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Feasibility study of reactive distillation for the production of propylene glycol monomethyl ether acetate through transesterification Xiaoda Wang, Qinglian Wang, Changshen Ye, Xiaolian Dong, and Ting Qiu Ind. Eng. Chem. Res., Just Accepted Manuscript • Publication Date (Web): 30 May 2017 Downloaded from http://pubs.acs.org on June 4, 2017

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Industrial & Engineering Chemistry Research

Feasibility study of reactive distillation for the production of propylene glycol monomethyl ether acetate through transesterification

Xiaoda Wang, Qinglian Wang, Changshen Ye, Xiaolian Dong, Ting Qiu* School of Chemical Engineering, Fuzhou University, Fuzhou 350108, Fujian, China

ABSTRACT This paper aims to present the feasibility of conducting the transesterification of propylene glycol ether (PM) with methyl acetate (MeAc) in a reactive distillation (RD) column to improve the reaction conversion. The essential thermodynamic and reaction kinetic data of the reaction system were measured for the feasibility study. There is only one azeotrope MeAc-MeOH in the reaction system, and the NRTL could

describe

well

the

thermodynamic

behavior

of

this

system.

The

transesterification of PM with MeAc is an endothermic reaction with activation energy E = 55.704 kJ·mol-1. The feasibility was analyzed by the residue curve maps (RCM), showing that full conversion of PM could be realized by RD with mole ratio of MeAc-PM larger than 2.882. The intensification effect was experimentally verified in a batch RD column. Finally, some important parameters are given through the conceptual design to develop continuous RD column for the transesterification of PM with MeAc.

1

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1 INTRODUCTION Propylene glycol methyl ether acetate (PMA) is a versatile solvent with numerous industrial applications, due to its outstanding advantages of high dissolving capacity, low causticity and toxicity, and excellent thermal stability

1-2

. One of the important

applications of PMA is to be used as solvent for the production of electronic products. In this application, the electronic grade PMA is necessary. With the rapid development of electronics industry, more and more electronic grade PMA is demanded.

The typical route for the synthesis of PMA is the esterification reaction of propylene glycol methyl ether (PM) and acetic acid in the presence of acid catalyst, as shown in Figure 1a. The complete conversion of PM could not be achieved in a batch reactor in the traditional technique, due to the limitation of chemical equilibrium. In order to improve the PM conversion, the technologies of reactive chromatography reaction distillation

4

3

and

have been applied to intensify this reaction. With these two

process intensification technologies, full conversion of PM could be realized. However, the acid catalyst would inevitably result in the formation of acidic PMA product, even though solid acid catalyst was used. The acidity of the electronic grade PMA is required to maintain below ultra-low value. It is improper to adjust the acidity of PMA by basic solutions, since unwanted impurities would be introduced through the solutions to populate the PMA product. It means that the PMA produced through the esterification route is difficult to satisfy the requirement of electronic grade chemical without complex post-processing. 2


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Producing the PMA via the reaction route which could be catalyzed by basic catalysts is an effective method to avoid yielding acidic PMA. A novel route for PMA synthesis through the transesterification of PM with methyl acetate (MeAc) is applied in this work, as shown in Figure 1b. Since the transesterification reaction can be catalyzed by basic catalysts, the acidity of PMA can be low enough to satisfy the product standard of electronic grade chemical without further treatment. Two kinds of basic catalysts, homogeneous and heterogeneous catalysts, are available for the transesterification reaction. Heterogeneous catalysts are preferred, because of their reusability and facile separation from the reaction mixture. Unfortunately, most of the basic heterogeneous catalysts with high activity suffer from the problem of low lifetime, due to their high sensitivity to moisture and poor mechanical stability

5-6

. Therefore, the basic

homogeneous catalyst was always chosen to catalyze the transesterification reaction. Sodium methoxide is the most widely used basic homogeneous catalyst for transesterification reaction because of its high activity and selectivity

7-8

. In present

work, sodium methoxide was chose as catalyst for the transesterification of PM with MeAc.

It is inappropriate to perform the PM-MeAc transesterification reaction in a batch reactor to achieve full conversion of PM, since this reaction is limited by the chemical equilibrium. For the chemical-equilibrium-limited reactions, the technology of reactive distillation (RD) has been proven to be an effective enhancement alternative 3



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with many advantages, such as significant capital saving, complete reactant conversion, significantly reduced catalyst requirement, heat integration and so on 9-12 . Up to now, several transesterification reactions limited by the chemical equilibrium have been successfully enhanced by the RD to produce some important chemicals, such as biodiesel, dimethyl carbonate, butyl butyrate, sec-butanol and acrylic acid 13-17. It has been experimentally shown that the complete conversion of transesterification reaction could be smoothly realized at suitable operation conditions with the application of RD technology, resulting in the efficiency improvement and cost reduction of the production process 13-17.

The development of RD for a special reaction system involves four stages

12

. Since

combining reaction and distillation in a RD column is not always suitable for all the chemical process based on simultaneous reaction and separation, the first stage is to analyze the feasibility, that is, to analyze the potential top and bottom products

12

.

Only when the RD is proved to be feasible, could the subsequent conceptual design, equipment selection and control strategy evaluation be conducted. Residue curve maps (RCM), which reflects the variation of liquid compositions in an isobaric open evaporation with time, is the widely used method for the feasibility analysis of RD, due to its simplicity and availability

18

. In order to obtain RCM, the vapor-liquid

equilibrium (VLE) data is essential. There is only one azeotrope MeAc-methanol (MeOH) in the quaternary system MeAc + MeOH + PM + PMA, according to the VLE data reported in the literatures 19-24 and measured in our experiment (Section 3.1). 4


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The boiling points of MeAc-MeOH azeotrope and PMA are lowest and highest, respectively, for the investigated quaternary system. However, whether the MeAc-MeOH azeotrope and pure PMA are the top and bottom products of a RD column needs further investigation, since the chemical reaction might have an effect on the thermodynamic behavior of reaction system, for example “destroying” the azeotrope 12 or forming reactive azeotrope 25. RCM can predict the effect of chemical reaction on the thermodynamic behavior 25.

In this work, we aim to study the feasibility of developing a RD process for the transesterification of PM with MeAc to provide an alternative for the production of electronic grade PMA. Firstly, the required thermodynamic and reaction kinetic data for this reaction system were determined to theoretically analyze the feasibility this RD process with residue curve maps (RCM). Then, the batch RD experiment was conducted for this this reaction to verify our theoretical analyzation. Finally, some important parameters were provided to design a RD column for the transesterification of PM with MeAc.

2 EXPERIMENTAL SECTION 2.1 Chemicals PM and PMA are purchased from Aladdin with a guaranteed purity of 99.5 mass%. MeOH and MeAc were purchased from Sinopharm with a guaranteed purity of 99.5 mass%. The basic sodium methoxide was used as catalyst. It was obtained from 5



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Fuchen Chemical Reagent Factory in Tianjin.

2.2 Apparatus and Procedure The VLE measurements were carried out in a modified Othmer equilibrium still at 101.3 kPa. A schematic of the apparatus is shown in Figure 2a. The phase equilibrium temperature was measured by a mercury thermometer with an accuracy of ± 0.1 K. In each run, about 40 ml of mixture was required. The mixture was heated by a heating bar. Both the liquid and vapor phases were continuously circulating in the still to ensure the establishment of vapor-liquid two-phase equilibrium. It took approximately 1 h for the investigated system to reach phase equilibrium. After the equilibrium had been established, both the liquid and condensed vapor phases were sampled at about 0.5 ml to analyze the compositions by a gas chromatograph. Similar apparatus and procedure was used in our previous work to measure the VLE data of other systems 19. The error for VLE experiment was the measurement deviations in vapor-liquid equilibrium temperature, atmospheric pressure and compositions of vapor and liquid phases.

The reaction kinetics for the transesterification of PM with MeAc was performed in a glass three-necked flask with a volume of 500 ml at 101.3 kPa, as shown in Figure 2b. Two necks of the flask were for measuring reaction temperature with a mercury thermometer and sampling with a syringe, respectively. The third neck was connected to a spherical condenser, to prevent the mass loss due to the component volatilization. 6


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The other end of the spherical condenser was connected to a glass desiccator filled with solid CaCl2 to prevent the entrance of water into the reaction system. A magnetic stirrer with the stirring rate up to 2000 rpm was applied to achieve the uniform mixing of the reactive mixture. The reaction temperature was maintained by immersing the glass three-necked flask into a thermostat-controlled water bath with accuracy of ± 0.3 K. To ensure the reaction occurred at desired temperature, the mixture of reactant MeAc and catalyst sodium methoxide was heated in the three-necked flask, while the reactant PM was heated in a preheater. After the reactants and catalyst were heated to the desired reaction temperature, the reactant PM was added into the three-necked flask. This point was defined as the time zero of the reaction. During each run of the reaction kinetics experiment, about 0.5 ml liquid sample was withdrawn from the flask at regular time intervals using a syringe. The samples were cooled rapidly to 263 K in a refrigerator to avoid any further reaction and analyzed within 30 minutes using the gas chromatography. All the experiments were continued until the chemical equilibrium was reached. It should be noted that the experimental apparatus shown in Figure 2b also used as batch reactor for the PM conversion comparison with batch RD column. The error for kinetic experiment was measurement deviations in reaction temperature, atmospheric pressure and compositions of reaction mixtures.

The schematic diagram of the batch RD column is shown in Figure 2c. The batch RD experiments were carried out at atmospheric pressure in a glass column with inner diameter of 22 mm. The column was consisted of three segments, in which the θ ring 7



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packing with size of 3 × 3 mm was incorporated. The total packing height is 2 m. The column was bound up by two layers of mineral wool to prevent heat loss. For evaporation, a 500 ml round-bottom flask immersed in a thermostat-controlled water bath was connected to the bottom end of the column. For reflux, a vertical, water-cooled total condenser was connected to the upper end of the column. The reflux ratio was controlled by a reflux ratio controller. The temperatures of the liquid phase in the round-bottom flask and the vapor phase entering into the condenser were measured by a mercury thermometer and a calibrated PT-100 thermocouples, respectively. The reactants PM and MeAc were fed into the round-bottom flask with solid

catalyst

sodium

methoxide

before

each

run

of

experiment.

The

transesterification reaction only took place in the round-bottom flask, since the catalyst only existed there. The total reflux operation was implemented at the beginning of experiments to realize a stable control. In the normal operation, both reactants and catalyst were not added, and the product MeOH was continuously removed from the column top with MeAc. The target product PMA was not removed out and accumulated in the round-bottom flask before the ending of the experiment. Liquid samples were withdrawn every 30 minutes with a syringe from the round-bottom flask as well as from the distillate stream to supervise the reaction extent. The experiment was stopped when no MeOH was detected in the distillate samples by gas chromatograph. A composition analysis on the liquid in the round-bottom flask was conducted to determine the PM conversion. The side reaction was ignored since no side-product was detected. The error for batch RD experiment 8


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was the measurement deviations in the temperatures of the liquid phase in the round-bottom flask and the vapor phase entering into the condenser, the atmospheric pressure and the concentrations of the top and bottom products.

2.3 Analytical Method A gas chromatograph (GC2014, Shimadzu Corporation) equipped with a flame ionization detector (FID) and a RTX-5 capillary column (30 m × 0.25 mm × 0.25 mm) was used to analyze all samples. Nitrogen gas with a purity of 99.99 mass% was used as carrier gas at the velocity of 1 ml/min. An optimized temperature control program was implemented (313 K for 3 min, a ramp of 40 K/min to 443 K, 443 K for 1 min). The temperature of the injection port and detector were controlled at 523 K and 553 K, respectively. 1 µl sample was injected each time. The compositions were determined by the internal standard method with isopropanol as internal standard substance.

3 RESULTS AND DISCUSSION 3.1 Vapor-Liquid Equilibrium (VLE) For the quaternary system of PMA + PM + MeAc + MeOH, the VLE data of the five binary systems MeAc + MeOH, MeOH + PM, PM + PMA, MeAc + PM, MeOH + PMA were available in literatures 19-24, but the VLE data of the PMA + MeAc system has not been reported yet. In order to complete VLE database for this quaternary system, the VLE data of the binary system PMA + MeAc were measured in this work. The measured data are shown in Table S1 in supporting information and illustrated 9



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graphically on a T-x-y diagram in Figure 3. xMeAc and yMeAc are the mole fraction of liquid and vapor phase, respectively. T is the phase equilibrium temperature. As shown in Figure 3, the phase equilibrium temperature decreases monotonously from 419 K (the boiling temperature of PMA) to 330 K (the boiling temperature of MeAc) with the mole fraction of MeAc increasing from 0 to 1. The monotonous variation of phase equilibrium temperature indicates that there is no azeotrope between MeAc and PMA.

Since errors are inevitable in VLE measurement experiments, it is necessary to carry out a thermodynamic consistency test to estimate the reliability of the measured VLE data. The semi-empirical test method proposed by Herington 26-27 was employed to conduct the thermodynamic consistency test. The values of liquid activity coefficient are necessary for the consistency test. They were calculated by the following equation 28

: ∧

Py φ v γi = s i s i pi φi xi

(1)

P represents the system pressure, and is equal to 101.3 kPa for our experiments. pis is the saturated vapor pressure of pure component i, and can be calculated by the Antoine equation:

log pis ( kPa ) = A −

B T (K) + C

(2)

The value of constants A, B and C for PMA and MeAc have been reported in literature ∧

, as listed in Table 1. φiv and φis are the vapor and liquid phase fugacity

19

10


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coefficients of component i, respectively. Fugacity coefficient represents the correlation of real substance to ideal gas. Since the VLE experiments were carried out ∧

at atmospheric pressure, φiv ≈ 1 and φis ≈ 1 28. According to the test method proposed by Herington 26-27, the parameter D-J is 2.11% for the experimental VLE data of PMA-MeAc system, which fulfills Herington’s original criterion: D-J < 10%. This means that the measured VLE data of the PMA-MeAc system are accurate enough.

The VLE data in Table S1 were used to correlate the interaction parameters of non-random two liquid (NRTL) model

29

by the simplex method with the following

expression as objective function: 2 2 (y T j ,exp − T j ,cal )  ( j ,exp − y j ,cal )  Fob = ∑  + 2 2   y T j =1 j ,cal j ,cal   n

(3)

where n is the number of the experimental data set. The subscripts “cal” and “exp” are the calculated and experimental results, respectively. The non-random parameter α of the NRTL model is generally between 0.2 and 0.47 20. It was set as 0.3, as it was done in our previous work

19

, to simply the correlation. The correlated binary interaction

parameters of NRTL model for the system of PMA + MeAc were given in Table 2. The fitting results were evaluated by the root mean square error (RMSE): n

RMSE =

∑(M j =1

j ,exp − M j ,cal )

N

2

( M = T or y )

(4)

The RMSE of phase equilibrium temperature and vapor composition are 0.496 and 0.0064, respectively. The small RMSE values show the high reliability of the NRTL 11



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model parameters correlated. The calculated and experimental vapor and liquid mole fraction are compared in Figure 3 for different temperatures. It can be seen that the calculated and experimental results are coincident with each other.

The VLE data for the quaternary system PMA + PM + MeOH + MeAc were measured at 101.3 kPa and the results are listed in Table S2. Table 1 and Table 2 list the Antoine equation parameters and the NRTL model parameters for this quaternary system, respectively. These parameters were used to calculate the vapor phase compositions and phase equilibrium temperatures at different liquid phase compositions. The calculative deviations from the experimental results are expressed by RMSE. The RMSE of T, yMeOH, yMeAc, yPM and yPMA are 2.14, 0.0757, 0.0327, 0.0421 and 0.0225. These small RMSE values show the reliability of the parameters in Tables 1 & 2.

3.2 Reaction Kinetics The reaction kinetics is essential to establish a model for the RD process. We applied the activity-based pseudo-homogeneous model to describe the reaction kinetics for the transesterification of PM with MeAc:

 aˆ aˆ − rPM = kf  aˆPM aˆ MeAc − PMA MeOH  K eq 

 aˆ PMA aˆ MeOH  −E   ˆ ˆ  = k0 exp    aPM aMeAc − K eq  RT   

  

(5)

In Eq. 5, kf is the forward reaction rate constant, k0 is the pre-exponential factor, E is the activation energy of the transesterification reaction, R (=8.314J·mol-1·K-1) is the ideal gas constant, and Keq is the activity-based chemical equilibrium constant. 12


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The reaction kinetics experiments were carried out at different temperatures to obtain the information about k0, E and Keq. Figure 4 shows the experimental results measured at the temperatures ranging from 303 K to 333 K. The experiments of reaction kinetics were not conducted above 333 K to prevent the vaporization of the reactant MeAc. Almost all the chemical equilibria were reached within 30 minutes at all temperature investigated, as illustrated in Figure 4. The increase of temperature is apparently favorable to accelerate the transesterification reaction. The equilibrium conversion of PM increases from 11.1% to 26.4% with temperature increasing from 303 K to 333 K, which implies that the transesterification of PM with MeOH is an endothermic reaction.

The equilibrium constant was expressed with the reaction standard Gibbs free energy

∆ r GmΘ :  ∆ GΘ K eq = exp  − r m  RT

  ∆ r H mΘ ∆ r S mΘ  exp = +  −  R    RT

(6)

where ∆ r H mΘ and ∆ r SmΘ are the reaction standard enthalpy and reaction standard entropy, respectively. The chemical equilibrium constant Keq for the transesterification of PM with MeAc was determined with the following expression:  aˆ aˆ  K eq =  PMA MeOH   aˆPM aˆ MeAc eq

(7)

The activity coefficient aˆi was calculated by the NRTL model with the interaction parameters in Table 2. According to Eq. 6, the logarithm of equilibrium constant is 13



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plotted as a function of the inverse absolute temperature, as shown in Figure 5a. The following expression was obtained with the least square method:

 4963.30  K eq = exp  − + 12.65  T  

(8)

∆ r H mΘ = 41.265 kJ·mol-1 and ∆ r SmΘ = 0.105 kJ·mol-1·K-1 were determined through Eq. 8 by assuming that they are independent of the reaction temperature.

The forward reaction rate constant of Eq. 5 were fitted with the data of reaction kinetics shown in Figure 4 by the least square method. For a complete mixed batch reactor with a constant reaction volume, the reaction rate of PM -rPM can be calculated through

− rPM =

N PM 0  dCPM    M cat  dt 

(9)

where NPM0 is the initial moles of PM, Mcat is the mass of catalyst, CPM is the conversion of PM, and t is the reaction time. An Arrhenius plot is shown in Figure 5b to express the temperature dependence of the forward reaction rate constant over the investigated range. A pre-exponential factor of k0 = 2.96 × 109 mol·s-1·kg (cat)-1 and an activation energy of E = 55.704 kJ·mol-1 were found for the transesterification reaction between PM and MeAc by the least square method. The activation energy is 66.503 kJ·mol-1, for the esterification of PM with acetate acid to produce PMA using ion-exchange resin Amberlyst 15 as catalyst 4. This shows that the transesterification reaction catalyzed by sodium methoxide is easier to occur than the esterification reaction catalyzed by Amberlyst 15 to produce PMA. With the information of k0, E 14


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and Keq, Eq. 5 can be applied to simulate the catalyzed transesterification of PM with MeAc.

The effect of initial MeAc-PM mole ratio on the reaction kinetics was investigated with MeAc-PM mole ratio from 2:1 to 4:1 at the catalyst concentration of 0.2% (wt) and the temperature of 323 K. To show the prediction accuracy of the derived kinetic model, these experimental results (symbols) are compared with the calculated results (solid lines), as shown in Figure 6. The calculated results are excellent agreement with the experimental results with a maximum RMSE of 2.78%.

3.3 Feasibility Analysis by RCM The feasibility analysis was conducted by the RCM 18, 30. The Damkohler number (Da) is an important parameter for the RCM model. By varying the value of Da, the effect of reaction extent on the thermodynamic behavior can be investigated 25. It is defined as the dimensionless ratio of characteristic liquid residence time (H0/V0) to characteristic reaction time (NPM0/kf,ref/Mcat):

Da =

H 0 / V0 N PM0 / ( kf ,ref M cat )

(10)

H0 and V0 are the initial molar liquid holdup and initial molar vapor rate, respectively. kf,ref is the forward reaction rate constant at reference temperature. The azeotropic temperature of MeAc-MeOH, 326.8K, is set as the reference temperature. The RCM can be mathematically described as 31:

15



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dxi = (1-D )( xi − yi ) + D ∑ ( vi , j − vT , j xi ) ℜ j dξ j =1

Page 16 of 46

(11)

xi and yi are the liquid and vapor phase mole fractions, respectively. dξ = (V / H ) dt is the dimensionless time. H and V are the instantaneous liquid holdup and the instantaneous vapor rate, respectively. vi is the stoichiometric coefficient of component i in the reaction. vT is the sum of the stoichiometric coefficients for the reaction, ℜ is the dimensionless reaction rate:

ℜ=

Cr r 1 Cp vp v = ∏ aˆr r − ∏ aˆp kf r =1 K eq p =1

(12)

DDa =Da/(1+Da) is a dimensionless parameter, which varies from zero to unity. In the case of DDa → 0, there is no reaction in the system. In the case of DDa → 1.0, the system is in the limit of reaction equilibrium.

Mathematically, the RCM model is a set of nonlinear differential equations. A complete residue curve can be obtained by numerically integrating these equations forward and backward in time, starting from an initial liquid composition. The numerical integration was performed by the MATLAB software with the ode45 solver. The thermodynamic and reaction kinetics parameters of the reactive system were needed when solving the RCMs model. The non-ideality of the reactive system studied was described by the NRTL model, and the required interaction parameters are listed in Tables 1 & 2. The information about the reaction kinetics is given in Section 3.2. The structure of the RCM is determined by the singular points, including unstable node, saddle point and stable node. The unstable and stable nodes are the 16


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origin and terminal of the residue curves, respectively. The unstable and stable nodes with physical meaning are considered to be the possible product composition of column top and bottom, respectively

32

. The singular points can be obtained by

solving Eq. 11 with dxi / dξ = 0 , and they depend on the value of DDa.

The RCM without reaction for the quaternary system PMA, PM, MeOH and MeAc is shown in Figure 7a. There are five singular points for this quaternary system at DDa = 0. The azeotrope between MeOH and MeAc, which is the minimum boiling point of this quaternary system, is the only unstable point. There are three saddle nodes, namely, the boiling points of pure PM, MeOH and MeAc. The boiling point of PMA, which is the maximum boiling point of this quaternary system, is the only stable node. All the trajectories start at the MeOH-MeAc azeotrope and end at the pure PMA. The terminal of the residue curves shows that, for the quaternary system of PMA + PM + MeOH + MeAc, it is possible to collect pure target product PMA from the bottom of a distillation column. In contrast, the obtainment of pure product MeOH by the simple distillation is impossible since a minimum azeotrope is formed between MeOH and MeAc.

The variation of singular points with the value of DDa is illustrated in Figure 7b. As soon as the reaction begins (DDa > 0), the MeOH-MeAc azeotrope changes from an unstable node to a saddle point. The unstable node has no physical meaning now, since some of its composition values are negative. Although the MeOH-MeAc 17



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azeotrope is no longer an unstable node, it remains the minimum boiling point of the reactive quaternary system. Thus, in a RD column for the transesterification of PM with MeOH, the MeOH-MeAc azeotrope is still a possible product of the column top. Except the MeOH-MeAc azeotrope, the other singular points continue their solution and stabilities at the whole range of DDa. The analysis on the variation of the singular points with DDa shows that it is possible to intensify the transesterification of PM with MeOH in a RD column with pure PMA as bottom product. The analysis in Figure 7b is for a fully RD column, but the analysis result is also suitable for a hybrid RD column. That is because the separating sections of the hybrid RD column mainly play the role of increasing the reactant concentration in the reactive section when they are added. The RD process for transesterification of PM with MeAc still suffers from the excessive consumption of MeAc, since the MeOH-MeAc azeotrope dose not disappear when the reaction occurs. In order to realize the complete PM conversion in a RD column, the minimum MeAc-PM mole ratio is 2.882 according to the mass conversation based on the azeotropic composition of MeOH-MeAc. The azeotrope MeOH-MeAc could be separated by extractive distillation with water as extraction agent 33.

3.4 Experimental Verification of RD Intensification Effect The PM conversions in the batch reactor (Figure 2b) and in the batch RD column (Figure 2c) are compared in Table 3 for the verification of RD intensification effect on the transesterification of PM with MeAc. As shown in Table 3, the PM conversion 18


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increases from 31.21% in the batch reactor to 66.70% in the batch distillation column at MeAc-PM mole ratio of 3:1, and from 38.94% to 75.35% at MeAc-PM mole ratio of 4:1. The low PM conversion in the batch reactor is attributed to the limit of chemical equilibrium. Although the increase of MeAc-PM mole ratio improves the PM conversion, it plays a limited role due to the accumulation of products in the batch reactor. The higher conversion of PM in the batch RD column should be contributed the fact that the chemical equilibrium is enormously shifted forward with the continuous removal of product MeOH from the RD column. The comparison in Table 3 indicates that the transesterification of PM with MeAc can be intensified by RD to improve the reactant conversion.

The variation of PM conversion with reaction time is illustrated in Figure 8 (the points marked as ■) for the batch RD column. At the beginning of the reaction, the PM transformed quickly due to the high reactant concentration and low product concentration. The PM conversion increases slightly with the progression of reaction, since the high concentration of PMA suppresses the forward reaction. Theoretically, PM can be completely transformed to PMA in a batch RD column above the MeAc-PM mole ratio of 2.882:1, due to the continuous removal of product MeOH. However, the full conversion of PM was not achieved, as shown in Figure 8 (the points marked as ■). It should be contributed to the over-withdrawal of reactant MeAc from the batch RD column at the later stage of the reaction. In this stage, the high concentration of PMA in the bottom of RD column suppresses the forward reaction, 19



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resulting in the extremely low formation rate of MeOH. In order to prevent the over-withdrawal of reactant MeAc which forms minimum azeotrope with product MeOH, the batch RD column should be operated at extremely low distillate volume due to the extremely low formation rate of MeOH, leading to the infinitely long operation time. Nevertheless, it is unpractical to conduct the batch RD experiment in infinitely long time. The effects of PM-MeAc mole ratio on PM conversion in the batch RD column were also investigated, as shown in Figure 8 (the points marked as ●). The PM conversion increases drastically from 66.70% to 89.0% with the MeAc-PM mole ratio increasing from 3:1 to 7:1. Above the mole ratio of 7:1, the PM conversion improves slightly. Obviously, complete PM conversion is still difficult to be realized in a batch RD column even at very high mole ratio of PM-MeAc, due to the high concentration of PMA in the bottom of RD column. The full conversion of PM is important for the simplification of the subsequent separation process. In order to achieve the complete conversion of PM, it is necessary to develop a continuous RD process for the transesterification of PM with MeAc to separate timely the products MeOH and PMA from the reactants.

For the achievement of continuous operation of RD column, the catalyst should be loaded in the column or continuously introduced into the column. However, neither of these two methods is feasible in present work, due to the usage of powder sodium methoxide as catalyst. Sodium methoxide is insoluble in the reactive mixture, so a mass of catalysts would accumulate at the packing surfaces if they were fed 20


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continuously with reactants into the column. The clustering of catalyst in the column would certainly results in the deterioration of vapor-liquid mass-transfer. On the other hand, the extremely small particle diameter of sodium methoxide hinders its direct loading in a column as the ion-exchange-resin catalyst. Although this kind of catalyst might hinder the continuous RD production of our transesterification system, the analysis results still show that it is thermodynamically feasible to produce PMA via the transesterification route in a RD column using basic catalyst. Heterogeneous solid catalysts are preferred for the RD process, because they can be easily loaded in the RD column. However, most of the basic heterogeneous catalysts with high activity suffer from either low lifetime or low activity 5-6. Therefore, the solid basic catalyst with both long lifetime and high activity should be developed in the future work to produce PMA via transesterification route in a RD column.

3.5 Conceptual Design for the Continuous RD Process If there is a kind of catalyst both suitable for the continuous operation of a RD column and sufficiently active for the PM-MeAc transesterification reaction, the RD process shown in Figure 9 can be applied as an alternative for the production of PMA. In this section, we would give some guidance through conceptual design for the future development of a continuous RD column for PMA production through the transesterification of PM with MeAc. The mathematical model and design procedure developed by Barbosa et al 34 and Qi et al 35 for a single-feed fully-reactive distillation column was used for the simplification of conceptual design. Supporting information 21



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lists the details about the mathematical model and design procedure.

Let us consider the following design problem: the MeAc-PM mixture (mole ratio of MeAc to PM = 3:1) is fed into a RD column, and the complete conversion of PM is expected. In order to simply the subsequent separation procedure, we want to obtain the highly purified product PMA as bottom, and the mixture of unreacted reactant MeAc and product MeOH as distillate. According to the design requirement, we set the desired compositions of bottom product as xMeOH, B = 0.0000, xMeAc, B = 0.0000, xPM, B = 0.0001 and xPMA, B = 0.9999. Through the mathematical model and design procedure given in the Supporting information, the transformed liquid composition profile are shown in Fig. 10a by specifying a reflux ratio of 0.8. Counting the stage numbers yields the total number of theoretical plate. The required number of reactive theoretical plate N decreases with the increase of reflux ratio r, as shown in Figure 10b. The dependence of N on r can been explained from two aspects. On the one hand, improving r would markedly increase the separation performance of reaction zone, which means that less theoretical plate is required to remove the reaction products from reaction zone. On the other hand, improving r would introduce more MeAc into the reaction zone to promote the formation of PMA due to the high concentration of MeAc in the distillate, which means that less reactive volume is required to realize the high conversion of PM. These two effects jointly lead to the reduction of N with increasing r. Both the minimum number of reactive theoretical plate Nmin (●) and minimum reflux ratio rmin (■) decreased with the increase of MeAc-PM feed ratio as 22


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illustrated in Figure 10c. The variation trend should be attributed to the facilitating effect of feed ratio on PMA formation. The Nmin is quite useful to estimate the minimum height of the reactive zone, if the height equivalent of theoretical plate (HETP) of the catalyst packing is known. The minimum reflux ratio can be used to calculate the actual reflux ratio by r = ηrmin, where η is a constant lying between 1.05 and 1.5 for conventional distillation. It is a trade-off between the capital and the operating costs for a reactive distillation column when an optimal value of η was chosen. Although the calculation results in Figure 10 are for a fully RD column and would be slightly different for a hybrid column, they are still meaningful in the primary design of a RD process.

4 CONCLUSION This work aims to study the feasibility of producing PMA by RD through the transesterification of PM with MeAc. The corresponding thermodynamic and reaction kinetic parameters are obtained to analyze the feasibility with RCM. The results of VLE experiment showed that there is only one azeotrope MeAc-MeOH in the quaternary system PMA + PM + MeOH + MeAc and that the NRTL model was suitable for the description of the non-ideality of this quaternary system. The reaction kinetics for the transesterification of PM with MeAc was determined with sodium-methoxide as catalyst, and the results indicated that the pseudo-homogeneous model was capable of representing the kinetic behavior of the reactive system investigated. The transesterification reaction of PM with MeAc is an endothermic 23



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reaction. The activation energy, standard enthalpy and standard entropy for this reaction are 55.704 kJ·mol-1, 41.265 kJ·mol-1 and 0.105 kJ·mol-1·K-1. According to the RCM analysis, it is feasible to conduct the PM-MeAc transesterification in a RD column with highly-purified PMA as bottom. Due to the existence of the azeotrope MeAc-MeOH, the mole ratio of MeAc to PM is required to be larger than 2.882 to realize the complete conversion of reactant PM. The feasibility was verified experimentally in a batch RD column. The minimum number of theoretical stage and minimum reflux ratio were provided through conceptual design to guide the development of a practical RD column for producing PMA through transesterification of PM with MeAc.

ASSOCIATED CONTENT Supporting Information VLE data of MeAc-PMA system at 101.3 kPa, VLE data of PMA + PM + MeOH + MeAc system at 101.3 kPa, mathematical model and design procedure for a single-feed fully-reactive distillation column

AUTHOR INFORMATION Corresponding Author *Email: [email protected]

ACKNOWLEDGMENTS 24


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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), the Education Department of Fujian Province (No. JAT160056), the Scientific Research Starting Foundation of Fuzhou University (No. XRC-1539) and the Valuable Instrument and Equipment Open Test Fund of Fuzhou University (Nos. 2017T033, 2017T035 and 2017T036).

NOMENCLATURE A, B, C

constants of the Antoine equation

CPM

conversion of PM

DDa

dimensionless parameter, DDa =Da/(1+Da)

Da

Damkohler number

E

activation energy, J·mol-1

Fob

objective function

H

instantaneous molar liquid holdup, mol

H0

initial molar liquid holdup, mol

J

parameter for thermodynamic consistency test

k0

pre-exponential factor, mol·s-1·kg (cat)-1

kf

forward reaction rate constant, mol·s-1·kg (cat)-1

kf,ref

forward reaction rate constant at reference temperature, mol·s-1·kg (cat)-1

Keq

activity-based chemical equilibrium constant

Mcat

mass of catalyst, kg 25



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n

number of experimental data set

N

number of reactive theoretical plate

Nmin

minimum number of reactive theoretical plate

NPM0

initial moles of PM, mol

P

system pressure, kPa

pis

saturated vapor pressure of pure component i, kPa

r

reflux ratio, mol·mol-1

~

r

transformed reflux ratio

rmin

minimum reflux ratio, mol·mol-1

−rPM

reaction rate of PM, mol·s-1·kg (cat)-1

R

ideal gas constant, J·mol-1·K-1

s

reboil ratio, mol·mol-1

~

s

transformed reboil ratio

t

time, s

T

temperature, K

Tmax

maximum boiling temperature for thermodynamic consistency test, K

Tmin

minimum boiling temperature for thermodynamic consistency test, K

V

instantaneous vapor rate, mol·s-1

V0

initial molar vapor rate, mol·s-1

vi

stoichiometric coefficient of component i in the reaction

vT

sum of the stoichiometric coefficients for the reaction

xi

mole composition i of liquid phase, mol·mol-1 26


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Xi

transformed liquid phase mole fraction of composition i

yi

mole composition i of vapor phase, mol·mol-1

Yi

transformed vapor phase mole fraction of composition i

Z

parameter for thermodynamic consistency test

∆ r GmΘ

reaction standard Gibbs free energy, J·mol-1

∆ r H mΘ

reaction standard enthalpy, J·mol-1

∆ r S mΘ

reaction standard entropy, J·mol-1·K-1

ℜ

dimensionless reaction rate

α

non-random parameter of NRTL model

aˆi

activity of component i

γi

activity coefficient of component i

∧

φiv

vapor phase fugacity coefficient of component i

φis

liquid phase fugacity coefficient of component i

ξ

dimensionless time

Abbreviations: RCM

residue curve map

RD

reactive distillation

RMSE

root mean square error

MeAc

methyl acetate

MeOH

methanol

NRTL

non-random two liquid 27



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PM

propylene glycol methyl ether

PMA

propylene glycol monomethyl ether acetate

VLE

vapor liquid equilibrium

Subscripts B

bottom

D

distillate

F

feed

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(13) Talebian-Kiakalaieh, A.; Amin, N. A. S.; Mazaheri, H., A review on novel processes of biodiesel production from waste cooking oil. Appl. Energ. 2013, 104, 683-710. (14) Zuo, C.; Ge, T.; Li, C.; Cao, S.; Zhang, S. Kinetic and reactive distillation for acrylic acid synthesis via transesterification. Ind. Eng. Chem. Res. 2016, 55, 8281-8291. (15) Fang, Y.; Xiao, W. Experimental and modeling studies on a homogeneous reactive distillation system for dimethyl carbonate synthesis by transesterification. Sep. Purif. Technol. 2004, 34, 255-263. (16) Wierschem, M.; Schlimper, S.; Heils, R.; Smirnova, I.; Kiss, A. A.; Skiborowski, M.; Lutze, P. Pilot-scale validation of enzymatic reactive distillation for butyl butyrate production. Chem. Eng. J. 2016, 312, 106-117. (17) Wang, H.; Wu, C.; Bu, X.; Tang, W.; Li, L.; Qiu, T. A benign preparation of sec-butanol via transesterification from sec-butyl acetate using the acidic imidazolium ionic liquids as catalysts. Chem. Eng. J. 2014, 246, 366-372. (18) Song, W.; Venimadhavan, G.; Manning, J. M.; Malone, M. F.; Doherty, M. F. Measurement of residue curve maps and heterogeneous kinetics in methyl acetate synthesis. Ind. Eng. Chem. Res. 1998, 37, 1917-1928. (19) Ye, C.; Dong, X.; Zhu, W.; Cai, D.; Qiu, T. Isobaric vapor–liquid equilibria of the binary mixtures propylene glycol methyl ether + propylene glycol methyl ether acetate, methyl acetate + propylene glycol methyl ether and methanol + propylene glycol methyl ether acetate at 101.3kPa. Fluid Phase Equilibr. 2014, 367, 45-50. 30


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(20) Tochigi, K.; Takahara, H.; Shiga, Y.; Kawase, Y. Isobaric vapor–liquid equilibria for water + propylene glycol monomethyl ether (PGME), water + propyleneglycol monomethyl ether acetate (PGMEA), and PGME + PGMEA at reduced pressures. Fluid Phase Equilibr. 2007, 260, 65-69. (21) Hsieh, C.; Lee, M.; Lin, H. Multiphase equilibria for mixtures containing acetic acid, water, propylene glycol monomethyl ether, and propylene glycol methyl ether acetate. Ind. Eng. Chem. Res. 2006, 45, 2123-2130. (22) Tu, C.; Wu, Y.; Liu, T. Isobaric vapor-liquid equilibria of the methanol, methyl acetate and methyl acrylate system at atmospheric pressure. Fluid Phase Equilibr.

1997, 135, 97-108. (23) Martin, M. C.; Mato, R. B. Isobaric vapor-liquid equilibrium for methyl acetate + methanol + water at 101.3 kPa. J. Chem. Eng. Data 1995, 40, 326-327. (24) Lin, Z. Y. Vapor-liquid equilibria for the binary systems of PGME + water and PGME + alcohols. Minghsin University of Science and Technology, 2009. (25) Okasinski, M. J.; Doherty, M. F. Thermodynamic behavior of reactive azeotropes. AIChE J. 1997, 43, 2227-2238. (26) Herington, E. F. G. A thermodynamic test for the internal consistency of experimental data on volatility ratios. NATURE 1947, 160, 610-611. (27) Herington, E. F. G., Tests for the Consistency of experimental isobaric vapor-liquid equilibrium data. J. Inst. Petrol. 1951, 37, 457-470. (28) Dodge, B. F. Chemical engineering thermodynamics; McGraw-Hill: New York, 1944. 31



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(29) Renon, H.; Prausnitz, J. M., Local compositions in thermodynamic excess functions for liquid mixtures. AIChE J. 1968, 14, 135-144. (30) Thiel, C.; Sundmacher, K.; Hoffmann, U. Residue curve maps for heterogeneously catalysed reactive distillation of fuel ethers MTBE and TAME. Chem. Eng. Sci. 1997, 52, 993-1005. (31) Venimadhavan, G.; Malone, M. F.; Doherty, M. F., Bifurcation study of kinetic effects in reactive distillation. AIChE J. 1999, 45, 546-556. (32) Steyer, F.; Qi, Z.; Sundmacher, K. Synthesis of cylohexanol by three-phase reactive distillation: influence of kinetics on phase equilibria. Chem. Eng. Sci. 2002, 57, 1511-1520. (33) Langston, P.; Hilal, N.; Shingfield, S.; Webb, S. Simulation and optimisation of extractive distillation with water as solvent. Chem. Eng. Process. 2005, 44, 345-351. (34) Barbosa, D.; Doherty, M. F., Design and minimum-reflux calculations for single-feed multicomponent reactive distillation columns. Chem. Eng. Sci. 1988, 43, 1523-1537. (35) Qi, Z.; Flockerzi, D.; Sundmacher, K. Singular points of reactive distillation systems. AIChE J. 2004, 50, 2866-2876.

32


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FIGURES

(a)

(b)

Figure 1 Two routes for the synthesis of PMA from PM. (a) Esterification of PM with acetic acid; (b) Transesterification of PM with MeAc.

33



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(a)

(b)

(c)

Figure 2 Schematic of the experimental apparatus for (a) VLE experiment, (b) reaction kinetic experiment, (c) batch RD experiment. 34


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Figure 3 T–x–y diagram for the binary system of MeAc-PMA at 101.3 kPa. “Exp” and “Cal” represent the experimental and calculated results, respectively. Uncertainty: u(P)= ±0.1 kPa; u(T) =±0.1 K; u(xMeAc) = u(yMeAc) =±0.0005.

35



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Figure 4 Effect of the reaction temperature on the PM conversion CPM. Initial mole ratio of MeAc to PM=2:1, concentration of catalyst = 0.2 (wt) %. Uncertainty: u(P)= ±0.1 kPa; u(T) =±0.15 K; u(xMeAc) = u(yMeAc) =±0.0004.

36


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(a)

(b)

Figure 5 (a) Equilibrium constant Keq as a function of reaction temperature T; (b) Forward reaction rate constant kf as a function of reaction temperature T.

37



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Figure 6 Effect of initial mole ratio of PM to MeAc on the PM conversion CPM. Catalyst concentration = 0.2 (wt) %, reaction temperature: 323.15K. “Exp” and “Cal” represent the experimental and calculated results, respectively. Uncertainty: u(P)= ±0.1 kPa; u(T) =±0.15 K; u(xMeAc) = u(yMeAc) =±0.0004.

38


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(a)

(b)

Figure 7 (a) Residue curve map (RCM) without reaction; (b) variation of significant points with DDa.

39



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Figure 8 Effect of reaction time t (■) and initial mole ratio of MeAc to PM (●) on PM conversion CPM. Uncertainty: u(P)= ±0.1 kPa; u(T) =±0.2 K; u(xMeAc) = u(yMeAc) =±0.0005.

40


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Figure 9 Continuous RD process for the production of PMA through the transesterification of PM with MeAc.

41



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(a)

(b)

(c)

Figure 10 Conceptual design of the continuous RD process for the transesterification of PM with MeAc. (a) Composition profile at reflux ratio of 0.8, feed MeAc-PM mole ratio of 3:1 and bottom composition of XMeAc,B = 0.9999 and XPM, B = 0.9999; (b) variation of required reactive-theoretical-plate number N with reflux ratio r; (c) variation of minimum reflux ratio rmin and minimum reactive-theoretical-plate number Nmin with feed MeAc-PM mole ratio. 42


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TABLES Table 1 Antoine equation parameters for the reaction system Component 19

PM PMA19 MeAc19 MeOH19

Antoine equation parameters

A

B

C

6.2428 6.2287 6.0078 7.0949

1323.40 1429.31 1076.05 1521.23

-80.66 -80.36 -61.91 -39.18

43



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Table 2 NRTL parameters of the reaction system NRTL parameters Binary system ∆g12 (J/mol)

∆g21 (J/mol)

PM (1) + PMA (2) 19

2739.4

-1878.6

MeAc (1) + PM (2) 19

4499.8

-2297.3

MeOH (1) + PMA (2) 19

3512.6

-1114.9

MeAc (1) + PMA (2)*

5026.2

-3318.1

MeAc (1) + MeOH (2) 23

1466.6

1617.8

MeOH (1) + PM (2) 24

948.6

599.4

α

0.3

*: the NRTL parameters for MeAc (1) + PMA (2) system was determined in present work.

∑τ G x G x = +∑ G x ∑G x ji

The NRTL model is

ln γ i

ji

j

j

ij

ki k

 ∑τ ljGlj xl τ − l  ij ∑k Gkj xk  

j

kj k

and

α ji = α ij

k

τ ji = ∆g ji / RT , G ji = exp ( −α jiτ ji )

j

44


  , where   

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Table 3 PM conversions in the batch reactor and in the batch RD column

Batch reactor

Batch RD column

Catalyst concentration

Initial mole ratio of

PM conversion

(wt)

MeAc to PM

(%)

0.2

3

31.21

0.2

4

38.94

0.2

3

66.70

0.2

4

75.35

45



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