Hydrodynamic and Mass Transfer Characteristics of Single and

Ind. Eng. Chem. Res. 2008, 47, 2829-2841

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Hydrodynamic and Mass Transfer Characteristics of Single and Multiple Impeller Hollow Self-Inducing Reactors Rupesh B. Kasundra, Anand V. Kulkarni, and Jyeshtharaj B. Joshi* Institute of Chemical Technology, UniVersity of Mumbai, Matunga, Mumbai-400 019, India

In the present work, hydrodynamics and mass transfer characteristics of a gas-liquid stirred tank provided with hollow self-inducing impellers were investigated. Critical impeller speed for the onset of gas induction, gas induction rate, power consumption, and mass transfer coefficient (kLa) were measured using two types of self-inducing impellers, viz., hollow pitched-blade downflow turbine and double-disk impeller. Experiments were carried out in 100 and 800 L tanks. Mass transfer coefficient was measured with steady-state hydrazine oxidation technique with air-water as the gas-liquid system. The impeller speed was varied from 2.5 to 14 rps, and power input per unit volume was varied from 0.5 to 10 kW/m3. Further, the performance of a multiimpeller hollow self-inducing system has been investigated. Four different hollow impellers in combination with three axial flow impellers have been studied in an 800 L tank using three impeller diameters and three liquid submergences. Suitable correlations have been proposed that also take into account all the published kLa data in the literature. 1. Introduction Many industrial processes involve the dispersion of a gas phase into a liquid phase. This operation is often conducted in gas-liquid stirred tanks. In the conventional gas-liquid stirred tanks, the gas is introduced through a sparger located under the impeller. In many gas-liquid reactions, pure gas is used, and when the per pass gas-phase conversion is low, it is desirable to operate the reactor in dead-end mode or to recycle the unreacted gas. This condition arises in a number of industrially important reactions such as hydrogenation, alkylation, ethoxylation, ammonolysis, oxidation with pure oxygen, hydrochlorination, etc.1 A self-inducing reactor is such a system that recycles the gas from head space to the liquid phase without the use of an external compressor. This means not only savings but also the operation of the system is safer, reliable, and relatively maintenance-free. This reactor also finds application where the gas is available at relatively low pressure. It is well-known that the impeller performance per unit power consumption strongly depends upon the impeller design and the geometric configuration of the tank. It is desirable to know the dependence of the global parameters such as the gas holdup, the power consumption, and the volumetric gas-liquid mass transfer coefficient (kLa) on the impeller design, tank geometry, power consumption, and properties of the gas-liquid system. In large-scale industrial reactors, a single self-inducing impeller is often not enough. This is because, for maximization of the gas induction rate, the impeller needs to be located closer to the top gas-liquid interface. Such a location provides gasliquid dispersion in a limited volume around the impeller, and most of the reactor volume remains devoid of the gas phase. This limitation can be overcome by providing one more impeller (parts A and B of Figure 1), which disperses the induced gas by the self-inducing impeller. In addition to the good gas dispersion, often there is a need for solid suspension and/or good heat transfer. In such cases, relatively higher circulation velocity is required in the reactor. These requirements can also be met by the second impeller. Performance of such a reactor greatly * To whom correspondence should be addressed. Phone: 91-22414 5616. Fax: 91-22-414 5614. E-mail: [email protected].

depends upon the designs of self-inducing as well as the second impeller. Also, the geometric parameters like location of the self-inducing impeller from the liquid surface and the location of the lower impeller from the bottom are of great importance. In the present work, first of all, the literature pertaining to the mass transfer characteristics of single- and multi-impeller self-inducing systems has been analyzed. Second, the performance of some hollow self-inducing impellers was investigated by measuring the rate of gas induction, the mass transfer coefficient, and the power consumption. An attempt has been made to develop some new designs of self-inducing impeller or to make some modifications in the existing designs so as to achieve a higher rate of gas induction with good gas distribution in the reactor. After analyzing performance of these impellers at a small scale (100 L), experiments were performed in an 800 L tank with single- and multi-impeller systems. Correlations have been proposed for the rate of gas induction (QG), the power consumption (P), and the mass transfer coefficient (kLa), which are expected to be useful in practice. 2. Previous Work During the past four decades, several investigations on different types of self-inducing impellers have been reported in the literature.2 One of the important characteristics of the self-inducing impellers is the critical impeller speed for the onset of gas induction, which refers to the minimum speed (NC) at which it creates sufficient pressure driving force to overcome the static head above it. Several studies1,3-10 have reported the mechanism of gas induction and developed correlations11-13 for the estimation of NC. They have also made measurements of QG, P, and kLa for different types of self-inducing impellers, and correlations have been proposed for these parameters. Details regarding the geometry of their reactors, the impeller type, the system employed, the measurement method, and the correlation proposed and its range of applicability in terms of power input are given in Table 1. It was thought desirable to compare all the proposed correlations. Figure 2A shows kLa as a function of specific power consumption for an air-water-type system, whereas Figure 2B presents kLa versus PG/V for nonaqueous systems.

10.1021/ie071392m CCC: $40.75 © 2008 American Chemical Society Published on Web 03/15/2008

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Ind. Eng. Chem. Res., Vol. 47, No. 8, 2008

Figure 1. Schematic representation of self-inducing reactor. (A) Single self-inducing impeller system and (B) self-inducing + dispersing multi-impeller system: (1) tank, (2) mechanical seal, (3) shaft, (4) impeller, (5) top liquid surface, (6) inlet for gas induction, (7) outlet for induced gas, and (8) dispersing impeller.

Figure 2 shows a large variation of kLa values for a specific power consumption and impeller speed. Figure 2A shows that Joshi and Sharma,1 Sawant et al.,7 Smith et al.,22 and Topiwala and Hamer4 have measured kLa over a wide range of specific power consumption. A variation in kLa values is due to the difference in physicochemical properties of the liquid phase under consideration and the impeller design. In view of these observations that the kLa depends upon the impeller design, it was thought desirable to develop even more efficient impeller designs as compared with those reported in the published literature. 3. Experimental Section Experiments were carried out in an 100 L, 0.5 m i.d. tank and an 800 L, 1 m i.d. tank. Figure 3 shows the schematic diagram of experimental setup. Two different types of hollow self-inducing impellers, viz., hollow pitched-blade downflow (PBTD) turbine and self-inducing Rushton type (double-disk impeller) were studied (Figure 4). Further, two blade angles of 45° and 60° were investigated for hollow PBTD. The diameter of the PBTD impeller was 0.25 m, and that of the double-disk impeller was 0.17 m. The impeller was located at 0.17 m from the tank bottom, and the liquid height was 0.5 m in all the experiments in the 100 L tank. For experiments in the 1 m diameter tank with hollow PBTD, the liquid height in the tank was varied between 0.6 and 0.8 m with a constant impeller clearance (C) of 0.33 m. Experiments of the multi-impeller self-inducing system were performed in an 800 L, 1.0 m i.d. tank. Figure 5 shows different impellers used for the multi-impeller study. Table 2 summarizes dimensions of hollow impellers used for the multi-impeller study. Dimensions and measured power number for axial flow impellers used in combination with the hollow impeller are listed in Table 3. For comparison of performance of different impeller combinations, submergence of the self-inducing (GI) impeller was maintained constant at 0.3 m from the free liquid surface. Clearance of the lower axial flow impeller was maintained at 0.33 m from the bottom of the tank. To study the effect of submergence, liquid height in the tank was varied at three levels, viz., 1, 0.9, and 0.8 m. Two impeller diameters of 0.33 and 0.5 m were employed. In order to study the effect of lower impeller design, PBTU45° and PBTD45° impellers were used in

combination with a 0.5 m six-pipe impeller. Table 4 gives details of the geometrical parameters covered with each combination. Mass transfer measurements (kLa) were carried out using the steady-state hydrazine oxidation technique reported by Onken et al.23 and Marquez et al.24 During the experiment, hydrazine solution (NH2NH2‚H2O) was fed continuously to the tank using a peristaltic pump. The hydrazine flow rate was measured with 1% accuracy. The concentration was measured by titrating with standardized iodine solution. Dissolved oxygen concentration was measured with a dissolved oxygen meter having an accuracy of 1% saturation concentration. A dissolved oxygen probe was located 0.3 m from the bottom in the case of the 0.5 m diameter tank and 0.33 m from the bottom in the case of the 1 m diameter tank. During the experiments, the flow rate of hydrazine was adjusted to maintain 40-60% of saturation concentration of oxygen at steady state. The value of kLa was calculated by knowing the overall rate of oxidation and the steady-state dissolved oxygen concentration using the following equation:

RAa ) QB[Bo] ) kLaV([A*] - [Ao])

(1)

Impeller rotation speed was measured with an electronic rpm meter with a magnetic proximity probe for detecting shaft rotation. The speed can be read with the accuracy of 1 rpm. The critical impeller speed for gas induction was measured by visually observing the smallest speed when air bubbles were induced near the impeller. In the case of the hollow PBTD, the surface aeration also occurs with gas induction. For accurate measurement of the critical impeller speed, a horizontal baffle was introduced near the gas-liquid interface and the surface aeration was blocked. The rate of gas induction was measured by using a precalibrated turbine-type anemometer with an accuracy of 2%. The rotating shaft was made leakproof by providing a stuffing-boxlike arrangement with air inlet. Because of the intrinsic fluctuations with the gas induction process, each flow reading was taken 10 times and reported as an average value. The power consumption was measured by a torque table that was restrained from rotating by a string, and the force on the string was then measured by connecting it to a cantilever-type load cell. The precalibrated load indicator displayed the load on the torque table. Ten readings were taken for each set, and an average was used for the calculation of power.

Ind. Eng. Chem. Res., Vol. 47, No. 8, 2008 2831 Table 1. Reported Mass Transfer Studies on Self-Inducing Impellers experimental details

author

impeller type

tank diam, T, and impeller diam, D (m)

Topiwala & Hamer4

pipe impeller

T ) 0.158 D ) 0.075

Joshi & Sharma1

pipe impeller (2, 4, & 6 pipes) flattened cylindrical

T ) 0.41, 0.57, 1 D ) 0.2-0.5

Sawant et al.7

Denver

Matsumura et al.14

Baczkiewicz and Michalski15

system

mass transfer measurement method

air-aq K2SO4 (0.5 M) dynamic oxygen probe fermentation broth distilled water response method air-water chemical air-sodium dithionite CO2-Na2CO3 + air-water + isopropyl NaHCO3 alcohol air-water + DEG

correlation

( )

T ) 0.1-0.38 (sq) air-sodium dithionite chemical D ) 0.07-0.115 CO2-Na2CO3 + NaHCO3 + kL independent of PG/V tricrecyl phosphate µ -1/3 FD 2/3 two impellers T ) 0.19, 0.242, air-aq sodium sulfite chemical sulfite kLa ) 3.08 × 103 µ and draft tube 0.316 (0.5 kmol/m3) oxidation Fg2 D/T ) 0.23, 0.25, -0.84 0.4 -0.43 υ 3 3 NDµ NDF S 0.26 D/T ) 0.44, 0.45, σ µg ND 0.46 Q 0.59 d 2.17 0.46 hollow pipe T ) 0.44 air-water + dynamic oxygen kLa ) 0.17(n)1.68 (i) VL D (2, 3, & 6 D ) 0.154, 0.198, acetic acid + probe response pipes) 0.242 ethyl alcohol method i ) no. of paddles, n ) rps

( ) ( ) ( )( ) ( ) ( ) ()

Rushton-type hollow shaft

T ) 0.07 D ) 0.032

H2- adiponitrile + Raney nickel doped catalyst

dynamic pressure method

PG VL

0.48

k La ∝

PG VL

0.48

T ) 0.05 D ) 0.029

() () PG VL

PG/V ) 2-50 kLa ) 0.015-0.15

kLa ) 0.07-0.4

kLa ) 0.02-0.4

0.74

Q VL

(VS)0.74

PG VL

0.76

Q max VL

at

Sh ) BRe1.45Sc0.5We0.5 PG k La ∝ VL

PG/V ) 0.4-15 kLa ) 0.04-0.3

() () () () ()

kLa ) 4 × 10-4

1.5-2

PG at < 4 kW/m3 VL

0.6-0.8

at

PG/V ) 0.8-10 kLa ) 0.01-0.5 at 20 °C kLa ) 0.1-1.6 at 80 °C

PG > 4 kW/m3 VL

H2-2 propanol H2- o-cresol


Sh ) 0.123Re0.44Sc0.5We1.271.1 G 104 < Sh < 5 × 105 7 × 103 < Re < 13 × 104 500 < Sc < 900 180 < We < 550 1.2 < G < 1.7

no power measurement kLa ) 0.04-0.92

air-water air-biomass

dynamic oxygen probe response method

Sh ) 1 - exp(-K Rek1 Frk2) Sh∞

PG/V ) 0.16-2 kLa ) 0.008-0.043

O2-cyclohexane N2-cyclohexane


no power measurement kLa ) 0.003-0.08

T ) 0.45 D ) 0.154

air-water


Sh ) 4.51 × 103We-0.21Fr0.92 × (1 + 1.867 × 103G) 2 100 < We < 13 300 1 < Fr < 3 PG 0.475 kLa ) 0.01 (VG)0.4 VL

radial flow gas inducing

T ) 0.6 D ) 0.2

air-water

dynamic oxygen probe response method (at low kLa) sulfite oxidation (at high kLa)

stator-rotor type single and multiple Rushton-type disk impeller

T ) 1, 1.5 D ) T/5, T/3

air-water

T ) 0.064 D ) 0.032

Turboxal-2

T ) 0.75 (sq) D ) 0.08 & 0.105

Hichri et al.17

self-gasinducing agitator

Heim et al.18

Tekie et al.11

for pipe T ) 0.3 six pipe D ) 0.125 disk impeller D ) 0.125 D ) 0.1, 0.15 six flat-blade T ) 0.127 turbine D ) 0.0635

Forrester et al.19

hollow pipe (half cut)

Poncing et al.20

Patil et al.12

Sardeing et al.21

( ) ( )

PG 0.55 (VG)0.5 V for VG < 0.005 m/s PG 0.55 kLa ) 3.26 × 10-3 (VG)0.25 V for VG > 0.005 m/s PG 0.5 kLa ) 0.0195 V kLa ) 6.8 × 10-3

kLa ) 0.83

Zieverink et al.13

PG/V ) 1.5-8 kLa ) 0.01-0.1

kLa increases with PG/V

kLa ) 0.22

Dietrich et al.16

power input per unit volume, PG/V (kW/m3) and mass transfer coefficient, kLa (s-1)

()

0.065(Fr* - Fr/C)1.1

k La )

1 + 0.132(Fr* - Fr/C)1.1 for coalescing system) 0.35(Fr* - Fr/C)1.6 k La ) 1 + 0.2(Fr* - Fr/C)1.6 (for noncoalescing system)

( )


kLaV(υ/g2)1/3

H2-acetone H2-toluene H2-sunflower oil H2-n-hexadecane


air-water


Sh ) 10aReb (Fr* - Fr/C)cScm a ) -3.35 ((0.77) b ) 1.50 ((0.14) c ) 1.12 ((0.11) m ) 0.81 ((0.09) no correlation proposed

D

3

) A′FrB′

VA V

C′

PG/V ) 0.32- 1.7 kLa ) 0.006-0.055 PG/V ) 0.45-4 kLa ) 0.01-0.8

PG/V ) 0.075-0.525 kLa ) 0.0015-0.058

no power measurement

PG/V ) 0.171 kLa ) 0.0015-0.0026

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of the higher D/T ratio, the P/V value for the hollow PBTD is higher than that for the double-disk impeller. Figure 8A shows variation of ratio of gassed to ungassed power consumption with impeller speed. The reduction in gassed power consumption as compared to ungassed power consumption is attributed to (a) cavity formation behind the impeller blade and (b) reduction in average density in the impeller region, due to recirculation of gas phase. Figure 8A shows that reduction in power consumption for a self-inducing turbine impeller is higher as compared to that in the case of a pipe impeller. In the case of the modified double disk with hydrofoil, power consumption falls sharply with impeller speed. Almost the same trend was observed with a hollow PBTD45° with hydrofoil; however, the extent of reduction in power consumption was less than that for the modified double-disk impeller. For the self-inducing impeller, the ratio of PG/P has been correlated by the following equation,

PG ) 1 - c(N - Ncr)b PO

Figure 2. Volumetric mass transfer coefficient: comparison of correlations reported in the literature for (A) aqueous systems and (B) organic systems.

4. Results and Discussion 4.1. Power Consumption. The power consumption by an impeller under turbulent conditions is given by the following equation:

NP )

P N D5FL 3

(2)

Power number (NP) in the absence of gas remains constant and depends on the impeller design and the geometric parameters of the vessel and the internals. The values of power number for different individual impellers are given in Table 5 , and those for multi-impeller systems are given in Table 6. As can be seen from Table 5, hollow PBTD60° has a slightly higher power number than the double-disk turbine because of the higher impeller-to-tank diameter ratio (D/T). Figure 6 shows the variation in gassed power input with aeration number (QG/ND3) under self-induction conditions. For the disk impeller, the power number drops sharply at an aeration number of 0.06. However, in the case of the modified double-disk impeller, a sharp fall in the power number was observed at higher gas-induction rate corresponding to aeration number of 0.07. In the case of hollow PBTD, no such sharp fall in power number was observed; instead, the fall in power number was seen to be gradual. Figure 7 shows the variation of power with impeller speed. Because

(3)

where c and b are empirical constants and N and NC are impeller speed and critical impeller speed for gas induction in rps, respectively. Correlation constants for different impeller combinations are listed in Table 6 along with power number for each impeller. The proposed correlation correlates all data within (10% of the experimental value, and the correlation coefficient was found to be 0.94 for eq 3. It was observed that the modified double-disk impeller has the highest power number among all the impellers investigated, whereas the pipe impeller has the minimum power number. At this stage, it is important to understand the behavior of PG/Po versus N for the self-inducing impeller as against the conventional impeller where gas is sparged separately. For the latter case, PG/Po versus N has been plotted in Figure 8B where QG is a parameter. In Figure 8B, curves 1, 2, and 3 represent the case of the sparged contactor whereas curve 4 is for the self-inducing case. It will be worthwhile to revisit the explanation for the behavior of PG/P versus N at one gas flow rate. As the gas flow rate increases, the PG/P versus N follows the three lines L1Q1, L2Q2, and L3Q3. As the impeller speed increases along the curve L1N1, the flow generated by the impeller increases continuously and provides increasing resistance to the gas flow. The gas holdup increases steadily, and the reduction in the power number occurs along the curve L1N1. It shows that the impeller action is not strong enough to effect any noticeable change in the flow behavior of the gas phase. Since bubble breakup does not occur through the impeller action in this region, the impeller is said to be flooded with gas, and the region L1N1 is called the flooding region. Impeller rotation generates a low-pressure region behind the impeller blades. The extent of pressure reduction depends on the kinetic head [(πND)2/2]. At a certain impeller speed, the reduction in pressure is sufficient to hold the gas phase in the low-pressure region, against the buoyancy force. The gas accumulated behind the impeller is referred to as a cavity. Cavity formation starts somewhere at the point M1, and its size increases with increasing impeller speed along the curve M1N1. At point N1, the cavity size reaches its maximum. The flow pattern under these conditions is shown in Figure 9B. It is seen that the impeller action distributes the gas bubbles in the radial direction above the plane of the impeller. Because of cavity formation, the intensity of eddy motion behind the impeller

Ind. Eng. Chem. Res., Vol. 47, No. 8, 2008 2833

Figure 3. Schematic diagram of experimental setup: (1) inlet for gas induction, (2) anemometer (turbine type), (3) stuffing box, (4) Thyristor controlled dc motor; (5) dissolved oxygen probe, (6) peristaltic pump, (7) dissolved oxygen meter, (8) torque cell, (9) hydrazine solution tank, (10) screw jack, (11) torque indicator, and (12) anemometer indicator.

decreases and, hence, the rate of turbulent energy dissipation decreases. Therefore, the impeller power consumption decreases along the curve M1N1. It may be reminded that the impeller motion generates shear stresses in the liquid and shear stress increases with an increase in the impeller speed. Corresponding to point N1, the shear stress field generated by the impeller is sufficiently strong to break up the cavity. The flow pattern is shown in Figure 9C. At this stage, the impeller action starts contributing to the bubble size. The gas bubbles now occupy more liquid volume, and the gas phase extends up to the impeller plane. As the impeller speed increases along the curve N1P1, the cavity breakup increases continuously and is practically completed at point P1. As a result of continuous reduction of cavity size along the curve N1P1, the power number increases continuously. At impeller speeds beyond point N1, the gas bubbles penetrate into the region below the impeller plane, and at a certain critical speed (point O1), they even reach the vessel bottom (Figure 9D). Under these conditions, the impeller action predominates over the effect of gas sparging. This means that both the average bubble size and the liquid flow in the vessel are controlled by the impeller action. The speed at point O1 is referred to as the critical impeller speed for gas dispersion. At impeller speed beyond point P1, the liquid circulation generated by the impeller increases continuously and the bubbles become entrained in the downflow above the impeller. Initially, only small bubbles are entrained by this downflow. However, at sufficiently high impeller speeds, larger bubbles are also entrained. The corresponding flow pattern is shown in Figure 9E. Because of gas recirculation into the impeller region, the gas holdup increases continuously along the curve P1Q1. This results in a reduction of dispersion density in the impeller region,

and the power number decreases along the curve P1Q1. This region is called the recirculation region. From Figure 8B, it can be seen that the values of critical impeller speed for loading, dispersion, and recirculation increase with an increase in QG, for instance, (N3 > N2 > N1, O3 > O2 > O1, P3 > P2 > P1, etc.). This particular feature can be used to explain the PG/P versus N behavior for self-inducing impellers. For this case, QG increases with N and, hence, the values of N at loading, dispersion, and recirculation increase continuously with an increase in N. Such a possible increase in QG with N modifies the PG/P behavior (line 4 in Figure 8B), and it becomes flatter as compared to the cases (lines 1, 2, and 3) where the gas flow rate is constant. This explains the nature of PG/P versus N in Figure 8A. 4.2. Rate of Gas Induction. The most important variables influencing the rate of gas induction are the impeller design, the impeller diameter, the impeller speed, the liquid submergence, and the physical properties of the liquid. The results of QG as a function of impeller speed for different impellers under consideration are reported in Figure 10. The gas-induction characteristic of any impeller depends on two main factors, the pressure driving force created by the impeller and the dispersing ability of the impeller. The impeller having the higher power number tends to give more pressure driving force. Gas dispersing ability of the impeller depends upon its ability for cavity breakage and the immediate transport of gas bubbles from the impeller region. The later feature is useful for keeping higher mixture density in the vicinity of the impeller, which in turn is useful for efficient pressure reduction (which is proportional to the mixture density) as well as for efficient generation of liquid circulation. The rate of gas induction for hollow PBTD 45° and PBTD 60° increases uniformly with impeller speed; however, for the case of the double-disk impeller and the modified double-

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Figure 4. Hollow impellers used for single-impeller experiments: (A) hollow PBTD impeller; (B) double-disk impeller, schematic; (C) double-disk impeller, plan; (D) modified double-disk impeller, schematic; and (E) modified double-disk impeller, plan.

disk impeller, rise is not uniform. This may be due to the cavity formation at the back of impeller blade at impeller speed around 11 rps. When the gas-induction rate is plotted against power input per unit volume (Figure 11), it shows a uniform increase in induction rate with power consumption. For the better understanding of cavity buildup, QG/N2 is plotted against impeller speed for the case of the disk impeller. The reason for plotting QG/N2 is that the driving force for gas induction increases proportional to N2, and if other resistances (like frictional loss and orifice loss due to gas flow) are assumed to

be negligible, QG should approximately increase with N2. Figure 12 suggests the selection of operating speed (in the QG/N2 versus N plot) in the range of maximum gas dispersion condition. The gas induction rate for the multi-impeller system is plotted in Figure 13. As observed from Figure 13, impellers having higher power numbers give higher induction rates at equal diameter. Another factor that decides the rate of gas induction is the impeller’s gas-dispersing ability. The impeller that produces high liquid flow helps in propelling dispersed gas bubbles. Therefore, PBTD45° and the modified double-disk


Figure 5. Impellers used for multi-impeller experiments: (A) modified double-disk impeller, (B) hollow PBTD45° impeller, (C) six-pipe impeller, (D) modified four-pipe impeller, (E) hydrofoil impeller, 6-bladed, and (F) PBTD45° impeller. Table 2. Dimensions of Hollow Impeller Employed for Multi-Impeller Study diameter, D (mm)

impeller hollow PBTD45°

blade width/pipe diameter, (mm)

330 500 330 330 500 330

modified double disk six-pipe modified four-pipe

blade thickness (mm)

diameter and no. of holes for gas induction

60 66 66 33

15 (hollow) 15 (hollow) 3 N.A.

Ø6 mm × 6 holes on each blade Ø6 mm × 6 holes on each blade Ø10 mm × 6 holes radially on hub Ø10 mm × 6 holes radially on hub

60 pipe diam. 27 mm

4

Ø10 mm × 4 holes radially on hub

Table 3. Dimensions and Power Number of Axial Flow Impellers Used for Multi-Impeller Study impeller

diameter, blade width blade thickness measured D (mm) (mm) (mm) power number

six-blade hydrofoil PBTD45° PBTU45°

330 330 330

55 99 99

3 3 3

0.4 2 2.2

impeller produce a higher rate of gas induction vis-a`-vis a pipe impeller of equal diameter. When the rate of gas induction is compared at equal power input per unit volume as shown in

Figure 14, hollow PBTD45° and modified double-disk impeller with hydrofoil give higher gas induction rates. Further, the pipe impeller of diameter 0.33 m with PBTD45° gives the minimum induction rate, whereas the pipe impeller of 0.5 m diameter with PBTD45° gives a relatively high gas-induction rate. It shows that the optimum impeller diameter for the pipe impeller is T/2. In the case of hollow PBTD45°, the rate of gas induction at 0.33 m diameter is higher than that at 0.25 and 0.5 m diameter. It shows that the optimum diameter for hollow PBTD45° impeller is T/3. Also, it was observed that, as the submergence

Table 4. Different Geometric Parameters Varied for Multi-Impeller Studies upper impeller (gas inducing)

diameter, D1 (m)

clearance, C1 (m)

lower impeller

diameter, D2 (m)

clearance, C2 (m)

liquid height, H (m)

submergence of GI impeller, S (m)

modified double disk six-pipe

0.33 0.33

0.7 0.7

hydrofoil (6-bladed) PBTD45°

0.33 0.33

0.33 0.33

0.5

0.7

0.33

0.7

0.33 0.33 0.33

0.33 0.33 0.33

0.3 0.3 0.2 0.1 0.3

hollow PBTD45°

PBTD45° PBTU45° hydrofoil (6-bladed)

1 1 0.9 0.8 1

hollow PBTD60° modified four-pipe

0.5 0.25 0.33

0.7 0.7 0.64

PBTD45° PBTD45° PBTU45°

0.33 0.25 0.33

0.33 0.33 0.33

1 0.9 1 1 0.94

0.3 0.2 0.3 0.3 0.3

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Figure 6. Gassed power characteristics of hollow self-inducing impellers: (9) hollow PBTD45°, (0) hollow PBTD60°, (2) double-disk impeller, and (O) modified double-disk impeller.

Figure 7. Variation of power input with impeller speed: (9) hollow PBTD45°, (0) hollow PBTD60°, (2) double-disk impeller, and (O) modified double-disk impeller. Table 5. Power Number for Self-Inducing Impellers Used for Single-Impeller Studies impeller type hollow° PBTD45 hollow° PBTD60 double disk modified double disk

impeller diameter, D (m)

tank diameter, T (m)

liquid height, H (m)

clearance from bottom, C (m)

power number, NP

0.25 0.25 0.25 0.25 0.17 0.17

0.5 1 0.5 1 0.5 0.5

0.5 0.8 0.5 0.8 0.5 0.5

0.17 0.33 0.17 0.33 0.17 0.17

2.8 2.2 4.7 3.4 4.5 3.8

increases, the rate of gas induction at equal power input decreases. 4.3. Fractional Gas Holdup. Fractional gas holdup is an indication of mass transfer performance of any reactor. Factors affecting the gas holdup are the gas-induction rate and the average liquid circulation velocity. The impeller combination that compliments both these factors results in increased holdup. Figure 15 shows variation of fractional gas holdup with PG/V for a multiple-impeller system. It was observed that the fractional gas holdup is a linear function of impeller speed in all combinations. The minimum gas holdup can be seen (Figure 15) for hollow 6-pipe 0.33 m + PBTD45° 0.33 m, whereas maximum holdup can be seen for 6-pipe 0.5 m + PBTU45° 0.33 m. This is because the flow pattern of the pipe impeller is

Figure 8. (A) Ratio of gassed to ungassed power input for multi-impeller system: (() modified double disk 0.33 m + hydrofoil 0.33 m, H ) 1 m; (9) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 1 m; (2) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.9 m; (×) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.8 m; (-) 6-pipe 0.5 m + PBTD45° 0.33 m, H ) 1 m; (4) 6-pipe 0.5 m + PBTU45° 0.33 m, H ) 1 m; (O) hollow PBTD45° 0.33 m + Hydrofoil 0.33 m, H ) 1 m; (+) hollow PBTD45° 0.33 m +hydrofoil 0.33 m, H ) 0.9 m; (0) hollow PBTD60° 0.25 m + PBTD45° 0.25 m, H ) 1 m; ()) hollow PBTD45° 0.5 m + PBTD45° 0.33 m, H ) 1 m; and (b) modified 4-pipe 0.33 m + PBTU45° 0.33 m. (B) Schematic representation of quality of gas-liquid dispersion: L, flooding; M, cavity formation; MN, cavity growth; N, loading; NOP, cavity breakup; O, complete dispersion; P, recirculation. Lines 1, 2, and 3 represent the conventional gas-liquid reactor where the gas flow rate increases from 1 to 3. Line 4 represents PG/P behavior for self-inducing impeller. Table 6. Correlation Constants for Power Consumption Correlation for Multi-Impeller System

impeller combination

power number of multiimpeller

modified double disk 0.33 + hydrofoil 6-pipe 0.33 + PBTD45° 6-pipe 0.5 + PBTD45° 6-pipe 0.5 + PBTU45° hollow PBTD60° 0.25 + PBTD45° 0.25 hollow PBTD45° 0.33 + hydrofoil hollow PBTD45° 0.5 + PBTD45° modified 4-pipe impeller + PBTU45°

5.8 2.68 2.62 2.82 5.4 1.76 3.1 4.9

correlation constants b c 0.51 0.61 1.16 0.96 0.24 0.74 1.19 0.53

0.25 0.08 0.08 0.08 0.34 0.15 0.13 0.21

radial and the upflow turbine impeller helps the pipe impeller in discharging liquid as well as readily takes away the induced


Figure 9. Schematic representation of quality of gas-liquid dispersion: (A) flooding, (B) flooding-loading transition, (C) loading, (D) completely dispersed condition, and (E) recirculation.

Figure 10. Gas-induction characteristics of self-inducing impellers: (9) hollow PBTD45°, (0) hollow PBTD60°, (2) double-disk impeller, and (O) modified double-disk impeller.

Figure 12. Variation of QG/N2 with impeller speed for double-disk impeller: (9) double-disk impeller and (0) modified double-disk impeller.

Figure 11. Rate of gas induction as a function of power consumption per unit volume of liquid: (9) hollow PBTD45°, (0) hollow PBTD60°, (2) double-disk impeller, and (O) modified double-disk impeller.

Figure 13. Variation of gas-induction rate with impeller speed for multiimpeller system: (() modified double disk, 0.33 m + hydrofoil, 0.33 m H ) 1 m; (9) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 1 m; (2) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.9 m; (×) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.8 m; (-) 6-pipe 0.5 m + PBTD45° 0.33 m, H ) 1 m; (4) 6-pipe 0.5 m + PBTU45° 0.33 m, H ) 1 m; (O) hollow PBTD45° 0.33 m + hydrofoil, 0.33 m, H ) 1 m; (+) hollow PBTD45° 0.33 m + hydrofoil, 0.33 m, H ) 0.9 m; (0) hollow PBTD60° 0.25 m + PBTD45° 0.25 m, H ) 1 m; ()) hollow PBTD45° 0.5 m + PBTD45° 0.33 m, H ) 1 m; and (b) modified 4-pipe 0.33 m + PBTU45° 0.33 m.

gas. As a result of this coordination in flow direction of both impellers, average circulation velocity is higher and, hence, the gas holdup is higher. In all the cases, as the inducing impeller diameter decreases (second impeller diameter is constant 0.33 m), the fractional gas holdup also decreases. The hierarchy is 6-pipe 0.5 m > hollow PBTD45° 0.5 m > 6-pipe 0.33 m > hollow PBTD45° 0.33 m. If the gas-inducing impeller diameter is kept constant at 0.33 m, then the modified 4-pipe gives

comparable holdup values as that for the 6-pipe 0.5 m impeller. The modified double-disk, 0.33 m, gas-inducing impeller gives

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Figure 14. Rate of gas induction as a function of power input per unit volume for multi-impeller system: (() modified double disk, 0.33 m + hydrofoil, 0.33 m, H ) 1 m; (9) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 1 m; (2) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.9 m; (×) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.8 m; (-) 6-pipe 0.5 m + PBTD45° 0.33 m, H ) 1 m; (4) 6-pipe 0.5 m + PBTU45° 0.33 m, H ) 1 m; (O) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 1 m; (+) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 0.9 m; (0) hollow PBTD60° 0.25 m + PBTD45° 0.25 m, H ) 1 m; ()) hollow PBTD45° 0.5 m + PBTD45° 0.33 m, H ) 1 m; and (b) modified 4-pipe 0.33 m + PBTU45° 0.33 m.

Figure 16. Volumetric mass transfer coefficient as a function of impeller speed: (9) hollow PBTD45°, T ) 0.5 m; (0) hollow PBTD60°, T ) 0.5 m; (2) double-disk impeller, T ) 0.5 m; (O) modified double-disk impeller, T ) 0.5 m; (s) hollow PBTD45°, T ) 1 m, H ) 0.66 m; ()) hollow PBTD45°, T ) 1 m, H ) 0.8 m; (+) hollow PBTD60°, T ) 1 m, H ) 0.7 m; and (×) hollow PBTD60°, T ) 1 m, H ) 0.8 m.

Figure 17. Volumetric mass transfer coefficient as a function of power input per unit volume for experiments in 0.5 m i.d. tank: (9) hollow PBTD45°, (0) hollow PBTD60°, (2) double-disk impeller, and (O) modified double-disk impeller. Figure 15. Variation of fractional gas holdup with power input per unit volume for multi-impeller system: (() modified double disk 0.33 m + hydrofoil 0.33 m, H ) 1 m; (9) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 1 m; (2) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.9 m; × 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.8 m; (-) 6-pipe 0.5 m + PBTD45° 0.33 m, H ) 1 m; (4) 6-pipe 0.5 m + PBTU45° 0.33 m, H ) 1 m; (O) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 1 m; (+) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 0.9 m; (0) hollow PBTD60° 0.25 m + PBTD45° 0.25 m, H ) 1 m; ()) hollow PBTD45° 0.5 m + PBTD45° 0.33 m, H ) 1 m; and (b) modified 4-pipe 0.33 m + PBTU45° 0.33 m.

higher gas holdup values as compared to the 6-pipe 0.33 m and the hollow PBTD 0.33 m impellers. 4.4. Volumetric Gas-Liquid Mass Transfer Coefficient (kLa). The values of kLa were measured in the 0.5 and 1 m i.d. vessels and for both the impeller types. The impeller speed was varied from 2.5 to 14 rps, and the power input per unit volume was varied from 0.5 to 10 kW/m3. The values of kLa are plotted in Figure 16 against the impeller speed (N). It can be seen that the value of kLa increases with an increase in N and a decrease in the submergence (h). Figures 17 and 18 show the mass transfer performance of different impellers for a given power

input. From Figure 16 it can be observed that the modified double-disk impeller gives the highest kLa for a given power input per unit volume. The higher mass transfer efficiency of the modified double-disk impeller is due to its good dispersing ability; the average bubble diameter is less than that in the case of the hollow PBTD impeller. In the case of conventional stirred tanks, the gas is sparged separately below the impeller and the superficial gas velocity can be independently varied. However, in the case of selfinducing reactors, superficial gas velocity is a function of impeller speed and, hence, power input per unit volume. One more feature to be noted is that the value of kLa is zero at the critical impeller speed. In view of these features, it was thought desirable that kLa can be better correlated with P/V above that at the critical impeller speed. The following correlation has been proposed for all the impellers,

kLa ) A

(( ) ( ) ) PG P V V cr

0.8

(4)


Figure 18. Volumetric mass transfer coefficient as a function of power input per unit volume at different submergences for experiments in 1 m i.d. tank: (O) hollow PBTD45°, H ) 0.66 m; (0) hollow PBTD45°, H ) 0.8 m; (×) hollow PBTD60°, H ) 0.7 m; and (+) hollow PBTD60°, H ) 0.8 m. Table 7. Constants in Eq 4 for Different Impeller Studied impeller

A

PBTD45 PBTD60 double disk modified double disk disk impeller18 concave blade impeller19

0.0307 0.0384 0.0737 0.0772 0.0243 0.0361

range of PG/VL, kW/m3 0.75-6.75 0.73-10 1.55-6 1.1-6.16 0.16-2 0.32-1.7

range of kLa, s-1

Figure 19. Volumetric mass transfer coefficient as a function of impeller speed for multi-impeller system: (() modified double disk 0.33 m + hydrofoil 0.33 m, H ) 1 m; (9) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 1 m; (2) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.9 m; (×) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.8 m; (-) 6-pipe 0.5 m + PBTD45° 0.33 m, H ) 1 m; (4) 6-pipe 0.5 m + PBTU45° 0.33 m, H ) 1 m; (O) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 1 m; (+) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 0.9 m; (0) hollow PBTD60° 0.25 m + PBTD45° 0.25 m, H ) 1 m; ()) hollow PBTD45° 0.5 m + PBTD45° 0.33 m, H ) 1 m; and (b) modified 4-pipe 0.33 m + PBTU45° 0.33 m.

0.01-0.13 0.02-0.21 0.06-0.24 0.04-0.25 0.008-0.043 0.006-0.055

Table 8. Constants for Mass Transfer Correlation for Multi-Impeller Studies impeller combination

A (for H ) 1 m)

modified double disk 0.33+ hydrofoil 6-pipe 0.33 + PBTD45° 6-pipe 0.5 + PBTD45° 6-pipe 0.5 + PBTU45° hollow PBTD60° 0.25 + PBTD45° 0.25 hollow PBTD45° 0.33 + hydrofoil hollow PBTD45° 0.5 + PBTD45° modified 4-pipe impeller + PBTU45°

0.039 0.025 0.027 0.034 0.024 0.039 0.031 0.038

where A is an empirical constant that depends on the impeller design. The above equation correlates kLa within (20% of that of experimental value. The correlation coefficient for all multiimpeller data was found to be 0.96. The present correlation considers the experimental data of Heim et al.18 and Forrester et al.19 The correlation coefficient was found to be 0.98 and 0.99 when the kLa value was predicted by our equation (eq 4) and that by Forrester et al.19, respectively. It is interesting to note that the exponent is 0.8 for all the impellers so that the correlation has a single parameter; the values have been reported in Table 7 for a single-impeller system and Table 8 for a multiimpeller system. For a multi-impeller system, variation of kLa as a function of impeller speed has been plotted in Figure 19. It can be observed that the trend is somewhat similar to that of gasinduction rate as a function of impeller speed. Figure 20 shows kLa for different impeller combinations plotted at equal power inputs per unit volume. It can be observed that the hollow PBTD45° with hydrofoil gives the highest kLa, when compared at equal power input vis-a`-vis other combinations studied. The modified double-disk impeller with hydrofoil impeller also gives good kLa. This is because of inherently good gas-dispersing ability of the disk impeller. The optimum impeller diameter for

Figure 20. Volumetric mass transfer coefficient as a function of power input per unit volume for multi-impeller system: (() modified double disk 0.33 m + hydrofoil 0.33 m, H ) 1 m; (9) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 1 m; (2) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.9 m; (×) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.8 m; (-) 6-pipe 0.5 m + PBTD45° 0.33 m, H ) 1 m; (4) 6-pipe 0.5 m + PBTU45° 0.33 m, H ) 1 m; (O) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 1 m; (+) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 0.9 m; (0) hollow PBTD60° 0.25 m + PBTD45° 0.25 m, H ) 1 m; ()) hollow PBTD45° 0.5 m + PBTD45° 0.33 m, H ) 1 m; and (b) modified 4-pipe 0.33 m + PBTU45° 0.33 m.

hollow PBTD45° was 0.33 m, i.e., D/T ) 0.33. In the case of the pipe impeller, the optimum diameter observed was 0.5 m. The effect of submergence on kLa was studied by changing submergence at 0.3, 0.2, and 0.1 m in the case of the six-pipe impeller with PBTD45°. It was found that kLa increases with an increase in the submergence though the rate of gas induction decreases with an increase in the liquid submergence. This is mainly because of the increase in power consumption with an increase in submergence. To correlate kLa with power consumption, a similar correlation as that for a single impeller was found to hold (eq 4). The

2840


the order was hydrofoil > PBTU > PBTD. As an overall result, for energy-efficient mass transfer, the combinations of MDD and hydrofoil as well as hollow PBTD and hydrofoil are recommended. Nomenclature

Figure 21. Overall transfer efficiency as a function of impeller speed for multi-impeller system: (() modified double disk 0.33 m + hydrofoil 0.33 m, H ) 1 m; (9) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 1 m; (2) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.9 m; (×) 6-pipe 0.33 m + PBTD45° 0.33 m, H ) 0.8 m; (-) 6-pipe 0.5 m + PBTD45° 0.33 m, H ) 1 m; (4) 6-pipe 0.5 m + PBTU45° 0.33 m, H ) 1 m; (O) hollow PBTD45° 0.33 m + hydrofoil 0.33, H ) 1 m; (+) hollow PBTD45° 0.33 m + hydrofoil 0.33 m, H ) 0.9 m; (0) hollow PBTD60° 0.25 m + PBTD45° 0.25 m, H ) 1 m; ()) hollow PBTD45° 0.5 m + PBTD45° 0.33 m, H ) 1 m; and (b) modified 4-pipe 0.33 m + PBTU45° 0.33 m.

correlation constant A for different impeller combinations are listed in Table 8. Figure 21 shows the overall oxygen transfer efficiency (OTE) as a function of impeller speed for different combinations studied. This efficiency is calculated as (kg of O2 transferred)/ (kWh consumed).

OTE )

kLa × 3.6 × ([A*] - [Ao]) P/V

(5)

In eq 5, [A*] was taken as 7.6 mg/L and [AO] as zero, with P/V in kW/m3. It was observed that OTE first increases with an impeller speed, reaches a maximum, and then again decreases. This type of trend indicates the existence of an optimum impeller speed where OTE is maximum. Also, it was observed that the impeller combination that gives higher kLa at given power input gives higher OTE. In the present work, we have investigated eight combinations of multi-impeller systems. Among these, the combination of 45° PBTD (as self-inducing impeller) and six-bladed hydrofoil impeller (as the dispersing impeller) was found to be the best. This is mainly because the hydrofoil impeller creates better flow at the same power consumption as compared with all the other impellers covered in this work. In view of these results, the abovementioned combination is recommended in practice. 5. Conclusions (1) A new design of self-inducing impeller has been developed (namely, modified double-disk impeller) which is relatively more energy efficient than all the impeller designs published in the literature. (2) For industrial size equipment, a multi-impeller system is needed for achieving both the objectives: rate of gas induction and distribution of gas bubbles throughout the reactor volume. (3) Several multi-impeller combinations were investigated. With regards to the objective of gas induction, the impellers can be arranged in the order of MDD > hollow PBTD > hollow pipe. For the objective of gas distribution (the second impeller),

[A*] ) saturation concentration of oxygen in liquid phase, kmol/m3 [Ao] ) concentration of dissolved oxygen in bulk liquid phase, kmol/m3 a ) effective average interfacial area, m2/m3 [Bo] ) concentration of hydrazine at inlet to the reactor, kmol/ m3 (P/V)cr ) power input per unit volume of liquid at critical impeller speed, kW/m3 QB ) hydrazine flow rate, m3/s C ) Height of impeller from the bottom of the vessel, m D ) impeller diameter, m Fr ) Froude number, DN2/g, dimensionless Fr* ) modified Froude number, D2N2/gh, dimensionless Frc* ) modified critical Froude number, D2Ncr2/gh, dimensionless g ) acceleration due to gravity, m/s2 h ) height of liquid above impeller, m kL ) liquid-side mass transfer coefficient, m/s kLa ) volumetric mass transfer coefficient, s-1 kLao ) volumetric mass transfer coefficient at critical impeller speed, s-1 N ) impeller speed, rps NA ) aeration number, QG/ND3, dimensionless NC ) critical impeller speed for the onset of gas induction, rps P/V ) power input per unit volume of liquid, kW/m3 PBTD ) pitched-blade downflow turbine PBTU ) pitched-blade upflow turbine PG ) power consumption in presence of gas, W PO ) power consumption in absence of gas, W Q ) gas flow rate, m3/s RAa ) specific rate of oxidation of hydrazine, kmol/s Re ) Reynolds number, D2NF/µ, dimensionless Sc ) Schmidt number, µ/DF, dimensionless Sh ) Sherwood number, kLa‚D2/D, dimensionless T ) tank diameter, m VA ) active volume, m3 VG ) superficial gas velocity, m/s Vg ) volume of gas, m3 We ) Weber number, N2FD3/σ, dimensionless D ) diffusivity of gas in liquid, m2/s G ) fractional gas holdup µ ) liquid viscosity, Pa s υ ) kinematic liquid viscosity, m2/s F ) liquid density, kg/m3 σ ) surface tension, N/m Literature Cited (1) Joshi, J. B.; Sharma, M. M. Mass transfer and hydrodynamic characteristics of gas inducing type of agitated contactors. Can. J. Chem. Eng. 1977, 65, 683-695. (2) Patwardhan, A. W.; Joshi, J. B. Design of gas-inducing reactors. Ind. Eng. Chem. Res. 1999, 38, 49-80. (3) Martin, G. Q. Gas inducing agitator. Ind. Eng. Chem. Process Des. DeV. 1972, 11, 397-404. (4) Topiwala, H. H.; Hamer, C. Mass transfer and dispersion properties in a fermenter with a gas-inducing impeller. Trans. Inst. Chem. Eng. 1974, 52, 113-120.

Ind. Eng. Chem. Res., Vol. 47, No. 8, 2008 2841 (5) Sawant, S. B.; Joshi, J. B. Critical impeller speed for onset of gas induction in gas-inducing types of agitated contactors. Chem. Eng. J. 1979, 18, 87-91. (6) Joshi, J. B. Modifications in the design of gas inducing impellers. Chem. Eng. Commun. 1980, 5, 109-114. (7) Sawant, S. B.; Joshi, J. B.; Pangarkar, V. G.; Mhaskar, R. D. Mass transfer and hydrodynamic characteristics of the Denver type of flotation cells. Chem. Eng. J. 1981, 21, 11-19. (8) Evans, G. M.; Rielly, C. D.; Davidson, J. F.; Carpenter, K. J. Hydrodynamic characteristics of gas inducing impeller. In Proceedings of the 7th European Conference on Mixing IV, Kiav, Brugge, Belgium, September 18-20, 1991; pp 515-523. (9) Saravanan, K.; Mundale, V. D.; Joshi, J. B. Gas inducing type mechanically agitated contactors. Ind. Eng. Chem. Res. 1994, 33, 22262241. (10) Deshmukh, N. A.; Patil, S. S.; Joshi, J. B. Gas induction characteristics of hollow self-inducing impeller. Trans. Inst. Chem. Eng. 2006, 84 (A2), 124-132. (11) Tekie, Z.; Li, J.; Morsi, B. I.; Chang, M. Gas-liquid mass transfer in cyclohexane oxidation process using gas-inducing and surface-aeration agitated reactors. Chem. Eng. Sci. 1997, 52, 1541-1551. (12) Patil, S. S.; Deshmukh, N. A.; Joshi, J. B. Mass-transfer characteristis of surface aerators and gas-inducing impellers. Ind. Eng. Chem. Res. 2004, 43, 2765-2774. (13) Zieverink, M. M.; Kreutzer, M. T.; Kapteijn, F.; Moulijn, A. Gasliquid mass transfer in benchscale stirred tankssFluid properties and critical impeller speed for gas induction. Ind. Eng. Chem. Res. 2006, 45, 4574458. (14) Matsumura, M.; Sakuma, H.; Yamagata, T.; Kobayashi, J. Gas entrainment in a new gas entraining fermentor. J. Ferment. Tech. 1982, 5, 457-467. (15) Baczkiewicz, J.; Michalski, M. Oxygen transfer during mixing of acetic acid fermentation medium with self-aspirating tube agitator. In Proceedings of the 6th European Conference on Mixing, Pavia, Italy, May 24-26, 1988; BHRA: Cranfield, U.K., 1988.

(16) Dietrich, E.; Mathieu, C.; Delmas, H.; Jenck, J. Raney-Nickel catalyzed hydrogenations: Gas-liquid mass transfer in gas-induced stirred slurry reactors. Chem. Eng. Sci. 1992, 47, 3597-3604. (17) Hitchri, H.; Accary, A.; Puaux, J. P.; Andrieu, J. Gas-liquid mass transfer coefficients in a slurry batch reactor equipped with a self-gasinducing agitator. Ind. Eng. Chem. Res. 1992, 31, 1864-1867. (18) Heim, A.; Kraslawski, A.; Rzyski, E.; Stelmach, J. Aeration of bioreactors by self-aspirating impellers. Chem. Eng. J. 1995, 58, 59-63. (19) Forrester, S. E.; Rielly, C. D.; Carpenter K. J. Gas-inducing impeller design and performance characteristis. Chem. Eng. Sci. 1998, 53, 603615. (20) Poncin, S.; Nguyen, C.; Midoux, N.; Breysse, J. Hydrodynamic and volumetric gas-liquid mass transfer coefficient of a stirred vessel equipped with a gas-inducing impeller. Chem. Eng. Sci. 2002, 57, 32993306. (21) Sardeing, R.; Xuereb, C.; Poux, M. Improvement of the performances of a gas-inducing system for application in wastewater treatment. Int. J. Chem. React. Eng. 2006, 4, Article A30. (22) Smith, J. M.; Van’t Riet, K.; Middleton, J. C. Scale-up of agitated gas-liquid reactors for mass transfer. In Proceedings of the 2nd European Conference on Mixing, Cambridge, U.K., 1977; Paper F4. (23) Onken, U.; Sick, R.; Weiland, P. Determination of gas-liquid mass transfer by oxidation of hydrazine. Chem. Eng. Sci. 1985, 40, 1990-1993. (24) Marquez, A. L.; Nguyen, C.; Poncin, S.; Wild, G.; Midoux, N. A Novel Hydrazine Oxidation Technique for the Determination of kLa in GasLiquid and Gas-Liquid-Solid Reactors. Chem. Eng. Sci. 1994, 49, 56675679.

ReceiVed for reView October 16, 2007 ReVised manuscript receiVed December 13, 2007 Accepted December 20, 2007 IE071392M

Hydrodynamic and Mass Transfer Characteristics of Single and

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