Novel Procedure for Production of Isopropanol by Transesterification

Article pubs.acs.org/IECR

Novel Procedure for Production of Isopropanol by Transesterification of Isopropyl Acetate with Reactive Distillation Ting Qiu,*,† Pei Zhang,† Jinbei Yang,†,‡ Li Xiao,† and Changshen Ye† †

College of Chemistry and Chemical Engineering, Fuzhou University, Fuzhou, Fujian 350108, China Department of Biological and Chemical Engineering, Fuqing Branch of Fujian Normal University, Fuzhou, Fujian 350300, China

‡

S Supporting Information *

ABSTRACT: A reactive distillation process is proposed for the production of isopropanol by transesterification of isopropyl acetate (IPAc) with methanol using sodium methoxide solution as catalyst. The reaction kinetics experiments were carried out to correlate the parameters in a simple homogeneous kinetic model. On the basis of the fixed point analysis of residue curve maps, the feasibility of a reactive distillation process for the transesterification of isopropyl acetate and methanol was analyzed, and the results showed that the process is feasible under certain conditions. Reactive distillation experiments were set up, and several operation conditions were varied, such as reboiler duty, reflux ratio, space velocity, catalytic loading, and reactant ratio. The results show that the proper reboiler duty, reflux ratio, space velocity, catalytic loading, and MeOH to IPAc molar ratio are 314− 365 kJ/h, 3.0, 0.73 m3/(m3 h), 0.4 wt %, and 2.5, respectively, and the conversion of IPAc is above 99%. Finally, the overall process flowsheet for synthesizing isopropanol is proposed.

1. INTRODUCTION Reactive distillation (RD) is an innovating process which combines both chemical reaction and physical separation into a single unit. The instantaneous product is promptly removed from the reaction section by distillation, which leads to higher conversion and selectivity for equilibrium-limited reactions and consecutive reactions, such as esterifications,1−4 etherifications,5−9 hydrations,10,11 hydrolysis,12−16 and transesterifications.17−19 This helps to reduce investment and operating costs significantly, and may be important for sustainable development because of lower consumption of resources. Despite the fact that reactive distillation was invented in 1921,20 its first industrial application took place in the 1980s.21 RD has been applied to interesting equilibrium reactions in the past. Especially, methyl acetate synthesis and hydrolysis have been investigated extensively, which serve as model systems for reactive distillation processes now.22,23 Isopropanol (IPA) is an excellent organic raw material and industrial solvent in the chemical industry. Mainly, it is used as pharmaceutical intermediates, an intermediate in chemical synthesis, or as an important solvent in the ink, cosmetics, and coatings industries.24 Conventionally, there are two commercial routes to produce isopropanol: direct hydration and indirect hydration of propylene.25 Both routes use propylene and water as raw materials. Indirect hydration is based on a two-step process in which an ester is formed and then hydrolyzed to isopropanol. Diisopropyl ether is the main byproduct. The advantages of indirect hydration are that it is less demanding on the purity of raw material propylene (mass fraction of 50% ∼ 90%) and propylene conversion is over 90%. However, it mainly suffers from the complexity of the process, a low selectivity, high energy requirement, and serious corrosion to equipment because of the use of sulfuric acid as catalyst. Since the 1980s, this route has been gradually phased out in China. Direct hydration of propylene is mainly used in the © 2014 American Chemical Society

manufacture of IPA because it avoids some corrosion and environment problems encountered in the indirect hydration routes. Diisopropyl ether is also the main byproduct. Improvements to the direct hydration routes have been made in recent years. Nevertheless, a complex distillation scheme is still required to recover IPA from the product stream, and the separation of IPA from the azeotropic mixture (IPA and water) is technically difficult and expensive.26,27 Therefore, a reactive distillation process is used to produce isopropanol by transesterification of isopropyl acetate (IPAc) with methanol using sodium methoxide solution as catalyst, which is a simpler and less expensive alternative to the conventional processes. IPAc comes from the reaction of acetic acid and propylene, which has been developed on a scale of 20 000 tons/year by Hunan Zhongchuang Chemical Co., Ltd. in 2006. The reaction has high conversion and selectivity; this reduces the cost of IPAc, which can then be used as a reactant for producing IPA through transesterification. However, no literature data are available for the transesterification of isopropyl acetate with methanol forming methyl acetate and isopropanol (eq 1). isopropyl acetate (IPAc) + methanol (MeOH) ↔ methyl acetate (MeOAc) + isopropanol (IPA)

(1)

This paper describes a systematic approach to a homogeneously catalyzed reactive distillation process for the production of IPA. Thermodynamic analysis of the considered system has been investigated by us.28 The process development starts with the reaction kinetic studies of the transesterification reaction under the conditions expected in the reactive Received: Revised: Accepted: Published: 13881

November 10, 2013 August 9, 2014 August 13, 2014 August 13, 2014 dx.doi.org/10.1021/ie5026584 | Ind. Eng. Chem. Res. 2014, 53, 13881−13891

Industrial & Engineering Chemistry Research

Article

2 K/min to 338.15 K held for 1 min, heat rate of 10 K/min to 353.15 K held for 1 min, heat rate of 25 K/min to 453.15 K. The injected volume was a liquid sample of 0.2 μL. The internal standard method was used with cyclohexanone as an internal standard. 2.4. Kinetic Measurements. The kinetic experiments were performed in a glass three-necked flask with a total volume of 1000 mL. One neck was used to measure the temperature with a thermometer, and another to withdraw samples with a syringe. A reflux condenser was connected to the third neck to avoid any loss of volatile components. A constant-temperature heating magnetic stirrer with a rotational speed of 600 rpm ensured the complete mixing and kept the temperature constant within ±0.1 K. The reactants (isopropyl acetate and methanol) were placed in the three-necked flask with a total volume of 500 mL and heated to the desired temperature. On the basis of the preliminary tests, no reaction took place without any catalyst present. To start the reaction, the homogeneous catalyst sodium methoxide was added to the reaction mixture, and time measurement was started. During the experiment, the liquid samples of about 1 mL were taken out from the reaction mixture by a syringe. The samples were cooled rapidly to 263.15 K to avoid further reaction and analyzed by gas chromatography. During one experiment, above 15 to 20 samples were taken. None of the run side products have been detected. Measurements were performed in the temperature range between 308.15 and 333.15 K. Besides the temperature of the experiments, the initial reactant ratio and the amount of catalyst were varied. All the experiments were continued until chemical equilibrium was reached. 2.5. Reactive Distillation Experiments. 2.5.1. Setup. Figure 1 shows a schematic diagram of the reactive distillation setup. The experiments in a laboratory scale were performed in

distillation column. The reactive kinetic experiments data were used to correlate the parameters in the homogeneous kinetic model. In the second step, the results of the thermodynamic and the kinetic data were used to analyze the reactive residue curve maps (RCMs); then, an RCM-based feasibility analysis were applied to the reactive distillation. Then, the experiments in a reactive distillation column were carried out in different operation conditions (e.g., reboiler duty, reflux ratio, space velocity, catalytic loading, reactant ratio). Finally, on the basis of the reactive distillation, a new process flowsheet for the synthesis of isopropanol was proposed.

2. EXPERIMENTAL SECTION 2.1. Chemicals. Isopropanol (>99.8% weight, analytical grade), methanol (>99.8% weight, analytical grade), and methyl acetate (>99.8% weight, analytical grade) were supplied by Sinopharm Chemical Reagent Co., Ltd. Isopropyl acetate (>99.5% weight, analytical grade) was supplied by Hunan Zhongchuang Chemical Co., Ltd. The purity of these chemicals was determined by gas chromatography. Distilled water was prepared in our laboratory. The system has three pairs of azeotropes, and the azeotropic data are shown in Table 1. Table 1. Boiling Points of Pure Components and Azeotropic Data at 1 atm28 name azeotrope 1 azeotrope 2 azeotrope 3 isopropanol isopropyl acetete methyl acetate methanol

composition

type

boiling point, Tb(K)

azeotropic composition

unstable

326.69

0.66/0.34

MeOAc/ MeOH MeOH/ IPAc IPA/IPAc IPA IPAC

saddle

333.40

0.89/0.11

stable stable stable

353.67 355.39 361.75

0.69/0.31

MeOAc

saddle

330.05

MeOH

saddle

337.85

2.2. Catalyst. The homogeneous catalyst sodium methoxide, a strong alkaline soluble in both methanol and isopropanol, was obtained from Zhejiang Lianshen Chemical Co., Ltd.; the catalyst is a mixture of CH3OH and CH3ONa with a CH3ONa content of about 30 wt % . In the following text, the mass or molar fraction of the catalyst always refers to the pure amount of catalyst, rather than that of the dilution. As a catalyst, sodium methoxide has the advantages of high activity and low consumption. In this reaction system, sodium methoxide can dissolve in methanol well. It is important to realize the continuous reaction distillation process and reach the high conversion rate. 2.3. Analytics. The samples of the kinetic and reactive distillation experiments were both analyzed by a gas chromatography GC7900 (Shanghai TianMei Scientific Instrument Co., Ltd.). Flame ionization detector (FID) was used together with a TM-5 capillary column (0.32 mm × 50 m) with 0.5 μm film thickness. Splitless injection was used, and the carrier gas was nitrogen at a flow-rate of 2 mL/min. Nitrogen (30 mL/min), hydrogen (20 mL/min), and dry air (300 mL/ min) were used as auxiliary gases for the flame ionization detector. The injector temperature and detector temperatures were both set at 523.15 K. The oven was operated at programmed temperature: 328.15 K held for 1 min, heat rate of

Figure 1. Diagram of reactive distillation experimental setup. 13882

dx.doi.org/10.1021/ie5026584 | Ind. Eng. Chem. Res. 2014, 53, 13881−13891


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a vacuum-insulated glass column with an inner diameter of 22 mm. A three-neck round-bottom flask with a temperatureadjustable electric jacket served as a reboiler. The heat duty at the reboiler was controlled by regulating transformer. The distillation column consisted of three parts: stripping section, reactive section, and rectifying section (heights, 600, 1000, and 400 mm, respectively). Stainless steel θ type packings (φ3 × 3 mm) were used as packing materials for all the three sections. The column was insulated by a vacuum jacket, pieces of asbestos cloth, and asbestos ropes simultaneously. Circulating water at a temperature of 278.15K served as a coolant for the condenser located at the top. The reflux ratio was controlled by the reflux splitter, which was a solenoid valve with a multitimer. Constant flow pumps (HL-2S) were employed as feed pumps, and all feeds were measured by determining the mass flow using balances with an accuracy of ±0.1%. Temperatures were measured by using mercury thermometers with an accuracy of 0.1 K. Thermometers were installed both at the three-neck flask in the reboiler and in the top of the column. To achieve the countercurrent flow in the reactive section of the column, the high-boiling reactant was usually fed into the column above the reactive section, and the low-boiling reactant was fed below the reactive section of the column. In this work, methanol was fed into the column below the reactive section and isopropyl acetate together with homogeneous catalyst sodium methoxide was fed above the reactive section of the column. 2.5.2. Procedure. A mixture of isopropanol and methanol was placed into the bottom flask and heated to boiling point. When the distillate appeared at the top, the feed mixture of isopropyl acetate and homogeneous catalyst sodium methoxide were introduced to the upper part of the reactive section and methanol at room temperature was introduced to the lower part of the reactive section by constant flow pumps. At the same time, liquid from the reboiler was removed by another constant flow pump. Then, the continuous reactive distillation operation was started. The liquid level in the reboiler was maintained by adjusting the constant-flux pump. The pump always kept the balance between the feeding and withdrawing production during the process. The liquid samples in the reboiler and decanter at the top of column were pumped out by the means of a syringe, immediately cooled to 263.15 K to avoid further reaction, and analyzed by gas chromatography. Every once in a while, samples at distillate and the bottom were taken out and analyzed, respectively. The continuous reactive distillation operation came to a stable state on condition that the error of analysis results of two successive samples was within 0.1%. Then, the outlet material at distillate and bottom were collected, weighed and analyzed, respectively.

Figure 2. Effect of catalyst loading on the conversion of IPAc. Conditions: catalyst, sodium methoxide; temperature, 313.15 K; MeOH/IPAc feed ratio, 1.0.

further increase of the catalyst loading, the conversion of IPAc changes little. Moreover the increment of catalyst loading leads to an increment of the reaction rate and a decrement of time to reach the reaction equilibration. The reason is that the greater the catalyst loading, the more available actives sites for the transesterification reaction. As a result, the practical catalyst loading is controlled between 0.2 and 0.4 wt %. 3.1.2. Effect of Temperature. The reaction temperature has a critical impact on the kinetic experiments. Activation energy of the IPAc transesterification can be obtained by studying the influence of temperature on the transesterification reaction. During the course of the experiments, the behavior of the IPAc transesterification was determined over the temperature range of 308.15−333.15 K. As can be seen in Figure 3, the reaction of MeOH with IPAc is relatively fast, requiring a period less than 15 min to achieve chemical equilibrium. The increment of temperature is apparently favorable to accelerate the transesterification reaction. However, the IPAc conversion decreases

3. RESULTS AND DISCUSSION 3.1. Reaction Kinetics. 3.1.1. Effect of Catalyst Loading. The transesterification reaction was catalyzed by a homogeneous catalyst, sodium methoxide, soluble in methanol. The effect of the catalyst loading on the transesterification reaction was studied in the experiments. In this paper, catalyst loading is defined as the mass percentage of the catalyst to the total feed mass. The catalyst loading was varied over a range of 0.06−0.4 wt % at the temperature of 313.15 K and MeOH/IPAc molar ratio of 1.0, and the results are shown in Figure 2. Figure 2 shows that the catalyst loading has considerable effect on the IPAc conversion until the catalyst loading of 0.2 wt %. With

Figure 3. Effect of temperature on the conversion of IPAc. Conditions: catalyst, sodium methoxide; MeOH/IPAc feed ratio, 1.0; catalyst loading, 0.3 wt %. The dots represent experiment results; the lines represent model results. 13883



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Table 2. Thermodynamic Dataa for the Isopropanol System28

weakly from 68% to 65% as the reaction temperature increases from 308.15 to 333.15 K. This indicates that the IPAc transesterification reaction is a marginal exothermal reaction. 3.1.3. Effect of Feed Ratio. Figure 4 illustrates the IPAc conversion variation with reaction time at the temperature of

Antoine Equation Coefficients Ai, Bi, and Ci Antoine constants component

A

B

C

temperature range (K)

methyl acetate(1) methanol(2) isopropanol(3) isopropyl acetate(4)

6.25449 7.09498 6.86634 6.46612

1189.608 1521.23 1360.183 1436.530

−50.035 −39.18 −75.557 −39.485

260−365 338−487 325−362 235−362

NRTL Parameters MeOAc(1) + MeOH(2) MeOAc(1) + IPA(3) MeOAc(1) + IPAc(4) MeOH(2) + IPA(3) MeOH(2) + IPAc(4) IPA(3) + IPAc(4) a

Aji (J/mol)

αij

1466.6 194.9 1616.8 1507.9 2525.75 1160.468

1617.8 1926.4 −1192.3 −1706.1 1767.46 632.2797

0.3 0.3 0.3 0.3 1.03 0.2

Values are calculated as follows:

ln P s = A − Figure 4. Effect of MeOH/IPAc feed ratio on the conversion of IPAc. Conditions: catalyst, sodium methoxide; temperature, 313.15 K; catalyst loading, 0.3 wt %. The dots represent experiment results; the lines represent model results.

Aij (J/mol)

B T+C

P s (kPa), T (K)

n

ln γi =

∑ j = 1 τjiGjixj n

∑k = 1 Gkixk

n

+

∑ j=1

n ⎛ ∑k = 1 xkτkjGkj ⎞ ⎜ ⎟ τ − n ij n ∑ j = 1 Gkjxk ⎜⎝ ∑k = 1 Gkjxk ⎟⎠

xjGij

where

313.15 K and the catalyst loading of 0.3 wt % under different feed molar ratio of MeOH to IPAc. As can be seen in Figure 4, the equilibrium conversion is rather sensitive to feed molar ratio. It can be seen that the equilibrium conversion of IPAc increased with the increase of the feed molar ratio of MeOH to IPAc. The equilibrium conversion of IPAc decreases from 0.913 to 0.665 when the feed ratio was changed from 3:1 to 1:1, and the time required to reach equilibrium conversion ranged from 15 to 120 min. Nevertheless, a suitable feed ratio should be determined in consideration of the energy requirement. 3.1.4. Chemical Equilibrium Constant. The molar-based chemical equilibrium constant of the transesterification reaction is defined by the following equation:29 Ka =

∏ (aieq)v

i

=

∏ (xieq)v ∏ (γieq)v i

i

= K xK γ

τij =

A ij‐A ji RT

Gij = exp(− αijτij)

A ij (J/mol)

According to eq 4, the reaction enthalpy can be calculated from the slope of the line that describes the relationship between the experimental values of ln Ka and 1/T. A plot of ln Ka versus 1/T is shown in Figure 5. As can be seen in Figure 5, the equilibrium constant as a function of temperature can be expressed by eq 5. ln K a = 7868.62/RT − 1.6395

(5)

Equation 5 describes the best fit of the experimentally obtained equilibrium constants. From eq 5, we estimate the

(2)

As can be seen from the above analysis, the transesterification of IPAc with MeOH catalyzed by sodium methoxide is a relatively fast reaction. This allowed us to calculate the equilibrium constant Ka from the equilibrium composition through eq 2. In this work, the liquid-phase activity coefficients γi were calculated with the nonrandom two-liquid(NRTL) model. The required thermodynamic data for the reaction including NRTL interaction parameters are presented in Table 2. The chemical equilibrium constant is a function only of temperature; the relation is presented in eq 3. d ln K a /dT = ΔHr0/RT 2

(3)

ΔH0r

where, is the standard enthalpy of reaction and T is the absolute temperature. When a small temperature interval is investigated, the standard enthalpy of the reaction can be assumed to be constant. Therefore, ln K a = ln K a(T 0) − (ΔHr0/R )(1/T − 1/T 0)

Figure 5. Temperature dependence of the chemical equilibrium constants: ln Ka versus 1000/T (■) and the best fit of our data ().

(4) 13884



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standard enthalpy of the reaction (ΔH0r ) to be −7.87 kJ/mol. The standard enthalpy of reaction can also be calculated from the standard enthalpy of formation, and the standard enthalpies of formation for the reactants and products in the liquid phase are listed in Table 3.30 Both the calculated and experimental Table 3. Standard Enthalpies of Formation component

ΔH0f (kJ/mol)

methyl acetate methanol isopropanol isopropyl acetate

−445.8 −239.1 −318.1 −518.9

values of the standard enthalpy of reaction are reported in Table 4. The value of ΔH0r is very small, which means there is Figure 6. Arrhenius diagram of the rate constants for the reaction of IPAc with MeOH k1 (■) and backward reaction k−1(●) of the homogeneous catalyzed reaction. The lines represent the results of the linear regression.

Table 4. Standard Enthalpies of Reaction source

ΔH0r (kJ/mol)

calculated from thermodynamic properties, Table 2 obtained by linear regression, Figure 5

−5.9 −7.87

are shown in Figures 3 and 4. There is very good agreement between the calculated values and experimental data. 3.2. Residue Curve Model. 3.2.1. Residue Curve Map. The residue curve map was introduced as a very useful tool for the conceptual design and analysis of both nonreactive and reactive distillation systems.31 It has been used by many researchers, such as Foucher et al.,32 Fidkowski et al.,33 Venimadhavan et al.,34 Ung and Doherty,35 Thiel et al.,36 Qi et al.,37 and so on. RCM represents the dynamic composition of the liquid phase in a batch RD process.38 The analysis of the location and stability of the singular points in a RCM generate valuable information on the target product of an RD process. The RCM model equation can be written as follows (for the detailed derivation of the model, see Venimadhavan et al.34 and Huang et al.39):

nearly no temperature dependence of the equilibrium constant. It also can be seen that the transesterification reaction is slightly exothermic because the equilibrium constant Ka decreases with increasing temperature. 3.1.5. Kinetic Model Parameters. The kinetic experimental data were used to establish a kinetic model, which could be used in the theoretical prediction of residue curve map. The kinetic constants have nothing to do with the catalyst loading, as in the previous studies; the catalyst loading should be greater than 0.2 wt %, and the calculated values are based on catalyst loading of 0.3 wt %. Transesterification reaction is regarded as a reversible second-order reaction. The kinetic model can be written as follows: r = −daIPAc /dt = k1aIPAcaMeOH − k −1aMeOAcaIPA

dxi /dt = (xi − yi ) + (k1/k1,ref )(vi − vT xi)Da 9

(6)

In eq 6, activities were used instead of concentrations or mole fractions, and ai is the activity of component i (with ai = γixi). The kinetic parameters k1 and k−1 can be fitted with the experimental data by numerically integrating the kinetic equations using a fourth-order Runge−Kutta method, and minimization of the mean square deviation between the calculated and the experimental mole fractions using the Nelder−Mead simplex method. The temperature dependence of the rate constant can be expressed by the following Arrhenius law. ki = ki0 exp( −EA , i /RT )

(8)

where xi and yi are the liquid and vapor phase mole fractions, respectively; vi is the stoichiometric coefficient of component i, and vT is the total mole change of reaction. The quantity k1,ref is the forward rate constant of reaction rate at the reference temperature, which is generally the lowest boiling point of the system. Da, the Damköhler number, is always regarded as a constant. Da = 0 represents no reaction taking place in the system, which equals a simple distillation process. Da → ∞ represents the reaction reaching chemical equilibrium, which equals a reactive distillation process of balance control. Da ∈ (0, ∞) represents the reaction not reaching chemical equilibrium, which is equal to a reactive distillation process of kinetic control. The 9 denotes the dimensionless reaction rate depending on the liquid phase composition, expressed in terms of activities ai = zγixi.

(7)

Four adjustable parameters and EA,−1) were fitted for the second-order model by the experimental kinetic data using eq 7, and the values are listed in Table S1 in Supporting Information. The results from the overall fit of the four parameters to all experimental data are shown in an Arrhenius diagram (as seen in Figure 6). From Figure 6, it can be seen that the temperature dependence of the rate constants can be described by Arrhenius law and that the simple homogeneous model is able to reproduce the kinetic data. The correlated parameters were taken back to the kinetic model to calculate the conversion of IPAc, and the calculated values were compared with the experimental data. The results (k01, k0−1,EA,1

9 = r / k1

(9)

where r is the rate of the reaction, which is calculated according to eq 6. 3.2.2. Analysis of Residue Curve Map. When the results of thermodynamic and kinetic parameters that are presented in this work were combined, the RCMs were calculated at different Da numbers by using eq 8, as shown in Figure 7. 13885



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IPA and Az. 3. As a result, if the conversion of the IPAc is low, the subsequent purification of IPA will be very difficult because there is an azeotrope between IPAc and IPA, and the design flowsheet of the overall process will be very complicated. Figure 8 shows the residue curve map of MeOAc−MeOH−IPA

Figure 8. RCM for methyl acetate−methanol−isopropanol system at P = 1 atm (○, unstable node; □, saddle node; ●, stable node).

system at 1 atm. As can be seen in Figure 8, all the residue curves originate from the MeOAc−MeOH azeotrope and end to IPA. Therefore, if there is no IPAc in the system, pure IPA can be obtained easily. Consequently, high purity IPA can be successfully produced through the reactive distillation technology when IPAc reacts completely in the reaction zone and a stripping section is set under the reaction section. 3.3. Reactive Distillation. During the experimental process, several operation conditions were varied: reboiler duty, reflux ratio, space velocity, catalytic loading, and feed molar ratio. The experimental method manipulated only one variable at a time by holding all the other variables constant. As a result, it does not guarantee the global optimal design, yet the approximate optimal design of the RD column can be achieved by this method.40 According to the results of the RCM analysis, we know that IPAc should react completely as much as possible so that the high purity of IPA can be obtained. Therefore, MeOH is excessive in order to increase isopropyl acetate conversion in the reactive distillation experiments. 3.3.1. Material Balance of Column. The material balance of the column is shown in Table S2 in Supporting Information. The reactive distillation operation condition was set: reboiler duty was 365 kJ/h, reflux ratio was 3, space velocity was 0.73 m3/(m3 h), catalytic loading was 0.4 wt %, MeOH to IPAc molar ratio was 2.7. As can be seen from Table S2 in Supporting Information, the total mass loss in the RD column exhibited minor loss (0.3%), which indicated that the RD column is mass balanced. Moreover, Table S2 shows that the conversion of isopropyl acetate and isopropanol yield is basically the same because of the absence of side reactions. 3.3.2. Effect of Reboiler Duty. The effect of the reboiler duty was investigated by fixing the reflux ratio, space velocity, catalytic loading, and feed molar ratio. Figure 9 shows the effect

Figure 7. RCMs for isopropanol system at P = 1 atm; (a) Da = 0 and (b) Da = 1.0 (○, unstable node; □, saddle node; ●, stable node).

Figure 7a shows the RCM for nonreactive distillation (Da =0) at 1 atm. There are three nonreactive binary azeotropes in the RCM. The azeotrope between MeOAc and MeOH (Az. 1) is the unstable node; the azeotrope between IPAc and MeOH (Az. 2) is a saddle node, whereas the azeotrope between IPAc and IPA (Az. 3) is a stable node. Pure MeOAc and MeOH are also saddle nodes, whereas IPAc and IPA are stable nodes. Most residue curves originate from the unstable node (Az. 1) and end at IPAc; a few of them originate from the unstable node (Az. 1) and end at IPA, and fewer of them originate from the unstable node (Az. 1) and end at IPAc-IPA azeotrope (Az. 3). The other residue curves originate from the unstable node, first approaching the saddle points, and finally converge to the stable node. The information on all singular points is listed in Table 1. The RCM model equation was used to predict residue curve maps at lower values of Da equal to 0.5, 1.0, and 2.0, while the structure of the residue curve map is qualitatively the same for all values of Da (e.g., Da = 1.0 in Figure 7b); therefore, we predict that IPA can be successfully produced by reactive distillation in both the equilibrium-controlled and kinetically controlled regimes. It also can be seen from Figure 7b that most of the residue curves end at IPAc and only a few of them end at 13886



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Figure 9. Effect of reboiler duty on (a) IPAc conversion and (b) composition profile. Conditions: catalyst, sodium methoxide; reflux ratio, 2.0; space velocity, 0.73 m3/(m3 h); catalytic loading, 0.4 wt %; MeOH/IPAc feed ratio, 2.5.

Figure 10. Effect of reflux ratio on (a) IPAc conversion and (b) composition profile. Conditions: catalyst, sodium methoxide; reboiler duty, 365 kJ/h; space velocity, 0.73 m3/(m3 h); catalytic loading, 0.4 wt %; MeOH/IPAc feed ratio, 2.5.

of reboiler duty on the IPAc conversion and composition profile. As shown in Figure 9a, when the reboiler duty of the RD column is increased from 186 to 602 kJ/h, the conversion of IPAc significantly increases from 73.7% to the maximum value of 99.5%. Increasing the reboiler duty leads to an increase of the conversion of IPAc until the reboiler duty of 314 kJ/h. With further increase of the reboiler duty, the conversion of IPAc decreases. Moreover, Figure 9b shows that with an increase in the reboiler duty from 186 to 314 kJ/h, the mass fraction of IPAc at the bottom of the RD column decreases and the mass fraction of IPAc at the top of RD column is close to zero. As a result, the transesterification reaction is favored in the RD column, and the conversion of IPAc increases. However, when the reboiler duty reaches 314 kJ/h, there is no IPAc in the bottom and the rate of the transesterification reaction then decreases; meanwhile, the mass fraction of IPAc at the top of RD column significantly increases. from 0.07% to 18.6%. A possible reason is that with higher reboiler duty, MeOH and IPA were streamed out from the column bottom to the catalytic section, inhibiting the forward reaction. On the other hand, if the reboiler duty was too small, IPAc can not vapor to the catalytic section from residue. Therefore, the optimum reboiler duty is between 314 and 365 kJ/h, and the conversion of IPAc is above 98%. 3.3.3. Effect of Reflux Ratio. The effect of the reflux ratio of the RD column is illustrated in Figure 10 using fixed reboiler

duty, space velocity, catalytic loading, and feed molar ratio. Figure 10 shows the effect of the reflux ratio on the IPAc conversion and composition profile. Reflux ratio is an important factor for the RD column duty and specification. Therefore, it needs to be optimized. It can be seen from Figure 10a that the conversion of IPAc increases from 82.2% to 98.9% with the increase of the reflux ratio from 1.0 to 5.0. Increasing the reflux ratio leads to an increase of the conversion of IPAc until the reflux ratio reaches 3.0. With further increase of the reflux ratio, the conversion of IPAc remained unchanged. Figure 10b shows that by increasing the reflux ratio, the mass fraction of IPAc at the top of the RD column first decreased and then remained unchanged; the mass fraction of IPAc at the bottom of RD column is close to zero. In this case, the rates of all feeds and outlets are constant in the column; therefore, with the increase of the reflux ratio, the loading of the reflux stream increases and more unconverted reactant IPAc recycles to the reaction zone in the column, promoting the transesterification reaction to the positive direction. However, further increase in the reflux ratio leads to increase of the column duty but no further increase in the conversion of IPAc. Therefore, the optimum reflux ratio is 3.0 for this reaction under the experimental conditions. 3.3.4. Effect of Space Velocity. Figure 11 shows the effect of space velocity of the RD column on the IPAc conversion and composition profile. In this paper, space velocity is defined as 13887



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Figure 11. Effect of space velocity on (a) IPAc conversion and (b) composition profile. Conditions: catalyst, sodium methoxide; reboiler duty, 365 kJ/h; reflux ratio, 3.0; catalytic loading, 0.4 wt %; MeOH/ IPAc feed ratio, 2.7.

Figure 12. Effect of catalytic loading on (a) IPAc conversion and (b) composition profile. Conditions: catalyst, sodium methoxide; reboiler duty, 365 kJ/h; reflux ratio, 3.0, space velocity, 0.73 m3/(m3 h); MeOH/IPAc feed ratio, 2.5.

the ratio of the volume flow of methanol and the volume of the reaction section (m3/m3 h). As can be seen in Figure 11a, the conversion of IPAc first increased and then decreased when the space velocity increased from 0.39 to 1.1 m3/m3 h. For the reactive distillation operation of the setup in this study, the space velocity should not be higher than 1.5 m3/(m3 h) to avoid flooding in the column. When the space velocity is too small, the contact area of vapor−liquid phase in the RD column decreases, leading to reduction in the catalytic efficiency and decline in the conversion of IPAc. When the space velocity is too high, the vapor and liquid load is increased too high and the residence time is too short to reach physical and chemical equilibrium. Because of the fast space velocity, IPAc has not time to react with MEOH and some unconverted reactants come out from the top and bottom of the RD column; the conversion of IPAc declines. As a result, the amount of IPAc at the top and bottom of the RD column is first decreased and then increased, as shown in Figure 11b. By comprehensive consideration of the different values, the space velocity of 0.73 m3/(m3 h) was deemed to be appropriate. 3.3.5. Effect of Catalytic Loading. The effect of the catalytic loading iss illustrated in Figure 12 at the fixed reboiler duty, reflux ratio, space velocity, and feed molar ratio. Figure 12 shows the effect of catalytic loading on the IPAc conversion and composition profile. As shown in Figure 12a, the conversion of IPAc increased when the catalytic loading increased from 0.07

to 0.8 wt %. However, with the increase of the catalytic loading from 0.4 to 0.8 wt %, the conversion of IPAc did not change significantly. Figure 12b shows that by increasing the catalytic loading, the mass fraction of IPAc at the top and bottom of RD column first decreased then remained unchanged and close to zero. This result indicates that an increase of the catalytic loading leads to increase of the transesterification reaction, resulting in the decrease of IPAc at the top and bottom of the RD column. In fact, if there is too much catalyst, this can easily lead to precipitation and fouling of the catalyst, resulting in increased recovery cost of the catalyst. The appropriate catalytic loading for this reaction was 0.4 wt %. 3.3.6. Effect of Feed Molar Ratio. The effect of the MeOH− IPAc molar ratio on IPAc conversion and composition profile was studied by fixing the reboiler duty, reflux ratio, space velocity, and catalytic loading, the results are shown in Figure 13. As shown in Figure 13a, the conversion of IPAc increases from 92.2% to 99.1% with the increase of the MeOH−IPAc molar ratio in a feed ratio range of 1.5−3.0. Increasing the MeOH to IPAc molar ratio leads to an increase of the conversion of IPAc until the molar ratio reaches 2.5. With further increase of molar ratio, the conversion of IPAc remained unchanged. Figure 13b shows that by raising the MeOH−IPAc molar ratio, the mass fraction of IPAc at the bottom of the RD column first decreases and then remained unchanged; the mass 13888



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lead to an increased amount of MeOH and the heat duty of following purification but no further increase in the conversion of IPAc. Under our experimental conditions, the MeOH−IPAc molar ratio in a feed ratio of 2.5 is the most suitable. From the experimental investigation done in this research work, the optimum reactive distillation operation conditions can be concluded as follows: reboiler duty is between 314 and 365 kJ/h; reflux ratio is 3.0; space velocity is 0.73 m3/(m3 h); catalytic loading is 0.4 wt %; and MeOH−IPAc molar ratio is 2.5. Under the optimum reactive distillation operation conditions, the conversion of IPAc is above 99%. To verify the reliability of the experimental data, a repeatability experiment was performed. Table S3 in Supporting Information shows the results of the repeatability experiment under the optimum operational conditions. 3.4. Concept of the Overall Process. According to the above experiments and analysis, it is reasonable and feasible for the production of isopropanol by transesterification of isopropyl acetate with methanol using sodium methoxide solution as catalyst. Because excess MeOH is required to drive the IPAc to react completely, some MeOH exists in both the overhead product and bottom product. Therefore, separating MeOH from the bottom product is necessary for achieving higher IPA purity. Figure 14 shows the overall process flowsheet of IPA synthesis, which consists of the RD column and product purification sections. On the top of the RD column, the mixture of MeOAc and MeOH and very little IPA and IPAc as byproduct can be obtained; at the bottom of the RD column, the mixture of IPA, MeOH, and sodium methoxide can be obtained by using the evaporator to recover the catalyst sodium methoxide at the bottom of RD column, then separating and recovering MeOH through a distillation column (T1). Finally, the bottom product of T1 is fed into another distillation column (T2) in the downstream for further purification and decolorization of IPA.

Figure 13. Effect of MeOH−IPAc molar ratio on (a) IPAc conversion and (b) composition profile. Conditions: catalyst, sodium methoxide; reboiler duty, 365 kJ/h; reflux ratio, 3.0; space velocity, 0.73 m3/(m3 h); catalytic loading, 0.4 wt %.

4. CONCLUSION This paper demonstrates the feasibility of the reactive distillation process to produce IPA via the transesterification of IPAc with MeOH. The reaction kinetics of IPA synthesis was studied, and the kinetic parameters for the simple homogeneous model based on activities were determined. An RCMbased feasibility analysis has been applied to the reactive distillation, which indicates that IPA can be successfully produced by reactive distillation in both the equilibrium-

fraction of IPA at the bottom of the RD column decreases with the increase of the molar ratio of the MeOH to IPAc. With the increasing of the MeOH to IPAc molar ratio, the amount of the MeOH increases and more MeOH recycles to the reaction zone in the column and promotes the transesterification reaction to the positive direction. Considering the actual production process, further increase in the feed molar ratio will

Figure 14. Overall process flowsheet for the synthesis of IPA. 13889



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(12) Abella, L. C.; Gaspillo, P. A. D.; Itoh, H.; Goto, S. Dehydration of tert-butyl alcohol in reactive distillation. J. Chem. Eng. Jpn. 1999, 32 (6), 742−746. (13) Qi, Z.; Sundmacher, K. Multiple product solutions of tert-butyl alcohol dehydration in reactive distillation. Ind. Eng. Chem. Res. 2006, 45 (5), 1613−1621. (14) Sander, S.; Flisch, C.; Geissler, E.; Schoenmakers, H.; Ryll, O.; Hasse, H. Methyl acetate hydrolysis in a reactive divided wall column. Chem. Eng. Res. Des. 2007, 85 (1), 149−154. (15) Pöpken, T.; Steinigeweg, S.; Gmehling, J. Synthesis and hydrolysis of methyl acetate by reactive distillation using structured catalytic packings: Experiments and simulation. Ind. Eng. Chem. Res. 2001, 40 (6), 1566−1574. (16) Lin, Y. D.; Chen, J. H.; Cheng, J. K.; Huang, H. P.; Yu, C. C. Process alternatives for methyl acetate conversion using reactive distillation. 1. Hydrolysis. Chem. Eng. Sci. 2008, 63 (6), 1668−1682. (17) Fang, Y. J.; Xiao, W. D. Experimental and modeling studies on a homogeneous reactive distillation system for dimethyl carbonate synthesis by transesterification. Sep. Purif. Technol. 2004, 34 (1), 255− 263. (18) Luo, H. P.; Xiao, W. D. A reactive distillation process for a cascade and azeotropic reaction system: Carbonylation of ethanol with dimethyl carbonate. Chem. Eng. Sci. 2001, 56 (2), 403−410. (19) Tobias, K.; Johannes, H.; Alexander, N.; Andrej, G. Transesterification of dimethyl carbonate with ethanol to form ethyl methyl carbonate and diethyl carbonate: A comprehensive study on chemical equilibrium and reaction kinetics. Ind. Eng. Chem. Res. 2011, 50 (19), 11073−11086. (20) Backhaus, A. A. Continuous process for the manufacture of esters. U.S. Patent, 1,400,849, 1921. (21) Huss, R. S.; Chen, F.; Malone, M. F.; Doherty, M. F. Reactive distillation for methyl acetate production. Comput. Chem. Eng. 2003, 27 (12), 1855−1866. (22) Tong, L.; Chen, L.; Ye, Y.; Qi, Z. Analysis of intensification mechanism of auxiliary reaction on reactive distillation: Methyl acetate hydrolysis process as example. Chem. Eng. Sci. 2014, 106, 190−197. (23) Han, S. J.; Jin, Y.; Yu, Z. Q. Application of a fluidized reactiondistillation column for hydrolysis of methyl acetate. Chem. Eng. J. (Amsterdam, Neth.) 1997, 66 (3), 227−230. (24) Inokuma, K.; Liao, J. C.; Okamoto, M.; Hanai, T. Improvement of isopropanol production by metabolically engineered Escherichia coli using gas stripping. J. Biosci. Bioeng. 2010, 110 (6), 696−701. (25) Xu, Y.; Chuang, K. T.; Sanger, A. R. Design of a process for production of isopropyl alcohol by hydration of propylene in a catalytic distillation column. Chem. Eng. Res. Des. 2002, 80 (6), 686− 694. (26) Mahajani, S. M.; Sharma, M. M.; Sridhar, T. Direct hydration of propylene in liquid phase and under supercritical conditions in the presence of solid acid catalysts. Chem. Eng. Sci. 2002, 57 (22), 4877− 4882. (27) Wang, S. J.; Wong, D. S. H. Control of reactive distillation production of high-purity isopropanol. J. Process Control 2006, 16 (4), 385−394. (28) Xiao, L.; Wang, Q. L.; Wang, H. X.; Qiu, T. Isobaric vapor− liquid equilibrium data for the binary system methyl acetate + isopropyl acetate and the quaternary system methyl acetate + methanol + isopropanol + isopropyl acetate at 101.3 kPa. Fluid Phase Equilib. 2013, 344, 79−83. (29) Ma, P. S. Chemical Engineering Thermodynamics; Chemical Industry Press: Beijing, 2009. (30) Poling, B. E.; Prausnitz, J. M.; O’Connell, J. P. The Properties of Gases and Liquids, 5th ed.; McGraw-Hill: New York, 2001. (31) Song, W.; Venimadhavan, G.; Manning, J.; Malone, M. F.; Doherty, M. F. Measurement of residue curve maps and heterogeneous kinetics in methyl acetate synthesis. Ind. Eng. Chem. Res. 1998, 37 (5), 1917−1928. (32) Foucher, E. R.; Doherty, M. F.; Malone, M. F. Automatic screening of entrainers for homogeneous azeotropic distillation. Ind. Eng. Chem. Res. 1991, 30 (4), 760−772.

controlled and kinetically controlled regimes. Several reactive distillation experiments in a laboratory scale column were performed with variations in several operating conditions, including the reboiler duty, reflux ratio, space velocity, catalytic loading and feed ratio. Finally, the overall process flowsheet of IPA synthesis was proposed, which can serve as the basis for the industrialized scale-up of the process.

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ASSOCIATED CONTENT

S Supporting Information *

Kinetic parameters for the kinetic model (Table S1), material balance of the column (Table S2), and results of repeatability experiment under optimum operation conditions (Table S3). This material is available free of charge via the Internet at http://pubs.acs.org.

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AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors thank the National Natural Science Foundation of China (21176049) and Natural Science Foundation of Fujian Province (2011J01038) for financial support.

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REFERENCES

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Novel Procedure for Production of Isopropanol by Transesterification

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