Improving the Performance of Extractive Dividing-Wall Columns with

Article pubs.acs.org/IECR

Improving the Performance of Extractive Dividing-Wall Columns with Intermediate Heating Jieping Yu,† San-Jang Wang,‡ Kejin Huang,*,† Yang Yuan,† Haisheng Chen,† and Li Shi† †

College of Information Science and Technology, Beijing University of Chemical Technology, Beijing 100029, P. R. China Center for Energy and Environmental Research, National Tsing Hua University, Hsinchu 300, Taiwan

‡

ABSTRACT: In spite of the fact that using extractive dividing-wall columns (EDWCs) lead to a considerably reduced reboiler heat duty in comparison with conventional extractive distillation column (CEDC) flowsheets, the involvement of a heavy entrainer frequently requires a relatively high-pressure steam for process heating and incurs consequently adverse steady-state economics. To address the potential deficiency, we proposed, in the present study, to facilitate the EDWC with intermediate heating because the great boiling-point difference between the components in the azeotropic mixture and entrainer results in steep temperature profiles and allows effectively recovering sensible heat from the recycled entrainer and using relatively lowpressure steam. When the extractive separation operation dominates the EDWC, intermediate heating should be arranged in the left side of the dividing wall and feed preheating is frequently the most favorable option. On the other hand, when the entrainer recovery operation dominates the EDWC, intermediate heating should be arranged below the lower end of the dividing wall, i.e., in the stripping section, and an intermediate reboiler is the favorable option. Under both circumstances, the heat recovery from the recycled entrainer should be considered prior to the use of available utilities, thereby permitting great improvement in the steady-state economics. The strategy features simplicity in principle and requires a relatively small number of trial and error searches. It is evaluated in terms of extractive separations of two binary azeotropic mixtures: dimethyl carbonate and methanol (with aniline as the entrainer) and acetone and methanol (with dimethyl sulfoxide as the entrainer). It is found that intermediate heating could substantially enhance the performance of the EDWC with the resultant steady-state economics overwhelmingly above that of the CEDC flowsheet. Even compared with the CEDC flowsheet reinforced with intermediate heating, the EDWC is still likely to yield comparable steady-state economics. These outcomes indicate that intermediate heating should be taken into account in the synthesis and design of the EDWC and the EDWC with intermediate heating should be regarded as a potential option for the extractive separations of binary azeotropic mixtures. low-grade utility as in the CEDC flowsheet, and this represents definitely a kind of potential drawback of process intensification, posing essentially a severe limitation to process applicability and flexibility. Recently, the EDWC has received increasing attention in the separation of binary azeotropic mixtures.7−15 Chien and his coworkers assessed the performance of the EDWC in terms of the separations of three binary azeotropic mixtures, including isopropyl alcohol and water (with dimethyl sulfoxide (DMSO) as the entrainer), dimethyl carbonate and methanol (with aniline as the entrainer), and acetone and methanol (with water or DMSO as the entrainer).16 Although the EDWC led to reduced total reboiler heat duty for all the three systems studied in comparison with the corresponding CEDC flowsheets, only did the one separating the mixture of acetone and methanol secure a reduction in the total reboiler heat duty and total annual cost (TAC) simultaneously (because low-pressure steam could be used as the hot utility for both the CEDC flowsheet and EDWC in this system). The increases in the TACs in the other two systems revealed the potentially adverse effect of mass and thermal coupling between the extractive

1. INTRODUCTION As one of the most promising technologies for ternary mixture separations, dividing-wall columns (DWCs) can be much more thermodynamically efficient and cost-effective than their conventional alternatives (e.g., the frequently adopted direct and indirect separation sequences).1−6 However, when the DWC is employed to the extractive separation of binary azeotropic mixtures (the process is termed the EDWC, hereafter, in the current work), in spite of the fact that reboiler heat duty can still be considerably reduced in comparison with that of conventional extractive distillation column (CEDC) flowsheets, the situation is actually twofold: (i) the EDWC is still more thermodynamically efficient and cost-effective than the CEDC flowsheet, and (ii) the EDWC is more thermodynamically efficient but less cost-effective than the CEDC flowsheet. Although many design and operating variables affect the steady-state performance of the EDWC, the occurrences of the second situation are aroused mainly by the necessity of exclusive utilizations of relatively high-grade utilities for reboiler heating (In contrast, in the CEDC flowsheet as shown in Figure 1a although the entrainer recovery distillation column requires relatively high-grade utilities, the extractive distillation column uses frequently lowgrade utilities). It is the mass and thermal coupling between the extractive separation operation and entrainer recovery operation involved in the EDWC that precludes the use of relatively © 2015 American Chemical Society

Received: Revised: Accepted: Published: 2709

August 6, 2014 December 9, 2014 February 24, 2015 February 24, 2015 DOI: 10.1021/ie503148t Ind. Eng. Chem. Res. 2015, 54, 2709−2723

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Industrial & Engineering Chemistry Research

EDWC, although favored with relatively high thermodynamic efficiency, appears frequently to be less cost-effective than the corresponding CEDC flowsheet. To overcome the drawback, one has therefore to find a way to use effectively low-grade utility in the EDWC as in the CEDC flowsheet. In the current work, we attempt to alleviate or even remove completely the drawback of the EDWC indicated above and our major objective is to extend its applicability and flexibility as per its operation characteristics. In the next section, a systematic strategy is proposed to enhance the steady-state economics of the EDWC with intermediate heating, followed by the evaluation studies in terms of the extractive separations of two binary azeotropic mixtures of dimethyl carbonate and methanol and acetone and methanol. In-depth comparisons are conducted between the EDWC and CEDC flowsheet with and without intermediate heating and the salient features of intermediate heating are analyzed. Some concluding remarks are finally given in the last section.

2. IMPROVING THE STEADY-STATE PERFORMANCE OF THE EDWC WITH INTERMEDIATE HEATING 2.1. Operating Characteristics of the EDWC. In the CEDC flowsheet as shown in Figure 1a, two distillation columns are included. One is an extractive distillation column (EDC), separating the light component from the top and the heavy component compounded with the entrainer from the bottom. The other is an entrainer recovery column (ERC), separating the heavy component from the top and the entrainer from the bottom. The entrainer recovered is then recycled to the EDC. Since the CEDC flowsheet deals actually with the separations of three components and includes two distillation columns to accomplish a given task, the DWC technique is then applied to introduce mass and thermal coupling between the two distillation columns involved and the attempt gives rise to the configuration of the EDWC as shown in Figure 1b. It is noted that the dividing wall runs from the top to a certain stage in the middle and generates two rectifying sections for purifying, respectively, the two components involved in the binary azeotropic mixture separated. Two condensers are needed at the top, serving to provide liquid flows for the two sections along the dividing wall. In terms of the vapor split ratio, defined by βV = V1/(V1 + V2), the operation of the EDWC can be categorized into three typical modes: (i) When βV > βU, the EDWC is dominated actually by the extractive separation operation in the left side of the dividing wall because most of utility is consumed in the purification of the light component. (ii) When βV < βL, the EDWC is dominated actually by the entrainer recovery operation in the right side of the dividing wall because most of utility is consumed in the purification of the heavy component. (iii) When βL < βV < βU, the EDWC works in a rather balanced mode between i and ii. Here, βU and βL are two characteristic parameters for the EDWC. Usually, setting βU = 0.8 and βL = 0.2 can yield reasonable results for process design and analysis. Since the vapor split ratio βV holds approximately an inequality relationship with feed composition as shown in eq 1, feed composition can sometimes be used as a rough index to determine the operation mode of the EDWC, especially at the early stage of process synthesis and design.

Figure 1. CEDC flowsheet versus EDWC: (a) CEDC flowsheet and (b) EDWC.

separation operation and entrainer recovery operation involved. Lately, Sun et al. studied the separation of an azeotropic mixture of benzene and cyclohexane with furfural as the entrainer.17 They found that the EDWC was not only more thermodynamically efficient but also more cost-effective than the corresponding CEDC flowsheet. Due to the exclusive use of medium-pressure steam in the EDWC (while the corresponding CEDC flowsheet used low-and medium-pressure steams in the extractive and entrainer recovery distillation columns, respectively), the degree of the reduction in the TAC was substantially smaller than that of the reduction in total reboiler heat duty, implying again the adverse effect of mass and thermal coupling between the extractive separation operation and entrainer recovery operation involved. Similar tendencies were also reported by some other researchers.10,18,19 Since an entrainer is chosen based on the thermodynamic properties of the azeotropic mixture separated (e.g., the elevation of the relative volatility between the two components contained and easy separation from them), its inclusion is quite likely to make infeasible the use of the relatively low-grade utility in the EDWC as in the CEDC flowsheet. This explains why the 2710

DOI: 10.1021/ie503148t Ind. Eng. Chem. Res. 2015, 54, 2709−2723

Article

Industrial & Engineering Chemistry Research βV = V1/(V1 + V2) ≈ D1(1 + R1)/(D1(1 + R1) + D2(1 + R 2)) (if constant vapor flow rates are applicable, then V1 ≈ D1(1 + R1) and V2 ≈ D2(1 + R 2)) > D1/⎛⎝D1 + D2⎞⎠ (since R1 > R 2 because the extractive separation operation is generally more difficult than the entrainer recovery operation under the framework of the EDWC) = Z1 (1)

Since the entrainer should be easily separated from the two components of the binary azeotropic mixture processed, it generally exhibits a much higher boiling-point than those of the two components and this characteristic leads to the formation of two steep temperature profiles in the extractive separation operation and entrainer recovery operation involved in the EDWC. These steep temperature profiles allow using effectively intermediate heating with the available low-grade steam as the hot utility, thereby being likely to reduce the operating cost of the EDWC. Moreover, the sensible heat of the entrainer recycled from the bottom of the EDWC should also be used as a heat source for intermediate heating and this represents essentially the major thrust for the enhancement of steady-state economics. Many studies already indicated the efficacy of this kind of heat integration in the various extractive separation operations.20,21 2.2. Improving the Performance of the EDWC with Intermediate Heating. As per the thermodynamic interpretation of distillation operation, intermediate heating should be arranged somewhere in the stripping section.22−25 For the EDWC, there exist actually two stripping sections, respectively, in the extractive separation operation and entrainer recovery operation involved although they share quite a large common section at the bottom. While the former has its stripping section from the bottom to the feed stage, the latter from the bottom to the lower end of the dividing wall. It is straightforward to understand that intermediate heating should be arranged to favor simultaneously both the extractive separation operation and entrainer recovery operation involved. In order to maximize the effect of intermediate heating (as a result, the consumption of relatively high-grade utilities can be reduced in the reboiler), one needs to determine with great caution the location of intermediate reboiler and this can be facilitated as per the operation characteristics of the EDWC outlined in the preceding subsection. For the EDWC operated in the first mode (i.e., βV > βU), it is preferable to arrange the intermediate reboiler near the feed (in the left side of the dividing wall) and this can alleviate the negative effect of a potential feed pinchpoint. In this situation, the intermediate reboiler needs not to affect directly the entrainer recovery operation (in the right side of the dividing wall) because only a small amount of utility is consumed there (note that the vapor split ratio can be adjusted to play a role in this aspect). The extreme situation happens when the intermediate reboiler is installed on the feed pipeline and the intermediate reboiler functions actually as a feed preheater. Figure 2a sketches such a configuration for the EDWC with a feed preheater (EDWC−FPH). Here, the two feed preheaters are indicated with bold lines with the recycled entrainer and low-pressure steam as hot utilities, respectively. The EDWC−FPH is characterized by the greatest temperature

Figure 2. Configurations for the EDWC with intermediate heating: (a) EDWC−FPH and (b) EDWC−IR.

driving forces between feed and the entrainer recovered from the bottom and between feed and the available low-grade utility. Because of the heat exchange with the feed (whose temperature is usually low), the recovered entrainer can, in general, be directly recycled into the top of the left side of the dividing wall, and the employment of a cooler to cool it is, in most cases, no longer necessary (as opposed to the CEDC flowsheet and EDWC shown in Figures 1a and b, respectively). It is noted that the feed is now divided into two portions (i.e., a vapor and a liquid ones) and fed onto different stages of the left side of the dividing wall. This delicate arrangement not only enhances the vapor−liquid phase mass transfer in the extractive separation operation but also entails an additional degree of freedom for process synthesis and design.26 For the EDWC operated in the rest modes (i.e., βL < βV < βU and βV < βL), it is preferable to arrange the intermediate reboiler near the lower end of the dividing wall. Figure 2b sketches such a configuration for the EDWC with an intermediate reboiler (EDWC−IR). Here, the two intermediate reboilers are indicated again with bold lines with the recycled entrainer and low-pressure steam as hot utilities, respectively. The intermediate reboiler can affect both the extractive separation 2711


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Figure 3. Systematic strategy for the development of the EDWC with intermediate heating.

operation and entrainer recovery operation involved with the vapor split ratio βV as an effective adjusting parameter in process synthesis and design. In comparison with the EDWC− FPH, its temperature driving forces could be severely confined between the liquid withdrawn from the common stripping

section (i.e., the section below the dividing wall) and the entrainer recovered from the bottom and between the liquid withdrawn from the common stripping section (i.e., the section below the dividing wall) and the available low-grade utility. Hence, it is sometimes beneficial to arrange further a feed 2712


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Industrial & Engineering Chemistry Research preheater in the EDWC−IR especially when βV or feed composition is around 0.5, because in these situations the great temperature driving force between feed and the entrainer from the intermediate reboiler can still justify the recovery of sensible heat. 2.3. Systematic Strategy for Developing the EDWC with Intermediate Heating. The unique operating characteristics of the EDWC outlined above remind us of the possibility that the process could be reinforced with intermediate heating no matter whether it is cost-effective or not as compared with the corresponding CEDC flowsheet. To enhance its steadystate economics, the designer should make full use of the sensible heat of the entrainer recovered from the bottom prior to using the available low-grade utility. A systematic strategy is thus derived for the synthesis and design of the EDWC−FPH/ EDWC−IR as shown in Figure 3. As can be seen, it consists generally of three main steps (indicated here by three curves with end arrows). In the first step, the discrimination is made between the EDWC−FPH and EDWC−IR and this can be conducted as per eq 1 with reference to the given operating conditions and product specifications. It should be indicated here that the early determination of process configuration favors greatly the search of optimum process design, leading to not only an easy-to-understand procedure but also a dramatic reduction in problem complexity and computational requirement in trial and error searches. In the second step, an initial design of the EDWC−FPH or EDWC−IR is generated and serves as a starting-point for process optimization. No special requirements are posed actually on this process design as long as it meets the given specifications on the two top and one bottom products through the adjustments of two reflux flow rates and reboiler heat duty. It helps to confine all the structural and operating variables in their feasible regions and thus improves the trial and error search efficiency. In the third step, the adjustment of the initial design of the EDWC−FPH or EDWC−IR is performed with the aim to minimize the TAC. Since many structural design variables are involved and they are in nature discontinuous, the employment of a multivariable search method will no doubt lead to a complicated procedure for process synthesis and design. To avoid the difficulty, in the current work we adopt instead the commonly use grid-search method. Despite the strategy appears to be quite cumbersome, it can locate the optimum solution fairly easily and quickly because of the simplicity of single variable based search method and the heuristics gained in the previous rounds of iterations. In the present study, the TAC, a summation of operating cost and annualized capital investment by a payback period of 3 years, is taken as the objective function for process development. While the operating cost includes steam cost, cooling water cost, and entrainer makeup cost, the annualized capital cost consists of column shells cost, column stages cost, intermediate and bottom reboiler cost, top condenser cost, and cooler cost. The prices of steam and cooling water are taken from the work fo Seider et al.,27 and the formulas for cost estimations are adopted from the work of Douglas (c.f., Appendix A).28 In particular, the dividing wall cost is assumed to be negligible as compared with the other equipment costs, and the EDWC shell cost is calculated with the method by Chien and his co-workers.16 The M&S index is assigned a value of 1536.5 in the current work.

3. EXAMPLE I: SEPARATING A BINARY AZEOTROPIC MIXTURE OF DIMETHYL CARBONATE (DMC) AND METHANOL (MEOH) WITH ANILINE AS THE ENTRAINER 3.1. Problem Description. As shown in Figure 4a, DMC and MEOH form an azeotrope at 13.52 mol % DMC under the atmospheric pressure. Figure 4 parts b and c describe, respectively, the VLE behaviors of MEOH/ANILINE and

Figure 4. T−x−y diagrams for binary mixtures in example I at 1 atm: (a) DMC/MEOH, (b) MEOH/aniline, (c) DMC/aniline. 2713


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Industrial & Engineering Chemistry Research DMC/ANILINE, and they show clearly the inclusion of ANILINE could favor the separation of DMC and MEOH. The operating conditions and design specifications are the same as in the work of Chien and his co-workers and reproduced here in Table 1 for quick reference.16 The normal boiling points of Table 1. Operating Conditions and Design Specifications of Example I parameter condenser pressure (atm) stage pressure drop (atm) feed flow rate (kmol/h) feed temperature (°C) feed composition (mole fraction) product specification (mole fraction)

value

DMC MEOH DMC MEOH aniline

1 0.0068 170.50 63.73 0.1467 0.8533 0.998 0.9999 0.99999

DMC, MEOH, and aniline are 90.22, 64.53, and 183.88 °C, respectively. One can readily note that a great temperature difference exists between the boiling points of the two components in the azeotropic mixture and entrainer, implying that intermediate heating can be an effective means to improve the steady-state economics of the EDWC. The commercial software ASPEN PLUS is employed to predict system performance and the UNIQ-RK thermodynamic model to represent vapor−liquid equilibrium relationship. All the relevant parameters are taken from the work of Hsu et al.29 3.2. CEDC Flowsheet. Hsu et al. once studied the extractive separation of a binary azeotropic mixture of DMC and MEOH via a CEDC flowsheet with aniline as the entrainer.29 With the same steady-state operating conditions and product specifications, we reproduce their optimum process design here as shown in Figure 5a. The reboiler heat duties are 3165.13 and 949.43 kW, respectively, for the extractive distillation column (CEDC−EDC) and entrainer recovery column (CEDC−ERC), resulting in a sum of 4114.56 kW. The temperature profiles of the CEDC−EDC and CEDC−ERC are shown in Figure 6a and b, respectively. While the bottom temperature of the CEDC−EDC is 163.79 °C, the bottom temperature of the CEDC−ERC rises to 188.06 °C. For the maintenance of a minimum temperature driving force (i.e., ≥10 °C) in the two reboilers, medium-pressure steam ($4.8/1000 lb) and high-pressure steam ($6.6/1000 lb) must be employed, respectively, as hot utilities in the CEDC− EDC and CEDC−ERC. 3.3. EDWC with and without Intermediate Heating. The initial design of the EDWC can simply be constructed in terms of the thermodynamic equivalent structure of the CEDC flowsheet obtained above. Namely, while the bottom section below the lower end of the dividing wall is assumed to have the same number of stages as in the stripping section of the CEDC−ERC, the two rectifying sections share the same number of stages as in the rectifying section of the CEDC− EDC. Optimization of this initial process design is then conducted, and this leads to the optimum EDWC as shown in Figure 5b. The reboiler heat duty is 3538.54 kW, which is and much lower than the total reboiler heat duty of the CEDC flowsheet, implying a great enhancement in thermodynamic efficiency by mass and thermal coupling between the CEDC− EDC and CEDC−ERC. The resultant temperature profiles of

Figure 5. Optimum designs for example I: (a) CEDC flowsheet, (b) EDWC, (c) EDWC−FPH.

the extractive separation operation (EDWC−EDC) and entrainer recovery operation (EDWC−ERC) involved are also shown in Figure 6a and b, respectively. Since the bottom temperature of the EDWC rises now to 193.40 °C, it only 2714


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for the former, and ΔT1, EDWC = 193.40 − 78.34 = 115.06 °C and ΔT2, EDWC = 193.40 − 115.15 = 78.25 °C for the latter. These characteristics imply the possibilities of improving their performance with intermediate heating. Since the vapor split ratio, βV = 0.9069 > Z1 = 0.8533 > 0.8, it is reasonable to conclude that the extractive separation operation dominates essentially the EDWC and intermediate heating should be included into process synthesis and design in the form of the EDWC−FPH. The initial process design is assumed to be the EDWC obtained above with the addition of a feed preheater. According to the systematic strategy proposed in section 2, the synthesis and design of the EDWDC−FPH is performed and the resultant optimum process design is sketched in Figure 5c. The feed preheater only uses the sensible heat of the entrainer recovered from the bottom of the EDWDC−FPH and the bottom reboiler heat duty turns to be 2316.20 kW. Which is considerably lower than that of the EDWC, implying a great enhancement in thermodynamic efficiency by intermediate heating. 3.4. CEDC Flowsheet versus EDWC with and without Intermediate Heating. The summary of all the process designs studied are presented in Table 2. In comparison with the CEDC flowsheet, although the EDWC achieves a 14.00% reduction in the total reboiler heat duty, the exclusively using of high-pressure steam results in a 17.88% increase in steam cost and consequently brings about a 5.41% increase in the TAC. With the consideration of intermediate heating, the EDWC− FPH becomes not only much more thermodynamic efficient but also more cost-effective than the CEDC flowsheet, securing, respectively, a reduction in the total reboiler heat duty, steam cost, and TAC by 43.71%, 22.84%, and 17.94%.

Figure 6. Temperature profiles of the CEDC flowsheet and EDWC for example I: (a) CEDC−EDC versus EDWC−EDC, (b) CEDC− ERC versus EDWC−ERC.

4. EXAMPLE II: SEPARATING A BINARY AZEOTROPIC MIXTURE OF ACETONE (ACE) AND METHANOL (MEOH) WITH DMSO AS THE ENTRAINER 4.1. Problem Description. As shown in Figure 7a, ACE and MEOH give an azeotrope at 22.74 mol % MEOH under the atmospheric pressure. Figure 7 parts b and c describe, respectively, the VLE behaviors of MEOH/DMSO and ACE/ DMSO and show clearly that the inclusion of DMSO can favor the separation of ACE and MEOH. The example is adopted again from Chien and his co-workers, and the operating conditions and product specifications are reproduced here in Table 3 for quick reference.16 The normal boiling points of the ACE, MEOH, and DMSO are found in the database of ASPEN PLUS and are 56.25, 64.48, and 190.74 °C, respectively. Again,

permits the use of high-pressure steam ($6.6/1000 lb) for reboiler heating, lowering inevitably the steady-state economics as compared with the CEDC flowsheet. While the EDWC− ERC displays a slightly greater temperature span than the CEDC−ERC, the EDWC−EDC exhibits a much greater one than the CEDC−EDC (it is essentially closely related to the mass and thermal coupling between the extractive separation operation and entrainer recovery operation involved). In particular, both the CEDC flowsheet and EDWC feature quite steep temperature profiles in their two stripping sections. The detailed temperature differences are ΔT1, CEDC = 163.79 − 79.82 = 83.97 °C and ΔT2, CEDC = 188.06 − 166.27 = 21.79 °C Table 2. Summary of All Process Designs Studied for Example I configurations TAC for column (103 $/y) TAC for reboiler (103 $/y) TAC for condenser (103 $/y) TAC for cooler (103 $/y) TAC for IR (103 $/y) capital cost (103 $/y) steam cost (103 $/y) cooling water cost (103 $/y) entrainer cost (103 $/y) operating cost (103 $/y) reboiler heat duty (kW) total TAC (103 $/y)

CEDC 434.89 166.52 145.00 80.39 0.00 826.80 720.03 13.67 4.75 738.44 4114.56 1565.24

EDWC

(0.00%)

(0.00%) (0.00%)

495.44 88.56 124.52 80.29 0.00 788.81 848.75 11.75 0.66 861.15 3538.54 1649.96

(+17.88%)

(−14.00%) (+5.41%) 2715

EDWC−FPH 419.68 68.52 125.14 0.00 107.54 720.88 555.56 7.67 0.34 563.57 2316.20 1284.45

(−22.84%)

(−43.71%) (−17.94%)

CEDC−FPH 353.31 132.59 144.95 0.00 100.61 731.46 531.98 9.64 4.80 546.42 2906.24 1277.88

(−26.12%)

(−29.37%) (−18.36%)


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Table 3. Operating Conditions and Design Specifications of Example II parameter condenser pressure (atm) stage pressure drop (atm) feed flow rate (kmol/h) feed temperature (°C) feed composition (mole fraction) product specification (mole fraction)

value

ACE MEOH ACE MEOH DMSO

1 0.0068 540 46.85 0.5 0.5 0.994 0.995 0.9999

4.2. CEDC Flowsheet. The optimum design of the CEDC flowsheet is shown in Figure 8a. The reboiler heat duties are 6871.95 and 5332.90 kW, respectively, for the CEDC−EDC and CEDC−ERC, summing into 12204.85 kW. Figure 9a and b depict the steady-state temperature profiles of the CEDC− EDC and the CEDC−ERC. Since their bottom temperatures are 124.39 and 194.72 °C, respectively, low-pressure steam ($3.0/1000 lb) should be used in the former and high-pressure steam ($6.6/1000 lb) in the latter, thereby being able to maintain a minimum temperature driving force of 10 °C in these two reboilers. 4.3. EDWC with and without Intermediate Heating. The initial process design is again constructed based on the thermodynamic equivalent structure of the obtained CEDC flowsheet and the resultant optimum EDWC is sketched in Figure 8b. The reboiler heat duty is 12009.34 kW and only slightly lower than the total reboiler heat duty of the CEDC flowsheet, implying a marginal enhancement in thermodynamic efficiency by mass and thermal coupling between the CEDC− EDC and CEDC−ERC. The steady-state temperature profiles of the EDWC−EDC and EDWC−ERC involved are also shown in Figure 9a and b, respectively. Note that the bottom temperature becomes now 199.56 °C, so only can highpressure steam be used here as a hot utility. Steep temperature profiles are still found in the two stripping sections of the CEDC flowsheet and EDWC. While the former exhibits temperature differences ΔT1,CEDC = 124.39 − 79.29 = 45.10 °C and ΔT2,CEDC = 194.72 − 120.50 = 74.22 °C, the latter ΔT1,EDWC = 199.56 − 78.15 = 121.41 °C and ΔT2,EDWC = 199.56 − 96.66 = 102.9 °C (Note also the fact that it is the mass and thermal coupling between the extractive separation operation and entrainer recovery operation involved that enables the ΔT1,EDWC to be much greater than the ΔT1,CEDC). It is thus reasonable to improve their performance with intermediate heating during process synthesis and design. Since the vapor split ratio of the resultant EDWC is 0.2 < 0.5 = Z1 < βV = 0.6788 < 0.8, it is reasonable to interpret that the EDWC works in a balanced mode between the extractive separation operation and entrainer recovery operation involved. Intermediate heating should therefore be included into process synthesis and design in the form of the EDWC−IR and the optimum process design is sketched in Figure 8c. Two intermediate reboilers (IR−I and IR−II) are installed between the 39th and 40th stages with the entrainer recovered from the bottom and low-pressure steam as hot utilities, respectively. Although the addition of the intermediate reboilers increases somehow capital cost, it leads to a considerable reduction of reboiler heat duty. The heat duty of bottom reboiler and IR−II sums now to 9626.62 kW, and it is much lower than that of the

Figure 7. T−x−y diagrams for binary mixtures in example II at 1 atm: (a) MEOH/ACE, (b) MEOH/DMSO, (c) ACE/DMSO.

a great temperature difference exists between the boiling points of the two components in the azeotropic mixture and entrainer, implying the possibility that intermediate heating can be an effective means to enhance the steady-state economics of the EDWC. The commercial software ASPEN PLUS is used to simulate the processes and the UNIQUAC model to describe the vapor−liquid phase behaviors. 2716


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Figure 9. Temperature profiles of the CEDC flowsheet and EDWC for example II: (a) CEDC−EDC versus EDWC−EDC, (b) CEDC− ERC versus EDWC−ERC.

designs studied are presented in Table 4. While the EDWC leads to a 1.60% smaller reboiler heat duty than the CEDC flowsheet, the former requires 51.03% more steam cost than the latter. The increased steam cost makes finally the former exhibit a 21.34% larger TAC than the latter. As far as the EDWC−IR is concerned, it achieves simultaneously a reduction in reboiler heat duty, steam cost, and TAC by 21.12%, 29.48%, and 15.49%, respectively.

5. DISCUSSION For the extractive separation of binary azeotropic mixtures, while the great difference in the boiling point temperatures between the two components in the azeotropic mixture and entrainer is a prerequisite for the selection of entrainer, it is also the primary reason for the EDWC to frequently have a higher TAC than the CEDC flowsheet, thereby restricting its applicability and flexibility. The steep temperature profile in the stripping sections is thus an inherent operating characteristic of the EDWC and allows using effectively intermediate heating to enhance its thermodynamic efficiency. It is also due to the same reason that the sensible heat of the entrainer recovered from the bottom of the EDWC should, prior to the other available utilities, be used as a hot utility for intermediate heating. This design strategy permits to fully tap the potential of intermediate heating and is likely to yield great enhancement of steady-state economics. The two examples studied in the current work have demonstrated clearly the feasibility and effectiveness of the systematic strategy proposed in the current work. Even compared with the CEDC flowsheets with intermediate heating (i.e., the CEDC−FPH/CEDC−IR as

Figure 8. Optimum designs for example II: (a) CEDC flowsheet, (b) EDWC, (c) EDWC−IR.

EDWC, implying again a great enhancement in thermodynamic efficiency by intermediate heating. 4.4. CEDC Flowsheet versus EDWC with and without Intermediate Heating. The summary of all the process 2717


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(−28.69%) (−23.61%) (−27.69%) (−22.81%) (−14.24%) (−19.29%) (−21.12%) (−15.49%) (−1.60%) (+21.34%) (0.00%) (0.00%)

830.30 326.84 362.79 220.46 0.00 1740.39 1907.28 40.04 3.66 1950.98 12204.84 3691.37 TAC for column (10 $/y) TAC for reboiler (103 $/y) TAC for condenser (103 $/y) TAC for cooler (103 $/y) TAC for IR (103 $/y) capital cost (103 $/y) steam cost (103 $/y) cooling water cost (103 $/y) entrainer cost (103 $/y) operating cost (103 $/y) reboiler heat duty (kW) total TAC (103 $/y)

CEDC

(0.00%)

767.08 216.36 355.08 220.09 0.00 1558.62 2880.53 39.39 0.73 2920.65 12009.34 4479.27

EDWC

(+51.03%)

880.99 92.79 418.15 112.75 238.03 1742.70 1344.96 31.45 0.29 1376.70 9626.62 3119.41

(−29.48%)

730.97 214.39 399.41 158.65 261.68 1765.11 1176.14 34.25 3.65 1214.03 10466.47 2979.14

(−38.33%)

909.91 61.65 425.99 0 370.09 1767.65 1051.64 28.78 1.33 1081.74 8824.79 2849.39

(−44.86%)

771.13 187.14 420.51 0.00 396.76 1775.54 1012.30 28.37 3.65 1044.31 8703.38 2819.85

(−46.92%)

shown in Figure 10a and b), the EDWC−FPH/EDWC−IR may not be inferior in steady-state economics. As shown in Table 2, the EDWC−FPH really presents a comparable TAC with the CEDC−FPH for example I. For example II, although the EDWC−IR appears to be slightly inferior to the CEDC−IR flowsheet (c.f., Table 4), it is mainly because the sensible heat of the entrainer recovered from the bottom has not been used to its fullest extent (for example, to further preheat the feed where a great temperature driving force is available there). If feed preheating is allowed in both processes, they are considered to have comparable steady-state economics. Figure 11a and b sketch the resultant optimum designs of the EDWC−IR−FPH and the CEDC−IR−FPH, respectively. As shown in Table 4, the EDWC−IR−FPH gives a rather comparable TAC with the CEDC−IR−FPH. It is worthwhile to compare here the effect of arranging intermediate heating to the CEDC flowsheet and EDWC. For example I, while the CEDC−FPH flowsheet cuts the TAC by 18.36% as compared with the CEDC flowsheet, the EDWC− FPH by 22.15% as compared with the EDWC. Similarly, for example II, while the CEDC−IR and the CEDC−IR−FPH flowsheets cut the TACs by 19.29%, and 23.61%, respectively, as compared with the CEDC flowsheet, the EDWC−IR and the EDWC−IR−FPH by 30.36%, and 36.39%, respectively, as compared with the EDWC. It appears to be more advantageous to consider intermediate heating in the EDWC than in the CEDC flowsheet. The phenomenon stems from the fact that intermediate heating reduces the consumption of relatively high-grade utility in the former but low-grade and relatively high-grade utilities in the latter. It is worth mentioning here an interesting phenomenon observed in the current study that different designs of the extractive separation process lead to different entrainer recycle flow rates from the extractive separation operations to the entrainer recovery operations involved. It can clearly be identified from the comparison of process designs shown in Figures 5, 8, 10, and 11. For instance, in example II, the entrainer recycle flow rates of the CEDC−IR and EDWC−IR are obviously smaller than those of the CEDC flowsheet and EDWC, respectively (c.f., Figures 8 and 10). The phenomenon also gives rise to different entrainer makeup flow rates between the process designs presented in the current work. It is considered here that the strong interactions between the extractive separation operation and entrainer recovery operation involved and the high degree of nonlinearity in the thermodynamic properties of the binary azeotropic mixture separated should be responsible for the occurrence of such a unique phenomenon. It is no doubt that the effectiveness of intermediate heating depends heavily on the relative cost between the low-grade and high-grade steams used (which varies accordingly to the changes in plant locations). In the case that the price difference between the low-grade and high-grade steams is great enough, it is certainly advantageous to enhance the steady-state performance of the EDWC−FPH and EDWC−IR with intermediate heating (as shown in the current work). Even in the case that the price difference is fairly small, intermediate heating can still be effective provided that a steep temperature profile is available in the stripping section of the EDWC. However, only should heat recovery from the recycled entrainer be considered in this situation. It should be borne in mind that intermediate heating can only be applied to the cases that the chosen entrainers make the

3

configurations

Table 4. Summary of All Process Designs Studied for Example II

EDWC−IR

CEDC−IR

EDWC−IR−FPH

CEDC−IR−FPH


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Figure 10. Optimum designs of the CEDC−FPH and CEDC−IR for examples I and II, respectively: (a) CEDC−FPH, (b) CEDC−IR.

EDWC less competitive than the CEDC flowsheet. For those situations that the chosen entrainers do not make the EDWC less competitive than the CEDC flowsheet, it cannot be used because there may be no steep temperature profiles to permit heat recovery from the recycled entrainers and use of relatively low-grade utility. Although intermediate heating makes the resultant EDWC−FPH and EDWC−IR (EDWC−IR−FPH) have more complicated structures than the CEDC flowsheet, it never means that they have inevitably complicated process dynamics and operation difficulties. Therefore, if operationally feasible, they should be preferred to the CEDC flowsheet with and without intermediate heating because at least the space needed to accommodate them is likely to be considerably reduced. The dynamics and control of the EDWC−FPH and

EDWC−IR (EDWC−IR−FPH) remain, therefore, to be an important issue to be addressed in the future.

6. CONCLUSIONS For the extractive separations of binary azeotropic mixtures, although the EDWC can be more thermodynamic efficient than the CEDC flowsheet, the chosen entrainers can frequently enable the former to be less economically favorable than the latter. In terms of the operating characteristics of the EDWC, i.e., the existence of steep temperature profiles, intermediate heating is proposed to enhance the steady-state economics of the EDWC and a systematic strategy is thus devised to facilitate process synthesis and design. In case that the EDWC is dominated by the extractive separation operation, intermediate heating should be arranged as a feed preheater because the 2719


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Figure 11. Optimum designs of the EDWC−IR−FPH and CEDC−IR−FPH for example II: (a) EDWC−IR−FPH, (b) CEDC−IR−FPH.

greatest temperature driving forces are reached with the entrainer recovered from the bottom and the available utilities. In case that the EDWC is dominated by the entrainer recovery operation or operates in a balanced mode between these two situations, intermediate heating should be arranged below the lower end of the dividing wall. Under both circumstances, the sensible heat recovered from the recycled entrainer and the energy supplied by low-grade utility can be maximized, permitting consequently great improvement in steady-state economics. Provided that a steep temperature profile is available in the stripping section of the EDWC, the strategy

can be used and effective irrespective of the relative cost between the low-grade and high-grade steams employed. In terms of the two example systems separating, respectively, two azeotropic mixtures of DMC and MEOH and MEOH and ACE, the feasibility and effectiveness of the systematic strategy proposed have been confirmed. It has been demonstrated that the steady-state economics of the EDWC can be substantially enhanced with intermediate heating. The EDWC with intermediate heating is not only much more thermodynamically efficient but also more cost-effective than the CEDC flowsheet. It can even yield comparable steady-state economics with the CEDC flowsheet reinforced also with intermediate heating. The 2720


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Industrial & Engineering Chemistry Research striking outcome reminds us of the fact that two schemes have essentially the same minimum utility requirement because the extremely different process configurations do not alter their separation routines. The systematic strategy proposed in the current work is considered to be of great significance to the synthesis and design of the EDWC. Future work will be centered on the dynamics and operation of the EDWC with intermediate heating.

where M&S indicates the Marshall & Swift Equipment Cost Index and FC = FmFp = 3.67. The stage cost is given by stage cost [$] =

heat exchanger cost [$] =

APPENDIX A: TAC ESTIMATIONS FOR THE CEDC FLOWSHEET AND EDWC The TAC is given by CC β

(A.8)

(A.1)

⎡$⎤ Cs ⎛ Q R ⎞⎛ h⎞ steam cost ⎢ ⎥ = ⎜ ⎟⎜8150 ⎟ y⎠ ⎣ y ⎦ 1000 lb ⎝ λV ⎠⎝

⎡$⎤ cooling water cost ⎢ ⎥ ⎣y⎦

(A.2)

where QR (Btu/h) is the heat duty of reboiler or intermediate reboiler, ΔTR (°F), the temperature driving force, and UR, the overall heat-transfer coefficient, which is assumed to be 250 Btu/(h·ft2·°F) for heating steams and 50 Btu/(h·ft2·°F) for the recycled entrainer. (2) The heat-transfer area for condenser or cooler (AC) is given by

A C[ft2] =

=

(A.11)

(A.3)

Notation

A = hypothetical component or heat-transfer area, ft2 B = hypothetical component or bottom product, kmol/h C = hypothetical component C1 = section of the EDWC, EDWC−FPH, or EDWC−IR C2 = section of the EDWC, EDWC−FPH, or EDWC−IR C3 = section of the EDWC, EDWC−FPH, or EDWC−IR Cs = saturated steam price, $ D = diameter, m or distillate product, kmol/h F = feed flow rate, kmol/h FL = liquid phase feed flow rate, kmol/h FV = vapor phase feed flow rate, kmol/h L = liquid flow rate, kmol/h LC = column length, ft NT = total number of stages QC = condenser heat duty or cooler heat duty, kW Qcon = condenser heat duty, kW Qcoo = cooler heat duty, kW QIR = intermediate reboiler heat duty, kW Qreb = reboiler heat duty, kW QR = reboiler heat duty or intermediate reboiler heat duty, kW R = reflux ratio T = temperature, K Tbot = temperature at bottom, K ΔT = temperature difference, K ΔTC = temperature driving force, K ΔTR = temperature driving force, K

(A.4)

where NT is the total number of stages. (4) Whereas the diameter of the column shell below the lower end of the dividing wall can be directly estimated with ASPEN PLUS, the equivalent diameter (De) of the column shell above the top part of the dividing wall is back-calculated as follows

De =

D12 + D2 2

(A.5)

where D1 and D2 are the diameters of the left and right sides of the dividing wall. The capital and operating costs are calculated according to the following expressions: The column cost is given by column cost [$] =

(A.10)

⎡$⎤ ⎛ kmol ⎟⎞⎛ h ⎞⎛ $ ⎞⎟ entrainer cost ⎢ ⎥ = ⎜FEM ⎜8150 ⎟⎜price h ⎠⎝ y ⎠⎝ kmol ⎠ ⎣y⎦ ⎝

where QC (Btu/h) is the heat duty of condenser or cooler, ΔTC (°F), the log-mean temperature driving force, and UC, the overall heat-transfer coefficient, which is assumed to be 150 Btu/(h·ft2·°F) for the former and 50 Btu/(h·ft2·°F) for the latter. (3) The column length (LC) is given by ⎛N ⎞ LC[ft] = 2.4⎜ T ⎟ ⎝ 0.5 ⎠

$0.03 ⎛ 1 gal ⎞⎛ Q C ⎞⎛ h⎞ ⎜ ⎟⎜ ⎟⎜8150 ⎟ 1000 gal ⎝ 8.314 lb ⎠⎝ 30 ⎠⎝ y⎠

The entrainer cost is given by

QC UCΔTC

(A.9)

where Cs ($) is the saturated steam price, and λV (Btu/lb), the latent heat of steam. The cooling water cost is given by

QR UR ΔTR

⎛ M&S ⎞ 0.65 ⎜ ⎟101.3A (2.29 + FC) ⎝ 280 ⎠

where FC = (Fd + Fp)Fm = (1.35 + 0) × 3.75 for reboiler or intermediate reboiler, and FC = (Fd + Fp)Fm = (1 + 0) × 3.75 for condenser or cooler. The steam cost is given by

where OC is the total operating cost; CC, the total installed capital cost; and β, the payback period. The equipment is sized as follows: (1) The heat-transfer area for reboiler or intermediate reboiler (AR) is given by AR [ft2] =

(A.7)

where FC = Fs + Ft + Fm = 1 + 0 + 1.7 = 2.7. The heat exchanger cost is given by

■

TAC = OC +

⎛ M&S ⎞ 1.55 ⎜ ⎟4.7D LCFC C ⎝ 280 ⎠

⎛ M&S ⎞ 1.066 ⎜ ⎟101.9D LC 0.802(2.18 + FC) C ⎝ 280 ⎠ (A.6) 2721


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Industrial & Engineering Chemistry Research UC = overall heat-transfer coefficient, Btu/(h·ft2·°F) UR = overall heat-transfer coefficient, Btu/(h·ft2·°F) V = vapor flow rate, kmol/h Z = feed composition

Programs Foundation of Ministry of Education of China (20100010110008), and Ministry of Science and Technology of Taiwan under Grant No. MOST 103-2221-E-233-005.

■

Abbreviations

ACE = acetone CEDC = conventional extractive distillation column CEDC−FPH = conventional extractive distillation column with feed preheater CEDC−IR = conventional extractive distillation column with intermediate reboiler CC = total installed capital cost, $ CL = cooler DMC = dimethyl carbonate DMSO = dimethyl sulfoxide DWC = dividing-wall column EDC = extractive distillation column EDWC = extractive dividing-wall column EDWC−FPH = extractive dividing-wall column with feed preheater EDWC−IR = extractive dividing-wall column with intermediate reboiler EF = entrainer feed EM = entrainer makeup ERC = entrainer recovery column FPH = feed preheater HP = high pressure IR = intermediate reboiler LP = low pressure MEOH = methanol MP = medium pressure OC = total operating cost, $/y TAC = total annual cost, $/y

(1) Ho, Y. C.; Ward, J. D.; Yu, C. C. Quantifying Potential Energy Savings of Divided Wall Columns Based on Degree of Remixing. Ind. Eng. Chem. Res. 2011, 50, 1473. (2) Errico, M.; Tola, G.; Rong, B. G.; Demurtas, D.; Turunen, I. Energy Saving and Capital Cost Evaluation in Distillation Column Sequences with a Divided Wall Column. Chem. Eng. Res. Des. 2009, 87, 1649. (3) Chen, W.; Huang, K.; Chen, H.; Xia, C.; Wu, G.; Wang, K. Design and Operation of Dividing-Wall Distillation Columns.1. Diminishing the Black-Hole Problem through Over-Design. Chem. Eng. Process. 2014, 75, 90. (4) Kiss, A. A.; Bildea, C. S. A Control Perspective on Process Intensification in Dividing-Wall Columns. Chem. Eng. Process. 2011, 50, 281. (5) Dejanović, I.; Matijašević, L.; Olujić, Z. Dividing Wall Column A Breakthrough towards Sustainable Distilling. Chem. Eng. Process. 2010, 49, 559. (6) Wu, Y.; Lee, H. Y.; Huang, H. P.; Chien, I. L. Energy-Saving Dividing-Wall Column Design and Control for Heterogeneous Azeotropic Distillation Systems. Ind. Eng. Chem. Res. 2014, 53, 1537. (7) Cristofer, B. B.; Juan, G.; Claudia, G. A.; Ana, L.; Adrián, B. P.; Abel, B. R. Extractive Dividing Wall Column: Design and Optimization. Ind. Eng. Chem. Res. 2010, 49, 3672. (8) Yildirim, Ö .; Kiss, A.; Kenig, E. Dividing Wall Columns in Chemical Process Industry: A Review on Current Activities. Sep. Purif. Technol. 2011, 80, 403. (9) Kiss, A.; Ignat, R. Innovative Single Step Bioethanol Dehydration in an Extractive Dividing-Wall Column. Sep. Purif. Technol. 2012, 98, 290. (10) Xia, M.; Yu, B.; Wang, Q.; Jiao, H.; Xu, C. Design and Control of Extractive Dividing-Wall Column for Separating Methylal− Methanol Mixture. Ind. Eng. Chem. Res. 2012, 51, 16016. (11) Kiss, A. Novel Applications of Dividing-Wall Column Technology to Biofuel Production Processes. J. Chem. Technol. Biotechnol. 2013, 88, 1387. (12) Xia, M.; Xin, Y.; Luo, J.; Li, W.; Shi, L.; Min, Y.; Xu, C. Temperature Control for Extractive Dividing-Wall Column with an Adjustable Vapor Split: Methylal/Methanol Azeotrope Separation. Ind. Eng. Chem. Res. 2013, 52, 17996. (13) Kiss, A. Advanced Distillation Technologies: Design. Control and Applications; Wiley, 2013. (14) Zhang, H.; Ye, Q.; Qin, J.; Xu, H.; Li, N. Design and Control of Extractive Dividing-Wall Column for Separating Ethyl Acetate− Isopropyl Alcohol Mixture. Ind. Eng. Chem. Res. 2014, 53, 1189. (15) Tututi-Avila, S.; Jiménez-Gutiérrez, A.; Hahn, J. Control Analysis of an Extractive Dividing-Wall Column Used for Ethanol Dehydration. Chem. Eng. Process. 2014, 82, 88. (16) Wu, Y.; Hsu, P.; Chien, I. L. Critical Assessment of the EnergySaving Potential of an Extractive Dividing-Wall Column. Ind. Eng. Chem. Res. 2013, 52, 5384. (17) Sun, L.; Wang, Q.; Li, L.; Zhai, J.; Liu, Y. Design and Control of Extractive Dividing Wall Column for Separating Benzene/Cyclohexane Mixtures. Ind. Eng. Chem. Res. 2014, 53, 8120. (18) Wang, S. J.; Huang, H. P.; Yu, C. C. Plantwide Design of Transesterification Reactive Distillation to Co-Generate Ethyl Acetate and n-Butanol. Ind. Eng. Chem. Res. 2010, 49, 750. (19) Kiss, A.; Suszwalak, D. Enhanced Bioethanol Dehydration by Extractive and Azeotropic Distillation in Dividing-Wall Columns. Sep. Purif. Technol. 2012, 86, 70. (20) Gil, I.; Botía, D.; Ortiz, P.; Sánchez, O. Extractive Distillation of Acetone/Methanol Mixture Using Water as Entrainer. Ind. Eng. Chem. Res. 2009, 48, 4858.

Greek Letters

β = payback period, y βV = vapor split ratio βL = characteristic value of vapor split ratio βU = characteristic value of vapor split ratio λV = latent heat of steam, Btu/lb

Subscripts

bot = bottom con = condenser coo = cooler C = condenser or cooler FPH = feed preheater IR = intermediate reboiler L = liquid phase or characteristic value of vapor split ratio reb = reboiler R = reboiler or intermediate reboiler U = characteristic value of vapor split ratio V = vapor phase or characteristic value of vapor split ratio

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REFERENCES

AUTHOR INFORMATION

Corresponding Author

*Phone: +86−10−64437805. Fax: +86−10−64437805. E-mail: [email protected] (K.H.). Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The research is financially supported by the National Science Foundation of China (21076015 and 21376018), Doctoral 2722


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Industrial & Engineering Chemistry Research (21) Tavan, Y.; Shahhosseini, S.; Hosseini, S. Feed-Splitting Technique in the Extractive Distillation of CO2−Ethane Azeotropic Process. Sep. Purif. Technol. 2014, 122, 47. (22) Agrawal, R.; Herron, D. M. Efficient Use of an Intermediate Reboiler or Condenser in a Binary Distillation. AIChE J. 1998, 44, 1303. (23) Agrawal, R.; Herron, D. M. Intermediate Reboiler and Condenser Arrangement for Binary Distillation Columns. AIChE J. 1998, 44, 1316. (24) Luyben, W. L. Design and Control of Distillation Columns with Intermediate Reboilers. Ind. Eng. Chem. Res. 2004, 43, 8244. (25) Malinen, I.; Tanskanen, J. Thermally Coupled Side-Column Configurations Enabling Distillation Boundary Crossing. 2. Effects of Intermediate Heat Exchangers. Ind. Eng. Chem. Res. 2009, 48, 6372. (26) Wankat, P. C.; Kessler, D. P. Two-Feed Distillation: SameComposition Feeds with Different Enthalpies. Ind. Eng. Chem. Res. 1993, 32, 3061. (27) Seider, W. D.; Seader, J. D.; Lewin, D. R.; Widagdo, S. Product and Process Design: Principles Synthesis. Analysis, and Evaluation; Wiley, Hoboken, NJ, 2010. (28) Douglas, J. M. Conceptual Design of Chemical Processes; McGrawHill: New York, 1988. (29) Hsu, K. Y.; Hsiao, Y. C.; Chien, I. L. Design and Control of Dimethyl Carbonate-Methanol Separation via Extractive Distillation in the Dimethyl Carbonate Reactive-Distillation Process. Ind. Eng. Chem. Res. 2010, 49, 735.

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Improving the Performance of Extractive Dividing-Wall Columns with

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