Ind. Eng. Chem. Res. 2008, 47, 8995–9005
8995
Microstructured Mesh Contactor for Asymmetric Transfer Hydrogenation with Simultaneous Stripping: Modeling and Experiments M. Zanfir, X. Sun, and A. Gavriilidis* Department of Chemical Engineering, UniVersity College London, Torrington Place, London WC1 7JE, United Kingdom
A two-dimensional model for a mesh contactor utilized for acetone stripping from binary mixtures acetone-isopropanol with nitrogen as inert gas, as well as for asymmetric transfer hydrogenation of acetophenone with isopropanol as the hydrogen donor in the presence of a homogeneous catalyst and simultaneous acetone removal by stripping has been formulated and solved. Experiments have been performed with a microstructured mesh contactor that consists of parallel metal plates, gaskets, and a microstructured mesh so that passages for gas and liquid phases are formed. The model gives good agreement with experimental data for both acetone removal from a binary mixture and asymmetric transfer of hydrogenation. Theoretical and experimental investigations show that the amount of acetone removed from the liquid phase is strongly dependent on the ratio of inlet gas to liquid flowrate. For a higher ratio, more efficient acetone removal from the liquid phase is achieved. If dry nitrogen is used as a stripping agent, simultaneous evaporation of acetone and isopropanol results in significant solvent depletion of the liquid phase. In this case, the gas phase becomes saturated with solvent (isopropanol) near the reactor entrance. Due to efficient removal of acetone produced as byproduct during the reaction, reaction time in the mesh reactor can be significantly decreased as compared to a laboratory batch reactor. 1. Introduction Mesh, and membrane contactors in general, are devices that allow two phases to come into direct contact with each other, for the purpose of mass transfer without dispersing one phase into the other. The concept of using membranes to bring two phases into contact1,2 has been applied to processes such as reverseosmosis,filtration,3 extraction,4 electrodialysis,pervaporation,5–9 evaporation,10 stripping,11 distillation,13–16 multiphase reactive systems,17–20 and absorption.12 Their use is seen as part of process intensification trends, boosting efficiency, saving energy, minimizing environmental impact, and increasing safety. In this context, membrane processes have the potential to replace conventional energy intensive techniques, accomplishing selective and efficient transport of specific components and improving performance of reactive processes. The main advantages of membrane contactors are the following: (a) High and constant specific interfacial area. Membrane contactors can supply twenty to one hundred times more surface area per volume than conventional equipment, such as packed towers. (b) Hydrodynamic independence of the fluid phases in contact. They can be operated over a wide range of flowrates without constraints such as flooding, channelling or backmixing. (c) High modularity and compatibility. Scale-up of membrane contactors can be accomplished by numberingup and can be easily customized and modified. The recognized disadvantages of membrane contactors are the following: (a) Additional resistance to mass transfer introduced by membrane. This typically occurs only when the resistance of the contained phase in the membrane pores is significant compared to the bulk phases resistances. (b) Operational constraints posed by phase breakthrough. Nonwetting fluid does not pass through the pores as long as the pressure is kept below a critical threshold, the breakthrough pressure. The Laplace equation indicates that smaller pore size is beneficial. The breakthrough pressure can * To whom correspondence should be addressed. E-mail: a.gavriilidis@ ucl.ac.uk. Tel.: +44 20 76793811. Fax: +44 20 73832348.
be drastically reduced in the presence of surfactants (because they reduce the surface tension) or irregularities of the pores (i.e., breakthrough happens at the location of the biggest pore). In order to minimize these disadvantages, an ideal membrane should combine high porosity and well-defined pore size together with short mass transfer paths. Recent developments in the area of microengineered structures for chemical processing21,22 make it possible to manufacture meshes from various materials (e.g., steel, silicon nitride) by techniques such as standard mask lithography or laser interference lithography.23,24 Thin meshes (1-5 µm thickness) with straight pores, micrometerrange pore size, and a regular arrangement can be obtained.25–28 Such meshes combine the advantage of minimizing the resistance in the mesh with high porosity and regular patterned pore structure having at the same time good mechanical strength. In the present paper, a two-dimensional model for a microstructured mesh contactor is formulated and its predictions compared with experimental data for gas stripping of acetone from a binary isopropanol-acetone mixture. Further, asymmetric hydrogenation of acetophenone to its chiral alcohol 1-phenylethanol and simultaneous stripping of the byproduct acetone is investigated theoretically and experimentally.
The asymmetric transfer hydrogenation of ketones with isopropanol as hydrogen donor in the presence of a homogeneous catalyst is of significant importance in the synthesis of pharmaceutically relevant substances. The use of hydrogen donor rather than molecular hydrogen avoids the risks and the constraints associated with hydrogen handling and requirements for pressure vessels. Asymmetric transfer hydrogenation using isopropanol as a hydrogen donor is an equilibrium process. The
10.1021/ie071653s CCC: $40.75 2008 American Chemical Society Published on Web 06/18/2008
8996 Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008
Figure 1. Microstructured mesh contactor design. (a) Magnified picture of the mesh. (b) Components of the contactor. (c) Flow arrangement: (1) top plate, (2) gasket, (3) mesh, (4) bottom plate.
rate of the reverse reaction of the desired enantiomer is faster than that of the undesired enantiomer. Acetone presence in the reaction therefore limits conversion and erodes selectivity. Experimental studies established the extent to which these processes occur.29 2. Reactor Design and Experimental Conditions The microstructured mesh contactor consisted of parallel metal plates, gaskets, and a microstructured mesh so that passages for gas and liquid phases were formed as shown in Figure 1. Each plate had width × length of 3 cm × 8 cm, while the gasket thickness which determines the thickness of gas and liquid layers on each side of the mesh plate was 200 µm. The inner width of the gasket was 2 cm. The top and bottom plates were made of stainless steal, and the gaskets, of copper. The inlets and outlets were 1/8 in. stainless steel tubing mounted on the top plate. The mesh was made of stainless steel 304 and had a thickness of 50 µm (Internetmesh). The average pore size was 76 µm, and a regular pore arrangement (see Figure 1a) corresponded to an open area of about 23%. The flowsheet of the experimental setup is shown in Figure 2. First, the gas-liquid mesh contactor was utilized to investigate acetone removal from binary mixtures acetone-isopropanol using nitrogen. The liquid was pumped through the contactor using a syringe pump (RAZEL A-99.FJZ). The range of liquid flowrates utilized was between 0.05-0.3 mL/min. Nitrogen was used as stripping agent. Its flowrate was controlled by a mass flow controller (Brooks 5850), and the gas flowrate range utilized was between 70-280 mL/min. A condenser was used to condense the volatile components carried by the gas phase leaving the mesh contactor. A U-tube manometer was used to measure the exit gas phase pressure. The pressure difference between the gas and liquid phases was controlled by a micrometering valve (Swagelok) mounted in the exit
gas stream, to prevent phase breakthrough. The reactor operated at almost atmospheric pressure. Breakthrough experiments, identified that the breakthrough of gas in the liquid phase occurred at a pressure difference PG - PL of about 700-800 Pa. In the stripping and reaction experiments, the pressure difference was maintained in the range 300-500 Pa. In order to keep isothermal conditions, the contactor was placed in a water bath. Subsequently, studies were performed to investigate the reactor performance for asymmetric transfer hydrogenation of acetophenone in the presence of the homogeneous catalyst, 1R,2Samino-indanol/pentamethylcyclopentadienylrhodium,30,31 one of a family of catalysts with the trademark CATHy. First, the catalyst solution was prepared in a round-bottom flask. Pentamethylcyclopentadienylrhodium chloride dimer (11.2 mg, 18 µmol) (Avecia), 1R,2S-aminoindanol (5.6 mg, 37 µmol) (Avecia), and isopropanol (125 mL) (Aldrich 99.5%) were introduced in the flask for catalyst preparation. Nitrogen (BOC, CP grade) bubbling through the solution with a flow rate of 40 mL/min was used to keep an inert blanket. The mixture was stirred overnight at room temperature. A deep red catalyst solution was obtained. A substrate solution was prepared from acetophenone (4.5 g, 37.1 mmol) (Aldrich 99%) in isopropanol (122 mL), and 3 mL of sodium isopropoxide solution was added. The two solutions were pumped by syringe pumps (RAZEL A-99.FJZ) and were mixed in a micromixer (IMM, Standard Slit Interdigital Micro Mixer) before entering the reactor where they flowed in a cocurrent arrangement. During mixing, sodium isopropoxide from the substrate solution came in contact with the catalyst solution and the homogeneous catalyst was activated. The nitrogen flowrate was controlled by a mass flow controller (Brooks 5850). The inlet acetophenone concentration was 0.14 mol/L, gas flowrate 70 mL/min, and inlet liquid flowrate in the range 0.026-0.4 mL/min. The ratio of substrate concentration to catalyst concentration was 1000, the reactor was operated isothermally at 30 °C, and the gas phase was bubbled in solvent before entering the reactor. The liquid stream was collected in a flask after leaving the reactor. A 0.4 mL sample was withdrawn by a syringe into a vial containing 10 µL of acetic acid (Aldrich 99.7%), which was used to stop further reaction, and then analyzed by an Agilent 6890 GC system. The inlet temperature was 250 °C, the inlet pressure, 20 psi, and the split ratio, 200:1. An injection volume of 0.2 µL was used in the autoinjector. A CYCLODEX-B capillary column (30.0 m × 250 µm × 0.25 µm) was employed for the separation of the acetophenone, (R)-1-phenylethanol and (S)-1-phenylethanol, to estimate conversion and enantioselectivity, along with a flame ionization detector at 250 °C. The oven temperature was kept at 110 °C. The internal normalization method (with acetophenone)32 was used for quantification. Acetone concentration was determined using a DB-624 capillary column (30.0 m × 530 µm × 3.00 µm) and a thermal conductivity detector at 250 °C, with the following oven temperature program: 60-100 °C at 10 °C/min; 100 °C for 4 min; 100-200 °C at 40 °C/min; 200 °C for 2 min. 3. Mathematical Model A two-dimensional model of the microstructured mesh reactor was formulated, validated against experimental data, and used to perform parametric studies. The model is based on the following assumptions: (a) The liquid phase fills the space between the bottom plate and the mesh. The liquid layer thickness is constant and equal to the gasket thickness
Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 8997
Figure 2. Schematic of experimental setup. Table 1. Definition of Dimensionless Coordinates and Variables name
definition
axial dimensionless coordinate transverse dimensionless coordinate average liquid dimensionless velocity dimensionless concentration Fourier number
ζ ) z/L ξ ) x/δ U ) uL/u0L 0 CiL ) ciL/cisopropanol Fo ) (D/δ2)(L/u0)
along with boundary conditions and additional equations are given below. 3.1. Mole Balances in the Liquid Phase. The material balance for a wide plate geometry for each component is as follows:
Figure 3. Model domains with dimensionless coordinates and boundaries.
along the whole reactor length. Constant densities and velocities are assumed in the transverse direction for simplification. (b) Mesh pores are completely filled by the liquid so that the gas-liquid interface is situated at the same location as the mesh-gas interface. (c) Evaporation of volatile components, isopropanol and acetone, is accounted for and results in the decrease of the overall flow rate of the liquid phase. Since the cross section is constant, this causes reduction of liquid velocity along the flow direction. (d) Since the gas phase is dilute in volatile components stripped from the liquid phase, the gas velocity along the flow direction is considered constant. (e) The boundary conditions for volatile components at the liquid-gas interface, are the Vapor-liquid equilibrium condition and the equality of molar fluxes transferred between phases, while for the nonvolatile components zero molar flux is considered. (f) The reactor is considered to operate isothermally, because the reaction heat is low. Furthermore, it is placed in a water bath maintained at constant temperature. (g) Pressure drop along the reactor is neglected. (h) It is assumed that the isopropanol-acetone vapor-liquid equilibrium is not significantly affected by the presence of the reagents in the reaction mixture. The reactor model is divided into three main domains: the gas phase, the mesh, and the liquid phase. The domains taken into consideration and their dimensionless coordinate and boundaries are shown in Figure 3. Mole balances in each domain
(
)
∂2cLi ∂2cLi ∂ (uLcLi ) ) DLi + + Ri (1) ∂z ∂xL ∂z2 Introducing the dimensionless coordinates and variables given in Table 1, eq 1 becomes
(
)
∂2CLi Ri L δ2L ∂2CLi ∂ (UL · CLi ) ) FoLi 2 + + 0 0 2 ∂ζ ∂ξ L ∂ζ cipa uL L
(2)
The kinetic mechanism utilized in the model consisted of 10 reversible reactions.33 The reaction steps and values of the rate constants are presented in Table 2. 3.2. Decrease of Liquid Flowrate due to Removal of Volatile Components. Significant evaporation of the solvent (isopropanol) together with acetone results in a decrease of the liquid flowrate. Consequently, since the liquid is considered to fill the cross section of the passage, solvent and acetone evaporation results in a decrease of average liquid velocity along the reactor: δL
duL P )dz RGT
(
2
∑D
G i
i)1
∂yi ∂xG
| ) xG)δG
Mi Fi
(3)
In dimensionless form, eq 3 becomes P dU )dζ RgT
(∑ 2
i)1
DGi
∂yi ∂ξG
| )
Mi L 0 F δ δ i ξG)1 L GuL
(4)
3.3. Mole Balances in the Mesh. The liquid reaction mixture wets the membrane. We stabilized the gas-liquid interface at the gas side of the mesh, because even though this is not desirable from the mass transfer point of view, it provides more stability to potential pressure difference disturbances. With a wetting fluid, it
8998 Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 Table 2. Kinetic Mechanism and Rate Constants for Asymmetric Transfer Hydrogenation of Acetophenone with 1R,2S-Aminoindanol/ pentamethylcyclopentadienylrhodium Catalyst, at 30 °C, Utilized in the Model
RhL + IPAa RhLH + acetone k2
k3 ) 308.23 L/(mol s)
RhL + suba SIM
∀ ξL ∈ (0 : 1)
ζ ) 1,
k2 ) 179.7 L/(mol s)
k3
∀ζ ∈ [0 : 1],
k5 ) 112.1 L/(mol s)
k5
RhL + suba RIM
k7
(10)
ξL ) 0
∂CLi )0 ∂ξL
(11)
k9
RhL ) active catalyst, RhLH ) catalyst intermediate, IPA ) isopropanol, sub ) substrate (acetophenone), SIM, RIM ) chiral intermediates, S_prod ) (S)-1-phenylethanol, R_prod ) (R)-1-phenylethanol. a
is difficult to stabilize the gas-liquid interface at the liquid side of the mesh. Thus, the mesh pores are filled with liquid. The mass transport takes place only in the transverse direction through the mesh, so that the mole balance is given by eq 5 ∂2cM i DLi ∂xM
+ Ri ) 0
(13)
ξG ) 0
∂yi )0 ∂ξG
(14)
∀ ξM ∈ (0 : 1)
∂CM i )0 ∂ζ
(15)
Wall
k10 ) 2.88 L/(mol s)
k10
∂yi )0 ∂ζ
∀ ξG ∈ (0 : 1)
ζ ) 1,
k9 ) 2.39 1/s
RIMa RhL + R_prod
(5)
∀ζ ∈ [0 : 1], Mesh. Inlet, outlet ζ)0
∂CM i )0 ∂ζ Liquid-Mesh Interface ∀ζ ∈ [0:1] ξL ) 1; ξM ) 0 • equality of concentration ∀ ξM ∈ (0 : 1)
ζ)1
Introducing the dimensionless variables and coordinates, eq 5 becomes
CLi ) CM i
(6)
3.4. Mole Balances in the Gas Phase. The gas phase remains dilute in volatile components transferred from the liquid phase. For example for an initial acetone concentration of 0.1 mol/L in the liquid phase, the mole fraction of acetone in the gas phase for equilibrium conditions does not exceed 1%, while for the situation when dry nitrogen is used, the mole fraction of isopropanol in the gas phase cannot exceed 6%. Thus, there is negligible variation in gas phase density and the average gas velocity is considered constant. Consequently, the material balance for the gas phase can be written in terms of mole balance
(
2
2
∂yi ∂ yi ∂ yi ) DGi + uG ∂z ∂z2 ∂xG
)
(
)
∂cLi ∂xL
|
xL)δL
) εDLi
∂cM i ∂xM
(7)
(8)
3.5. Boundary Conditions. The following boundary conditions are considered (see also Figure 3): Liquid Phase. Inlet
|
xM)0
which, in dimensionless form, becomes ∂CLi ∂ξL
|
)ε ξL)1
δL ∂CM i δ ∂ξM
|
(18) ξM)0
Gas-Mesh Interface ∀ζ ∈ [0:1], ξG ) ξM ) 1 • vapor-liquid equilibrium condition for volatile components (i ) isopropanol, acetone), γixipvap i ) ΦiyiP
where i represents the volatile components: isopropanol and acetone. Introducing the dimensionless variables given in Table 1, eq 7 becomes δG2 ∂2yi ∂2yi ∂yi ) FoGi 2 + ∂ζ L ∂ζ2 ∂ξG
DLi
(16)
(17)
• equality of molar fluxes 2 Ri δM ∂2CM i + 0 )0 ∂ξM c DL ipa i
(12)
Outlet
k8 ) 2176 L/(mol s)
k8
(9)
∂CLi )0 ∂ζ
∀ ξG ∈ (0 : 1) yi ) 0 or yi ) yeq
ζ ) 0,
k7 ) 11.32 1/s
SIMa RhL + S_prod
, u)1
Gas Phase. Inlet
k6 ) 9.50 1/s
k6
c0ipa
Wall
k4 ) 0.0033 1/s
k4
c0i
Outlet
k1 ) 0.877 L/(mol s)
k1
∀ ξL ∈ (0 : 1) CLi )
ζ ) 0,
(19)
Since the total pressure, P, is low, the gas phase fugacity coefficient Φ has values close to unity.34 Vapor-liquid equilibrium data available in the literature for the binary mixture acetone-isopropanol35 were correlated using a Redlich-Kister type correlation and Wilson equation34 in order to obtain the activity coefficient of acetone in isopropanol. The average value of acetone activity coefficient for mole fraction lower than 0.001 obtained from these two approaches was γacetone ) 3. For solvent, the liquid isopropanol mole fraction is close to 1; hence, its activity coefficient can be considered to be close to unity, γisopropanol ) 1. • equality of molar fluxes for volatile components
|
∂cM i εDLi )∂xM xL)δM
P G ∂yi D RgT i ∂xG
|
Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 8999 Table 3. Model Parameters and Corresponding Values Used in the Calculations xG)δG
which, in dimensionless form, becomes ε
∂CM i ∂ξM
|
)ξM)1
DGi δM ∂yi P (RgT)c0ipa DLi δG ∂ξG
|
(20) ξG)1
• zero flux for nonvolatile components (i * isopropanol; acetone) ∂CM i ∂ξM
|
)0
(21)
ξM)1
3.6. Additional Equations. The inlet liquid velocity and gas velocity are calculated based on inlet volumetric flowrates and the area of the cross-section of the passages for the liquid and gas phase, respectively: u0L )
FL δLw
(22)
uG0 )
FG δGw
(23)
For reaction simulations, acetophenone conversion as a function of reactor length has been calculated by averaging the number of moles of acetophenone in the cross section of the liquid phase passage, thus
Xacetophenone ) 1 -
1 δL
∫
δL
0
(uLcacetophenone) dxL
u0Lc0acetophenone
(24)
which can be further rearranged using the dimensionless variables as Xacetophenone ) 1 -
∫
1 (UCacetophenone) dξL 0 0 0 cacetophenone/cisopropanol
(25)
The model parameters along with the values used in the calculations are given in Table 3. The semiempirical method of Fuller36 was used to calculate the diffusion coefficients in the gas phase and the method of Wilke and Chang37 for the diffusion coefficients in the liquid phase. The diffusion coefficients of phenylethanol and all intermediate species were considered similar to the acetophenone diffusion coefficient in isopropanol. 3.7. Solution of the Mathematical Model. The mathematical model consists of a system of partial differential equations (PDEs) and algebraic equations 1–25. It was solved utilizing the gPROMS modeling environment,38 which is an equationbased system that facilitates easy implementation of discretised models for dynamic or steady-state processes. For the current work, the backward finite difference method of second order was employed for integration along both coordinates. Each domain was discretised using 50 elements in the axial direction (ζ) and 30 elements in the transverse direction (ξ). Further increase of the number of elements (i.e., by 10 in each direction) did not make a difference in the solution. 4. Results and Discussion 4.1. Acetone Removal from a Binary Isopropanol-Acetone Mixture. The mesh contactor was first utilized to investigate its performance in removing acetone from a binary mixture
parameter name
symbol
value
plate length plate width liquid passage thickness gas passage thickness mesh thickness mesh porosity inlet liquid flowrate inlet gas flowrate acetone vapor pressure (30 °C) isopropanol vapor pressure (30 °C) diffusion coefficient in gas phase acetone-nitrogen isopropanol-nitrogen diffusion coefficient in liquid phase acetone-isopropanol acetophenone-isopropanol
L w δL δG δM ε FL FG vap pacetone vap pisopropanol DG
60 mm 20 mm 200 µm 200 µm 50 µm 23% 0.05-0.3 mL/min 70-280 mL/min 37 720 Pa 8 087 Pa
DL
1.15 × 10-5 m2/s 1.11 × 10-5 m2/s 6.4 × 10-10 m2/s 4.4 × 10-10 m2/s
acetone-isopropanol using nitrogen as stripping agent. The experimental studies investigated the influence of gas flowrate, liquid flowrate, and inlet acetone concentration. The cases considered are summarized in Table 4. For each case, Table 4 shows inlet conditions given as inlet liquid and gas flowrates and acetone inlet concentration together with the outlet acetone concentration and outlet liquid flowrate reported experimentally and obtained through model simulation. The influence of gas flowrate is addressed through cases 1-4, the influence of liquid flowrate is investigated in cases 5-8, and the influence of inlet acetone concentration is given through cases 9-12, while in cases 13-14 the gas phase has been saturated with solvent (isopropanol) before entering the contactor in order to minimize solvent evaporation. 4.1.1. Model Validation. In order to compare the model performance with experimental data, parity plots are presented for acetone outlet concentrations and outlet liquid flowrate in Figure 4. The model simulations give slightly lower outlet acetone concentrations, than the ones determined experimentally. This may be due to transverse velocity gradients due to laminar flow which are not considered in the model as well as the liquid flowing diagonally from inlet to outlet, creating stagnation zones in the corner regions.39 However, since no adjustable parameters were used, the model agreement with the experimental data for a wide rage of operating conditions is considered as satisfactory. The relative amount of acetone removed calculated from F0Lc0acetone - Foutlet coutlet L acetone F0Lc0acetone
(26)
is given in Table 4 together with the ratio FG/FL of the gas flowrate over the liquid flowrate. Figure 5 shows the dependence of the relative amount of acetone removed as a function of gas/ liquid flowrate ratio FG/FL. It can be seen that the amount of acetone removed increases as this ratio increases. Amounts of acetone removed at equal FG/FL ratio are very similar. For example, for FG/FL ) 700, which occurs for cases 1, 6, and 9-12, acetone removed was about 74% regardless of inlet acetone concentration or particular values of inlet gas and liquid flowrates. For cases 13-14, when the gas phase is saturated with isopropanol, there is negligible depletion of the liquid flowrate as compared to the cases where dry nitrogen is used. Thus, slightly shorter residence time for the liquid phase is obtained leading to a smaller amount of acetone removed, ca. 69%.
9000 Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 Table 4. Cases studied for acetone removal from binary mixtures acetone-isopropanol. Comparison of theoretical and experimental results. experimental
1 2 3 4 5 6 7 8 9 10 11 12 13a 14a a
modeling
inlet liquid flowrate, mL/min
inlet gas flowrate, mL/min
inlet acetone concentration, mol/L
outlet acetone concentration, mol/L
outlet liquid flowrate, mL/min
0.1
70 140 210 280 70
0.1123
70
0.1086 0.2273 0.3237 0.4133 0.1022
0.0384 0.0300 0.0270 0.0238 0.0170 0.0368 0.0617 0.0740 0.0376 0.0756 0.1156 0.1545 0.0436 0.0446
0.0779 0.0634 0.0537 0.0441 0.0279 0.0789 0.1785 0.2778 0.0775 0.0765 0.0768 0.0751 0.0978 0.0970
0.05 0.10 0.20 0.30 0.1
0.1
70
0.1066
acetone removed, %
outlet acetone concentration mol/L
outlet liquid flowrate, mL/min
ratio FG/FL
acetone removed %
73.36 83.06 87.09 90.65 91.10 72.77 48.35 35.73 73.16 74.55 72.57 71.92 58.26 57.66
0.0360 0.0232 0.0159 0.0118 0.0193 0.0345 0.0574 0.0699 0.0345 0.0732 0.1039 0.1323 0.0317 0.0317
0.0824 0.0652 0.0480 0.0309 0.0326 0.0824 0.1822 0.2821 0.0820 0.0816 0.0811 0.0800 0.0990 0.0990
700 1400 2100 2800 1400 700 350 233 700 700 700 700 700 700
73.58 86.53 93.20 96.75 88.20 73.34 50.95 38.35 73.94 73.72 73.97 74.39 69.28 69.28
The gas phase was saturated with isopropanol before entering the mesh contactor.
Figure 4. Parity plots for (a) outlet acetone concentration and (b) outlet liquid flowrate.
Figure 6. (a) Average dimensionless acetone concentration. (b) Dimensionless liquid flowrate as a function of dimensionless axial coordinate, for cases 1-4 of Table 4.
∫ c ) [∫ C
acetone )
1 δL
δL
0
1
0
Figure 5. Dependence of the relative amount of acetone removed on gas and liquid flowrates ratio, obtained by the model, for cases 1-14 of Table 4.
4.1.2. Influence of Gas Flowrate. Average acetone concentration in the cross section was calculated by
acetone(xL, z)
acetone(ξL, ζ)
dxL
]
0 dξL cisopropanol
(27)c¯
The average dimensionless acetone concentration in the cross section, jcacetone/c0acetone, and the dimensionless liquid flowrate, FL/F0L, are shown as a function of the axial dimensionless coordinate, ζ, in Figure 6, for cases 1-4 given in Table 4. The results suggest that most of solvent (isopropanol) evaporation takes place near the contactor entrance. Afterward, the gas phase becomes saturated in isopropanol and the solvent evaporation is less significant. As a result of solvent evaporation, the liquid flowrate decreases along the contactor and near the entrance the average acetone concentration shows a peak since fast evaporation of the solvent results in concentration increase in the liquid phase. The higher the gas flowrate used for stripping, the higher
Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 9001
Figure 7. Comparison of average dimensionless acetone concentration and dimensionless liquid flowrate as a function of dimensionless axial coordinate, for dry nitrogen as the gas phase (case 1) and gas phase saturated initially with isopropanol (case 13) (see Table 4).
Figure 9. Mass transfer driving force in the gas phase related to acetone mole fraction at the gas-mesh interface, as a function of the dimensionless axial coordinate, for cases 1 and 4 (see Table 4).
the amount of solvent evaporated and consequently the depletion of the liquid flowrate. Thus, higher nitrogen flowrate results in a larger enrichment in acetone near the contactor entrance, illustrated in Figure 6a as a higher peak. However, increasing the gas flowrate results ultimately in better removal of acetone from the liquid phase, consequently the outlet acetone concentration is lower as the nitrogen flowrate increases. If the gas phase is saturated with solvent (isopropanol) before entering (cases 13 and 14 in Table 4), evaporation of isopropanol inside the contactor and consequently liquid flowrate depletion is prevented. Thus, no peak in average acetone concentration along the contactor is observed. Figure 7 presents a comparison of average dimensionless acetone concentration in the cross section, jcacetone/c0acetone, and dimensionless liquid flowrate, FL/ F0L, for case 1, when dry nitrogen is used, and case 13, when
nitrogen is first saturated with isopropanol and similar inlet liquid and gas flowrate are used. Acetone concentration profiles across the liquid film, mesh, and gas film are presented in Figure 8 for different axial positions, as a comparison between cases 1 and 4. As expected, the profiles suggest that the dominant resistance for mass transfer is located in the liquid contained in the mesh pores, while in the gas phase mass transfer resistance is negligible. For higher nitrogen flowrate, case 4, the gas phase is more dilute in acetone lowering the acetone concentration at the gas-mesh interface and thus maintaining a higher driving force for mass transfer in the first half of the contactor. After the dimensionless axial coordinate ζ ) 0.4, the gas phase concentration is very close to its value at the interface. Figure 9 shows the driving force for acetone mass transfer in the gas phase expressed as ∆Y(%)
Figure 8. (a and b) Acetone concentration profiles across the liquid film and mesh for cases 1 and 4. (c and d) Acetone mole fraction, %, in the gas phase across the gas film for cases 1 and 4 (see Table 4).
9002 Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008
Figure 10. Average dimensionless acetone concentration as a function of axial location for cases 2 and 5 (see Table 4). wall wall ) [(yinterface acetone - yacetone)/yacetone] × 100, along the contactor. It is found that this driving force is lower than 1% for ζ > 0.4 and the rate of mass transfer decreases significantly. 4.1.3. Influence of Liquid Flowrate. For similar contactor geometry and ratio of gas and liquid inlet flowrates, the liquid flowrate affects the residence time of the liquid phase in the contactor. A comparison between cases 2 and 5 is made in order to explain the influence of liquid flowrate on the performance of acetone removal from the liquid phase. In case 5, the gas and liquid flowrates utilized are half of that in case 2, and this results in doubling the residence time of the liquid phase in the contactor. However, as can be seen in Table 4, for a double residence time, the amount of acetone removed increase only slightly from 86.53% in case 2 to 88.20% in case 5. Acetone concentration averaged across the liquid film is presented as a function of the contactor dimensionless axial coordinate for cases 2 and 5 in Figure 10. It can be seen that for case 5 (at higher residence time) a similar acetone concentration in the liquid phase as the outlet acetone concentration for case 2 is achieved at the dimensionless axial coordinate ζ ) 0.5. The amount of acetone removed in the second half of the contactor for case 5 is not significant since the system reaches its thermodynamic limitation, and increasing the contactor length beyond 30 mm (ζ ) 0.5) does not provide any benefit. 4.1.4. Influence of Inlet Acetone Concentration. The results given in Table 4 indicate that the amount of acetone removed, defined by eq 26, is virtually independent of inlet acetone concentration. Moreover, the profiles of average acetone concentration across the liquid film related to the inlet concentration along the contactor for cases 9-13 almost overlap (not shown here). 4.2. Asymmetric Transfer Hydrogenation in the Mesh Reactor with Acetone Removal. Efficient removal of acetone from the liquid phase shortens the residence time required to achieve high conversion. Table 5 summarizes the investigated cases and gives values for outlet conversion and acetone
Figure 11. Parity plots for cases r1-r5 (see Table 5), [inlet substrate concentration 0.14 mol/L, substrate/catalyst ) 1000, inlet gas flowrate 70 mL/min] to compare theoretical and experimental results in terms of (a) outlet conversion and (b) outlet acetone concentration.
concentration obtained experimentally and by model simulation. The liquid residence time is calculated from the total liquid flowrate and the reactor volume filled by the liquid (0.24 mL). There is some reaction taking place in the tubing upstream and downstream of the mesh reactor (total volume 0.073 mL). However, this is not considered in the calculations, because the reaction in these regions occurs without acetone removal. A parity plot is given in Figure 11 for conversion and acetone concentration respectively, showing that the model gives good agreement with experimental data. 4.2.1. Influence of Inlet Liquid Flowrate at Constant Gas Flowrate. Conversion evolution along the reactor dimensionless axial coordinate, ζ, is shown in Figure 12a, while Figure 12b shows average acetone concentration in the cross section of the liquid passage as a function of conversion for cases r1-r5 given in Table 5. As inlet liquid flowrate decreases, the outlet conversion obtained is higher. Two factors contribute to this: first, the lower liquid flowrate results in a higher residence time of the liquid phase, and second, since the gas flowrate is kept at 70 mL/min, as the inlet flowrate decreases, the ratio FG/FL increases. As discussed in the previous section, this favors a better acetone removal from the liquid phase and consequently higher conversion can be achieved. This is confirmed by acetone concentration profiles given in Figure 12b. For low conversions (Xacetophenone < 0.4), acetone concentration profiles as function of conversion practically overlap for all cases considered. However, for higher conversion, (Xacetophenone > 0.4), as the ratio FG/FL increases, (from case r1 to r5) better acetone removal from the liquid phase is achieved.
Table 5. Cases Studied for Asymmetric Hydrogenation in Mesh Reactor, Comparison of Theoretical and Experimental Results experimental
modeling
inlet substrate inlet liquid inlet gas liquid residence outlet acetone outlet acetone concentration, mol/L flowrate, mL/min flowrate, mL/min time, s conversion concentration, mol/L conversion concentration, mol/L r1 r2 r3 r4 r5
0.14
0.400 0.200 0.100 0.050 0.026
70
36 72 144 288 554
0.2186 0.3767 0.5911 0.8447 0.9133
0.2016 0.0282 0.0333 0.0276 0.0159
0.1948 0.3698 0.6631 0.9700 0.9990
0.2100 0.0340 0.0440 0.0329 0.0130
Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 9003
Figure 12. Effect of liquid flowrate on conversion. (a) Acetophenone conversion as a function of the reactor dimensionless coordinate for cases r1-r5. (b) Average acetone concentration in the liquid passage cross section as a function of conversion, for cases r1-r5 (see Table 5).
In order to illustrate the relevance of acetone concentration gradients across the liquid film and mesh, Figure 13 shows acetone concentration in the liquid phase as a function of dimensionless axial coordinate, for different transverse locations, namely wall (ξL) 0), liquid-mesh interface (ξL ) 1, ξM ) 0), and mesh-gas interface (ξM ) ξG ) 1). The acetone concentration profiles suggest that significant transverse mass transfer resistances exist along the reactor and these are more significant at high liquid residence time. However, although acetone concentration gradients in the transverse direction are significant, this does not affect significantly the reactant concentration in the cross section, as demonstrated in Figure 14, where acetophenone concentration across the liquid film and mesh, at dimensionless coordinate ζ ) 0.3 for case r5 is given. This corresponds to the highest acetone concentration gradient in the cross section. In order to provide an understanding about the magnitude of the concentration gradients in the cross section, a relative difference in concentration was calculated as: Γ(%) )
c
wall
-c
gas-mesh interface
cwall
× 100
(28)
Thus, for case r5 at ζ ) 0.3, Γ(acetophenone) ) 3.6% while Γ(acetone) ) 83.2%. This indicates that, despite significant acetone concentration gradients in the cross section, the reactant concentration gradients in the cross section are not significant. 4.2.2. Influence of Inlet Gas Flowrate at Constant Liquid Flowrate. In order to understand how the gas flowrate influences the reactor performance for similar residence time for the liquid phase (constant inlet liquid flowrate), the model was used to simulate two cases derived from previously investigated cases r3 and r4 by increasing the inlet gas flowrate by 4 times. Table 6 specifies the inlet conditions in terms of substrate concentration and gas and liquid flowrates together
Figure 13. Effect of liquid flowrate on axial concentration profiles. Acetone concentration as a function of dimensionless axial coordinate for (1) wall (ξL ) 0); (2) liquid-mesh interface (ξL ) 1, ξM ) 0); (3) mesh-gas interface (ξM ) ξG ) 1), for (a) case r1 and (b) case r5 (see Table 5).
Figure 14. Acetophenone concentration profile across the liquid film and mesh, at axial dimensionless coordinate ζ ) 0.3, in case r5 (see Table 5).
with outlet conversion and acetone concentration. Even though the results suggest that increasing the gas flowrate results in slightly better removal of the acetone, the impact on acetophenone conversion is not significant. For illustration, Figure 15 compares the acetone concentration as a function of conversion for cases r4 and r4a. The fact that increase of FG/FL does not affect conversion is due to the low level of acetone concentration which is kept below 0.05 mol/L. 4.3. Comparison to Batch Reactor. Asymmetric transfer hydrogenation of acetophenone in a batch laboratory reactor using an initial liquid load of 250 mL and a continuous nitrogen flowrate of 800 mL/min bubbled through the liquid phase, kept isothermal at 30 °C, was simulated in ref 40, while relevant experimental results have been reported in ref 29. Figure 16a presents a comparison in terms of conversion obtained as a function of reaction time for the batch reactor and outlet conversion as a
9004 Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 Table 6. Cases Studied Theoretically for Asymmetric Hydrogenation in the Mesh Reactor modeling
r3a r4a
inlet substrate concentration, mol/L
inlet liquid flowrate, mL/min
inlet gas flowrate, mL/min
liquid residence time, s
conversion
outlet acetone concentration mol/L
0.14
0.10 0.05
280
144 288
0.6658 0.9745
0.0373 0.0218
function of liquid residence time for the mesh reactor (see also Table 5). As demonstrated theoretically and experimentally, high conversion can be achieved in a mesh reactor for much shorter residence time. Examining the acetone concentration profiles as a function of conversion presented in Figure 16b, it can be seen that in the batch reactor despite utilization of nitrogen as stripping agent, acetone concentration builds up as reaction progresses and reaches values up to 0.09 mol/L. In contrast, in the mesh reactor the acetone
is removed more efficiently as reaction progresses and its concentration is kept below 0.04 mol/L, which results in shorter reaction time needed to achieve the same conversion. The improved performance of the mesh reactor as compared to the batch reactor is due to the high amount of gas/liquid volume contacted. In the cases considered, for the mesh reactor the ratio FG/FL ) 700, while for the laboratory 250 mL batch reactor the ratio FGτR/VL ) 20 for the reaction time to achieve ca. 50% conversion. Increasing gas/liquid volume ratio in batch reactors is limited by impeller flooding, while contacting the gas and the liquid phase by means of a thin mesh decouples the gas and liquid flowrates and eliminates such a constraint. 5. Conclusions
Figure 15. Average acetone concentration as a function of conversion, for cases r4 and r4a (see Tables 5 and 6), corresponding to an inlet gas flowrate of 70 and 280 mL/min, respectively.
A two-dimensional mathematical model with no adjustable parameters was formulated to simulate a mesh reactor for a gas-liquid reacting system. The microstructured mesh reactor consists of parallel metal plates, gaskets, and a microstructured mesh so that passages for gas and liquid phases are formed. First, acetone removal from binary mixtures of acetone-isopropanol was investigated. The model gives good agreement with experimental data. Theoretical and experimental investigation revealed that the amount of acetone removed from the liquid phase is strongly dependent on the ratio of inlet gas and liquid flowrate, FG/FL. For higher FG/FL ratios, more efficient acetone removal from the liquid phase is achieved. If dry nitrogen is used as stripping agent, simultaneous evaporation of acetone and isopropanol results in significant solvent depletion and decrease of liquid flowrate. It was shown that the gas phase becomes saturated with solvent (isopropanol) near the reactor entrance. In order to avoid solvent evaporation which can lead to undesirable concentration of the liquid phase and solvent loss, the gas phase can be saturated with isopropanol before entering the reactor. The reactor model was further used to investigate asymmetric transfer hydrogenation of acetophenone. Due to efficient removal of acetone produced as byproduct during the reaction, high conversion (>90%) can be achieved for residence time of about 10 min, which is much shorter than the time necessary to achieve similar conversion in a laboratory batch reactor. Acknowledgment Financial support by the Foresight LINK Programme is gratefully acknowledged. We would like to thank Dr. John Blacker (Avecia) for providing the catalyst and for useful discussions. Notation
Figure 16. Comparison of a laboratory batch reactor with mesh reactor performance for asymmetric transfer hydrogenation of acetophenone. (a) Conversion as function of liquid phase reaction/residence time. (b) Acetone concentration profile as a function of conversion (case r5, Table 5).
c ) concentration in liquid phase, mol/L jc ) concentration in liquid phase averaged in the cross section as defined by eq 27, mol/L C ) dimensionless concentration as defined in Table 1 D ) diffusion coefficient, m2/s F ) volumetric flowrate, m3/s Fo ) Fourier number as defined in Table 1 L ) reactor length, m M ) molecular weight, kg/mol pvap ) vapor pressure, Pa P ) total pressure, Pa R ) reaction rate, mol/(L s)
Ind. Eng. Chem. Res., Vol. 47, No. 23, 2008 9005 Rg ) universal gas constant, J/(mol K) T ) temperature, K u ) velocity, m/s U ) dimensionless velocity as defined in Table 1 V ) volume, m3 w ) width of gas or liquid passage of contactor, m x ) mole fraction in liquid phase, – X ) conversion, y ) mole fraction in gas phase z ) axial coordinate, m Greek Symbols γ ) activity coefficient Γ ) relative difference in concentration as defined by eq 28 δ ) thickness of gaskets, gas, and liquid films, m ε ) mesh open area, % ζ ) dimensionless axial coordinate, as defined in Table 1 ξ ) dimensionless transverse coordinate, as defined in Table 1 F ) liquid density, kg/m3 τR) residence time for batch reactor, s Φ ) fugacity coefficient Subscripts G ) gas phase ipa ) isopropanol L ) liquid phase M ) mesh Superscripts 0 ) inlet conditions i ) component i G ) gas phase interface ) values at a two-phase interface L ) liquid phase M ) mesh
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ReceiVed for reView December 04, 2007 ReVised manuscript receiVed March 7, 2008 Accepted March 11, 2008 IE071653S
