980 0.979 0.968
cold systems G ) G ) G ) G ) G ) G )
0.82PTm0.212VS0.569 0.81PTm0.192VS0.563 0.73PTm0.178VS0.553 0.69PTm0.169VS0.545 0.68PTm0.152VS0.555 0.71PTm0.132VS0.569
R2 0.980 0.992 0.993 0.989 0.992 0.995
4. Conclusions Polypropylene particles of 3.8 mm diameter and ∼1.8 mm height have been used in the present work. The gas holdup and shaft power draw in the presence of different concentrations of buoyant particles have been measured in a cold- and hot-sparged vessels of 0.476 m diameter with four heaters, four baffles, and a dished base. The volume is 0.145 m3. The absolute value of the gassed-to-ungassed power demand (RPD) is generally higher in hot operation than in systems working at room temperature. With volumetric concentrations of buoyant particles in the range up to 15%, there is little change in the relative power demand, RPD, in a hot-sparged system. As the volumetric solid concentration Cv of buoyant particles is increased, the gas holdup in a hot-sparged system tends to rise, which is opposite to the trend at room temperature. The increase in gas holdup with increasing Cv becomes weaker as more gas is introduced into the system. Moreover, the correlations of gas holdup in a hot-sparged system show that the exponents for both power input and gas superficial velocity are reduced as Cv is increased. This means that, at higher solid concentrations, there is less influence of power input and gas rate on the gas holdup. The correlation G ) 0.36PTm0.15VS0.50‚ (1 + Cv)0.6 is suggested as representing the influence of power input, gas superficial velocity, and solid concentration in a hotsparged gas-liquid-buoyant particle system. Experiments carried out in a tank with a multiple-impeller agitator confirm the rapidity of mass transfer between the surrounding liquid and the air and the consequent internal saturation of the sparged air. When allowing for the contribution of the vapor in the hot-sparged system, at elevated temperatures (80.5-82.5 °C), the gas holdup is reduced to between twothirds and one-half of that at room-temperature values for a given total gas flow rate and power input. The changes of physical properties of gas and liquid, including viscosity, density, and interfacial tension, the different bubble sizes, and the different bubble dynamics in hot and cold liquids offer possible reasons for these differences, though the influence of liquid viscosity is believed to be decisive. Acknowledgment The authors sincerely acknowledge the financial support of National Nature Science Foundation of China (No. 20576009). Nomenclature Cv ) volumetric solid concentration, m3‚(m3 suspension)-1 D ) diameter of impeller, m Flg ) gas flow number Flg ) Qg/ND3 for cold-sparged system Flg ) Qg+v/ND3 for hot-sparged system Fr ) Froude number, Fr ) N2D/g H ) height of liquid in tank (general), m H0 ) height of the liquid without gas input, m HG ) height of the liquid during dispersion, m HS ) height from the sparger to the free surface, m
Ind. Eng. Chem. Res., Vol. 46, No. 20, 2007 6611
NP ) power number, NP ) P/FLN3D5 NPg ) aerated power number Pm ) specific shaft energy dissipation rate, W‚kg-1 PEm ) specific potential energy dissipation rate, W‚kg-1 PTm ) mean total specific energy dissipation rate, W‚kg-1 Qg ) inlet gas flow rate, m3‚s-1 Qg+v ) total gas flow rate including both air and vapor, m3‚s-1 R2 ) correlation coefficients T ) diameter of the tank, m T0 ) ambient temperature of the air, °C T* ) equilibrium temperature in hot-sparged system, °C V ) volume of the liquid in the tank, m3 VS ) superficial gas velocity, m‚s-1 G ) gas holdup AbbreViations HEDT ) half-elliptical hollow-blade disc turbine WHU ) wide-blade, up-pumping hydrofoil PP ) polypropylene Literature Cited (1) Bao, Y.; Hao, Z.; Gao, Z.; Shi, L.; Smith, J. M. Suspension of buoyant particles in a three phase stirred tank. Chem. Eng. Sci. 2005, 60, 2283-2292. (2) Smith, J. M.; Gao, Z. Power demand of gas dispersing impellers under high load conditions. Trans. Inst. Chem. Eng. 2001, 79A, 575-580. (3) Zhao, D.; Gao, Z.; Mu¨ller-Steinhagen, H.; Smith, J. M. Liquid-phase mixing times in sparged and boiling agitated reactors with high gas loading. Ind. Eng. Chem. Res. 2001, 40, 1482-1487. (4) Gao, Z.; Smith, J. M.; Mu¨ller-Steinhagen, H. Gas Dispersion in Sparged and Boiling Reactors. Chem. Eng. Res. Des. 2001, 79, 973-978. (5) Smith, J. M.; Gao, Z.; Mu¨ller-Steinhagen, H. The effect of temperature on the void fraction in gas-liquid reactors. Exp. Therm. Fluid Sci. 2004, 28, 473-478. (6) Gao, Z.; Smith, J. M.; Mu¨ller-Steinhagen, H. Void fraction distribution in sparged and boiling reactors with modern impeller configuration. Chem. Eng. Process. 2001, 40, 489-497. (7) Nienow, A. W.; Konno, M.; Bujalski, W. Studies on three-phase mixing: A review and recent results. Chem. Eng. Res. Des. 1986, 64, 3542.
(8) Bao, Y.; Hao, Z.; Gao, Z.; Shi, L.; Smith, J. M.; Thorpe, R. B. Gas dispersion and solid suspension in a three-phase stirred tank with multiple impellers. Chem. Eng. Commun. 2006, 193, 801-825. (9) Bao, Y.; Gao, Z.; Hao, Z.; Long, J.; Shi. L.; Smith, J. M.; Kirkby, N. F. Effects of equipment and process variables on the suspension of buoyant particles in gas-sparged vessels. Ind. Eng. Chem. Res. 2005, 44, 7899-7906. (10) Nocentini, M.; Pinelli, D.; Magelli, F. Dispersion coefficient and settling velocity of the solids in agitated slurry reactors stirred with multiple Rushton turbines. Chem. Eng. Sci. 2002, 57, 1877-1884. (11) Dohi, N.; Matsuda, Y.; Itano, N.; Minekawa, K.; Takahashi, T. Suspension of solid particles in multi-impeller three-phase stirred tank reacotors.. Can. J. Chem. Eng. 2001, 79, 107-111. (12) Moucha, T.; Linek, V.; Prokopova´, E. Gas hold-up, mixing time and gas-liquid volumetric mass transfer coefficient of various multipleimpeller configurations: Rushton turbine, pitched blade and techmix impeller and their combinations. Chem. Eng. Sci. 2003, 58, 18391846. (13) Hari-Parjitno, D.; Mishra, V. P.; Takenaka, K.; Bujalski, W.; Nienow, A. W.; Mckemmie, J. Gas-liquid mixing studies with multiple up- and down-pumping hydrofoil impellers: Power characteristics and mixing time. Can. J. Chem. Eng. 1998, 76, 1056-1068. (14) Dohi, N.; Matsuda, Y.; Shimizu, K.; Minekawa, K.; Kawase, Y. An experimental investigation into vapour dispersion and solid suspension in boiling stirred tank reactors. Chem. Engi. Process. 2002, 41, 267-279. (15) Wang, K.; Feng, L. F. Mixing Equipment Design; Machine Press: Beijing, China, 2000; p 294. (16) Himmelblau, D. M. Basic principles and calculations in chemical engineering, 6th ed.; Prentice Hall International, Inc.: Upper Saddle River, NJ, 1996; p 669. (17) Vlaev, S. D.; Valeva, M. D.; Mann, R. Some effects of rheology on the spatial distribution of gas hold-up in a mechanically agitated vessel. Chem. Eng. J. 2002, 87, 21-30. (18) Paglianti, A.; Takenaka, K.; Bujalski, W.; Takahashi, K. Estimation of gas hold-up in aerated vessels. Can. J. Chem. Eng. 2000, 78, 386-392.
ReceiVed for reView April 3, 2007 ReVised manuscript receiVed July 20, 2007 Accepted July 20, 2007 IE070481W
