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Analysis of intensification mechanism of auxiliary reaction on reactive distillation: Methyl acetate hydrolysis process as example
Chemical Engineering Science 106 (2014) 190–197
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Analysis of intensification mechanism of auxiliary reaction on reactive distillation: Methyl acetate hydrolysis process as example Liwei Tong a, Lifang Chen a, Yinmei Ye b, Zhiwen Qi a,n a Max Planck Partner Group at the State Key Laboratory of Chemical Engineering and School of Chemical Engineering, East China University of Science and Technology, Shanghai 200237, China b School of Resources and Environmental Engineering, East China University of Science and Technology, Shanghai 200237, China
H I G H L I G H T S
Concept of reactive distillation directly intensified by an auxiliary reaction is presented. RCMs is developed to analyze the intensification mechanism with two reactions and multiple components. The methyl acetate hydrolysis intensified by methanol dehydration is studied as example. A novel process of MeOAc hydrolysis is designed with 100% conversion of MeOAc. No additional H2O is needed in feed while high purity DME and HAc are products.
art ic l e i nf o
a b s t r a c t
Article history: Received 6 July 2013 Received in revised form 11 October 2013 Accepted 20 November 2013 Available online 3 December 2013
A reactive distillation (RD) process directly intensified by an auxiliary reaction is presented. The methyl acetate (MeOAc) hydrolysis intensified by methanol (MeOH) dehydration is studied as an example where the latter serves as the auxiliary reaction. A residue curve maps (RCMs) is developed to analyze the mechanism of intensification for the case with two reactions and multiple components. The analysis of thermodynamics and RCMs reveals that the reaction of MeOH dehydration is the controlling step which limits the conversion of MeOAc. Based on the chemical equilibrium and kinetically controlled design, a novel process is developed to hydrolyze MeOAc and MeOH in the process of polyvinyl alcohol. Compared to traditional processes, no additional water is required to feed into the RD column and the process is significantly simplified. With equal mole of MeOAc and MeOH as feed and a set of pre-reactor, closely 100% conversion of MeOAc and MeOH can be achieved and high purity dimethyl ether and acetic acid are the products in the RD column. & 2013 Elsevier Ltd. All rights reserved.
Keywords: Reactive distillation Process intensification Auxiliary reaction Methyl acetate Hydrolysis
1. Introduction Reactive distillation (RD) provides an attractive alternative for process intensification with advantages of simultaneous reaction and separation (Kaymak and Luyben, 2004). One of the ideal process configurations of RD consists of a column. The light and heavy reactants are fed at the lower and upper parts of the reactive zone while the heavy and light components are bottom and top products, respectively. Typical examples mainly include the productions of ethers of gasoline additives and ester such as methyl acetate, n-butyl acetate and amyl acetate (Sundmacher and Kienle, 2003).
n
Corresponding author. Fax: þ86 21 6425 3528. E-mail address: [email protected] (Z. Qi).
0009-2509/$ - see front matter & 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.ces.2013.11.036
However, the advantages of RD may be inhibited by their complicated physical properties of the reaction systems. One of the most common problems is the azeotropic mixture in the system. High conversion can be blocked when the reactants exist in the lightest azeotropic mixtures since they are likely to flee out of the column without taking reactions. Consequently, the configuration of the RD column has to be modified and extra equipment is required to recycle the reactants back to the RD column, as referred to the productions of ethyl acetate and isopropyl acetate (Lai et al., 2007). Another problem of RD is the excessive amount of reactants remaining in the system. In the case of a reversible reaction with very low chemical equilibrium constant (Keq), one of the reactants has to be excessive to improve the conversion of another reactant though RD can break through the limitation of chemical equilibrium. Further separation is thus required to deal with the excessive reactant in the downstream operation. In the case of liquid phase split, the reactants may not directly contact
L. Tong et al. / Chemical Engineering Science 106 (2014) 190–197
with the catalyst, and mass transfer resistances significantly limit the reaction rate. In this contribution, a concept of RD intensified by an auxiliary reaction is proposed to solve the above bottlenecks. By introducing an auxiliary reaction, the characteristics of original system such as thermodynamics, reaction rates and alternative products are modified, which will intensify the system and simplify the RD process. The indirect hydration of cyclohexene to cyclohexanol using a reactive entrainer is a typical example of intensification (Steyer et al., 2008; Katariya et al., 2009). As the liquid–liquid phase splitting in the system of cyclohexene/water limits the reaction rate, formic acid as the reactive entrainer is introduced into the system. Cyclohexene first reacts with formic acid forming formic acid cyclohexyl ester, and then this ester can be split with water into cyclohexanol and formic acid. The reaction route offers the advantage of overcoming reaction rate limits and having no significant amounts of side products. Another example is the process of methyl acetate (MeOAc) hydrolysis which will be investigated in this work. The effect of intensification on reactive distillation process has been presented in our previous work (Tong et al., 2013). However, the mechanism of intensification remains unclear since there is no existing suitable method to explore it. Thus, in this work, a novel residue curve maps (RCMs) tool is developed to explicitly explain the intensification mechanism. Then, the controlling step of the whole process is figured out by analyzing the thermodynamics and RCMs of the involved reaction systems. Finally, a novel process of hydrolysis of MeOAc is developed based on the equilibrium and kinetically controlled design. Compared with conventional processes, the novel process can be significantly simplified with 99.78% conversion of MeOAc. Furthermore, high purity dimethyl ether (DME) as an additional product can be obtained through methanol (MeOH) dehydration.
Fig. 1. Intensification mechanism of the auxiliary reaction.
total reflux design and removal of the light product (MeOH), heavy product (HAc), and excess reactant (H2O) from the bottoms of the column are reasonable choices (Fuchigami, 1990). This implies that high energy consumption is required at the condenser and reboiler in RD column. The second problem of MeOAc hydrolysis is related to the downstream separation process. Because of the low chemical equilibrium constant Keq, high feed ratio of H2O/MeOAc (normally 8–15) (Kim and Roh, 1998; Wang et al., 2001; Xiao et al., 2001; Lee, 2002) is employed to enhance the conversion of MeOAc as high as possible. However, the binary vapor–liquid equilibrium (VLE) diagram for the H2O-HAc system (see Fig. 1 of Lin et al., 2008) shows a tangent pinch point existing near the pure H2O vertex. This suggests that high reflux ratio or reboiler duty is required to remove HAc from H2O. Moreover, the excessive amount of H2O as feed in RD column would also cause high energy consumptions in the downstream separation process. Therefore, it is essential to develop an approach to intensify the process of MeOAc hydrolysis, namely, to obtain high MeOAc conversion with low energy requirement and “neat” operation. MeOAc þ H2 O 3 MeOH þ HAc K eq ¼
2. The process of MeOAc recovery in PVA production The PVA effluent contains a large amount of MeOAc as byproduct, around 1.68 times of the PVA product by weight (Zeng and Wu, 2009). Since the utility of MeOAc is limited, the major approach is to hydrolyze MeOAc back to HAc and MeOH (Eq. (1)), as raw materials for the PVA plant. Usually, two major problems are arisen in the process due to the system characteristics. The first is two minimum-boiling binary azeotropes formed by (1) MeOAc (44.88 mol%) and MeOH at 131.03 1C, and (2) MeOAc (74.74 mol%) and H2O at 138.05 1C. The order of boiling point temperature for pure components and azeotropes of the system at 1.0 MPa is illustrated in Table 1. Theoretically, pure light product (MeOH) can be obtained at the top of a single RD column if all MeOAc is consumed (Tung and Yu, 2007). However, MeOH is a saddle in the RCM diagram (Tang et al., 2005). With a low equilibrium constant (Keq 0.013, Eq. (2)), it is unlikely to obtain high purity product of MeOH under the “neat” operation. Thus, Table 1 Singular points in Fig. 2 (at 1.0 MPa). Name
Type
DME Unstable Azeotrope 2 Unstable/ Saddle MeOH Saddle Azeotrope 1 Saddle MeOAc Saddle H2O Saddle HAc Stable
Temp. (1C)
Composition (mole fraction) MeOH HAc DME
44.68 130.76
0 0 0.4488 0
0 0 0.5512 0
1 0
137.01 138.71 143.91 180.32 214.08
0 0.7474 1 0 0
1 0 0 0 0
0 0 0 0 0
0 0 0 0 1
½MeOH½HAc ½MeOAc½H 2 O
ð1Þ ð2Þ
To reach this goal, MeOH dehydration, as an auxiliary reaction, is introduced to intensify the MeOAc hydrolysis. As indicated in Fig. 1, the MeOH produced by the hydrolysis reaction is further dehydrated into H2O which can circularly react in the hydrolysis and thus enhance the conversion of MeOAc. In the system, H2O acts as an inner cycling component and the two reactions are “selffeeding”. As a result, no extra H2O is needed and MeOH is in-situ transferred to DME. The energy consumption of complicated process to separate the excess H2O in the downstream could be significantly saved. In the research of Dirk-Faitakis et al. (2009), it proved that these two reactions can proceed co-currently in a RD column to produce both DME and HAc. A batch catalytic distillation experiment showed that the ion exchange resin (Amberlyst 35) was active for both reactions without any initial water present in the feed. However, simulations of continuous RD were based on the assumption of both physical and chemical equilibrium, which was an extreme situation and it is usually unrealistic in practice. Thus, the kinetic models should be introduced to further investigate the intensification mechanism and a process based on kinetically controlled reaction condition will be developed.
3. Analysis of intensification mechanism
MeOAc H2O
0 0.2526 0 1 0
191
Reactive distillation integrates chemical reaction and distillation in one unit. Thus, intensification on reactive distillation process could be realized by improving either of them, which is the motivation of this work. The kinetic models used in this work are from Song et al. (1998) (see Section 3.1) where Wilson model is used for predicting the liquid activity coefficient of the VLE calculation and the model parameters are taken from literature (see Table 3 of Song et al., 1998). The vapor phase dimerization
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(for HAc) is also taken into account using the Hayden-O’Connell second virial coefficient model (Hayden and O’Connell, 1975) and the model parameters are from Aspen Plus database.
The model equations for RCMs are presented in Eq. (10) (for the detailed derivation of the model, refer to Venimadhavan et al., 1994,1999a,1999b).
3.1. Kinetic models
N dxi ¼ ð1 DÞðxi yi Þ þ D ∑ ðνi;j νT;j xi Þℜj dξ j¼1
For the reaction of MeOAc hydrolysis, the kinetic model can be referred to its reversible reaction, namely, MeOAc synthesis. Song et al. (1998) and Pöpken et al. (2000) established the kinetic model with Amberlyst 15 as catalyst, and the MeOH dehydration, as a side reaction, was also considered in the work of Song et al. (1998). Therefore, the kinetic models of Song et al. (1998) are employed in this work since the kinetic parameters for both reactions are obtained in the same experimental condition. The reaction kinetics for MeOAc hydrolysis in a Langmuir– Hinshelwood–Hougen–Watson rate model is r1 ¼
Wks;1 ðαMeOAc αH2 O ðαMeOH αHAc =ÞK eq1 Þ
ð1 þ K MeOAc αMeOAc þ K H2 O αH20 O þ K MeOH αMeOH þ K HAc αHAc Þ2
Inðks;1 =WÞ ¼ 24:64
6287:7 TðKÞ
782:98 K eq1 ¼ 1=exp 0:83983 þ TðKÞ
ð3Þ
ð4Þ ð5Þ
with K H2 O ¼ 10.50, KMeOH ¼4.95, KHAc ¼ 3.18, KMeOAc ¼ 0.82. ks,1 is the apparent forward reaction rate constant with units of (mol of mixture) (mol of H þ ions) 1 min 1, Keq1 stands for the chemical equilibrium constant for MeOAc hydrolysis, W represents the concentration of catalyst with units of (mol of H þ ions) (mol of mixture) 1, Ki is the adsorption equilibrium constant of component i. In the kinetic model for MeOH dehydration, it is assumed that DME is extremely volatile and has no significant presence on the catalyst surface, as represented in Eq. (6): r2 ¼
Wks;2 K 2MeOH ðα2MeOH ðαDME αH2 O =K eq2 ÞÞ ð1 þ K H2 O αH2 O þK MeOH αMeOH Þ2
K eq2 ¼ exp
ΔG 0
ð6Þ
!
RTðKÞ
Inðks;2 =WÞ ¼ 27:40
; ΔG0 ¼ 2:4634 1:5167 10 3 TðKÞ 10654:0 TðKÞ
ð7Þ
ð8Þ
where ΔG0 is free energy of the reaction in kcal/mol, ks,2 has the same units as ks,1, Keq2 is the chemical equilibrium constant for MeOH dehydration (Nisoli et al., 1997). The adsorption equilibrium constants are assumed to have the same values as above.
ði ¼ 1; …NC; j ¼ 1; ………NRÞ
ð10Þ
where xi and yi are the liquid and vapor phase mole fractions, respectively; dξ the dimensionless time with dξ ¼(V/H)dt, H the instantaneous liquid holdups in the still moles, V the instantaneous vapor rates, t the time, vi the stoichiometric coefficient of component i in the reaction, vT the sum of the stoichiometric coefficients for the reaction, ℜj the normalized reaction vector with ℜj ¼ ðr j =kf ;ref Þ, rj the rate of an individual reaction j and D¼
Da 1 þ Da
ð11Þ
where D is the dimensionless parameter of Da, which varies from zero to unity. In the case of D¼ 0, there is no reaction in the system so the effect of distillation can be investigated alone. In contrast, in the case of D-1.0, the system is in the limit of reaction equilibrium where the effect of reaction can be studied. As there are multiple components (i.e., MeOAc, H2O, MeOH, HAc and DME) and two reactions (MeOAc hydrolysis and MeOH dehydration) in the system, the traditional topological diagram in literature is not suitable. Therefore, in this work, a novel method of RCMs expression is developed for analysis of the intensification mechanism with two reactions and multiple components. For an equilibrium RD, the concept of reactive distillation lines (Bessling et al., 1997) on the basis of the transformation of concentration coordinates (Barbosa and Doherty, 1987) is used to study the feasibility of a reactive distillation. However, in the case of a kinetically controlled RD, the RCMs with five components are difficult for visualization in a usual tetrahedron. The color of curve is used to express additional information to solve this problem. The components of A, B, C and Dþ E are treated as four vertexes of a tetrahedron while the color of curve is represented by the mole fraction ratio of D/(D þ E). Therefore, the information of five components could be clearly read from the RCMs. It is essential to investigate which reaction rate is faster since the two reactions are simultaneously taking place and they are dependent on each other by “self-feeding”. Thus, the chemical equilibrium surface (CES) or curve is also illustrated in the RCMs. By observing the trajectories of RCMs and the CES at different values of D in Eq. (10), the reaction reaching chemical equilibrium can be confirmed. 3.3. Distillation intensification effect (D ¼ 0)
3.2. Residue curve maps with two reactions and multiple components One widely used tool for synthesis and analysis of both nonreactive and reactive distillation systems is RCMs which is the phase plane of the liquid compositions in an isobaric open evaporation. An important parameter in the study of RCMs for reactive system is the Damköhler number (Da) defined by the ratio of a characteristic residence time to a characteristic reaction time. Here, it is defined as Da ¼
H 0 =V 0 1=ðkf ;ref WÞ
ð9Þ
where H0 is the initial liquid holdup in the still in mole, V0 is the initial vapor rate from the still, kf,ref is the reference forward reaction rate constant (defined at a convenient temperature such as the lowest boiling temperature in the mixture).
In the conceptual design of RD, one of the ideal cases is that all reactants are intermediate components while the desired products should be either the lightest or the heaviest ones in the system. However, the system of MeOAc hydrolysis is not satisfied with such conditions. Fig. 2a displays the RCMs for MeOAc system without reaction (D¼0). In the RCMs, all of the trajectories start at the MeOAc/MeOH azeotrope (Az 2, the lightest boiler and unstable node), go via the intermediate boilers (saddle), and end at the pure HAc vertex (stable node). Thus, MeOH as a desired product is an intermediate component though HAc as another desired product is the heaviest one. Worse yet, part of MeOAc would flee out of the column without having reaction since it exists in Az 2. Once the MeOH dehydration is activated, the new product DME would replace Az 2 to be the lightest one in the system. The overall reaction in Fig. 1 implies that reactants are the intermediate components while products are the heaviest and lightest components,
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Fig. 2. RCMs of different systems at 1.0 MPa. (a) simple distillation for MeOAc hydrolysis system (D ¼ 0) and (b) simple distillation for simultaneous MeOAc hydrolysis and MeOH dehydration (D ¼ 0) □ saddle node 〇 unstable node ● stable node.
Fig. 3. Effect of temperature on chemical equilibrium constant Keq3 ¼ Keq1 Keq2.
respectively, which is ideal for RD. Fig. 2b shows the RCMs for the systems of MeOAc and DME without reaction (D¼0) where the color of curves represents the mole fraction ratio of H2O/(MeOHþH2O), and the (MeOHþH2O) is considered as one component. Az 1 and Az 2 at the edge of MeOAc/(MeOHþH2O) consist of MeOAc/H2O and MeOAc/ MeOH (see Table 1), respectively. Compared with Fig. 2a, all trajectories end at the same point (pure HAc) but start at DME (the lightest component). Therefore, MeOAc is forced to stay in the middle of the RD column by the effect of distillation which would promote its conversion. 3.4. Reaction intensification effect (D-1.0) The MeOAc hydrolysis is an endothermic reaction with the reaction heat ofΔH 0298 ¼ 10:87 kJ=mol while the MeOH dehydration is an exothermic reaction (ΔH 0298 ¼ 35:27 kJ=mol). Hence, the thermal coupling of the two reactions is helpful for keeping temperature stable in the reactive section. The effect of temperature on the chemical equilibrium constant is depicted in Fig. 3. As seen, the chemical equilibrium constant of the overall reaction Keq3 (Eq. (12)) gradually declines by rising the temperature implying that low temperature is favorable for the overall reaction. However, in the work of Song et al. (1998), it was found that the MeOH dehydration was not significant at low temperature (40–50 1C). Nevertheless, at higher temperature, i.e., 85 115 1C, significant amount of DME were formed. As the MeOH dehydration should be activated on purpose in this work, the reaction temperature over 85 1C is taken into account. K eq3 ¼
½DME½HAc ¼ K eq1 K eq2 ½MeOAc½MeOH
ð12Þ
Fig. 4. Intensification of reaction by introduction of MeOH dehydration at 120 1C (D-1.0). single MeOAc hydrolysis ( □ dashed line); simultaneous MeOAc hydrolysis and MeOH dehydration ( 〇 straight line).
Fig. 4 shows the difference of RCMs between the case of single MeOAc hydrolysis and the case of simultaneous MeOAc hydrolysis and MeOH dehydration at different initial points (i.e., H2O/ MeOAc). For the case of single reaction, as the RCMs is at chemical equilibrium condition (D-1.0), which is also the limit case, the trajectories (dashed line) swiftly close to the CES and intersect at the points marked by □, go along with the CES and end at the pure HAc vertex. In the case of two reactions, the xi (components of MeOAc, H2O, MeOH and HAc) calculated from Eq. (10) are reunified by summation of compositions of these four compounds so that the trajectories can track inside of the tetrahedron. The trajectories (solid line) also approach to the CES and intersect at the points marked by 〇. However, the locations of intersection are different from those in the case of single reaction. For instance, at the initial point with 0.5 MeOAc and 0.5H2O, P1 and P2 are the intersections of trajectories with the CES for two reactions (solid line) and single reaction (dashed line), respectively. It is found that P1 is more far away from the edge of MeOAc-H2O (i.e., closer to edge of MeOH-HAc). This implies that the auxiliary reaction (MeOH dehydration) makes the chemical equilibrium of MeOAc hydrolysis move to the forward direction since MeOH is converted into H2O by MeOH dehydration. 3.5. Controlling step The simultaneous MeOAc hydrolysis and MeOH dehydration are consecutive and inner-cycled reactions as indicated in Fig. 1.
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Thus, it is important to figure out which reaction is the controlling step. The components xi (i.e., MeOAc, H2O, MeOH, HAc and DME) in Eq. (10) calculated at simultaneous reactions is used to acquire the RCMs in the system of MeOAc hydrolysis and MeOH dehydration, respectively. For the system of MeOAc hydrolysis, the components xi (i.e., MeOAc, H2O, MeOH and HAc) calculated by Eq. (10) are reunified as in Fig. 4. Fig. 5 shows the RCMs at different values of D. From D¼ 0.1 to D ¼0.7, as seen, all trajectories go along with the CES, which means the rate of MeOAc hydrolysis is fast enough to reach its chemical equilibrium even at the low value of D. In the case of a single MeOH dehydration, the components xi (i.e., H2O, MeOH and DME) calculated by Eq. (10) are reunified as in Fig. 4 to ensure the trajectories go inside of the triangle. The RCMs at different values of D is illustrated in Fig. 6. It is found that all trajectories stay far away from the chemical equilibrium curve at the low D. Only when D is relatively high (i.e., D¼ 0.7), the trajectory would approach the chemical equilibrium curve. This is attributed to the low reaction rate of MeOH dehydration at the low D. As a sequence, the two reaction rates are not synchronous, and
MeOAc hydrolysis reaches its chemical equilibrium in advance. The conversion of MeOAc mainly depends on the reaction rate of MeOH dehydration because it is the only H2O supplier in the system. A high reaction rate of MeOH dehydration could produce excess H2O that is equivalent to giving a high feed ratio of H2O/MeOAc as feed. In contrast, the whole process will be limited by the low rate of MeOH dehydration.
4. Novel processes of MeOAc hydrolysis As MeOH dehydration can supply H2O for the system, it is unnecessary to add additional H2O into the RD column. Hence, a mixed MeOAc and MeOH feed is an alternative for MeOAc hydrolysis in RD process. As mentioned before, in the research of Dirk-Faitakis et al. (2009), the RD process was studied assuming that both physical and chemical equilibrium were reached. However, the equilibrium design is a critical case for the conceptual design of RD. Therefore, in the following section, both equilibrium and kinetically controlled RD are taken into account. 4.1. Equilibrium design In the conceptual design of a RD process, the first step in the procedure is to develop a spectrum of designs at both phase and reaction equilibria. The concept of reactive distillation line proposed by Bessling et al. (1997) is usually used to study the feasibility of RD process. Fig. 7 depicts the reactive distillation line for the system MeOAc, H2O, MeOH, HAc and DME where HAc and MeOH are the reference components, giving X3 ¼xMeOAc þ xHAc and X4 ¼xH2O þ (1/2)xHAc þ(1/2)xMeOH. The MeOAc hydrolysis and MeOH dehydration simultaneously react in the system. As shown in Fig. 7, the system has no reactive distillation borders and all reactive distillation lines start from the HAc vertex and end at the DME vertex. Thus, the process of RD is feasible and probably an economic operation since both products are connected by a reactive distillation line, the products are (unstable and stable) nodes in a reactive distillation line diagram, and the boiling point difference between the products is large. The feed ratio of MeOAc/ MeOH is equal to one so that pure DME and HAc could be the top and bottom product, respectively. This is because, in the overall reaction, DME and HAc as products are the lightest and heaviest component, while MeOAc and MeOH as reactants are both the
Fig. 5. RCMs in chemical equilibrium surface of the MeOAc hydrolysis system at 120 1C, for different values of D.
Fig. 6. RCMs in chemical equilibrium curve of the MeOH dehydration system at 120 1C, for different values of D.
Fig. 7. Reactive distillation line for MeOAc/DME system at 0.5 MPa. Concentration profile of a RD column with 30 stages. The input data for the simulation are molar ratio MeOAc/MeOH¼ 1 as feed; reflux ratio¼1.5; feed stage¼15; column pressure¼ 0.5 MPa (The marks are the same as in Fig. 2).
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the equilibrium design can be modified to achieve the desired products in a kinetically controlled column. An equilibrium stage model of RD was applied and implemented in gPROMS for kinetically controlled design and simulation. The first column simulation is performed at the reaction equilibrium conditions. Further simulations are performed at successively lower Da until a realistic value is achieved. Here, Da is defined as Da ¼
Fig. 8. Equilibrium design at 0.5 MPa for MeOA hydrolysis by feed of MeOAc and MeOH.
intermediate components. As a result, MeOAc and MeOH will be kept inside of the RD column all the time until getting 100% consumed if they are the equal mole in the feed, which is proved by the concentration profile line in Fig. 7. The final equilibrium design for MeOAc hydrolysis is presented in Fig. 8. The mole fraction of DME and HAc are over 0.99 in the distillate and bottom, respectively. The operation pressure in the RD column is set to 0.5 MPa so as to ensure the reaction temperature over 85 1C (as mentioned in Section 3.3). There are 30 stages in total, each with an equilibrium chemical reaction (the stage of condenser and reboiler has no reaction, so the effective number of chemical equilibrium stage is 28). The lowest reflux ratio corresponding to 30 stages is 1.5, which is chosen as the base-case design value. The designed configurations are similar to the results in Dirk-Faitakis et al. (2009) (see Table S1 in the supporting information).
W T H RT kf ;ref FT
ð13Þ
where FT is the total feed flow rate to the column; WT is the total concentration of catalyst with units of (mol of H þ ions) (mol of mixture) 1; H RT is the total holdup in the reactive zone. The comparison between the reaction equilibrium design and a kinetic simulation (Da¼100) is shown in Fig. 9. It shows that the performance of the kinetic condition approaches the reaction equilibrium design (at very high Da of 1000). For the kinetic simulation, the conversion of MeOAc is 87.56% since the amount of catalyst loading is not enough (corresponding to Da¼100). However, the equilibrium design gives a better performance of RD column. High purity DME and HAc as the products at the top and bottom of RD column are obtained while the high concentration of MeOAc is forced to stay in the reactive zone. The concentration of H2O in the reactive zone is kept at a low level since H2O is supplied by the auxiliary dehydration of MeOH and simultaneously consumed by MeOAc hydrolysis. Fig. 9 also implies that the reaction zone could be in the middle of column (e.g. 5–25 stage) since it performs essentially no reaction outside of this range. As defined in Eq. (13), high Da means a high catalyst loading for a specified FT, which may not be proper for realistic design of RD column. In order to avoid a high catalyst loading, a high reflux ratio may be employed. Thus, the effect of reflux ratio on the conversion of MeOAc will be studied, as reported in Fig. 10. For these calculations, the distillate flow rate remains a constant of 1.25 mol/s (this implies bottom rate is also a constant) and the reaction zone is stage 5–25. As shown in Fig. 10, higher conversions can be obtained at higher Da (i.e., higher catalyst loading) at the same reflux ratio. However, at Da4 100, the gain in conversion is very weak compared to the required increase in Da. Therefore,
4.2. Kinetically controlled design As the RD design based on the chemical equilibrium condition very often makes it practically unrealistic, it is necessary to conduct an analysis of influence of kinetic effect through, e.g. bifurcation analysis of the steady-state solutions of Eq. (10) with D as parameter. The advanced process simulator gPROMS is applied in this work. It can export the mathematical information of models such as Jacobian matrix by Foreign Process Interface (F.P.I) (gPROMS, 2003), which makes it available to compile a continuation algorithm such as arc-length continuation to trace any bifurcated branches. Meanwhile, the stability of the solutions could be verified by analysis of Jacobian matrix. Thus, an arclength continuation based on gPROMS platform is established to calculate all branches. The starting points for the analysis are the singular points for the nonreactive mixture, which are the pure components and the nonreactive azeotropes at D ¼0. It was found that the saddle node of azeotropes and MeOH would cease to be solutions as soon as the reactions were initiated while other fixed points remained the same through the entire kinetic regime up to the reaction equilibrium limit (see Fig. S1 in the supporting information). Therefore, there are no any kinetic and reactive azeotropes in the reacting system. The desired separation (as described in the previous section) that was found to be feasible at reaction equilibrium is also feasible for kinetic condition, and
Fig. 9. Comparison of RD column profiles between the simulation (Da ¼ 100) and equilibrium design.
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Table 2 Comparison of simulation results for single RD column and plus pre-reactor.
Fig. 10. Influence of reflux ratio on the conversion of MeOAc and Vcat/Vstage for different values of Da. (a) design parameters (b) reflux ratio vs. conversion and Vcat/Vstage.
further improving the performance of the column can be realized by increasing the reflux ratio. Fig. 10 also shows the influence of reflux ratio on Vcat/Vstage for different Da. Here, Vcat is the volume of catalyst per stage; Vstage is the volume of stage and V stage ¼
L HETP SD
ð14Þ
where L is the liquid flow rate of stage in the column; SD is the spraying density of stage in the column (set to 10 m3/m2/h); HETP is the height equivalent of theoretical stage. In a realistic design of RD column with ion exchange resin as catalyst, the CDTech bales are widely used packing. The HETP is set to 0.5 m, and the value of Vcat/Vstage has to be less than 0.25. Therefore, the catalyst loading of 6.9 kg/stage (Da¼100) and the reflux ratio of 4.2 are recommended to achieve 100% conversion of MeOAc. It is noticeable that the reflux ratio of 4.2 would lead to high energy consumptions in the reactive distillation process. This is because the MeOH dehydration as the controlling step requires a long residence time (i.e., high reflux ratio) to get 100% conversion of MeOAc. In order to reduce the reflux ratio and save energy of the reactive distillation column, a pre-reactor could be introduced, and the feed (MeOAc:MeOH¼1:1) is designed for the RD column. Table 2 compares the performance of RD column between single RD column and plus pre-reactor where the overall reaction got
Operation parameters
Single RD column
Plus pre-reactor
Total stages Rectifying zone Reactive zone Stripping zone Column pressure (Mpa) Reflux ratio (mol) Damköhler number (Da) catalyst loading (kg/stage) Feed stage Feed flowrate (mol/s)
30 1–4 5–25 26–30 0.5 4.2 100 6.90 15 2.5
30 1–4 5–25 26–30 0.5 2.1 100 6.90 15 2.5
xF (mole fraction of feed) MeOAc H2O MeOH HAc DME Top product flow rate (mol/s)
0.5 0.0 0.5 0.0 0.0 1.25
0.3595 0.1679 0.0237 0.1405 0.3084 1.25
xD (mole fraction of top product) MeOAc H2O MeOH HAc DME Bottom product flow rate (mol/s)
0.0044 o1.0 10 5 0.0080 0.0009 0.9946 1.25
0.0031 o 1.0 10 5 0.0025 o 1.0 10 5 0.9943 1.25
xB (mole fraction of bottom product) MeOAc H2O MeOH HAc DME MeOAc Conversion (%) MeOH Conversion (%) Condenser heat duty (kW) Reboiler heat duty (kW)
0.0052 0.0044 o1.0 10 5 0.9904 o1.0 10 5 99.04 99.90 124.58 168.99
0.0057 0.00309 o 1.0 10 5 0.9912 o 1.0 10 5 99.78 99.08 74.26 114.85
equilibrium in the pre-reactor. In the pre-reactor, 28.12% conversion of MeOAc can be achieved. Accordingly, the heat duty of the condenser and the reboiler of the RD column are saved by 40.56% and 32.04%, respectively. This is because MeOH has been partly converted into H2O in the pre-reactor, which is equivalent to a promoted reaction rate in a single RD column. Compared with traditional process (see Fig. 1 of Tong et al., 2013), the downstream separation are totally saved while high purity of DME (4 0.99) and HAc (4 0.99) are produced in the RD column.
5. Conclusions The concept of reactive distillation directly intensified by an auxiliary reaction is presented. A novel RCMs method is developed for the case with two reactions and multiple components, which is applied to analyze the intensification mechanism of the auxiliary reaction. As an example, the MeOAc hydrolysis intensified by the reaction of MeOH dehydration in the RD process is analyzed. Based on the analysis of thermodynamics and RCMs of the system, the reaction rate of MeOH dehydration is determined as the controlling step to block 100% conversion of MeOAc. With equal mole of MeOAc and MeOH as feed, a novel process has been designed to hydrolyze MeOAc in the PVA production. As the limit of MeOH dehydration, it is not economical to give 100% conversion of MeOAc and MeOH in a single RD column. Thus, a pre-reactor is set to compensate for the low rate of MeOH dehydration. Compared with traditional processes, two distillation columns in the downstream separation are saved while high purity of DME and HAc are the products in the RD column.
L. Tong et al. / Chemical Engineering Science 106 (2014) 190–197
Notation ai dj D Da FT ΔG0 H HETP H0 H RT ks,1 ks,2 kf,ref Keq1 Keq2 Ki L NC NR rj R ℜ SD T V0 V Vcat Vstage W WT xi Xi yi
activity of component i index of rate amplification for reaction j dimensionless parameter of Da Damköhler number total feed flow rate to RD column, mol/s free energy of the reaction, kcal/mol molar liquid holdup, mol height equivalent of theoretical stage, m initial molar liquid holdup, mol total molar liquid holdup in reactive zone, mol forward reaction rate constant of MeOAc hydrolysis, (mol of mixture) (mol H þ ions) 1 min 1 forward reaction rate constant of MeOH dehydration, (mol of mixture) (mol H þ ions) 1 min 1 forward reaction rate constant at reference temperature, (mol of mixture) (mol H þ ions) 1 min 1 chemical equilibrium constant of MeOAc hydrolysis chemical equilibrium constant of MeOH dehydration adsorption equilibrium constant for component i liquid flowrate of stage in RD column, m3/h number of components number of reactions reaction rate for reaction j, mol/s universal gas constant (1.987 cal/mol K) normalized reaction rate, dimensionless spraying density of stage in RD column, m3/m2/h reaction temperature, K initial vapor molar flow rate, mol/s vapor molar flow rate, mol/s volume of catalyst loading per stage, m3 volume of stage in the RD column, m3 concentration of catalyst, mol H þ ions/mol of mixture total concentration of catalyst in RD column, mol H þ ions/mol of mixture liquid phase mole fraction of component i transformed liquid phase mole fraction of component i vapor phase mole fraction of component i
Greek letters
νi νT ξ
stoichiometric coefficient of component i total mole change of reaction dimensionless time, dξ ¼ (V)/(H)dt
Abbreviations CES DME HAc MeOAc MeOH PVA RCMs RD
chemical equilibrium surface dimethyl ether acetic acid methyl acetate methanol polyvinyl alcohol residue curve maps reactive distillation
Acknowledgment The authors gratefully acknowledge the support from the Major State Basic Research Development Program of China (973 Program
197
2012CB720502), National High Technology Research and Development (863 Program 2013AA040306), National Natural Science Foundation of China (NSFC21076074), Shanghai Pujiang Talents Program (10PJ1402400), and B08021.
Appendix A. Supporting information Supplementary data associated with this article can be found in the online version at http://dx.doi.org/10.1016/j.ces.2013.11.036. References Barbosa, D., Doherty, M.F., 1987. A new set of composition variables for the representation of reactive phase diagrams. In: Proceedings of the Royal Society of London A, 413, 459. Bessling, B., Schembecker, G., Simmrock, K.H., 1997. Design of processes with reactive distillation line diagrams. Ind. Eng. Chem. Res. 36, 3032. Dirk-Faitakis, C.B., An, W.Z., Lin, T.B., Chuang, K.T., 2009. Catalytic distillation for simultaneous hydrolysis of methyl acetate and etherification of methanol. Chem. Eng. Process.: Process Intensif. 48, 1080. Fuchigami, Y., 1990. Hydrolysis of methyl acetate in distillation column packed with reactive packing of ion exchange resin. J. Chem. Eng. Jpn. 23 (3), 354. gPROMS Advanced User Guide, Release 2.2, 2003. Process Systems Enterprise Ltd.: London, U.K. Hayden, J.G., O’Connell, J.P., 1975. A generalized method for predicting second virial coefficients. Ind. Eng. Chem. Process Des. Dev. 14, 209. Katariya, A., Freund, H., Sundmacher, K., 2009. Two-step reactive distillation process for cyclohexanol production from cyclohexene. Ind. Eng. Chem. Res. 48, 9534. Kaymak, D.B., Luyben, W.L., 2004. Quantitative comparison of reactive distillation with conventional multiunit reactor/column/recycle systems for different chemical equilibrium constants. Ind Eng. Chem. Res. 43 (10), 2493. Kim, K.J., Roh, H.D., 1998. Reactive distillation process and equipment for the production of acetic acid and methanol from methyl acetate hydrolysis. US Patent 5770770. Lai, I.K., Hung, S.B., Hung, W.J., Yu, C.C., Lee, M.J., Huang, H.P., 2007. Design and control of reactive distillation for ethyl and isopropyl acetates production with azeotropic feeds. Chem. Eng. Sci. 62, 878. Lee, M.M., 2002. Method and apparatus for hydrolyzing methyl acetate. US Patent 20020183549A1. Lin, Y.D., Chen, J.H., Cheng, J.K., Huang, H.P., Yu, C.C., 2008. Process alternative for methyl acetate conversion using reactive distillation. 1. Hydrolysis. Chem. Eng. Sci. 63, 1668. Nisoli, A., Malone, M.F., Doherty, M.F., 1997. Attainable regions for reaction with separation. A.I.Ch.E. J. 43 (2), 374. Pöpken, T., Götze, L., Gmehling, J., 2000. Reaction kinetics and chemical equilibrium of homogeneously and heterogeneously catalyzed acetic acid esterification with methanol and methyl acetate hydrolysis. Ind. Eng. Chem. Res. 39 (7), 2601. Song, W., Venimadhavan, G., Manning, J.M., Malone, M.F., Doherty, M.F., 1998. Measurement of residue curve maps and heterogeneous kinetics in methyl acetate synthesis. Ind. Eng. Chem. Res. 37 (5), 1917. Steyer, F., Freund, H., Sundmacher, K., 2008. A novel reactive distillation process for the indirect hydration of cyclohexene to cyclohexanol using a reactive entrainer. Ind. Eng. Chem. Res. 47, 9581. Sundmacher, K., Kienle, A. (Eds.), 2003. Wiley-VCH, Weinheim, Germany. Tang, Y.T., Chen, Y.W., Huang, H.P., Yu, C.C., Hung, S.B., Lee, M.J., 2005. Design of reactive distillations for acetic acid esterification. A.I.Ch.E. J. 51 (6), 1683. Tong, L.W., Wu, W.G., Ye, Y.M., Chen, L.F., Wozny, G., Qi, Z.W., 2013. Simulation study on a reactive distillation process of methyl acetate hydrolysis intensified by reaction of methanol dehydration. Chem. Eng. Process.: Process Intensif. 67, 111. Tung, S.T., Yu, C.C., 2007. Effects of relative volatility ranking to the design of reactive distillation columns. A.I.Ch.E. J. 53 (5), 1278. Venimadhavan, G., Buzad, G., Doherty, M.F., Malone, M.F., 1994. Effect of kinetics on residue curve maps for reactive distillation. A.I.Ch.E. J. 40, 1814. Venimadhavan, G., Malone, M.F., Doherty, M.F., 1999a. A novel distillate policy for batch reactive distillation with application to the production of butyl acetate. Ind. Eng. Chem. Res. 38 (3), 714. Venimadhavan, G., Malone, M.F., Doherty, M.F., 1999b. Bifurcation study of kinetic effects in reactive distillation. A.I.Ch.E. J. 45, 547. Wang, J., Ge, X., Wang, Z., Jin, Y., 2001. Experimental studies on the catalytic distillation for hydrolysis of methyl acetate. Chem. Eng. Technol. 24 (2), 155. Xiao, J., Liu, J., Li, J., Jiang, X., Zhang, Z., 2001. Increase MeOAc conversion in PVA production by replacing the fixed bed reactor with a catalytic distillation column. Chem. Eng. Sci. 56 (23), 6553. Zeng, H., Wu, Y.X., 2009. The whole flow simulation of methyl acetate hydrolysis system in PVA plant. Comput. Appl. Chem. 26 (10), 1338. (in Chinese).
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Chemical Engineering Science journal homepage: www.elsevier.com/locate/ces
Analysis of intensification mechanism of auxiliary reaction on reactive distillation: Methyl acetate hydrolysis process as example Liwei Tong a, Lifang Chen a, Yinmei Ye b, Zhiwen Qi a,n a Max Planck Partner Group at the State Key Laboratory of Chemical Engineering and School of Chemical Engineering, East China University of Science and Technology, Shanghai 200237, China b School of Resources and Environmental Engineering, East China University of Science and Technology, Shanghai 200237, China
H I G H L I G H T S
Concept of reactive distillation directly intensified by an auxiliary reaction is presented. RCMs is developed to analyze the intensification mechanism with two reactions and multiple components. The methyl acetate hydrolysis intensified by methanol dehydration is studied as example. A novel process of MeOAc hydrolysis is designed with 100% conversion of MeOAc. No additional H2O is needed in feed while high purity DME and HAc are products.
art ic l e i nf o
a b s t r a c t
Article history: Received 6 July 2013 Received in revised form 11 October 2013 Accepted 20 November 2013 Available online 3 December 2013
A reactive distillation (RD) process directly intensified by an auxiliary reaction is presented. The methyl acetate (MeOAc) hydrolysis intensified by methanol (MeOH) dehydration is studied as an example where the latter serves as the auxiliary reaction. A residue curve maps (RCMs) is developed to analyze the mechanism of intensification for the case with two reactions and multiple components. The analysis of thermodynamics and RCMs reveals that the reaction of MeOH dehydration is the controlling step which limits the conversion of MeOAc. Based on the chemical equilibrium and kinetically controlled design, a novel process is developed to hydrolyze MeOAc and MeOH in the process of polyvinyl alcohol. Compared to traditional processes, no additional water is required to feed into the RD column and the process is significantly simplified. With equal mole of MeOAc and MeOH as feed and a set of pre-reactor, closely 100% conversion of MeOAc and MeOH can be achieved and high purity dimethyl ether and acetic acid are the products in the RD column. & 2013 Elsevier Ltd. All rights reserved.
Keywords: Reactive distillation Process intensification Auxiliary reaction Methyl acetate Hydrolysis
1. Introduction Reactive distillation (RD) provides an attractive alternative for process intensification with advantages of simultaneous reaction and separation (Kaymak and Luyben, 2004). One of the ideal process configurations of RD consists of a column. The light and heavy reactants are fed at the lower and upper parts of the reactive zone while the heavy and light components are bottom and top products, respectively. Typical examples mainly include the productions of ethers of gasoline additives and ester such as methyl acetate, n-butyl acetate and amyl acetate (Sundmacher and Kienle, 2003).
n
Corresponding author. Fax: þ86 21 6425 3528. E-mail address: [email protected] (Z. Qi).
0009-2509/$ - see front matter & 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.ces.2013.11.036
However, the advantages of RD may be inhibited by their complicated physical properties of the reaction systems. One of the most common problems is the azeotropic mixture in the system. High conversion can be blocked when the reactants exist in the lightest azeotropic mixtures since they are likely to flee out of the column without taking reactions. Consequently, the configuration of the RD column has to be modified and extra equipment is required to recycle the reactants back to the RD column, as referred to the productions of ethyl acetate and isopropyl acetate (Lai et al., 2007). Another problem of RD is the excessive amount of reactants remaining in the system. In the case of a reversible reaction with very low chemical equilibrium constant (Keq), one of the reactants has to be excessive to improve the conversion of another reactant though RD can break through the limitation of chemical equilibrium. Further separation is thus required to deal with the excessive reactant in the downstream operation. In the case of liquid phase split, the reactants may not directly contact
L. Tong et al. / Chemical Engineering Science 106 (2014) 190–197
with the catalyst, and mass transfer resistances significantly limit the reaction rate. In this contribution, a concept of RD intensified by an auxiliary reaction is proposed to solve the above bottlenecks. By introducing an auxiliary reaction, the characteristics of original system such as thermodynamics, reaction rates and alternative products are modified, which will intensify the system and simplify the RD process. The indirect hydration of cyclohexene to cyclohexanol using a reactive entrainer is a typical example of intensification (Steyer et al., 2008; Katariya et al., 2009). As the liquid–liquid phase splitting in the system of cyclohexene/water limits the reaction rate, formic acid as the reactive entrainer is introduced into the system. Cyclohexene first reacts with formic acid forming formic acid cyclohexyl ester, and then this ester can be split with water into cyclohexanol and formic acid. The reaction route offers the advantage of overcoming reaction rate limits and having no significant amounts of side products. Another example is the process of methyl acetate (MeOAc) hydrolysis which will be investigated in this work. The effect of intensification on reactive distillation process has been presented in our previous work (Tong et al., 2013). However, the mechanism of intensification remains unclear since there is no existing suitable method to explore it. Thus, in this work, a novel residue curve maps (RCMs) tool is developed to explicitly explain the intensification mechanism. Then, the controlling step of the whole process is figured out by analyzing the thermodynamics and RCMs of the involved reaction systems. Finally, a novel process of hydrolysis of MeOAc is developed based on the equilibrium and kinetically controlled design. Compared with conventional processes, the novel process can be significantly simplified with 99.78% conversion of MeOAc. Furthermore, high purity dimethyl ether (DME) as an additional product can be obtained through methanol (MeOH) dehydration.
Fig. 1. Intensification mechanism of the auxiliary reaction.
total reflux design and removal of the light product (MeOH), heavy product (HAc), and excess reactant (H2O) from the bottoms of the column are reasonable choices (Fuchigami, 1990). This implies that high energy consumption is required at the condenser and reboiler in RD column. The second problem of MeOAc hydrolysis is related to the downstream separation process. Because of the low chemical equilibrium constant Keq, high feed ratio of H2O/MeOAc (normally 8–15) (Kim and Roh, 1998; Wang et al., 2001; Xiao et al., 2001; Lee, 2002) is employed to enhance the conversion of MeOAc as high as possible. However, the binary vapor–liquid equilibrium (VLE) diagram for the H2O-HAc system (see Fig. 1 of Lin et al., 2008) shows a tangent pinch point existing near the pure H2O vertex. This suggests that high reflux ratio or reboiler duty is required to remove HAc from H2O. Moreover, the excessive amount of H2O as feed in RD column would also cause high energy consumptions in the downstream separation process. Therefore, it is essential to develop an approach to intensify the process of MeOAc hydrolysis, namely, to obtain high MeOAc conversion with low energy requirement and “neat” operation. MeOAc þ H2 O 3 MeOH þ HAc K eq ¼
2. The process of MeOAc recovery in PVA production The PVA effluent contains a large amount of MeOAc as byproduct, around 1.68 times of the PVA product by weight (Zeng and Wu, 2009). Since the utility of MeOAc is limited, the major approach is to hydrolyze MeOAc back to HAc and MeOH (Eq. (1)), as raw materials for the PVA plant. Usually, two major problems are arisen in the process due to the system characteristics. The first is two minimum-boiling binary azeotropes formed by (1) MeOAc (44.88 mol%) and MeOH at 131.03 1C, and (2) MeOAc (74.74 mol%) and H2O at 138.05 1C. The order of boiling point temperature for pure components and azeotropes of the system at 1.0 MPa is illustrated in Table 1. Theoretically, pure light product (MeOH) can be obtained at the top of a single RD column if all MeOAc is consumed (Tung and Yu, 2007). However, MeOH is a saddle in the RCM diagram (Tang et al., 2005). With a low equilibrium constant (Keq 0.013, Eq. (2)), it is unlikely to obtain high purity product of MeOH under the “neat” operation. Thus, Table 1 Singular points in Fig. 2 (at 1.0 MPa). Name
Type
DME Unstable Azeotrope 2 Unstable/ Saddle MeOH Saddle Azeotrope 1 Saddle MeOAc Saddle H2O Saddle HAc Stable
Temp. (1C)
Composition (mole fraction) MeOH HAc DME
44.68 130.76
0 0 0.4488 0
0 0 0.5512 0
1 0
137.01 138.71 143.91 180.32 214.08
0 0.7474 1 0 0
1 0 0 0 0
0 0 0 0 0
0 0 0 0 1
½MeOH½HAc ½MeOAc½H 2 O
ð1Þ ð2Þ
To reach this goal, MeOH dehydration, as an auxiliary reaction, is introduced to intensify the MeOAc hydrolysis. As indicated in Fig. 1, the MeOH produced by the hydrolysis reaction is further dehydrated into H2O which can circularly react in the hydrolysis and thus enhance the conversion of MeOAc. In the system, H2O acts as an inner cycling component and the two reactions are “selffeeding”. As a result, no extra H2O is needed and MeOH is in-situ transferred to DME. The energy consumption of complicated process to separate the excess H2O in the downstream could be significantly saved. In the research of Dirk-Faitakis et al. (2009), it proved that these two reactions can proceed co-currently in a RD column to produce both DME and HAc. A batch catalytic distillation experiment showed that the ion exchange resin (Amberlyst 35) was active for both reactions without any initial water present in the feed. However, simulations of continuous RD were based on the assumption of both physical and chemical equilibrium, which was an extreme situation and it is usually unrealistic in practice. Thus, the kinetic models should be introduced to further investigate the intensification mechanism and a process based on kinetically controlled reaction condition will be developed.
3. Analysis of intensification mechanism
MeOAc H2O
0 0.2526 0 1 0
191
Reactive distillation integrates chemical reaction and distillation in one unit. Thus, intensification on reactive distillation process could be realized by improving either of them, which is the motivation of this work. The kinetic models used in this work are from Song et al. (1998) (see Section 3.1) where Wilson model is used for predicting the liquid activity coefficient of the VLE calculation and the model parameters are taken from literature (see Table 3 of Song et al., 1998). The vapor phase dimerization
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(for HAc) is also taken into account using the Hayden-O’Connell second virial coefficient model (Hayden and O’Connell, 1975) and the model parameters are from Aspen Plus database.
The model equations for RCMs are presented in Eq. (10) (for the detailed derivation of the model, refer to Venimadhavan et al., 1994,1999a,1999b).
3.1. Kinetic models
N dxi ¼ ð1 DÞðxi yi Þ þ D ∑ ðνi;j νT;j xi Þℜj dξ j¼1
For the reaction of MeOAc hydrolysis, the kinetic model can be referred to its reversible reaction, namely, MeOAc synthesis. Song et al. (1998) and Pöpken et al. (2000) established the kinetic model with Amberlyst 15 as catalyst, and the MeOH dehydration, as a side reaction, was also considered in the work of Song et al. (1998). Therefore, the kinetic models of Song et al. (1998) are employed in this work since the kinetic parameters for both reactions are obtained in the same experimental condition. The reaction kinetics for MeOAc hydrolysis in a Langmuir– Hinshelwood–Hougen–Watson rate model is r1 ¼
Wks;1 ðαMeOAc αH2 O ðαMeOH αHAc =ÞK eq1 Þ
ð1 þ K MeOAc αMeOAc þ K H2 O αH20 O þ K MeOH αMeOH þ K HAc αHAc Þ2
Inðks;1 =WÞ ¼ 24:64
6287:7 TðKÞ
782:98 K eq1 ¼ 1=exp 0:83983 þ TðKÞ
ð3Þ
ð4Þ ð5Þ
with K H2 O ¼ 10.50, KMeOH ¼4.95, KHAc ¼ 3.18, KMeOAc ¼ 0.82. ks,1 is the apparent forward reaction rate constant with units of (mol of mixture) (mol of H þ ions) 1 min 1, Keq1 stands for the chemical equilibrium constant for MeOAc hydrolysis, W represents the concentration of catalyst with units of (mol of H þ ions) (mol of mixture) 1, Ki is the adsorption equilibrium constant of component i. In the kinetic model for MeOH dehydration, it is assumed that DME is extremely volatile and has no significant presence on the catalyst surface, as represented in Eq. (6): r2 ¼
Wks;2 K 2MeOH ðα2MeOH ðαDME αH2 O =K eq2 ÞÞ ð1 þ K H2 O αH2 O þK MeOH αMeOH Þ2
K eq2 ¼ exp
ΔG 0
ð6Þ
!
RTðKÞ
Inðks;2 =WÞ ¼ 27:40
; ΔG0 ¼ 2:4634 1:5167 10 3 TðKÞ 10654:0 TðKÞ
ð7Þ
ð8Þ
where ΔG0 is free energy of the reaction in kcal/mol, ks,2 has the same units as ks,1, Keq2 is the chemical equilibrium constant for MeOH dehydration (Nisoli et al., 1997). The adsorption equilibrium constants are assumed to have the same values as above.
ði ¼ 1; …NC; j ¼ 1; ………NRÞ
ð10Þ
where xi and yi are the liquid and vapor phase mole fractions, respectively; dξ the dimensionless time with dξ ¼(V/H)dt, H the instantaneous liquid holdups in the still moles, V the instantaneous vapor rates, t the time, vi the stoichiometric coefficient of component i in the reaction, vT the sum of the stoichiometric coefficients for the reaction, ℜj the normalized reaction vector with ℜj ¼ ðr j =kf ;ref Þ, rj the rate of an individual reaction j and D¼
Da 1 þ Da
ð11Þ
where D is the dimensionless parameter of Da, which varies from zero to unity. In the case of D¼ 0, there is no reaction in the system so the effect of distillation can be investigated alone. In contrast, in the case of D-1.0, the system is in the limit of reaction equilibrium where the effect of reaction can be studied. As there are multiple components (i.e., MeOAc, H2O, MeOH, HAc and DME) and two reactions (MeOAc hydrolysis and MeOH dehydration) in the system, the traditional topological diagram in literature is not suitable. Therefore, in this work, a novel method of RCMs expression is developed for analysis of the intensification mechanism with two reactions and multiple components. For an equilibrium RD, the concept of reactive distillation lines (Bessling et al., 1997) on the basis of the transformation of concentration coordinates (Barbosa and Doherty, 1987) is used to study the feasibility of a reactive distillation. However, in the case of a kinetically controlled RD, the RCMs with five components are difficult for visualization in a usual tetrahedron. The color of curve is used to express additional information to solve this problem. The components of A, B, C and Dþ E are treated as four vertexes of a tetrahedron while the color of curve is represented by the mole fraction ratio of D/(D þ E). Therefore, the information of five components could be clearly read from the RCMs. It is essential to investigate which reaction rate is faster since the two reactions are simultaneously taking place and they are dependent on each other by “self-feeding”. Thus, the chemical equilibrium surface (CES) or curve is also illustrated in the RCMs. By observing the trajectories of RCMs and the CES at different values of D in Eq. (10), the reaction reaching chemical equilibrium can be confirmed. 3.3. Distillation intensification effect (D ¼ 0)
3.2. Residue curve maps with two reactions and multiple components One widely used tool for synthesis and analysis of both nonreactive and reactive distillation systems is RCMs which is the phase plane of the liquid compositions in an isobaric open evaporation. An important parameter in the study of RCMs for reactive system is the Damköhler number (Da) defined by the ratio of a characteristic residence time to a characteristic reaction time. Here, it is defined as Da ¼
H 0 =V 0 1=ðkf ;ref WÞ
ð9Þ
where H0 is the initial liquid holdup in the still in mole, V0 is the initial vapor rate from the still, kf,ref is the reference forward reaction rate constant (defined at a convenient temperature such as the lowest boiling temperature in the mixture).
In the conceptual design of RD, one of the ideal cases is that all reactants are intermediate components while the desired products should be either the lightest or the heaviest ones in the system. However, the system of MeOAc hydrolysis is not satisfied with such conditions. Fig. 2a displays the RCMs for MeOAc system without reaction (D¼0). In the RCMs, all of the trajectories start at the MeOAc/MeOH azeotrope (Az 2, the lightest boiler and unstable node), go via the intermediate boilers (saddle), and end at the pure HAc vertex (stable node). Thus, MeOH as a desired product is an intermediate component though HAc as another desired product is the heaviest one. Worse yet, part of MeOAc would flee out of the column without having reaction since it exists in Az 2. Once the MeOH dehydration is activated, the new product DME would replace Az 2 to be the lightest one in the system. The overall reaction in Fig. 1 implies that reactants are the intermediate components while products are the heaviest and lightest components,
L. Tong et al. / Chemical Engineering Science 106 (2014) 190–197
193
Fig. 2. RCMs of different systems at 1.0 MPa. (a) simple distillation for MeOAc hydrolysis system (D ¼ 0) and (b) simple distillation for simultaneous MeOAc hydrolysis and MeOH dehydration (D ¼ 0) □ saddle node 〇 unstable node ● stable node.
Fig. 3. Effect of temperature on chemical equilibrium constant Keq3 ¼ Keq1 Keq2.
respectively, which is ideal for RD. Fig. 2b shows the RCMs for the systems of MeOAc and DME without reaction (D¼0) where the color of curves represents the mole fraction ratio of H2O/(MeOHþH2O), and the (MeOHþH2O) is considered as one component. Az 1 and Az 2 at the edge of MeOAc/(MeOHþH2O) consist of MeOAc/H2O and MeOAc/ MeOH (see Table 1), respectively. Compared with Fig. 2a, all trajectories end at the same point (pure HAc) but start at DME (the lightest component). Therefore, MeOAc is forced to stay in the middle of the RD column by the effect of distillation which would promote its conversion. 3.4. Reaction intensification effect (D-1.0) The MeOAc hydrolysis is an endothermic reaction with the reaction heat ofΔH 0298 ¼ 10:87 kJ=mol while the MeOH dehydration is an exothermic reaction (ΔH 0298 ¼ 35:27 kJ=mol). Hence, the thermal coupling of the two reactions is helpful for keeping temperature stable in the reactive section. The effect of temperature on the chemical equilibrium constant is depicted in Fig. 3. As seen, the chemical equilibrium constant of the overall reaction Keq3 (Eq. (12)) gradually declines by rising the temperature implying that low temperature is favorable for the overall reaction. However, in the work of Song et al. (1998), it was found that the MeOH dehydration was not significant at low temperature (40–50 1C). Nevertheless, at higher temperature, i.e., 85 115 1C, significant amount of DME were formed. As the MeOH dehydration should be activated on purpose in this work, the reaction temperature over 85 1C is taken into account. K eq3 ¼
½DME½HAc ¼ K eq1 K eq2 ½MeOAc½MeOH
ð12Þ
Fig. 4. Intensification of reaction by introduction of MeOH dehydration at 120 1C (D-1.0). single MeOAc hydrolysis ( □ dashed line); simultaneous MeOAc hydrolysis and MeOH dehydration ( 〇 straight line).
Fig. 4 shows the difference of RCMs between the case of single MeOAc hydrolysis and the case of simultaneous MeOAc hydrolysis and MeOH dehydration at different initial points (i.e., H2O/ MeOAc). For the case of single reaction, as the RCMs is at chemical equilibrium condition (D-1.0), which is also the limit case, the trajectories (dashed line) swiftly close to the CES and intersect at the points marked by □, go along with the CES and end at the pure HAc vertex. In the case of two reactions, the xi (components of MeOAc, H2O, MeOH and HAc) calculated from Eq. (10) are reunified by summation of compositions of these four compounds so that the trajectories can track inside of the tetrahedron. The trajectories (solid line) also approach to the CES and intersect at the points marked by 〇. However, the locations of intersection are different from those in the case of single reaction. For instance, at the initial point with 0.5 MeOAc and 0.5H2O, P1 and P2 are the intersections of trajectories with the CES for two reactions (solid line) and single reaction (dashed line), respectively. It is found that P1 is more far away from the edge of MeOAc-H2O (i.e., closer to edge of MeOH-HAc). This implies that the auxiliary reaction (MeOH dehydration) makes the chemical equilibrium of MeOAc hydrolysis move to the forward direction since MeOH is converted into H2O by MeOH dehydration. 3.5. Controlling step The simultaneous MeOAc hydrolysis and MeOH dehydration are consecutive and inner-cycled reactions as indicated in Fig. 1.
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Thus, it is important to figure out which reaction is the controlling step. The components xi (i.e., MeOAc, H2O, MeOH, HAc and DME) in Eq. (10) calculated at simultaneous reactions is used to acquire the RCMs in the system of MeOAc hydrolysis and MeOH dehydration, respectively. For the system of MeOAc hydrolysis, the components xi (i.e., MeOAc, H2O, MeOH and HAc) calculated by Eq. (10) are reunified as in Fig. 4. Fig. 5 shows the RCMs at different values of D. From D¼ 0.1 to D ¼0.7, as seen, all trajectories go along with the CES, which means the rate of MeOAc hydrolysis is fast enough to reach its chemical equilibrium even at the low value of D. In the case of a single MeOH dehydration, the components xi (i.e., H2O, MeOH and DME) calculated by Eq. (10) are reunified as in Fig. 4 to ensure the trajectories go inside of the triangle. The RCMs at different values of D is illustrated in Fig. 6. It is found that all trajectories stay far away from the chemical equilibrium curve at the low D. Only when D is relatively high (i.e., D¼ 0.7), the trajectory would approach the chemical equilibrium curve. This is attributed to the low reaction rate of MeOH dehydration at the low D. As a sequence, the two reaction rates are not synchronous, and
MeOAc hydrolysis reaches its chemical equilibrium in advance. The conversion of MeOAc mainly depends on the reaction rate of MeOH dehydration because it is the only H2O supplier in the system. A high reaction rate of MeOH dehydration could produce excess H2O that is equivalent to giving a high feed ratio of H2O/MeOAc as feed. In contrast, the whole process will be limited by the low rate of MeOH dehydration.
4. Novel processes of MeOAc hydrolysis As MeOH dehydration can supply H2O for the system, it is unnecessary to add additional H2O into the RD column. Hence, a mixed MeOAc and MeOH feed is an alternative for MeOAc hydrolysis in RD process. As mentioned before, in the research of Dirk-Faitakis et al. (2009), the RD process was studied assuming that both physical and chemical equilibrium were reached. However, the equilibrium design is a critical case for the conceptual design of RD. Therefore, in the following section, both equilibrium and kinetically controlled RD are taken into account. 4.1. Equilibrium design In the conceptual design of a RD process, the first step in the procedure is to develop a spectrum of designs at both phase and reaction equilibria. The concept of reactive distillation line proposed by Bessling et al. (1997) is usually used to study the feasibility of RD process. Fig. 7 depicts the reactive distillation line for the system MeOAc, H2O, MeOH, HAc and DME where HAc and MeOH are the reference components, giving X3 ¼xMeOAc þ xHAc and X4 ¼xH2O þ (1/2)xHAc þ(1/2)xMeOH. The MeOAc hydrolysis and MeOH dehydration simultaneously react in the system. As shown in Fig. 7, the system has no reactive distillation borders and all reactive distillation lines start from the HAc vertex and end at the DME vertex. Thus, the process of RD is feasible and probably an economic operation since both products are connected by a reactive distillation line, the products are (unstable and stable) nodes in a reactive distillation line diagram, and the boiling point difference between the products is large. The feed ratio of MeOAc/ MeOH is equal to one so that pure DME and HAc could be the top and bottom product, respectively. This is because, in the overall reaction, DME and HAc as products are the lightest and heaviest component, while MeOAc and MeOH as reactants are both the
Fig. 5. RCMs in chemical equilibrium surface of the MeOAc hydrolysis system at 120 1C, for different values of D.
Fig. 6. RCMs in chemical equilibrium curve of the MeOH dehydration system at 120 1C, for different values of D.
Fig. 7. Reactive distillation line for MeOAc/DME system at 0.5 MPa. Concentration profile of a RD column with 30 stages. The input data for the simulation are molar ratio MeOAc/MeOH¼ 1 as feed; reflux ratio¼1.5; feed stage¼15; column pressure¼ 0.5 MPa (The marks are the same as in Fig. 2).
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the equilibrium design can be modified to achieve the desired products in a kinetically controlled column. An equilibrium stage model of RD was applied and implemented in gPROMS for kinetically controlled design and simulation. The first column simulation is performed at the reaction equilibrium conditions. Further simulations are performed at successively lower Da until a realistic value is achieved. Here, Da is defined as Da ¼
Fig. 8. Equilibrium design at 0.5 MPa for MeOA hydrolysis by feed of MeOAc and MeOH.
intermediate components. As a result, MeOAc and MeOH will be kept inside of the RD column all the time until getting 100% consumed if they are the equal mole in the feed, which is proved by the concentration profile line in Fig. 7. The final equilibrium design for MeOAc hydrolysis is presented in Fig. 8. The mole fraction of DME and HAc are over 0.99 in the distillate and bottom, respectively. The operation pressure in the RD column is set to 0.5 MPa so as to ensure the reaction temperature over 85 1C (as mentioned in Section 3.3). There are 30 stages in total, each with an equilibrium chemical reaction (the stage of condenser and reboiler has no reaction, so the effective number of chemical equilibrium stage is 28). The lowest reflux ratio corresponding to 30 stages is 1.5, which is chosen as the base-case design value. The designed configurations are similar to the results in Dirk-Faitakis et al. (2009) (see Table S1 in the supporting information).
W T H RT kf ;ref FT
ð13Þ
where FT is the total feed flow rate to the column; WT is the total concentration of catalyst with units of (mol of H þ ions) (mol of mixture) 1; H RT is the total holdup in the reactive zone. The comparison between the reaction equilibrium design and a kinetic simulation (Da¼100) is shown in Fig. 9. It shows that the performance of the kinetic condition approaches the reaction equilibrium design (at very high Da of 1000). For the kinetic simulation, the conversion of MeOAc is 87.56% since the amount of catalyst loading is not enough (corresponding to Da¼100). However, the equilibrium design gives a better performance of RD column. High purity DME and HAc as the products at the top and bottom of RD column are obtained while the high concentration of MeOAc is forced to stay in the reactive zone. The concentration of H2O in the reactive zone is kept at a low level since H2O is supplied by the auxiliary dehydration of MeOH and simultaneously consumed by MeOAc hydrolysis. Fig. 9 also implies that the reaction zone could be in the middle of column (e.g. 5–25 stage) since it performs essentially no reaction outside of this range. As defined in Eq. (13), high Da means a high catalyst loading for a specified FT, which may not be proper for realistic design of RD column. In order to avoid a high catalyst loading, a high reflux ratio may be employed. Thus, the effect of reflux ratio on the conversion of MeOAc will be studied, as reported in Fig. 10. For these calculations, the distillate flow rate remains a constant of 1.25 mol/s (this implies bottom rate is also a constant) and the reaction zone is stage 5–25. As shown in Fig. 10, higher conversions can be obtained at higher Da (i.e., higher catalyst loading) at the same reflux ratio. However, at Da4 100, the gain in conversion is very weak compared to the required increase in Da. Therefore,
4.2. Kinetically controlled design As the RD design based on the chemical equilibrium condition very often makes it practically unrealistic, it is necessary to conduct an analysis of influence of kinetic effect through, e.g. bifurcation analysis of the steady-state solutions of Eq. (10) with D as parameter. The advanced process simulator gPROMS is applied in this work. It can export the mathematical information of models such as Jacobian matrix by Foreign Process Interface (F.P.I) (gPROMS, 2003), which makes it available to compile a continuation algorithm such as arc-length continuation to trace any bifurcated branches. Meanwhile, the stability of the solutions could be verified by analysis of Jacobian matrix. Thus, an arclength continuation based on gPROMS platform is established to calculate all branches. The starting points for the analysis are the singular points for the nonreactive mixture, which are the pure components and the nonreactive azeotropes at D ¼0. It was found that the saddle node of azeotropes and MeOH would cease to be solutions as soon as the reactions were initiated while other fixed points remained the same through the entire kinetic regime up to the reaction equilibrium limit (see Fig. S1 in the supporting information). Therefore, there are no any kinetic and reactive azeotropes in the reacting system. The desired separation (as described in the previous section) that was found to be feasible at reaction equilibrium is also feasible for kinetic condition, and
Fig. 9. Comparison of RD column profiles between the simulation (Da ¼ 100) and equilibrium design.
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Table 2 Comparison of simulation results for single RD column and plus pre-reactor.
Fig. 10. Influence of reflux ratio on the conversion of MeOAc and Vcat/Vstage for different values of Da. (a) design parameters (b) reflux ratio vs. conversion and Vcat/Vstage.
further improving the performance of the column can be realized by increasing the reflux ratio. Fig. 10 also shows the influence of reflux ratio on Vcat/Vstage for different Da. Here, Vcat is the volume of catalyst per stage; Vstage is the volume of stage and V stage ¼
L HETP SD
ð14Þ
where L is the liquid flow rate of stage in the column; SD is the spraying density of stage in the column (set to 10 m3/m2/h); HETP is the height equivalent of theoretical stage. In a realistic design of RD column with ion exchange resin as catalyst, the CDTech bales are widely used packing. The HETP is set to 0.5 m, and the value of Vcat/Vstage has to be less than 0.25. Therefore, the catalyst loading of 6.9 kg/stage (Da¼100) and the reflux ratio of 4.2 are recommended to achieve 100% conversion of MeOAc. It is noticeable that the reflux ratio of 4.2 would lead to high energy consumptions in the reactive distillation process. This is because the MeOH dehydration as the controlling step requires a long residence time (i.e., high reflux ratio) to get 100% conversion of MeOAc. In order to reduce the reflux ratio and save energy of the reactive distillation column, a pre-reactor could be introduced, and the feed (MeOAc:MeOH¼1:1) is designed for the RD column. Table 2 compares the performance of RD column between single RD column and plus pre-reactor where the overall reaction got
Operation parameters
Single RD column
Plus pre-reactor
Total stages Rectifying zone Reactive zone Stripping zone Column pressure (Mpa) Reflux ratio (mol) Damköhler number (Da) catalyst loading (kg/stage) Feed stage Feed flowrate (mol/s)
30 1–4 5–25 26–30 0.5 4.2 100 6.90 15 2.5
30 1–4 5–25 26–30 0.5 2.1 100 6.90 15 2.5
xF (mole fraction of feed) MeOAc H2O MeOH HAc DME Top product flow rate (mol/s)
0.5 0.0 0.5 0.0 0.0 1.25
0.3595 0.1679 0.0237 0.1405 0.3084 1.25
xD (mole fraction of top product) MeOAc H2O MeOH HAc DME Bottom product flow rate (mol/s)
0.0044 o1.0 10 5 0.0080 0.0009 0.9946 1.25
0.0031 o 1.0 10 5 0.0025 o 1.0 10 5 0.9943 1.25
xB (mole fraction of bottom product) MeOAc H2O MeOH HAc DME MeOAc Conversion (%) MeOH Conversion (%) Condenser heat duty (kW) Reboiler heat duty (kW)
0.0052 0.0044 o1.0 10 5 0.9904 o1.0 10 5 99.04 99.90 124.58 168.99
0.0057 0.00309 o 1.0 10 5 0.9912 o 1.0 10 5 99.78 99.08 74.26 114.85
equilibrium in the pre-reactor. In the pre-reactor, 28.12% conversion of MeOAc can be achieved. Accordingly, the heat duty of the condenser and the reboiler of the RD column are saved by 40.56% and 32.04%, respectively. This is because MeOH has been partly converted into H2O in the pre-reactor, which is equivalent to a promoted reaction rate in a single RD column. Compared with traditional process (see Fig. 1 of Tong et al., 2013), the downstream separation are totally saved while high purity of DME (4 0.99) and HAc (4 0.99) are produced in the RD column.
5. Conclusions The concept of reactive distillation directly intensified by an auxiliary reaction is presented. A novel RCMs method is developed for the case with two reactions and multiple components, which is applied to analyze the intensification mechanism of the auxiliary reaction. As an example, the MeOAc hydrolysis intensified by the reaction of MeOH dehydration in the RD process is analyzed. Based on the analysis of thermodynamics and RCMs of the system, the reaction rate of MeOH dehydration is determined as the controlling step to block 100% conversion of MeOAc. With equal mole of MeOAc and MeOH as feed, a novel process has been designed to hydrolyze MeOAc in the PVA production. As the limit of MeOH dehydration, it is not economical to give 100% conversion of MeOAc and MeOH in a single RD column. Thus, a pre-reactor is set to compensate for the low rate of MeOH dehydration. Compared with traditional processes, two distillation columns in the downstream separation are saved while high purity of DME and HAc are the products in the RD column.
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Notation ai dj D Da FT ΔG0 H HETP H0 H RT ks,1 ks,2 kf,ref Keq1 Keq2 Ki L NC NR rj R ℜ SD T V0 V Vcat Vstage W WT xi Xi yi
activity of component i index of rate amplification for reaction j dimensionless parameter of Da Damköhler number total feed flow rate to RD column, mol/s free energy of the reaction, kcal/mol molar liquid holdup, mol height equivalent of theoretical stage, m initial molar liquid holdup, mol total molar liquid holdup in reactive zone, mol forward reaction rate constant of MeOAc hydrolysis, (mol of mixture) (mol H þ ions) 1 min 1 forward reaction rate constant of MeOH dehydration, (mol of mixture) (mol H þ ions) 1 min 1 forward reaction rate constant at reference temperature, (mol of mixture) (mol H þ ions) 1 min 1 chemical equilibrium constant of MeOAc hydrolysis chemical equilibrium constant of MeOH dehydration adsorption equilibrium constant for component i liquid flowrate of stage in RD column, m3/h number of components number of reactions reaction rate for reaction j, mol/s universal gas constant (1.987 cal/mol K) normalized reaction rate, dimensionless spraying density of stage in RD column, m3/m2/h reaction temperature, K initial vapor molar flow rate, mol/s vapor molar flow rate, mol/s volume of catalyst loading per stage, m3 volume of stage in the RD column, m3 concentration of catalyst, mol H þ ions/mol of mixture total concentration of catalyst in RD column, mol H þ ions/mol of mixture liquid phase mole fraction of component i transformed liquid phase mole fraction of component i vapor phase mole fraction of component i
Greek letters
νi νT ξ
stoichiometric coefficient of component i total mole change of reaction dimensionless time, dξ ¼ (V)/(H)dt
Abbreviations CES DME HAc MeOAc MeOH PVA RCMs RD
chemical equilibrium surface dimethyl ether acetic acid methyl acetate methanol polyvinyl alcohol residue curve maps reactive distillation
Acknowledgment The authors gratefully acknowledge the support from the Major State Basic Research Development Program of China (973 Program
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2012CB720502), National High Technology Research and Development (863 Program 2013AA040306), National Natural Science Foundation of China (NSFC21076074), Shanghai Pujiang Talents Program (10PJ1402400), and B08021.
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