High Conversion of Methyl Acetate Hydrolysis in a Reactive Dividing

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High conversion of methyl acetate hydrolysis in a reactive dividing wall column by weakening the self-catalyzed esterification reaction Xiaoda Wang, Hongxing Wang, Jinyi Chen, Weiyue Zheng, and Ting Qiu Ind. Eng. Chem. Res., Just Accepted Manuscript • DOI: 10.1021/acs.iecr.7b01907 • Publication Date (Web): 21 Jul 2017 Downloaded from http://pubs.acs.org on July 31, 2017

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Industrial & Engineering Chemistry Research

High conversion of methyl acetate hydrolysis in a reactive dividing wall column by weakening the self-catalyzed esterification reaction

Xiaoda Wang, Hongxing Wang, Jinyi Chen, Weiyue Zheng, Ting Qiu * Engineering Research Center of Reactive Distillation Technology (Fujian Province), School of Chemical Engineering, Fuzhou University, Fuzhou 350108, Fujian, China Corresponding Author: Ting Qiu; Email address: [email protected]

ABSTRACT We would show the superiority of reactive dividing wall column (RDWC) to the single reactive distillation (RD) column in improving the conversion of reactant, taking the hydrolysis of methyl acetate (MA) as example. It is difficult to achieve above 99% conversion for MA in a traditional reactive distillation column (TRDC), due to the existence of self-catalyzed methanol (MeOH)-acetic acid (HAc) esterification reaction in the column bottom. In this work, more than 99% conversion of MA hydrolysis was realized experimentally in a RDWC by separate MeOH from hydrolysis mixture. The effects of several operation parameters on hydrolysis conversion were systematically investigated, including feed water-MA mole ratio, heat duty, mole flow rate of feed MA and vapor distribution. The simulation results by Aspen Plus showed that RDWC has several better performances in MA hydrolysis than TRDC, including lower energy consumption, less water-MA mole ratio and 1

ACS Paragon Plus Environment


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larger production capacity. With the increase of MA conversion, the superiorities became more obvious. All these superiorities should be contributed to the weaker self-catalyzed MeOH-HAc esterification reaction in the RDWC.

KEYWORDS: reactive dividing wall column; hydrolysis; methyl acetate; selfcatalyzed reaction.

1 INTRODUCTION Reactive dividing wall columns (RDWC) can be understood as a combination of reactive distillation and dividing wall columns. During the last ten years, the research on RDWC has shown its ability to reduce the energy consumption and equipment cost 1-10

. In present work, we would not only demonstrate its ability to save energy, but

also its superiority to the single RD column in improving the conversion of reactant. The hydrolysis of methyl acetate (MA) would be taken as example.

MA is produced in great amounts as byproduct in purified terephthalic acid and polyvinyl alcohol plants. Due to its low additional value as solvent, MA is always hydrolyzed to the more valuable chemicals methanol (MeOH) and acetic acid (HAc). The fixed-bed reactor (FBR) technology is the first industrialized technology for MA hydrolysis. The equilibrium conversion of MA is very low in the FBR, since the MA hydrolysis is a reversible reaction with small equilibrium constant. It is impossible to separate unreacted MA from reaction mixture via a distillation column due to the existence of the minimum azeotrope between MeOH and MA. The FBR for MA

2


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hydrolysis was followed by several distillation columns to separate and purify the reactants and products. Figure 1a shows the flow sheet for the FBR technology.

Reactive distillation (RD) is a process where reaction and separation take place instantaneously in the same zone of a single distillation column. This technology has been considered to be an excellent alternative for improving the conversion of equilibrium-limited chemical reaction, since the products can be separated continuously from reactants as the reaction proceeds 11-12. The conventional multiunit reactor/column/recycle system is also an alternative to equilibrium-limited chemical reaction.

It

has

been

proved

that

replacing

the

conventional

multiunit

reactor/column/recycle system with a RD column could reduce costs by a factor of 23, due to more effective energy utilization and less equipment investment

13

. The

liquid-phase reaction system where the reactant is the low-boiler is most unfavorable system for RD, since the volatilization of reactant leads to its concentration reduction in liquid. Nevertheless, the saving of energy and equipment costs could even be realized for such an unfavorable reaction system

13

. Fuchigami

14

was the first to

apply the RD technology to hydrolyze MA. It is impossible to conduct MA hydrolysis in a reactive distillation column with stoichiometric methyl acetate and water as feeds, methanol as overhead and acetic acid as bottom, since the reaction could not make the minimum MA-MeOH azeotrope disappear. Therefore, the RD column they adopted was operated at total reflux on the top with the products and unreacted reactants withdrawn from the bottom. It was reported that the conversion of MA could be remarkably improved and the heat requirement was estimated to be about half that of the FBR technology. Such a RD column is always followed by a distillation column to separate MeOH from hydrolysis mixture, as illustrated in Figure 1b. We defined 3



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this RD column as traditional reactive distillation column (TRDC) and the flow sheet in Figure 1b as TRDC technology in present work, for the convenience of statement. Xiao et al. 15 investigated in detail the effects of several operational parameters on the MA conversion, following the work of Fuchigami 14. Their experiment results showed that a high feed ratio of water to MA is necessary for the high MA conversion 15. The work of Kim et al. 16 showed that considerable MA conversion could also be achieved with the TRDC even though MA-MeOH azeotrope was used as feed instead of pure MA. Lee 17 and Lin et al. 18 used a fixed bed reactor as the reflux drum of the TRDC to reduce the total cost. Instead of operating the RD column at total reflux, Han et al. 19

withdrew the products from both ends of the column. Pöpken et al. 20 found that the

equilibrium-stage concept is sufficient to model MA hydrolysis in such a RD column. In comparison with the RD column Han et al. 19 applied, the TRDC has the advantage of improving the MA conversion but the disadvantage of losing the separating capacity

21

. In addition, a RD process of MA hydrolysis intensified by MeOH

dehydration was proposed to avoid the separation of hydrolysis mixture

22

, but its

industrial operability need to be further examined.

Among the MA hydrolysis technologies mentioned above, the one Fuchigami proposed

14

attracted the most attention due to its prominent industrial applicability.

However, there are two shortcomings for this configuration. The first one is that the separation of hydrolysis mixture remains an energy-intensive process, due to the total reflux operation and the excessive supply of water. Several energy-saving technologies have been applied to TRDC and considerable energy saving was achieved 23-26. The other shortcomings is that high MA conversion (>99%) is difficult to be realized 14-15, 27. The low MA conversion leads to the complex separation process 4


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of the hydrolysis mixture. Up to now, the effective solution to this shortcoming is limited. In present work, the RD column Fuchigami 14 designed was coupled with the MeOH distillation column following it, forming a RDWC, to achieve high MA conversion (>99%). Figure 1c illustrates the schematic of this RDWC, which has divided overhead section and common bottom section. Although Li et al.

25

has

studied the issues of design and control for MA hydrolysis in such a RDWC by Aspen Plus, they did not pay any attention to its ability to realize high MA hydrolysis conversion.

In this work, we firstly analyzed the reason for the difficulty in realizing high MA conversion (>99%) in TRDC and explained why the RDWC could be applied to overcome this difficulty. Then, the effects of several operation parameters on MA conversion were investigated experimentally to verify the feasibility of achieving more than 99% MA conversion in a RDWC. Finally, the superiority of RDWC to TRDC in MA hydrolysis was showed by the simulation results of Aspen Plus.

2 Problem analysis and solution It mainly contains water, MeOH and HAc in the hydrolysis mixture withdrawn from the bottom of the TRDC at >95% conversion of MA

14

. The hydrolysis mixture

contains 20 to 30wt% HAc at different feed mole ratio of water to MA, with the temperature ranging from 80 to 90℃ 14. Since the MA hydrolysis to MeOH and HAc is an acid-catalyzed reversible reaction, the high concentration of HAc would inevitably catalyze the esterification reaction of MeOH with HAc in the hydrolysis mixture due to the high concentration of MeOH and HAc and the high temperature, 5



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resulting in the formation of MA

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14, 28

. Although the self-catalyzed reaction of HAc

and MeOH proceeds very slowly, the formation amount of MA could not be neglected due to the long residence time of hydrolysis mixture in the bottom of TRDC. Therefore, it is difficult to realize the complete conversion of MA in TRDC due to the existence of the self-catalyzed MeOH-HAc esterification reaction.

Fuchigami pointed that 99.5% MA conversion could be obtained in a lab-scale RD column by greatly reducing the volume of the column bottom

14

. However, it is

impractical to greatly reduce the bottom volume for a plant-scale reactive distillation column, since the small bottom volume is unfavorable for the stable liquid-level control. Another method to increase the MA conversion is to improve the feed mole ratio of water to MA 15. It could be deduced from the work of Xiao et al. 15 that waterMA mole ratio higher than 7 is required to achieving more than 99% MA conversion at their operation conditions. However, high water-MA mole ratio would lead to the high energy consumption of the subsequent water-HAc separation process. Therefore, more exercisable and economical methods should be developed to suppress the selfcatalyzed MeOH-HAc esterification reaction in TRDC to realize the high hydrolysis conversion (>99%) of MA.

An effective method to suppress the self-catalyzed MeOH-HAc esterification reaction in TRDC is to remove MeOH or HAc from the hydrolysis mixture. There is no azeotrope in the ternary system of MeOH-water-HAc, as shown in Table 1. Additionally, MeOH has the minimum boiling point in this ternary system. Therefore, the RDWC with common bottom section is a good alternative to separate MeOH from hydrolysis mixture. The configuration of the RDWC we used is illustrated in Figure 6


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1c. A main column and a rectifier are integrated in one column, as shown in Figure 1c. The catalyst packing was loaded in the main column. It should be operated at total reflux and divided from the rectifier on the top to avoid drawing the azeotrope of MeOH-MA from the RDWC. The high conversion of MA can be achieved in the main column. The rectifier is used to separate momentarily the product MeOH from the hydrolysis mixture, to avoid the occurrence of the self-catalyzed MeOH-HAc esterification reaction as far as possible. The rectifier is operated at partial reflux with highly purified MeOH continuously taken from the top.

Except for RDWC, another solution to suppress the self-catalyzed reaction in TRDC is to draw a vapor stream below the catalytic section of TRDC to remove MeOH from the bottom. The drawn stream mainly contains MeOH and water, which could be separated by a distillation column. Such method has been applied by Gao et al

24

to

reduce the energy consumption of the TRDC technology. Although both TRDC with side-draw and RDWC have the ability to suppress the self-catalyzed reaction, RDWC would be a better alternative from the point of cost saving. Firstly, RDWC can save column equipment cost, due to the integration of two columns into one shell 8. Secondly, only one re-boiler is required for the RDWC shown in Figure 1c, while the side-draw-TRDC technology needs two re-boilers, one for the side-draw-TRDC and the other for the distillation column following the side-draw-TRDC. In conclusion, the RDWC was applied for our further study.

7



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3 EXPERIMENTAL SECTION 3.1 Chemicals and analytical method MA (mass purity ≥ 98%), MeOH (mass purity ≥ 99.5%) and HAc (mass purity ≥ 99.5%) were purchased from Shanghai Aladdin Bio-Chem Technology Co., LTD. Deionized water was prepared in our laboratory. C100E, a strong acid cation exchange resin functionalized with sulfonic groups, was purchased from Purolite Co., LTD. (China) and used as catalyst for MA hydrolysis in present work. The relevant characteristics of C100E resin are summarized in Table S1 in Supporting Information.

MA and MeOH were quantitatively analyzed by a gas chromatography (GC-2014, Shimadzu) equipped with a hydrogen flame ionization detector and a nonpolar capillary column (50 m × 0.32 mm × 0.5 µm). Nitrogen gas was used as the carrier gas at 0.1 MPa. Both the temperatures of the injector and the detector were set at 523.15 K. The column temperature was first kept at 313.15 K, and then increased at 10 K·min-1 to 393.15 K. After kept at 393.15 K for 1 minute, it increased at 20K·min-1 to 453.15K. Finally, the column temperature increased at 30 K·min-1 to 523.15 K and kept for 2 minutes. N, N-dimethylformamide was used as internal standard for the quantitative analysis of MA and MeOH. The concentration of HAc was detected by the method of acid-base titration with sodium hydroxide solution as standard solution and phenolphthalein as indicator. The water content was detected by a Karl-Fischer titration (KLS-411, INESA Scientific Instrument Co.,Ltd.).

8


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3.2 Apparatus and procedure The schematic of the lab-scale RDWC is illustrated in Figure 2. Two parallel glass columns representing the divided parts of RDWC were connected to a common stripping section through a Y-type junction. The two parallel columns were 1.5 m high with internal diameters of 24 mm. The left column named as main column was equipped with 1-meter-high catalytic packing in the upper section and 0.5-meter-high Ɵ ring packing in the lower section, playing the role of reactive distillation column. The catalytic packing was a series of catalyst bundles that consists of metal wave mesh and catalyst particles encapsulated with nylon cloth bags, as show in Figure 2. The detailed structure of the catalyst capsules can be found in references 29. The right column named as MeOH rectifier was filled with Ɵ ring packing for the separation of MeOH from the hydrolysis mixture. The common stripping section was 0.5 m high with an internal diameter of 35 mm and packed with 3 × 3 mm Ɵ ring. Two condensers equipped with reflux ratio controller were placed on the upper-ends of these two parallel columns, respectively. The reflux volume of the main column could be measured through a specially designed structure which is the combination of a valve and a scaled branch tubule 15. A three-necked round-bottom glass flask with a volume of 1200 ml was used as column bottom. The middle neck of the flask was connected to the common stripping section. The other two necks were prepared for withdrawing sample and measuring reaction temperature, respectively. The temperatures of the reflux liquid into the columns, the liquid at the bottom of the catalytic packing and the liquid in the three-necked glass flask were measured by calibrated PT-100 thermocouples.

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The liquid in the flask was heated by an electric heating jacket to produce vapor phase for the RDWC. The distribution of vapor phase in main column and rectifier was adjusted by the valve in the Y-junction. The main column was operated at total reflux, while the rectifier was operated at partial reflux. The reactants, water and MA, were separately supplied into the main column by two peristaltic pumps from the top and bottom of the reaction zone, respectively. The products were withdrawn from the top of the rectifier and the bottom of the RDWC. All the experiments were conducted at atmospheric pressure.

4 MODELING All the simulations were performed with the RADFRAC module by the commercial simulation software Aspen Plus 7.2. This module is based on the equilibrium-stage model for solving the mass balance, phase equilibrium, summation, and energy balance (MESH) equations. The theoretical plates are numbered from top to bottom, with the total condenser being the first stage and the reboiler being the last stage.

4.1 Vapor-liquid equilibrium Accurate phase equilibrium data is necessary for the simulation of reactive distillation process. The UNIQUAC

30

and Hayden–O’Connell (HOC)

31

models were used to

describe the liquid non-ideality and the dimerization of acetic acid in vapor phase, respectively 18-19, 25-26. All the model parameters were taken from Aspen Plus 7.3 data bank.

10


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4.2 Reaction kinetics The reaction kinetics for MA hydrolysis on catalyst CE-100 was determined experimentally in our laboratory. The experimental data were used to fit the parameters of the activity-based pseudo-homogeneous (PH) kinetic model. The determination method, experimental results and model fitting are given in Supporting Information. The PH model is written as:

rMA = −

1 dCMA = k +α MAα Water − k -α MeOHα HAc W dt

(1)

with

 −70.767 ×103  k + = 1.644 ×108   RT  

(2)

 −51.782 ×103  k − = 3.083 × 106   RT  

(3)

where rMA is the reaction rate of MA catalyzed by ion exchange resin with the unit of mol/g(cat)/min, W is the catalyst concentration with unit of g/m3, CMA is the molar concentration with unit of mol/m3, k+ is the rate constant of MA hydrolysis with units of mol/g(cat)/s, k- is the rate constant of MeOH-HAc esterification with units of mol/g(cat)/s, α is the liquid activity, R (=8.314J/mol/K) is the ideal gas constant and T is the reaction temperature in Kelvin. The reaction rate on native catalyst particle is different from that in catalyst capsule due to the concentration gradient between inside and outside of the capsule. An effectiveness factor η is defined to describe the difference 32:

η=

reaction rate in catalyst capsule (rcap ) reaction rate in batch reactor (rMA )

11


(4)


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The magnitude of this effectiveness factor depends on the mass transfer driving force and resistance between inside and outside of the capsule. In a special RD loaded with catalytic capsule, the mass transfer driving force and resistance are related with the reactant concentrations outside the capsule and the liquid film thickness on the capsule, respectively. These two factors could be varied by the molar ratio of water to MA and the flux of liquid phase, respectively. The dependence of effectiveness factor on water-MA molar ratio and liquid flux has been correlated for MA hydrolysis in catalytic capsule 24:

η =0.3431 + 311.3034L-0.1573 lnRm

(5)

where L is the flux of liquid phase with unit of m3/m2/h and Rm is the feed mole ratio of water-MA. Therefore, the actual reaction rate of MA in catalyst capsule could be calculated by the reaction kinetics model with effectiveness factor rcap = ηrMA. Since the reaction kinetics model in Aspen Plus does not contain an effectiveness factor, the reaction kinetics model with effectiveness factor was incorporated into the process simulator using an additional FORTRAN subroutine. In the simulation, this kinetic model was set in the theoretical plates where the catalyst capsule was loaded.

The kinetic model for the self-catalyzed esterification reaction of MeOH with HAc was given by Pöpken et al 28:

rMA,self =

dxMA + − = α HAc ( kself α MAα water − kself α MAα HAc ) dt

(6)

with

 −6.35 ×104  + kself = 5.11× 105 exp   RT  

12


(7)

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k

− self

 −8.02 ×104  = 9.83 ×10 exp   RT   6

(8)

+ − and k self are the rate constants of where xMA is the mole fraction of MA, and k self

forward and backward reaction for the self-catalyzed MeOH-HAc esterification reaction with unit of 1/s, respectively. In the simulation, the self-catalyzed kinetic model was set in the theoretical plates where the ϴ ring packing was loaded.

4.3 Liquid holdup and separation efficiency The liquid holdup in catalytic capsule should be specified by using the RADFRAC module to simulate the reactive dividing wall distillation process, since the liquid holdup has an influence on the extent of reaction. The liquid holdup of catalytic capsule has previously been measured by Xu et al. 33: hd = 0.0336u 0.0109 L0.429

(7)

where hd is the liquid holdup with unit of m3/ m3 and u is the gas superficial velocity with unit of m/s. The liquid holdups in the stripping section of main column, the public stripping section, the MeOH rectifier and the reboiler were also specified, since the self-catalyzed esterification reaction of MeOH with HAc will occur there. The liquid volume in the reboiler was 900 ml in our experiments. The liquid holdup for the Ɵ ring packing was specified to be 8% of its superficial volume

34

. The height

equivalent to a theoretical plate (HETP) of the catalytic capsule was set as 0.13 m according to our experimental measurement. The HETP of the 3×3 mm Ɵ ring packing was specified as 0.06m according to reference 35.

13



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5 RESULTS AND DISCUSSION 5.1 Effects of process parameters on MA hydrolysis in RDWC 5.1.1 Feed water-MA molar ratio Rm The effect of Rm on MA conversion in RDWC is shown in Figure 3a. It indicates that a high feed ratio leads to a high MA conversion. Similar results were also reported for the MA hydrolysis in TRDC

15

. The increase of conversion with Rm should be

attributed to the fact that the MA hydrolysis is a reversible reaction. The increase of Rm not only facilitates the MA hydrolysis in the region packed catalyst, but also suppresses the self-catalyzed MeOH-HAc esterification reaction in the region where ϴ ring packing is loaded. In the TRDC for MA hydrolysis, Rm is the most important operating variable to improve the MA conversion 15, 24. In the RDWC, Rm also plays a significant role. As shown in Figure 3a, the mole ratio of water-MA merely rises from 3 to 4, while the MA conversion increases by about 10% (from 90.2% to 98.8%). With the further increase of Rm, 99.5% MA conversion can be achieved. Although increasing the water-MA mole ratio is favorable for improving the MA conversion, more energy consumption is required to separate water from HAc. Therefore, a suitable feed mole ratio of water-MA is expected to simultaneously realize high MA conversion and low energy consumption.

5.1.2 Heat duty He The effect of He on MA conversion is given in Figure 3b. It can be seen that the MA conversion can be improved from 88.0% to 99.5% with the heat duty increased from 0.06 kW to 0.09 kW. Since the main column of the RDWC was operated at total 14


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reflux, increasing the heat duty leads to the increase of reflux volume. Increasing the reflux volume would improve the separation performance of the main column, which means that more hydrolysis products can be removed from the reaction zone to increase the MA conversion. The increase of the reflux liquid will also introduce more MA into the reaction zone, since the reflux liquid is an azeotrope in which the mass concentration of MA is about 90%. With more reflux liquid introduced into the reaction zone, the concentration of MA would be increased while the concentration of the MeOH and HAc will be decreased, promoting the growth of MA conversion. Moreover, the increase of heat duty can improve the separation performance of the rectifier, so MeOH can be more thoroughly separated from the hydrolysis mixture in the column bottom to inhibit the occurrence of the self-catalyzed esterification reaction. All these effects together give rise to the increase of MA conversion. Of course, it is uneconomical to further increase the heat duty as the required MA conversion is achieved.

5.1.3 Feed MA mole flow rate Sv The effect of Sv on MA conversion in RDWC is illustrated in Figure 3c. It can be seen that the MA conversion reduces from 99.6% to 98.5% with Sv rising from 0.6757 mol·h-1 to 1.0811 mol·h-1. The similar variation tendency was also observed in other configurations of RD column for MA hydrolysis 5. The declining MA conversion with increasing Sv should be contributed to the shorter residence time of reactants in the reaction zone of the RDWC at larger Sv. Besides, larger Sv means more feed throughput, inevitably inducing the reduction of the separation efficiency of the main column. As a result, more hydrolysis products stay in the reaction zone, inhibiting the 15



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reaction of water with MA. Both these two effects lead to the reduction of MA conversion with rising Sv. Although rising Sv is adverse to high MA conversion, the column equipment is always expected to be operated at higher Sv to achieve larger production capacity and create more economic benefits. Therefore, as long as the MA conversion is satisfied, larger Sv is expected.

5.1.4 Mass ratio of vapor entering into MeOH rectifier to feed MA Sm Since it is difficult to experimentally determine the vapor distribution ratio in RDWC, we replaced this parameter by the mass ratio of vapor entering into MeOH rectifier to feed MA Sm. The mass flow-rate of the vapor entering into the rectifier was calculated through multiplying the mass flow-rate of the liquid drawn from the top of the MeOH rectifier with the reflux ratio of the MeOH rectifier. The effect of Sm on MA conversion is displayed in Figure 3d. It can be seen that the MA conversion grows from 94.5% to 99.3% with Sm decreased from 3.9 to 1.6. The explanations for effect of Sm on MA conversion are similar to those for the effect of heat duty on MA conversion. The results in Figure 3d indicate that the vapor distributor of the RDWC should be carefully designed to bring more vapors into the main column to improve the MA conversion. However, it is unadvisable to operate the RDWC at too small Sm. With less vapor entering into the rectifier, the separation performance of the rectifier is likely to reduce. This will result in the un-complete separation of MeOH from the hydrolysis mixture. The coexistence of MeOH and HAc in the RDWC is adverse to realize complete conversion of MA. Therefore, it is advised to operate the RDWC for MA hydrolysis at small Sm, only if the separation requirement of the rectifier is satisfied. 16


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5.1.5 Response surface methodology (RSM) analysis In this section, the response surface methodology (RSM) analysis was applied to optimize the operation conditions based on a 3-factor-3-level Box-Behnken design (BBD). Four manipulated variables including Rm, He, Sv and Sm were studied in Sections 5.1-5.4. Since Sv impacts only the column sizing in industrial application, it was ignored in the RSM analysis to produce a three-variable test matrix. The results of BBD are listed in Table 2. The MA conversion obtained by Aspen Plus were fitted with a second order polynomial equation:

Y = 98.03 + 8.31X 1 + 9.57 X 2 − 4.01X 3 − 0.71X 1 X 2 +1.30 X 1 X 3 + 6.48 X 2 X 3 − 6.12 X 12 − 7.54 X 22 − 2.08 X 32

(8)

Y is the MA conversion. X1, X2 and X3 represent coded values of variables Rm, He and Sm, respectively. The determination coefficient R2 of Eq. 8 is 0.9710, indicating that Eq.8 is able to predict within the range of experimental variables. The analysis of variance (ANOVA) for the response surface quadratic equation is presented in Table 3. The p-value is used to evaluate the importance of each model term. The model term is significant if the p-value is less than 0.05. As shown in Table 3, X1, X2, X3, X2X3, X12, X22 are significant model terms. The model term is not significant if the p-value is greater than 0.1. Table 3 presents that X1X2 and X1X3 are insignificant model terms. Moreover, smaller p-value indicates more significance of corresponding model term. By comparing the p-values of model terms X1, X2 and X3, it could be concluded that the importance of variables Rm and He are comparative and they are more significant than variable Sm. Optimizing operation condition was performed on the basis of Eq. 8. The optimal operation conditions were Rm = 4.79, He = 0.09, Sm = 1.58 with 99% MA conversion as optimization goal. The conversion obtained by the Aspen Plus

17



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simulation was 98.7% at the optimal conditions. The small deviation indicated the validity of Eq. 8.

5.2 Comparison between RDWC and TRDC 5.2.1 Simulation validation The performances of TRDC and RDWC were compared by the commercial software Aspen Plus V7.3. In order to ensure the credibility of the comparison, the experimental data in Section 5.1 were used to verify the simulation results. The simulation was performed with the RADFRAC module, since this module is capable of describing the vapor-liquid equilibrium behavior in the reactive distillation column for MA hydrolysis

20, 24

. Pöpken et al.

24

showed that the assumption of chemical

equilibrium on each theoretical stage is unreasonable for the MA hydrolysis catalyzed by the heterogeneous ion-exchange resin catalyst, so the reaction kinetic model should be considered in the simulation. The kinetic models given in Section 4.2 were applied in the simulation. The comparison between the simulation results and the experimental data is shown in Figures 3a-d for MA conversion. An error of less than 2% was found in the comparison between the simulation results and experimental data, which verified the creditability of the models used.

5.2.2 Superiority of RDWC to TRDC for MA hydrolysis MA hydrolysis in TRDC was simulated under the conditions similar to those for MA hydrolysis in RDWC. In the simulation of TRDC, some parameters should be carefully specified to make the simulation results comparable to those of RDWC. 18


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Detailed specifications of these parameters are given in Supporting Information.

Table 4 lists the performance comparison between RDWC and TRDC for MA conversion of 99% and 99.5%. It should be acknowledged that much excess water is still required to obtain high methyl acetate conversions, even though the RDWC is used. Nevertheless, the water required for the RDWC technology is lower than that for the TRDC technology, especially at high MA conversion, as shown in Table 4. The less consumption of water in RDWC also implies the less energy consumption in the following water-HAc separation process. The energy consumed by the RDWC is lower than that by the TRDC at the same Rm and Sv. The technology with more energy saving is preferred in such an era of energy-shortage. With the same He and Rm, Sv for the RDWC is larger than that for the TRDC. It suggests that the RDWC has larger production capacity. With the MA conversion increased from 99% to 99.5%, He and Rm increase and the Sv decreases, both for the RDWC and TRDC. That is because the self-catalyzed esterification reaction need to be further suppressed. The increase of He and Rm and the decrease of Sv are smaller for the RDWC than the TRDC. It reveals that the RDWC performs better in improving the product purity of MeOH, which is always distillated from top of the MeOH rectifier with the unreacted MA. In short, the RDWC shows superior performance to TRDC for MA hydrolysis in reducing energy and reactant consumptions and in improving production capability and product quality.

There are two reasons for the better performance of the RDWC in MA hydrolysis. One of the reasons is that the integration of reactive distillation column and distillation column makes the energy utilization of RDWC more efficient 5. Li et al. 25 confirmed this reason with the Aspen Plus by neglecting the self-catalyzed MeOH19



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HAc esterification reaction. They show that 20% energy saving can be achieved for MA hydrolysis with the application of RDWC. The other reason is that the separation of MeOH from the hydrolysis mixture weakens the self-catalyzed MeOH-HAc esterification reaction in RDWC. This reason would be verified by comparing the effects of self-catalyzed MeOH-HAc esterification reaction on MA hydrolysis in RDWC and TRDC.

Figure 3 (dotted lines) illustrates the effect of self-catalyzed MeOH-HAc esterification reaction on the simulation results for the RDWC. The MA conversion is over-estimated if the self-catalyzed reaction was not considered in the simulation. In order to show the effect of the self-catalyzed MeOH-HAc esterification reaction, we defined the relative simulation deviation of MA conversion:

RSD=

X MA+self − X MA-self X MA+self

× 100

(9)

XMA+self and XMA-self are the MA conversions for the simulations with and without considering self-catalyzed esterification reaction, respectively. The relative simulation deviation is less than 1% and decreases with the increase of Rm and He and with the decrease of Sv, as presented in Figure 4 (the solid lines). It has been analyzed in Section 5.1 that either increasing Rm and He or decreasing Sv is conducive to promote the MA hydrolysis, that is, to suppress the MeOH-HAc esterification, in the RDWC. As a result, the self-catalyzed MeOH-HAc esterification in the column bottom is also suppressed, leading to the decrease of simulation deviation. Figure 5 illustrates the effect of self-catalyzed MeOH-HAc esterification reaction on the simulation results for the TRDC. Similar to those for RDWC, the MA conversion is over-estimated without considering the self-catalyzed reaction and the deviation decreases with the 20


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increase of Rm and He and with the decrease of Sv. Figure 4 (dotted lines) shows that the simulation deviation would approach 10% at some conditions, indicating that it is necessary to take the self-catalyzed esterification into account for the accurate simulation of TRDC. The simulation deviation for the RDWC is much smaller than those for the TRDC. It could be concluded that the self-catalyzed MeOH-HAc esterification reaction is stronger in TRDC than in RDWC. This conclusion was verified by the results in Figure 6, which shows the comparison of MA formation rate in the stripping sections of RDWC and TRDC. The MA formation rates in the MeOH rectifier of RDWC and in the MeOH distillation column following the TRDC are not discussed here, since they are ignorable in comparison with those in the stripping sections of RDWC and TRDC. It could be seen in Figure 6 (the top horizontal axis) that the formation rate of MA mainly occurs on the last theoretical plate (column bottom) either in TRDC or in RDWC. More importantly, the formation rate of MA on the last theoretical plate in RDWC is about 25 times higher than that in TRDC, as shown in Figure 6 (the bottom horizontal axis). The formation rate of MA on the last theoretical plate is closely connected with the concentration of reactants. Figure 7 compares the concentration profiles in the stripping section of RDWC and TRDC. The MeOH mole concentration on the last theoretical plate in TRDC is about 0.17, while the corresponding concentration in RDWC is about 0.00002. The much higher MeOH concentration in the bottoms of TRDC than that of RDWC results in much stronger self-catalyzed esterification reaction in TRDC. The extremely low MeOH concentration in the bottom of RDWC should be contributed to the fact that the rectifier of RDWC could separate MeOH from the bottom.

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6 CONCLUSION It is difficult to achieve above 99% conversion for methyl acetate (MA) in traditional reactive distillation column (TRDC), due to the existence of self-catalyzed methanol (MeOH)-acetic acid (HAc) esterification reaction in the column bottom. The RDWC has the ability to suppress this self-catalyzed reaction by removing MeOH from hydrolysis mixture, to realize more than 99% MA conversion. The experiment of MA hydrolysis in a RDWC was performed. The effects of feed water-MA mole ratio, heat duty, mole flow rate of feed MA and mass ratio of vapor entering into the MeOH rectifier to feed MA on hydrolysis conversion were investigated. More than 99% MA conversion could be achieved at suitable operation conditions. The MA hydrolysis in RDWC was simulated by Aspen Plus V 7.3 with the equilibrium stage model to describe the vapor-liquid equilibrium behavior and the pseudo-homogeneous kinetic mode to describe the reaction behavior. An error less than 2% was found between the simulation and experimental results for the RDWC, which indicates that the models used are accurate enough. The simulation results of RDWC and TRDC shows the RDWC is superior to the TRDC in reducing energy and reactant consumption and improving the production capability for MA hydrolysis. With the MA conversion increased from 99% to 99.5%, the superiority of RDWC to TRDC becomes more obvious. If the self-catalyzed esterification reaction was not taken into account, the MA conversion would be overestimated by 1% and 10% at certain conditions for the simulation of RDWC and TRDC, respectively. It indicates that the self-catalyzed reaction is non-ignorable in the simulation, especially for the TRDC. The much higher simulation deviation also suggests that the better performances of the RDWC should be contributed to the separation of MeOH from the hydrolysis mixture to suppress the self-catalyzed MeOH-HAc esterification reaction. 22


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ACKNOWLEDGMENTS We acknowledge the financial support for this work from the National Natural Science Foundation of China (No. 9153410), the Natural Science Foundation of Fujian Province (Nos. 2016J05036 and 2016J01689).

ASSOCIATED CONTENT Supporting Information Reaction kinetic for MA hydrolysis on C100E, Experimental measurement of HETP for catalyst capsule, Simulation of TRDC by Aspen Plus

AUTHOR INFORMATION Corresponding Author *Email: [email protected]

NOMENCLATURE CMA

molar concentration of MA, mol/m3

hd

liquid holdup of catalytic capsule, m3/ m3

He

heat duty, kW

k+

rate constant of MA hydrolysis kinetic model for ion exchange resin, 23



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mol/g(cat)/s k-

rate constant of MeOH-HAc esterification reaction kinetic model for ion exchange resin, mol/g(cat)/s

+ k self

rate constant of forward reaction for the self-catalyzed MeOH-HAc esterification reaction, 1/s

− k self

rate constant of backward reaction for the self-catalyzed MeOH-HAc esterification reaction, 1/s

L

flux of liquid phase, m3/m2/h

R

idea gas constant, 8.314J/mol/K

rcap

reaction rate of MA catalyzed by catalytic capsule, mol/g(cat)/min

Rm

feed mole ratio of water-MA, mol/mol

rMA

reaction rate of MA catalyzed by ion exchange resin, mol/g(cat)/min

rMA,self

reaction rate of MA for the self-catalyzed reaction, mol/g(cat)/min

Sm

mass ratio of vapor entering into MeOH rectifier to feed MA, g/g

Sv

feed mole flow rate of MA, mol/h

T

reaction temperature, K

u

gas superficial velocity, m/s

W

catalyst concentration, g/m3

xMA

mole fraction of MA, mol/mol

XMA

MA conversion

α

liquid activity

η

effectiveness factor of catalytic capsule

Abbreviations HAc

acetic acid 24


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HETP RD

height equivalent to a theoretical plate, m reactive distillation

RDWC reactive dividing wall column RSD TRDC MA MeOH

relative simulation deviation, % traditional reactive distillation column methyl acetate methanol

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2016, 39, 815-833. (13) Kaymak, D.B.; Luyben, W.L. Quantitative comparison of reactive distillation with conventional multiunit reactor/column/recycle systems for different chemical equilibrium constants. Ind. Eng. Chem. Res. 2004, 43, 2493-2507. (14) Fuchigami, Y. Hydrolysis of methyl acetate in distillation column packed with reactive packing of ion exchange resin. J. Chem. Eng. Jpn. 1990, 23, 354-359.

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(31) Hayden, J. G.; O'Connell, J. P. A generalized method for predicting second virial coefficients. Ind. Eng. Chem. Proc. Des. Dev. 1975, 14, 209-216. (32) Xu, X.; Zheng, Y. X.; Zheng, G. W. Kinetics and effectiveness of catalyst for synthesis of methyl tert-butyl ether in catalytic distillation. Ind. Eng. Chem. Res. 1995, 34, 2232-2236. (33) Xu, X.; Zhao, Z.; Tian, S. Study on catalytic distillation processes: Part III: Prediction of pressure drop and holdup in catalyst bed. Chem. Eng. Res. Des. 1997, 75, 625-629. (34) Iwai, Y.; Yamanishi, T.; Okuno, K.; Yokogawa, N.; Tsuchiya, H.; Yoshida, H.; Kveton, O. K. Design study of feasible water detritiation systems for fusion reactor of ITER scale. J. Nucl. Sci. Technol. 1996, 33, 981-992. (35) Kaba, A.; Akai, R.; Yamamoto, I.; Kanagawa, A. Measurement of HETP of SUS dixon ring and porcelain packing in small-scale water distillation column for H2OHTO isotope separation. J. Nucl. Sci. Technol. 1998, 25, 825-830.

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Tables Table 1 Boiling temperatures of pure compounds and azeotrope for the quaternary system of MeOH-HAc-water-MA at 101.3kPa Compound MeOH-MA MA-water MA MeOH Water HAc

Mass fraction (%) 33.56-66.44 91.55-8.45 100 100 100 100

Temperature (K) 326.72 329.81 330.95 337.85 373.15 391.05

30


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Table 2 BBD design and results for the optimization of operation conditions.

Run 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17

Rm (mol/mol) 5 5 5 5 3 5 5 7 7 7 3 7 5 5 5 3 3

Variable He (kw) 0.06 0.09 0.06 0.12 0.06 0.09 0.09 0.06 0.12 0.09 0.09 0.09 0.09 0.09 0.12 0.12 0.09

Sm (g/g) 1.5 2.0 2.5 2.5 2.0 2.0 2.0 2.0 2.0 2.5 2.5 1.5 2.0 2.0 1.5 2.0 1.5

Conversion (%) by Aspen Plus by Eq. 8 90.87 89.31 98.03 98.02 65.10 66.35 98.89 98.44 67.69 65.77 98.03 98.02 98.03 98.02 83.60 83.81 99.63 99.54 98.89 95.42 77.55 76.21 99.51 99.84 98.03 98.02 98.03 98.02 98.77 96.51 86.56 86.34 83.35 85.82

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Table 3 ANOVA for the fitted polynomial quadratic model

Source Model X1 X2 X3 X1X2 X1X3 X2X3 X12 X22 X32

Sum of Squares 2043.45 552.41 733.27 128.54 2.01 6.71 167.71 157.73 239.30 18.24

df 9 1 1 1 1 1 1 1 1 1

Mean Square 227.05 552.41 733.27 128.54 2.01 6.71 167.71 157.73 239.30 18.24

F-value

p-value

26.02 63.31 84.04 14.73 0.23 0.77 19.22 18.08 27.43 2.09

0.0001 < 0.0001 < 0.0001 < 0.0064 0.6458 0.4097 0.0032 0.0038 0.0012 0.1915

32


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Table 4 Performance comparison between the RDWC and TRDC

MA conversion XMA

99%

Technology Rm (mol/mol)a Relative variation of Rm (%) He (kW) b Relative variation of He (%) Sv (h-1) c Relative variation of Sv (%)

RDWC 4.25

99.5% TRDC 4.9

RDWC 5.2

0.0843

0.0860

-13.26 0.0660

-52.72

-21.72 0.1535

TRDC 11 0.1272 -32.40

0.1425 7.69

0.1023

0.0904 13.13

a

the fixed parameters are listed in Figure 3a; water mass concentration in product MeOH xw=0.0032 and 0.0036 for XMA=99% and 99.5%, respectively.

b

the fixed parameters are listed in Figure 3b; water mass concentration in product MeOH xw=0.0029 and 0.0025 for XMA=99% and 99.5%, respectively.

c

the fixed parameters are listed in Figure 3c; water mass concentration in product MeOH xw=0.0036 and 0.0029 for XMA=99% and 99.5%, respectively.

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Figures

(a)

(b)

(c) Figure 1 Development of the technologies for MA hydrolysis. (a) fixed bed reactor (FBR) technology; (b) traditional reactive distillation column (TRDC) technology; (c) reactive dividing wall distillation column (RDWC) technology. 34


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Figure 2 Schematic of the experimental setup

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(a)

(b)

(c)

(d)

Figure 3 Effect of process parameters on MA hydrolysis in RDWC. (a) Feed water-MA molar ratio Rm; (b) Heat duty He; (c) Feed mole flow rate of MA Sv; (d) Mass ratio of side product to feed MA Sm.

36


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(a)

(b)

(c)

Figure 4 Comparison of relative simulation deviation induced by the self-catalyzed MeOHHAc esterification reaction between TRDC and RDWC at different process parameters. (a) Feed water-MA molar ratio Rm; (b) Heat duty He; (c) Feed mole flow rate of MA Sv. 37



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(a)

(b)

(c) Figure 5 Effect of process parameters on MA hydrolysis in TRDC. (a) Feed water-MA molar ratio Rm; (b) Heat duty He; (c) Feed mole flow rate of MA Sv.

38


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Figure 6 Comparison between TRDC and RDWC for MA formation rate. Rm=5mol/mol, He=0.1kW, Sv=0.8108mol/h, Sm=1g/g (for RDWC only).

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Figure 7 Mole concentration of each component in the stripping section of TRDC and RDWC. Rm=5mol/mol, He=0.1kW, Sv=0.8108mol/h, Sm=1g/g (for RDWC only).

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Graphic for manuscript 500x400mm (156 x 156 DPI)


High Conversion of Methyl Acetate Hydrolysis in a Reactive Dividing

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