3590
Ind. Eng. Chem. Res. 1996, 35, 3590-3596
SEPARATIONS Characteristics of a Rotating Packed Bed Hwai-Shen Liu,* Chia-Chang Lin, Sheng-Chi Wu, and Hsien-Wen Hsu† Department of Chemical Engineering, National Taiwan University, Taipei, Taiwan, ROC
The mass transfer (stripping of ethanol) and pressure drop (water-air) of a rotating packed bed were investigated under 13-273 equiv of gravitational force with two packings. The results indicated that the mass-transfer coefficient (KGa) was enhanced due to the centrifugal force as compared to a conventional packed bed. An empirical correlation was also proposed, suggesting that the KGa value depends on the GrG number to the power of 0.25. As to pressure drop, the gas rate seems to be a more important factor than the liquid rate because of less liquid holdup under centrifugal force. A semiempirical equation was also developed to correlate the pressure drop data with good agreement. Introduction For the sake of enhancing mass transfer, Ramshaw and Mallinson (1981) have developed the rotating packed bed which replaces gravity with centrifugal force and named it “Higee” (for high gravity). This device is applicable to absorption, distillation, stripping, extraction, and other separation processes. Because of the enhancement of mass transfer, the size of the equipment can be greatly reduced as compared with the conventional packed bed. Therefore, the rotating packed bed is characterized as easy operation and high efficiency. On the theoretical side, Tung and Mah (1985) applied penetration theory to derive a formula for the liquidfilm gas-liquid mass-transfer coefficient in the absorption process and also combined an empirical correlation of the interfacial area in conventional packed bed to describe the behavior of liquid mass transfer in the Higee process. They found fair agreement with the limited data reported by Ramshaw and Mallinson (1981). Munjal et al. (1989a,b) used the chemical absorption to obtain the data for the liquid-film gasliquid mass-transfer coefficient and gas-liquid interfacial area. They also proposed a model to describe these data reasonably well. Keyvani and Gardner (1989) published experimental results for the gas-liquid mass-transfer rate, pressure drop, residence time distribution, and power consumption. Kumar and Rao (1990) reported a model to describe their pressure drop data. Further, they used the model of Tung and Mah (1985) to predict the liquid-film gas-liquid masstransfer coefficient for chemical absorption and found poor agreement with their experimental data. Singh et al. (1992) also investigated this system to remove the volatile organic compounds (VOCs) from groundwater and presented the results which included pressure drop, mass transfer, power consumption, and fouling of the packing. Basˇic and Dudukovic (1995) determined the liquid holdup and the degree of anisotropy of liquid distribution using conductance measurement. It indicated that the liquid flow in this system cannot be * Corresponding author. Fax: 886-2-3623040. E-mail:
[email protected]. † Present address: Department of Chemical Engineering, National University of Singapore, Singapore.
S0888-5885(96)00183-2 CCC: $12.00
Figure 1. Schematic diagram of the rotating packed bed.
interpreted by the classical theory of film flow on the particle scale. Lockett (1995) examined this system concerning its flooding behavior, and Wallis’ equation can correlate their results well. An extensive review of conventional packed beds can be found by Charpentier (1976). However, very few experimental results for gas mass transfer of the Higee system have been published to date. Our main goal is to study the effect of the operating parameters on the overall volumetric gas-film gas-liquid mass-transfer coefficients. These results are also compared with those from gravity flow. We also provide an empirical correlation for gas-film gas-liquid mass-transfer coefficients (KGa). Moreover, the pressure drop of the rotating packed bed was investigated and correlated with modified theory for gravity flow. Beside these, two packings are investigated side by side to realize the effect of the packing on both the mass transfer and pressure drop. Experimental Setup and Procedure Figure 1 shows a simplified schematic diagram of a rotating packed bed which consists of a rotor and a stationary housing. Liquid flows outward from the inner edge of the rotor due to the centrifugal force. Gas flows inward countercurrently from the outer edge of the packed bed by the pressure-driving force. For visual observation, the system is made of transparent acrylic. The packing is arranged randomly in the rotor. Two shapes of plastic grains are used as packings in this system. The first one is rectangular © 1996 American Chemical Society
Ind. Eng. Chem. Res., Vol. 35, No. 10, 1996 3591
Figure 2. Effect of rotor speed on pressure drop for the rectangular packing.
with a size of 5 × 5 × 2.8 mm3, specific surface area of 524 m2/m3, and voidage of 0.533. The second one is elliptic cylindrical packing with a size of 3 × 2.6 × 3 mm3, specific surface area of 1027 m2/m3, and voidage of 0.389. The axial height of the packed bed is 2 cm. The inner radius and the outer radius of the packed bed are 4.5 and 7 cm, respectively. As a result, the depth (radial height) of the packed bed is 2.5 cm. In this system, the packed bed can be operated from 400 to 2500 rpm, which provides gravitational force of 10g to 402g based on the arithmetic mean radius. At ambient conditions, the pressure drop of gas across the bed is measured with a manometer for various operating variables with the air-water system. The overall volumetric gas-film mass-transfer coefficient (KGa) is determined by stripping ethanol solution with air for various operating conditions. The ethanol concentration was measured by gas chromatography (Perkin-Elmer autosystem) with a FID and Supelco SPB-1 column. Helium (20 mL/min) was used as the carrier gas with oven, injector, and detector temperatures of 85, 250, and 300 °C, respectively. Results and Discussion Pressure Drop. The effect of the rotor speed on the pressure drop is shown in Figure 2 for the rectangular packing. When there is no liquid in the rotor (dry bed), the relationship between the pressure drop and the rotor speed is somewhat linear for different gas flow rates. However, the pressure drop initially decreases with the rotor speed when liquid is introduced to the rotor. After the rotor speed exceeds about 1295 rpm, the pressure drop maintains a constant. This indicates that high rotor speed may reduce liquid holdup to make gas pass easily. At high rotor speed and low liquid flow rate, the pressure drop of the wet bed can be lower than that of the dry one. This phenomenon becomes more obvious at higher gas flow rates. This trend is quite different from the conventional packed bed, where the wet bed gives a higher pressure drop. The reason for this phenomenon may be that liquid, in some way, assists gas to flow through the bed more easily due to the centrifugal force. As a result, the pressure drop is reduced. However, the pressure drop does not vary with rotor speed significantly with the elliptic cylindrical packing, as shown in Figure 3. This may imply that
Figure 3. Effect of rotor speed on pressure drop for the elliptic cylindrical packing.
Figure 4. Effect of gas flow rate on pressure drop for the rectangular packing.
the elliptic cylindrical packing would hold less liquid under the centrifugal field within the experimental conditions. Therefore, the pressure drop is not influenced by the liquid flow rate. As shown in Figure 4, the pressure drop is measured as a function of gas flow rate at a constant rotor speed (870 rpm) and different liquid flow rates for the rectangular packing. It is not surprising that the pressure drop is strongly dependent on the gas flow rate. At higher gas flow rates, the pressure drop is slightly affected by the liquid flow rate since gas is the dominant factor. Also from Figure 4, it can be seen that the pressure drop is not affected as much by the liquid rate as by the gas rate. This phenomenon is more significant for the elliptic cylindrical packing, as shown in Figure 5, since it has less liquid holdup as mentioned previously. To correlate the pressure drop of the rotating packed bed, we may modify the theoretical equation (Billet and Schultes, 1991) which was derived from the shear force/ pressure equilibrium and the Newtonian friction law for the conventional packed bed:
1 at ∆Pd ) Fv2 3 fd(r2 - r1) 2
(1)
3592 Ind. Eng. Chem. Res., Vol. 35, No. 10, 1996
Figure 5. Effect of gas flow rate on pressure drop for the elliptic cylindrical packing.
where ∆Pd is the pressure drop of the dry rotating packed bed, at is the total surface area of packings per unit packed volume, is the void fraction, the gas capacity factor Fv is defined as the product of average superficial gas velocity and the square root of gas density, and fd is the resistance coefficient which depends on the gas flow rate and the rotor speed and must be determined from the experimental data. Because the cross-sectional area in which fluid flows is not a constant, the average of superficial velocity is defined as
vavg )
Q dr ∫rr 2πrz
1 r2 - r1
2
(2)
1
where Q is the volumetric flow rate. If Q represents the flow rate of liquid, vavg is the average of the superficial liquid velocity and denoted as vL. In a similar way, vG is the average of the superficial gas velocity. With the experimental results and regression analysis, the correlations of fd can be obtained as follows:
Figure 6. Comparison of experimental and calculated pressure drop.
∆Pw )
at Fv2 f (r - r1) 2 ( - h )3 w 2 L
where ∆Pw is the pressure drop of the wet rotating packed bed; fw is the wetted resistance coefficient which depends on the gas flow rate, the liquid flow rate, and the rotor speed; and hL is the liquid holdup which may be neglected as compared with . Using the experimental results and regression analysis, we can get the correlations of fw as follows:
for rectangular packing fw ) 2.928 × 1012ReGp-1.409ReLp0.596Reω-1.844 + 1518
fd ) 7.786ReGp1.404 + 10 582ReGp-0.776 + 3.233 × 10 ReGp -3
-1.519
when Reω < 2.29 × 104 (6)
fw ) 7142ReGp-2.044ReLp0.22Reω0.243 + 2085
for rectangular packing
(5)
when Reω > 2.29 × 104 (7)
for elliptic cylindrical packing Reω
1.637
(3)
for elliptic cylindrical packing fd ) 19.14ReGp1.092 + 1529ReGp-1.016 + 1.227 × 10-4ReGp-1.618Reω1.605 (4) ReGp denotes the gas Reynolds number (4RhvG/νG) with Rh denoting hydraulic radius which is equal to /at, and Reω denotes the rotational Reynolds number (ωravg2/νG). Both correlation coefficients from the regression analysis are above 0.99. The pressure drop in the dry rotating packed bed for two packings can be predicted well by these correlations, as shown in Figure 6. For the wet rotating packed bed, part of void space where gas flows through is occupied by liquid. Thus, the effective void fraction should be equal to the void fraction of the dry bed deducted by the liquid holdup. Hence, the pressure drop can be further modified as
fw ) 72.93ReGp-1.143ReLp0.042Reω0.32 + 97.84ReGp0.536 (8) where ReL denotes the liquid Reynolds number (4RhvL/ νL). All correlation coefficients from the regression analysis are above 0.97. Figure 6 shows the comparison between the calculated and experimental ∆Pw values. It can be seen that the experimental results lie within 20% of the values calculated by these correlations. Less accuracy of prediction for the wet rotating packed bed may attribute to the ignorance of hL. As a result, these empirical equations can be used to represent the behavior of a pressure drop in the rotating packed bed. Mass Transfer. Similar to the design equation drived by Singh et al. (1992) for liquid mass transfer in the rotating packed bed, the design equation for gas mass transfer can be derived using mass balance and the transfer unit concept:
Gm zKGapt
[
]
Ind. Eng. Chem. Res., Vol. 35, No. 10, 1996 3593
Y1 - HcX2 ln (1 - S) +S Y2 - HcX2 ) π(r22 - r12) (9) S-1
where Gm is the molar flow rate of gas, Pt is the total pressure in the system, z is the axial height of the packing, Hc is the Henry’s law constant expressed as the ratio of the mole fraction in the gas phase to that in the liquid phase, and Y1 and Y2 are the mole fraction of ethanol in the outlet and inlet air streams, respectively. Because the inlet air stream is free of ethanol, Y2 is taken as zero. X2 is the mole fraction of ethanol in the outlet water stream; r1 and r2 are the inner radius and outer radius of packing respectively. The stripping factor S is defined as HcGm/Lm, where Lm is the molar flow rate of liquid. In eq 9, the first term on the left-hand side (Gm/ zKGaPt) is defined as ATUG (the area of transfer units on the gas side), and the remainder of the left-hand side denotes NTUG (the number of transfer units on the gas side). The Henry’s law of ethanol is required for evaluation of KGa and can be found from Liu and Hsu (1991). Therefore, KGa may be determined from the experimental data with Y1 and X2. The effect of the rotor speed on KGa at a fixed gas mass flux is shown in Figure 7. For both packings, KGa increases with an increase in the rotor speed for both liquid mass fluxes with a similar trend. This suggests that the liquid mass flux does not affect the influence of the centrifugal force on KGa. The difference between two packings notes that the liquid mass flux does not affect KGa for the rectangular packing as much as for the elliptic cylindrical packing. This may also be due to less liquid holdup for the elliptic cylindrical packing as discussed in the section of pressure drop. Liquid holdup affects KGa because of its effect on the gas-liquid interfacial area. Consequently, the elliptic cylindrical packing does not give much difference in KGa for both liquid mass fluxes. Figure 8 shows the effect of the rotor speed on KGa at a fixed liquid mass flux. From Figure 8, it can be seen that the effect of rotor speed on KGa of high gas flux is more obvious than that of low gas flux for the rectangular packing. However, for the elliptic cylindrical packing, the effect of the rotor speed on KGa is quite similar for both gas fluxes. It can also be noted that the rectangular packing gives higher KGa than the elliptic cylindrical packing at high gas flux. This suggests that the rectangular packing possesses a better degree of gas-liquid contact than the elliptic cylindrical packing at high gas flux. Figure 9 shows KGa as a function of gas mass flux at a constant liquid mass flux. For the rectangular packing, KGa increases as the gas mass flux increases for both rotor speeds with different trends. It can be noted that, for the rotating packed bed, the effect of gas flux on KGa varies for different rotor speeds. Therefore, the rotating packed bed can be operated at an appreciable rotor speed to enhance mass transfer. From Figure 9, it can be noted that, for the elliptic cylindricalpacking, KGa is not so strongly dependent on gas mass flux as the rectangular packing. This phenomenon may be caused by the good gas distribution of the rectangular packing. As shown in Figure 10, KGa is measured as a function of liquid mass flux at a constant gas mass flux. For the rectangular packing, KGa increases with the liquid mass flux for both rotor speeds with a similar trend. This
Figure 7. Variation of KGa with rotor speed at a constant gas mass flux.
Figure 8. Variation of KGa with rotor speed at a constant liquid mass flux.
Figure 9. Variation of KGa with gas mass flux.
indicates that the rotor speed slightly affects the influence of the liquid mass flux on KGa. This characteristic is very much different from the elliptic cylindrical packing. For the elliptic cylindrical packing, KGa does not depend on the liquid mass flux at lower rotor speed.
3594 Ind. Eng. Chem. Res., Vol. 35, No. 10, 1996
liquid-film mass-transfer coefficient (kL) is calculated by the model of Tung and Mah (1985):
()
kLdp 2 × 31/3 at ScL1/2ReLa1/3 ) DL π a
1/3
GrL1/6
(11)
The gas-film mass-transfer coefficient (kG) is estimated by the following two correlations from Onda et al. (1968) and Shulman et al. (1995), respectively:
kG ) 5.23ReGa0.7ScG1/3(atdp)-2 atDG
(12)
kGdp ) 1.195(1 - )0.36ReGd0.64ScG0.33 DG
(13)
The gas-liquid interfacial area can also be evaluated by the two following correlations from Onda et al. (1968) and Puranik and Vogelpohl (1974), respectively:
Figure 10. Variation of KGa with liquid mass flux.
(
()
σc a ) 1 - exp -1.45 at σ
0.75
ReLa0.1FrL-0.05WeL0.2
()
a σ ) 1.045ReLa0.041WeL0.133 at σc
)
(14)
-0.182
- 0.229 + 0.091 ln(WeL/FrL) (15)
Figure 11. Comparison of rotating packed bed and conventional packed bed for mass-transfer performance.
However, at higher rotor speed, KGa shows a minimum within the experimental conditions. For comparison, the KGa of the conventional packed bed from Houston and Walker (1950) is shown in Figure 11 along with some data of this investigation. The conventional packed bed of Houston and Walker (1950) is 1.94 ft in height and 12 in. in diameter and deals with the absorption of ethanol from air by water. It is noted from Figure 11 that, although the rotating packed bed is operated at lower gas mass flux, KGa is higher than that of the conventional packed bed for liquid mass flux. It can be realized that the rotating packed bed at lower gas mass flux can provide the same KGa of the conventional packed bed of higher gas rate. That is to say, the rotating packed bed has better mass transfer than the conventional packed bed. We attempt to predict the KGa for the rotating packed bed using existing correlations; it is convenient to use the two-film theory:
1 H 1 ) + KGa kGa kLa
(10)
where H is the Henry’s law constant expressed as the ratio of the partial pressure in the gas phase to the mass concentration in the liquid phase. In this equation, the
As a result, the value of KGa can be obtained by eq 10 with eqs 11-15. Unfortunately, all four combinations underpredict KGa for the rotating packed bed, as shown in Figure 12, since these estimations of kG do not include the effect of the centrifugal force. Since the above existing correlations are unable to predict KGa for the rotating packed bed, a correlation is developed by assuming that KGa is dependent on the following parameters: gas mass flux, gas density, gas viscosity, liquid mass flux, liquid density, liquid viscosity, total specific surface area of packings, diameter of packings, and centrifugal acceleration. The following equation is obtained from the regression of experimental data:
KGaRT DGat2
) 3.111 × 10-3ReGa1.163ReLa0.631GrG0.25 (16)
As shown in Figure 13, the experimental results lie within (30% of the values estimated by eq 16. The ranges of the dimensionless groups in this correlation are 0.208 < KGaRT/DGat2 < 2.997, 1.82 < ReGa < 8.211, 1.341 < ReLa < 6.083, 3.41 × 104 < GrG < 1.70 × 106. It is found that KGa varies with the centrifugal force to the 0.25 power, which is within the range estimated by Ramshaw and Mallinson (1981). This correlation can be more general by introducing the Schmidt number. However, it is not included since only one solute was involved in this study. Conclusion This study investigated the characteristics of a rotating packed bed, including pressure drop and mass transfer with the comparison of two packings. For both shapes of packings, the results indicated that pressure drop is highly dependent on the gas flow rate. At high rotor speed, due to low liquid holdup, the pressure drop is affected strongly by the gas flow rate and slightly by
Ind. Eng. Chem. Res., Vol. 35, No. 10, 1996 3595
relations for the pressure drop in the rotating packed bed were also developed. The mass-transfer coefficients in the rotating packed were presented for two packings of various operating conditions. The results showed that KGa increases with the gas rate, the liquid rate, and the rotor speed. Moreover, different shapes of packings give different mass-transfer performances in the rotating packed bed. Based on these results, it indicates that kG is influenced by the centrifugal force to some extent. Therefore, the rotating packed bed could provide better mass transfer than the conventional packed bed due to the effect of the centrifugal force. A reasonable correlation for KGa in the rotating packed bed was also presented. Acknowledgment The financial support from National Science Council is greatly acknowledged. Nomenclature
Figure 12. Comparison of experimental and predicted KGa using existing correlations.
Figure 13. Comparison of experimental and calculated KGa by eq 16.
the liquid flow rate. With low rotor speed, the liquid flow rate affects the pressure drop more strongly for the rectangular packing than for the elliptic cylindrical packing. This is because the liquid holdup for the elliptic cylindrical packing is lower than that for the rectangular packing. It can be realized that different shapes of packings would provide different flow patterns in the rotating packed bed. An unusual result is also noted that the pressure drop is lower for the wet bed than the dry one. The satisfactory semiempirical cor-
a ) gas-liquid interfacial area (m2/m3) ac ) centrifugal acceleration (m/s2) at ) total specific surface area of the packing (m2/m3) DG ) gas diffusivity (m2/s) DL ) liquid diffusivity (m2/s) Fv ) gas capacity factor (kg1/2/(m1/2 s)) fd ) resistance coefficient fw ) wetted resistance coefficient G ) volumetric flow rate of gas (m3/s) G′ ) gas mass flux (kg/(m2 s)) Gm ) molar flow rate of gas (mol/s) g ) gravitational acceleration (m/s2) H ) Henry’s law constant, eq 10, (atm‚m3/mol) Hc ) Henry’s law constant, eq 9 hL ) liquid holdup KG ) overall gas-film mass-transfer coefficient (mol/ (atm‚m2‚s)) KGa ) overall volumetric gas-film mass-transfer coefficient (mol/(atm‚m3‚s)) kG ) gas-film mass-transfer coefficient (mol/(atm‚m2‚s)) kL ) liquid-film mass-transfer coefficient (m/s) L ) volumetric flow rate of liquid (m3/s) L′ ) liquid mass flux (kg/(m2 s)) Lm ) molar flow rate of liquid (mol/s) ∆P ) pressure drop (Pa) Pt ) total pressure in the system (atm) Rh ) hydraulic radius (m) Q ) volumetric flow rate (m3/s) r ) radius of the packed bed (m) r1 ) inner radius of the packed bed (m) r2 ) outer radius of the packed bed (m) S ) stripping factor v ) superficial fluid velocity (m/s) Y1 ) mole fraction of a solute in the outlet air stream Y2 ) mole fraction of a solute in the inlet air stream X2 ) mole fraction of a solute in the outlet water stream z ) axial height of the packing (m) Greek Letters ) voidage (m3/m3) µG ) gas viscosity (kg/ms) µL ) liquid viscosity (kg/ms) FL ) liquid density (kg/m3) σ ) liquid surface tension (N/m) σc ) critical surface tension (N/m) νG ) dynamic gas viscosity (m2/s) νL ) dynamic liquid viscosity (m2/s) ω ) angular velocity (rad/s)
3596 Ind. Eng. Chem. Res., Vol. 35, No. 10, 1996 Dimensionless Groups ((L′)2a
FrL ) Froude number t/FLg) GrG ) gas Grashof number (dp2ac/νG2) GrL ) liquid Grashof number (dp2ac/νL2) ReGa ) gas Reynolds number for mass transfer (G′/atµG) ReGd ) gas Reynolds number for mass transfer (G′dp/µG) ReGp ) gas Reynolds number for pressure drop (4RhvG/νG) ReLa ) liquid Reynolds number for mass transfer (L′/atµL) ReLp ) liquid Reynolds number for pressure drop (4RhvL/ νL) Reω ) rotational Reynolds number (ωravg2/νG) ScG ) gas Schmidt number (νG/DG) ScL ) liquid Schmidt number (νL/DL) WeL ) Weber number ((L′)2/FLatσ)
Literature Cited Basˇic, A.; Dudukovic, M. P. Liquid Holdup in Rotating Packed Beds: Examination of the Film Flow Assumption. AIChE J. 1995, 41, 301. Billet, R.; Schultes, M. Modeling of Pressure Drop in Packed Columns. Chem. Eng. Technol. 1991, 14, 89. Charpentier, J. C. Recent Progress in Two Phase Gas-Liquid Mass Transfer in Packed bed. Chem. Eng. J. 1976, 11, 161. Houston, R. W.; Walker, C. A. Absorption in Packed Towers-Effect of Molecular Diffusivity on Gas Film Coefficient. Ind. Eng. Chem. 1950, 42, 1105. Keyvani, M.; Gardner, N. C. Operating Characteristics of Rotating Beds. Chem. Eng. Prog. 1989, 85, 48. Kumar, P. M.; Rao, D. P. Studies on a High-Gravity Gas-Liquid Contactor. Ind. Eng. Chem. Res. 1990, 29, 917. Liu, H. S.; Hsu, H. W. Gas Stripping during Ethanol Fermentation in a Batch Reactor. CIChE J. 1991, 22, 357. Lockett, M. J. Flooding of Rotating Structured Packing and Its Application to Conventional Packed Columns. Trans. Ind. Chem. Eng. 1995, 73, 379.
Munjal, S.; Dudukovic, M. P.; Ramachandran, P. A. Mass Transfer in Rotating Packed Beds: I. Development of Gas-Liquid and Liquid-Solid Mass-Transfer Coefficients. Chem. Eng. Sci. 1989a, 44, 2245. Munjal, S.; Dudukovic, M. P.; Ramachandran, P. A. Mass Transfer in Rotating Packed Beds: II. Experimental Results and Comparison with Theory and Gravity Flow. Chem. Eng. Sci. 1989b, 44, 2257. Onda, K.; Takeuchi, H.; Okumoto, Y. Mass Transfer Coefficients between Gas and Liquid Phases in Packed Columns. J. Chem. Eng. Jpn. 1968, 1, 56. Puranik, S. S.; Vogelpohl, A. Effective Interfacial Area in Irrigated Packed Columns. Chem. Eng. Sci. 1974, 29, 501. Ramshaw, C.; Mallinson, R. H. Mass Transfer Process. U.S. Patent 4,383,255, 1981. Shulman, H. L.; Ullrich, C. F.; Proulx, A. Z.; Zimmerman, J. O. Performance of Packed Column II. Wetted and Effective Interfacial Areas, Gas- and Liquid-Phase Mass Transfer Rates. AIChE J. 1955, 1, 253. Singh, S. P.; Wilson, J. H.; Counce, R. M.; Villiers-Fisher, J. F.; Jennings, H. L.; Lucero, A. J.; Reed, G. D.; Ashworth, R. A.; Elliott, M. G. Removal of Volatile Organic Compounds from Groundwater Using a Rotary air Stripper. Ind. Eng. Chem. Res. 1992, 31, 574. Tung, H. H.; Mah, R. S. H. Modeling of Liquid Mass Transfer in HIGEE Separation Process. Chem. Eng. Commun. 1985, 39, 147.
Received for review March 29, 1996 Revised manuscript received May 31, 1996 Accepted June 2, 1996X IE960183R
X Abstract published in Advance ACS Abstracts, August 15, 1996.
