Application of Vapor Recompression to Heterogeneous Azeotropic

Article pubs.acs.org/IECR

Application of Vapor Recompression to Heterogeneous Azeotropic Dividing-Wall Distillation Columns Li Shi,† Kejin Huang,*,† San-Jang Wang,*,‡ Jieping Yu,† Yang Yuan,† Haisheng Chen,† and David S. H. Wong§ †

College of Information Science and Technology, Beijing University of Chemical Technology, Beijing 100029, People’s Republic of China ‡ Center for Energy and Environmental Research, National Tsing Hua University, Hsinchu 300, Taiwan § Department of Chemical Engineering, National Tsing Hua University, Hsinchu 300, Taiwan ABSTRACT: Although heterogeneous azeotropic dividing-wall distillation columns (HADWDCs) can achieve considerable reductions in capital investment and operating cost as compared to conventional heterogeneous azeotropic distillation systems, their steady-state performance is likely to be greatly enhanced by the application of vapor recompression heat pump (VRHP) technology owing to their unique process configuration of one top condenser and two bottom reboilers and favorable steadystate behavior of relatively small top-to-bottom temperature elevations. In the current work, the feasibility and effectiveness of reinforcing the HADWDC with the VRHP (HADWDC-VRHP) are examined. A systematic procedure is developed to effectively determine the optimum combination of the VRHP and the HADWDC in order to minimize the total annual cost. Two example systems for the separation of isopropyl alcohol and water and pyridine and water with cyclohexane and toluene as entrainers, respectively, are studied to evaluate the steady-state economics of the HADWDC-VRHP. It is shown that the proposed procedure can yield the most effective connection between the VRHP and the HADWDC (in terms of the number, locations, and heat duties of intermediate heat exchangers employed between them), and the resultant HADWDC-VRHP requires a considerably lower operating cost than the corresponding HADWDC. Although additional capital investment is needed, the reduction in operating cost generally leads to a reasonable payback time. These outcomes reflect the potential economic advantages of employing the VRHP in the synthesis and design of the HADWDC. cost-effective than the CHADS.11−13 Owing to the separation of binary close-boiling and azeotropic mixtures, the HADWDC is commonly characterized by a relatively small top-to-bottom temperature elevation despite the addition of an entrainer, thereby establishing a favorable situation for the application of vapor recompression heat pump (VRHP) technology to enhance further its steady-state performance. Nonetheless, very few studies have identified and focused on this issue thus far. The VRHP represents an especially useful and effective technology for enhancing the operation efficiency of conventional distillation columns where thermal heat is generally supplied to the bottom reboiler at the highest temperature and withdrawn from the top condenser at the lowest temperature.14−16 Relevant studies can be dated back to the early 1940s and their primary objective is to develop a thermodynamically efficient and yet cost-effective combination of these two units. Because the VRHP imposes a stringent constraint on the temperature elevation from heat source to heat sink, its feasible connections to distillation columns are greatly confined and needed to be frequently determined with close reference to the operation characteristics of distillation columns.17,18 Thus, far, various types of process schemes have been developed, and they can be generally divided into three major categories: VRHP-assisted

1. INTRODUCTION Heterogeneous azeotropic distillation systems are commonly used in chemical process industry to separate binary close-boiling and azeotropic mixtures by adding an entrainer to alter the relative volatility between the involved components.1 Figure 1a shows a conventional heterogeneous azeotropic distillation system (CHADS), which includes an azeotropic distillation column (ADC) and an entrainer recovery distillation column (ERDC). The ADC separates one component from the bottom and an azeotropic mixture of the separated components and the added entrainer from the top. The azeotropic mixture splits into two phases in the decanter with the organic phase refluxed to the ADC and the aqueous phase fed to the ERDC. In the ERDC, the other component is extracted from the bottom and the entrainer is withdrawn from the top and recycled back to the ADC. Thus, far, numerous studies have investigated the design and control of the CHADS.2−7 With the introduction of thermal coupling between the ADC and the ERDC, heterogeneous azeotropic dividing-wall distillation columns (HADWDCs) have been developed recently, as shown in Figure 1b.8−10 Thus, the ADC and the ERDC are combined into a single shell with one condenser at the top and two reboilers at the bottom. The dividing wall runs from the bottom to a certain stage in the middle and generates one common section at the top for enriching the overhead vapor near the azeotropic composition and two separate sections at the bottom for purifying, respectively, the two components involved in the mixture. Some studies have already confirmed that the HADWDC is much more thermodynamically efficient and © 2015 American Chemical Society

Received: Revised: Accepted: Published: 11592

August 9, 2015 October 23, 2015 November 3, 2015 November 3, 2015 DOI: 10.1021/acs.iecr.5b02929 Ind. Eng. Chem. Res. 2015, 54, 11592−11609

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Industrial & Engineering Chemistry Research

Therefore, the potential of such a combination merits an indepth investigation. In the present study, we primarily explore the feasibility and effectiveness of applying the VRHP to the HADWDC. A systematic procedure is developed to determine their optimum combination and operating conditions. Two representative example systems with different operation characteristics are used to evaluate the steady-state performance of the HADWDC reinforced with the VRHP (i.e., the proposed HADWDCVRHP). One is the separation of isopropyl alcohol (IPA) and water using cyclohexane (CYH) as an entrainer and the other is the separation of pyridine (PIE) and water using toluene (TUE) as an entrainer. The steady-state economics of the HADWDC and the HADWDC-VRHP are compared in detail, and finally, the insights gained remarks are summarized.

2. APPLICATION OF VAPOR RECOMPRESSION TO THE HADWDC 2.1. Operation Characteristics of the HADWDC. As mentioned in the previous section, the HADWDC possesses two unique features. The first is its process configuration, i.e., one top condenser and two bottom reboilers, due to the thermal coupling between the ADC and the ERDC, which enables the HADWDC to be generally characterized by a much greater heat duty in the top condenser than in each of the two bottom reboilers. With the application of the VRHP, the overhead vapor can be recompressed to a high pressure and temperature and used to provide heat to the two bottom reboilers or two stripping sections along the dividing wall (in terms of intermediate heat exchangers), thereby enhancing the steady-state performance of the HADWDC considerably. The second is its steady-state behavior, i.e., a relatively small top-to-bottom temperature elevation (in spite of the addition of an entrainer). This behavior is closely related to the very similar thermodynamic properties of the components contained in the separated close-boiling or azeotropic mixtures and particularly favors the application of the VRHP to the HADWDC because of the low compression work required. These two features make the HADWDC an extremely favorable candidate for reinforcement with the VRHP. Therefore, the objective of the present study is to assess the feasibility and effectiveness of this process intensification strategy. 2.2. Potential Configurations of the HADWDC-VRHP. Owing to the existence of two stripping sections along the dividing wall, there exist three major types of combinations of the VRHP and the HADWDC as shown in Figure 2. Because these combinations are strongly dependent on the thermodynamic properties of the separated mixture and the costs of available utilities, for the sake of generality, they are indicated uniformly in terms of one/two intermediate heat exchangers connected to a certain stage in the stripping section of the ADC and/or the ERDC (highlighted in the figure with bold lines). The situations in which the VRHP is directly linked to the bottom reboilers are not shown here because they can be considered as the specific arrangements of the intermediate heat exchangers. When the reboiler heat duty of the ERDC is considerably smaller than that of the ADC (i.e., QREB,ERDC/QREB,ADC < εL), the VRHP should be arranged to recompress the overhead vapor and provide heat to the stripping section of the ADC (Figure 2a). This type of process configurations is referred to as the HADWDCVRHP(ADC), hereinafter. When the reboiler heat duty of the ERDC is somewhat comparable to that of the ADC (i.e., εL < QREB,ERDC/QREB,ADC < εU), the VRHP should be arranged to

Figure 1. Schemes for heterogeneous azeotropic distillation systems: (a) CHADS and (b) HADWDC.

distillation columns,19 internally heat-integrated distillation columns,20−22 and externally heat-integrated distillation columns.23,24 There exist a number of comprehensive overviews that elaborate the advantages and disadvantages of these process schemes.25−28 Recently, the VRHP has been applied to aid more complicated distillation columns, including reactive distillation columns and dividing-wall distillation columns. To avoid unfavorable operating conditions caused by wide-boiling reacting mixtures, Kumar et al. proposed multistage vapor recompression with intermediate heat exchangers for reactive distillation columns.29 Luo et al. used the VRHP with an intermediate heat exchanger to facilitate the steady-state performance of an extractive dividing-wall distillation column for purifying bioethanol fuel and achieved a substantial reduction in total annual cost.30 The relatively small top-to-bottom temperature elevation of the HADWDC is likely to make the VRHP an even more attractive means to improve its steady-state performance. 11593

DOI: 10.1021/acs.iecr.5b02929 Ind. Eng. Chem. Res. 2015, 54, 11592−11609

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Figure 2. Potential schemes of the HADWDC-VRHP: (a) HADWDC-VRHP(ADC), (b) HADWDC-VRHP(ADC-ERDC), and (c) HADWDCVRHP(ERDC).

set as 0.3 and 3, respectively, to facilitate the synthesis and design of the HADWDC-VRHP. 2.3. Systematic Procedure for the Synthesis and Design of the HADWDC-VRHP. The key aspect of the process synthesis and design include the effective determination of the configuration and operating conditions of the HADWDCVRHP. The decision variables include the locations of feed (NF) and aqueous phase reflux (NA); the number of stages in the common section above the dividing wall (NADW); the number of stages along the dividing wall (NDW,ADC and NDW,ERDC, which may take different values in the present work); the liquid split ratio (βL, which is defined as the ratio between the liquid flow rates to the ERDC side of the dividing wall and from the top common section); and the number, locations (NIHE,ADC and/or NIHE,ERDC), and heat duties (QIHE,ADC and/or QIHE,ERDC) of the

recompress the overhead vapor and release heat to the stripping sections of the ADC and the ERDC simultaneously (Figure 2b). This type of process configurations is referred to as the HADWDC-VRHP(ADC-ERDC), hereinafter. Finally, when the reboiler heat duty of the ERDC is considerably greater than that of the ADC (i.e., QREB,ERDC/QREB,ADC > εU), the VRHP should be arranged to recompress the overhead vapor and relinquish heat to the stripping section of the ERDC (Figure 2c). This type of process configuration is referred to as the HADWDC-VRHP(ERDC) hereinafter. Note that εL and εU are two important parameters that define the operation characteristics of the HADWDC. They are strongly dependent on the selected operating conditions and thermodynamic properties of the components involved in the separated binary close-boiling or azeotropic mixtures. In most cases, they can be 11594


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Figure 3. Systematic procedure for the synthesis and design of the HADWDC-VRHP.

(indicated in the figure with the bold shadowed blocks); this can greatly reduce the complexities and computation intensity involved in the process synthesis and design. An initial process design, which is estimated and selected as a starting point for process evolvement, serves to confine all the structural and operating decision variables in their feasible regions, consequently facilitating the search for optimum process design.

intermediate heat exchangers (IHEs) connecting the VRHP to the stripping sections of the ADC and/or the ERDC. A systematic procedure is developed for the synthesis and design of the HADWDC-VRHP in terms of its inherent operation characteristics, as shown in Figure 3. With regard to the reboiler heat duties of the ADC and the ERDC, the connection between the HADWDC and the VRHP should be determined first 11595


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Industrial & Engineering Chemistry Research Table 1. Operating Conditions and Design Specifications of Example I parameter condenser pressure (atm) stage pressure drop (atm) feed flow rate (mol/min) feed temperature (K) feed composition (mole fraction) product specification (mole fraction)

value

IPA water IPA water

1.05 0.0079/0.012 1365.274 355.85 0.610157 0.389843 0.999999 0.999

The heat duties of the two bottom reboilers are adjusted to maintain the two bottom products at their specifications.

Figure 5. Temperature profiles of the HADWDC for example I.

Figure 4. Residue curve map and optimum design of the HADWDC for example I. 11596


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Figure 6. Relationship between the TAC and relevant decision variables for example I.

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Figure 7. Optimum design of the HADWDC-VRHP(ADC-ERDC) for example I.

Figure 8. Comparison of steady-state behaviors between the HADWDC and the HADWDC-VRHP(ADC-ERDC). 11598


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Industrial & Engineering Chemistry Research Table 2. Comparison between All the Process Designs Studied for Example I process design relevant cost

HADWDC 3

capital cost for column (10 $) capital cost for reboilers (103 $) capital cost for condenser (103 $) capital cost for decanter (103 $) capital cost for preheater (103 $) capital cost for compressors (103 $) steam cost (103 $/year) cooling water cost (103 $/year) electricity cost (103 $/year) TUC (kW) CI (103 $) OC (103 $/year) TAC (103 $/year) payback time (βpbt) (year)

352.98 124.34 414.41 159.40 0.00 0.00 1304.18 44.29 0.00 3277.99 1051.13 1348.47 1698.85

HADWDC-VRHP(ADC-ERDC) 354.29 413.68 276.74 159.36 24.64 1280.92 110.72 15.33 149.56 1080.58 2509.63 275.61 1112.15 1.36

(0.00%) (0.00%) (0.00%) (0.00%)

(−67.04%) (+138.76%) (−79.56%) (−34.54%)

I-HADWDC-VRHP(ADC-ERDC) 352.98 415.40 251.75 159.40 26.00 1844.70 10.03 14.03 223.49 1224.09 3050.23 247.55 1264.29 1.82

(−62.46%) (+190.19%) (−81.64%) (−25.58%)

Figure 9. Typical design of the HADWDC-VRHP for example I.

The commonly used grid-search method is adopted to adjust all the decision variables because of its simplicity (in principle) and relatively low computational requirements. However, an exceptional case occurs in the determination of the intermediate heat exchangers connecting the VRHP to the ADC and/or the ERDC. Their locations and heat duties are sequentially determined within a single iteration loop because of their intimate relationships (also indicated with bold shadowed blocks). Such special treatment allows for effective tapping of the full potential of the HADWDC-VRHP. In the present study, the minimization of total annual cost (TAC) is taken as the objective function for the process synthesis

and design. The TAC consists of operating cost (OC) and capital investment (CI) annualized by a payback time, βpbt = 3 (c.f., eq 1). The OC includes the costs of steam, cooling water, electricity, and entrainer makeup, whereas the CI includes the costs of column shells, stages, intermediate heat exchangers, reboilers, condenser, decanter, compressors, and preheater. The formulas for these cost estimations are taken from the work of Douglas and presented in Appendix A for quick reference.31 In particular, the capital cost of the dividing wall is considered to be negligible as compared to the other capital costs. For comparative analysis of the operation efficiency of the HADWDCVRHP, we assume a ratio of 3 to convert compressor power 11599


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Industrial & Engineering Chemistry Research 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 (K) feed composition (mole fraction) product specification (mole fraction)

value

PIE water PIE water

1.05 0.008 1000 298.15 0.1 0.9 0.999 0.999

TAC = OC + CI/βpbt

(1)

Q UCONS = Q REB + Q PRE + 3Q COMP

(2)

3. EXAMPLE I: APPLICATION OF VAPOR RECOMPRESSION TO THE HADWDC FOR SEPARATING IPA AND WATER WITH CYH AS ENTRAINER 3.1. Problem Description. The example system is adopted from the work of Wang et al., and the operating conditions and product specifications are reproduced in Table 1 for quick reference.10 The product purities of IPA and water are set as 99.9999 and 99.9 mol %, respectively, and the temperature of the

(QCOMP) into thermal heat (c.f., eq 2). The value of M&S is set as 1536.5 here.

Figure 10. Residue curve map and optimum design of the HADWDC for example II. 11600


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Industrial & Engineering Chemistry Research decanter is set as 313.15 K. Figure 4a shows the residue curve map of the ternary mixture. IPA and water form a binary azeotrope at 352.96 K with a composition of 70 mol % of IPA and 30 mol % of water. IPA and CYH form a binary azeotrope at 342.22 K with a composition of 39 mol % of IPA and 61 mol % of CYH. CYH and water form a binary azeotrope at 342.59 K with a composition of 70 mol % of CYH and 30 mol % of water. A minimum-boiling ternary heterogeneous azeotrope is also formed at 336.75 K with a composition of 25 mol % of IPA, 53 mol % of CYH, and 22 mol % of water, which indicates that the inclusion of CYH could favor the separation of IPA and water.10 The normal boiling points of IPA, CYH, and water are 355.20, 353.93, and 373.17 K, respectively; the relatively small temperature differences imply that the VRHP could be effective in improving the operation efficiency of the HADWDC. The commercial software ASPEN PLUS is employed to perform steady-state simulation, and the NRTL thermodynamic model is adopted to estimate vapor−liquid and liquid−liquid equilibrium relationships. All thermodynamic parameters are taken from the work of Wang et al.2 The unit costs of lowpressure steam (at 5 bar and 433.15 K) and cooling water (at 303.15 K) are 13.28 and 0.354 $/GJ, respectively. The annual operating time is assumed to be 8322 h. 3.2. Synthesis and Design of the HADWDC. Using the same steady-state operating conditions and product specifications as Wang et al., we reproduce their optimum process design as shown in Figure 4b.10 While the reboiler heat duties of the ADC and the ERDC are 1652.01 and 1625.98 kW, respectively, the condenser heat duty is −4176.33 kW, which is considerably greater than the summation of the formers. The temperature profiles of the HADWDC are delineated in Figure 5. Note that the process design leads to a top temperature of 337.9 K and two bottom temperatures of 359.9 and 376.8 K, respectively, for the ADC and the ERDC. The top-to-bottom temperature elevations are 22.0 K in the ADC and 38.9 K in the ERDC, implying again that the VRHP could be effective in enhancing the steady-state performance of the HADWDC. 3.3. Synthesis and Design of the HADWDC-VRHP. Because 0.3 < QREB,ERDC/QREB,ADC = 1625.98/1652.01 = 0.984 < 3, the VRHP should be mounted, respectively, to the ADC and the ERDC of the HADWDC, i.e., the HADWDC-VRHP(ADCERDC) configuration should be adopted. In this example, the efficiencies of the compressors and motors are assumed to be 0.8 and 0.9, respectively. The unit cost of electricity, adopted from the work of Turton et al., is 16.28 $/GJ.32 A temperature difference of 10 K is assumed between the recompressed overhead vapor and the stages to be heated in the stripping sections of the ADC and the ERDC. The synthesis and design of the HADWDC-VRHP(ADC-ERDC) is conducted according to the systematic procedure proposed in the previous section. Figure 6 shows the relationship between the TAC and the relevant decision variables. One stage should be included in the common section above the dividing wall (c.f., Figure 6a), whereas 19 and 9 stages should be included in the ADC and the ERDC along the dividing wall, respectively (c.f., Figure 6b,c). Because the intermediate heating lowers the driving forces of mass transfer, appropriate compensation should be made; this is accomplished through the inclusion of an additional stage in the ERDC. The liquid split ratio is found to be 9 × 10−4, which is 3 times as high as that of the HADWDC (c.f., Figure 6d). The processed feed is introduced into the top stage of the common section (c.f., Figure 6e), whereas the aqueous phase reflux is

Figure 11. Temperature profiles of the HADWDC for example II.

introduced into stage 2 of the ERDC. No changes are found before and after the combination of the HADWDC and the VRHP (c.f., Figure 6f). It is economically favorable to employ the recompressed overhead vapor to provide heat to both the ADC and the ERDC. The recompressed overhead vapor is condensed in the bottom reboiler (stage 19) of the former, and it should be condensed in an intermediate heat exchanger located at stage 6 of the latter (c.f., Figure 6g,h). The relatively high steam price makes it economically favorable to connect directly the VRHP to the bottom reboiler of the ADC in this case. The reboiler heat duty of the ADC is completely provided by the recompressed overhead vapor, whereas the heat duty of the intermediate heat exchanger of the ERDC is set as 1372 kW because of the potential constraint of stage drying (c.f., Figure 6i,j). Here, the constraint is avoided by confining the intermediate heat exchanger to maximally vaporize 70% of the liquid flow rate leaving the selected stage. The resultant optimum design of the HADWDC-VRHP(ADC-ERDC) is shown in Figure 7. Note that preheating of the overhead vapor is necessary to prevent its condensation during recompression.17 The overhead vapor is recompressed through a two-stage compressor in the ADC, whereas the recompressed overhead vapor from the first stage is used to relinquish heat to stage 6 of the ERDC. The intermediate heating lowers the bottom reboiler heat duty to 255.10 kW. The inclusion of the VRHP leads to additional utility consumption, including the shaft power of the compressors (251.58 and 15.85 kW) and the heat duty of the preheater (23.19 kW). Figure 8 compares the steady-state behaviors of the HADWDC and the HADWDC-VRHP(ADC-ERDC). Here, the dashed lines represent the behaviors of the HADWDC, and the solid lines represent the behaviors of the HADWDCVRHP(ADC-ERDC). In the case of the liquid composition profiles (c.f., Figure 8a,b), fairly small differences are observed in the ADC, whereas fairly large ones are observed in the ERDC. The arrangement of the VRHP narrows the pinch zone below the position for introducing the aqueous phase reflux; this indicates the favorable effect of intermediate heating on the separation operation. As for the vapor and liquid flow rates (c.f., Figure 8c,d), negligible differences are again observed in the ADC. The application of the VRHP considerably reduces the heat loads and the flow rates of vapor and liquid below the stage with the intermediate heat exchanger in the ERDC. Although different diameters can, in principle, be assigned to the top and bottom sections in the ERDC, it is sometimes reasonable to adopt a uniform one in order to simplify the resultant process 11601


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Figure 12. Relationship between the TAC and relevant decision variables for example II.

denoted as the I-HADWDC-VRHP(ADC-ERDC), which is also summarized in Table 2. This design is characterized by recompressing the overhead vapor through a two-stage compressor, and the recompressed overhead vapor from the first and second stages is directly condensed in the two bottom reboilers of the ADC and the ERDC, respectively. In comparison with the I-HADWDC-VRHP(ADC-ERDC), the HADWDC-VRHP(ADC-ERDC) achieves substantial reductions of 51.43% and 8.96% in the CI and TAC, respectively, albeit with an increase of 2.08% in the OC (with regard to the HADWDC). The CI and TAC are reduced mainly because the compression ratio of the HADWDC-VRHP(ADC-ERDC) is lower than that of the

design. With deliberate arrangement of internal structures, the flooding or weeping phenomenon of the ERDC in the HADWDC-VRHP(ADC-ERDC) can be avoided. 3.4. HADWDC versus HADWDC-VRHP (ADC-ERDC). Table 2 shows an in-depth comparison of the steady-state economics of the HADWDC and the HADWDC-VRHP(ADCERDC). In comparison with the HADWDC, the HADWDCVRHP(ADC-ERDC) achieves considerable reductions in the OC, total utility consumption (TUC), and TAC (79.56%, 67.04%, and 34.54%, respectively) at the expense of a 138.76% increase in the CI. This leads to a short payback time of only 1.36 years. Figure 9 shows a typical design of the HADWDC-VRHP, 11602


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Figure 13. Optimum design of the HADWDC-VRHP(ERDC) for example II.

Figure 14. Comparison of steady-state behaviors between the HADWDC and the HADWDC-VRHP(ERDC).

I-HADWDC-VRHP(ADC-ERDC), whereas the increase in the OC is due to a certain amount of overhead vapor condensed

directly in the trim-condenser of the HADWDC-VRHP(ADC-ERDC). The substantial differences in their steady-state 11603


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Industrial & Engineering Chemistry Research Table 4. Comparison between All the Process Designs Studied for Example II process design relevant cost capital cost for column (103 $) capital cost for reboilers (103 $) capital cost for condenser (103 $) capital cost for decanter (103 $) capital cost for compressors (103 $) steam cost (103 $/year) makeup cost (103 $/year) cooling water cost (103 $/year) electricity cost (103 $/year) TUC (kW) CI (103 $) OC (103 $/year) TAC (103 $/year) payback time (βpbt) (year)

HADWDC 981.14 789.36 1087.90 198.50 0.00 1052.89 3.79 31.40 0.00 11518.70 3056.90 1084.29 2107.05

(0.00%) (0.00%) (0.00%) (0.00%)

HADWDC-VRHP(ERDC) 1029.75 1559.93 693.67 196.48 1632.41 389.56 3.63 8.41 141.30 5435.10 5112.24 539.27 2246.98 3.77

(−52.81%) (+67.24%) (−50.27%) (+6.64%)

II-HADWDCVRHP(ERDC) 979.13 1563.01 709.50 198.37 1797.89 397.23 3.56 8.94 159.00 5662.10 5247.90 565.17 2318.03 4.22

II-HADWDCVRHP(ADC-ERDC)

(−50.84%) (+71.67%) (−47.88%) (+10.01%)

1118.50 1779.29 699.91 199.35 1677.52 392.58 3.71 8.44 140.21 5456.20 5474.57 541.23 2369.80 4.45

(−52.63%) (+79.09%) (−50.08%) (+12.47%)

respectively, whereas the top temperature of the common section above the dividing wall is 360.6 K, resulting in two top-to-bottom temperature elevations of 33.2 and 16.4 K, respectively. The relatively small temperature spans thus permit the application of the VRHP to the HADWDC. 4.3. Synthesis and Design of the HADWDC-VRHP. Because QREB,ERDC/QREB,ADC = 8914.82/2603.95 = 3.424 > 3, the VRHP should be mounted to the ERDC of the HADWDC, i.e., the HADWDC-VRHP(ERDC) configuration should be adopted. As in example I, the efficiencies of the compressor and motor are assumed as 0.8 and 0.9, respectively. According to Seider et al., the unit cost of electricity is taken to be 0.04 $/(kW·h).34 As in example I, a temperature difference of 10 K is assumed between the recompressed overhead vapor and the stage to be heated in the stripping section of the ERDC. The synthesis and design of the HADWDC-VRHP(ERDC) is carried out according to the systematic procedure proposed in the current work. Figure 12 shows the relationship between the TAC and the relevant decision variables. Five stages are accommodated in the common section above the dividing wall (c.f., Figure 12a), whereas both the ADC and the ERDC should have 12 stages along the dividing wall (c.f., Figure 12b,c). The liquid split ratio is found to be 0.192, which is slightly higher than that of the HADWDC (c.f., Figure 12d). Both the feed and the aqueous phase reflux are introduced into stage 9 of the ERDC (c.f., Figure 12e,f). It is economically favorable to use the recompressed overhead vapor to provide heat to the ERDC, and the intermediate heat exchanger should be located at stage 12 (c.f., Figure 12g). All the latent heat of the recompressed overhead vapor is released here; thus, the heat duty of the intermediate heat exchanger is around 7166 kW (c.f., Figure 12h). The resultant optimum design of the HADWDC-VRHP(ERDC) is shown in Figure 13. Owing to the inclusion of the VRHP, the reboiler heat duty of the ERDC is reduced to 1815.34 kW (from 8914.82 kW of the HADWDC). However, this is achieved at the expense of 390.10 kW of shaft power of the compressor. Even though the ERDC has a rather small top-tobottom temperature elevation, in this situation, it is not economically justifiable to employ the recompressed overhead vapor to provide heat directly to the bottom reboiler because of the relatively high compressor cost and relatively low steam price. Figure 14 compares the steady-state behaviors of the HADWDC and the HADWDC-VRHP(ERDC). Here, the

economics reflect the importance of effectively determining the optimum combination of the HADWDC and the VRHP.

4. EXAMPLE II: APPLICATION OF VAPOR RECOMPRESSION TO THE HADWDC FOR SEPARATING PIE AND WATER WITH TUE AS ENTRAINER 4.1. Problem Description. The example system is taken from the work of Wu et al., and the operating conditions and product specifications are reproduced in Table 3 for quick reference.12 The product purities are set as 99.9 mol % for PIE in the ADC and 99.9 mol % for water in the ERDC, and the temperature of the decanter is set as 313.15 K. Figure 10a shows the residue curve map of the ternary mixture. PIE and water form a binary azeotrope at 368.04 K with a composition of 23 mol % of PIE and 77 mol % of water. PIE and TUE form a binary azeotrope at 383.30 K with a composition of 23 mol % of PIE and 77 mol % of TUE. TUE and water form a binary azeotrope at 357.68 K with a composition of 44 mol % of TUE and 56 mol % of water. Because the TUE-water azeotrope is located in the heterogeneous region with the lowest boiling point, the addition of TUE is favorable for the separation of PIE and water.12 The normal boiling points of PIE, TUE, and water are 388.31, 383.83, and 373.17 K, respectively; the relatively small temperature differences imply that the VRHP could be employed to improve the steady-state performance of the HADWDC. As in example I, the commercial software ASPEN PLUS is employed for rigorous steady-state simulation, and the NRTL thermodynamic model is adopted for the representation of vapor−liquid and liquid−liquid equilibrium relationships. All the thermodynamic parameters are taken from the work of Pommier et al.33 The unit costs of low-pressure steam (at 50 psig and 420.15 K) and cooling water (at 305.37 K) are 6.60 $/1000 kg and 0.02 $/m3, respectively. The annual operating time is assumed to be 8150 h. 4.2. Synthesis and Design of the HADWDC. Using the same steady-state operating conditions and product specifications as Wu et al., we reproduce their optimum process design as shown in Figure 10b.12 The condenser heat duty is −9417.03 kW and the two reboiler heat duties are 2603.95 and 8914.82 kW, respectively, for the ADC and the ERDC. The temperature profiles of the HADWDC are shown in Figure 11. The bottom temperatures of the ADC and the ERDC are 393.8 and 377.0 K, 11604


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Figure 15. Two typical designs of the HADWDC-VRHP for example II: (a) II-HADWDC-VRHP(ERDC), and (b) II-HADWDC-VRHP(ADC-ERDC).

pinch zone can be readily observed by the inclusion of intermediate heating. As in example I, the application of the VRHP considerably reduces the heat loads below the stage with the intermediate heat exchanger in the ERDC (c.f., Figure 14c,d).

dashed lines represent the behaviors of the HADWDC, and the solid lines represent the behaviors of the HADWDC-VRHP(ERDC). Although slight differences are observed in the liquid composition profiles (c.f., Figure 14a,b), the shrinkage of the 11605


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Industrial & Engineering Chemistry Research Although a uniform diameter is employed here, the flooding or weeping phenomenon of the ERDC in the HADWDCVRHP(ERDC) can also be avoided by deliberate arrangement of internal structures. 4.4. HADWDC versus HADWDC-VRHP (ERDC). Table 4 shows a detailed comparison of the steady-state economics of the HADWDC and the HADWDC-VRHP(ERDC). The HADWDC-VRHP(ERDC) achieves considerable reductions of 50.27% and 52.81%, respectively, in the OC and TUC (as compared with the HADWDC), highlighting the advantages of employing the VRHP to reinforce the HADWDC. Although the HADWDC-VRHP(ERDC) leads to an increase of 67.24% in the CI, it increases the TAC by only 6.64%. Moreover, even though the HADWDC-VRHP(ERDC) entails a relatively long payback time of 3.77 years, it can provide significant economic benefits from a long-term perspective (such as 10 years). Figure 15a,b show two typical designs of the HADWDC-VRHP, denoted as the II-HADWDC-VRHP(ERDC) and II-HADWDC-VRHP(ADC-ERDC), respectively, which are summarized in Table 4. The II-HADWDC-VRHP(ERDC) is characterized by recompressing the overhead vapor and releasing its latent heat to the bottom reboiler of the ERDC, whereas the II-HADWDCVRHP(ADC-ERDC) is characterized by recompressing the overhead vapor and releasing its latent heat to two intermediate heat exchangers at stages 9 and 14 of the ADC and the ERDC, respectively. The HADWDC-VRHP(ERDC) achieves reductions of 4.43%, and 3.37% in the CI and TAC as compared to the II-HADWDC-VRHP(ERDC) (with regard to the HADWDC), as well as reductions of 11.85% and 5.83% in the same aspects as compared to the II-HADWDC-VRHP(ADC-ERDC) (with regard to the HADWDC). As in example I, these comparisons reflect the importance of effectively determining the optimum combination of the HADWDC and the VRHP.

corroborated the feasibility and effectiveness of the systematic procedure proposed in the current work.

6. CONCLUSIONS The HADWDC features one top condenser and two bottom reboilers in process configuration and relatively small top-tobottom temperature elevation in steady-state behavior. These two characteristics make it especially favorable for the application of the VRHP to enhance its steady-state performance. The potential configurations of the resultant HADWDC-VRHP were presented and a systematic procedure was developed for process synthesis and design. Two example systems for separating two binary azeotropic mixtures of IPA and water and PIE and water with CYH and TUE as entrainers, respectively, were studied to assess the steady-state performance of the HADWDC-VRHP. It was shown that the systematic procedure proposed can effectively determine the optimum combination of the HADWDC and the VRHP in terms of the number, locations, and heat duties of the intermediate heat exchangers employed between these two units. The HADWDC-VRHP was found to entail a considerably smaller OC and TUC than the HADWDC. Although the proposed combination is strongly dependent on detailed utility costs and operating conditions, it is likely to result in a substantial reduction in the TAC. These striking outcomes highlight the potential advantages of applying the VRHP to the HADWDC. The insights gained from this study may be considered to be of general significance to further intensification of the HADWDC. Although the combination of the HADWDC and the VRHP can greatly enhance the steady-state performance of the HADWDC-VRHP, the resultant complicated process dynamics and operation problems may hinder its applications. Therefore, it is imperative to examine these issues in great detail in the future.

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APPENDIX A: TAC ESTIMATIONS FOR ALL PROCESS DESIGNS STUDIED (1) The reboiler heat-transfer area (AREB) is given by

5. DISCUSSION The two example systems presented herein show that the inherent features of the HADWDC make it a good candidate for the application of the VRHP, thereby representing effective means of process intensification. The relatively small top-tobottom temperature elevation guarantees the reasonable application of the VRHP to the HADWDC, and the unique process configuration of one top condenser and two bottom reboilers ensures a substantial enhancement of the steady-state economics, i.e., the economic feasibility of the application of the VRHP to the HADWDC. It is worth analyzing the significant difference in the steady-state economics arising from the applications of the VRHP to examples I and II. Although the outcomes are related to the different thermodynamic properties of the separated mixtures to some extent, it is essentially dominated by the sharply different ratio between the prices of electricity and steam, implying that the VRHP is more favorably applied to the HADWDC where relatively high-price steam and relatively low-price electricity have been employed as utilities. In order to gain further insights into the steady-state performance of the HADWDC-VRHP, one needs to make a detailed comparison against the HADWDC with the ratio between the unit prices of electricity and steam as a varying parameter. Because of existence of three major kinds of combinations between the VRHP and the HADWDC, it appears rather cumbersome and time-consuming to conduct the synthesis and design of the HADWDC-VRHP (in comparison with the other heat-pump assisted distillation systems). The studies of the two example systems have

AREB [ft2] =

Q REB UREBΔTREB

(A.1)

where QREB (Btu/h) is the reboiler heat duty, ΔTREB (°F), the temperature driving force, and UREB, the overall heat-transfer coefficient and assumed to be 250 Btu/(h·ft2·°F) here. (2) The condenser heat-transfer area (ACON) is given by A CON [ft2] =

Q CON UCONΔTCON

(A.2)

where QCON (Btu/h) is the condenser heat duty, ΔTCON (°F), the log-mean temperature driving force, and UCON, the overall heattransfer coefficient and assumed to be 150 Btu/(h·ft2·°F) for condensing zone and 50 Btu/(h·ft2·°F) for subcooling zone here. (3) The preheater heat-transfer area (APRE) is given by APRE [ft2] =

Q PRE UPREΔTPRE

(A.3)

where QPRE (Btu/h) is the preheater heat duty, ΔTPRE (°F), the log-mean temperature driving force, and UPRE, the overall heattransfer coefficient and assumed to be 10 Btu/(h·ft2·°F) here. (4) The column length (LC) is given by ⎛ N − 1⎞ ⎟ LC [ft] = 2.4⎜ T ⎝ 0.5 ⎠ 11606

(A.4) DOI: 10.1021/acs.iecr.5b02929 Ind. Eng. Chem. Res. 2015, 54, 11592−11609

Article

Industrial & Engineering Chemistry Research where NT is the total number of stages. (5) The equivalent diameter (De) of the column shell along the dividing wall is given by Ae = ADW,ADC + ADW,ERDC

compressor cost [$] =

(A.11)

where FC = 1.0 for a centrifugal gas compressor. (12) The theoretical horsepower (Hp) for a centrifugal gas compressor is given by

(A.5.1)

⎛ πDe 2 ⎜⎜Ae = , ADW,ADC 4 ⎝ =

πDDW,ADC2 4

, ADW,ERDC

πDDW,ERDC 2 ⎞ ⎟⎟ = 4 ⎠ De =

Vv = 2·Q D·20 (min)

(A.12.3)

(A.13)

⎡ $ ⎤ cooling water cost ⎢ ⎥ ⎣ year ⎦ =

(A.7)

C W ⎛ 1 gal ⎞⎛ 1 ⎞⎛ Q C ⎞⎛ hr ⎞ ⎜ ⎟⎜ ⎟⎜HA ⎟⎜ ⎟ 1000 gal ⎝ lb ⎠⎝ C H ⎠⎝ ΔTD ⎠⎝ year ⎠

(A.14)

where CW is the cooling water price, and ΔTD, the temperature difference between inlet and outlet cooling water, and CH, the specific heat of water. (15) The electricity cost is given by

where M&S is the Marshall and Swift Equipment Cost Index. FC = FmFp = 1 for example I and FC = FmFp = 3.67 for example II. (8) The stage cost is given by

⎡ $ ⎤ ⎛Q ⎞⎛ hr ⎞ electricity cost ⎢ ⎥ = C E⎜ COMP ⎟⎜HA ⎟ ⎝ ε ⎠⎝ year ⎠ ⎣ year ⎦

(A.8)

where FC = Fs + Ft + Fm = 2 for example I and FC = Fs + Ft + Fm = 2.7 for example II. (9) The heat exchanger cost is given by

(A.15)

where CE is the electricity price, QCOMP, the shaft power of the compressor, and ε, the motor efficiency. (16) The entrainer cost is given by

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

⎡ $ ⎤ entrainer cost ⎢ ⎥ ⎣ year ⎦

(A.9)

where FC = (Fd + Fp)Fm = (1.35 + 0) × 1.00 for the reboilers of example I and FC = (Fd + Fp)Fm = (1.35 + 0) × 3.75 for the reboilers of example II, FC = (Fd + Fp)Fm = (1 + 0) × 1 for the condenser of example I and FC = (Fd + Fp)Fm = (1 + 0) × 3.75 for the condenser of example II, FC = (Fd + Fp)Fm = (1.35 + 0) × 1.00 for the preheater. (10) The decanter cost is given by

⎛ kmol ⎞⎟⎛ hr ⎞⎛⎜ $ ⎞⎟ = ⎜FEN ⎟ price ⎜HA ⎝ hr ⎠⎝ year ⎠⎝ kmol ⎠

(A.16)

where FEN is the flow rate of entrainer. NOTATION Variables

A = hypothetical component ACON = heat-transfer area of a condenser, m2 APRE = heat-transfer area of a preheater, m2 AREB = heat-transfer area of a rebolier, m2 B = hypothetical component C = hypothetical component CE = electricity price, $/(kW·h)

⎛ M&S ⎞ 1.066 ⎜ ⎟101.9D L D0.802 D ⎝ 280 ⎠ (2.18 + FC)

Cv

where Cs is the saturated steam price, λV, the latent heat of steam, and HA, the operating hours per year. (14) The cooling water cost is given by

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

⎛ M&S ⎞ 1.55 ⎜ ⎟4.7D LCFC C ⎝ 280 ⎠

Cp

⎡ $ ⎤ Cs ⎛ Q REB ⎞⎛ hr ⎞ steam cost ⎢ ⎥= ⎜ ⎟⎜HA ⎟ ⎣ year ⎦ 1000 lb ⎝ λV ⎠⎝ year ⎠

where VV is the volume of the decanter and the ratio between LD and DD is 3. (7) The column cost is given by

decanter cost [$] =

(A.12.2)

where Pin and Pout are the inlet and outlet pressures, respectively, Q, the vapor flow rate into the compressor, μ, the isentropic coefficient, which can be calculated by heat capacity coefficients, Cp and Cv. (13) The steam cost is given by

(A.6.2)

heat exchanger cost [$] =

⎛ Cp ⎞ ⎛ Cp ⎞ γ=⎜ − 1⎟ /⎜ ⎟ ⎝ Cv ⎠ ⎝ Cv ⎠

(A.5.3)

(A.6.1)

D 2 Vv = π D L D 4

stage cost [$] =

(A.12.1)

μ=

where DDW,ADC and DDW,ERDC are the diameters of the ADC and the ERDC along the dividing wall, respectively, De, the equivalent diameter of the column shell along the dividing wall. (6) The diameter (DD) and length (LD) of a decanter are given by

column cost [$] =

⎤ ⎡⎛ P ⎞ γ ⎛ 3.03 × 10−5 ⎞ out HP = ⎜ ⎟ − 1⎥ ⎟PinQ ⎢⎜ ⎥⎦ ⎢⎣⎝ Pin ⎠ γ ⎠ ⎝

(A.5.2)

DDW,ADC 2 + DDW,ERDC 2

⎛ M&S ⎞ 0.82 ⎜ ⎟658.3Q (2.11 + FC) COMP ⎝ 280 ⎠

(A.10)

where FC = FmFp = 1. (11) The compressor cost is given by 11607


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Industrial & Engineering Chemistry Research Abbreviations

CH = specific heat of water, Btu/(lb·°F) CP = heat capacity coefficient, J·K−1 Cs = saturated steam price, $/kg Cv = heat capacity coefficient, J·K−1 CW = cooling water price, $/m3 CI = capital investment, $ DADW = diameter of the common section above the dividing wall, m DC = diameter of a distillation column, m DD = diameter of a decanter, m DDW,ADC = diameter of the ADC along the dividing wall, m DDW,ERDC = diameter of the ERDC along the dividing wall, m De = equivalent diameter of the column shell along the dividing wall, m FEN = feed flow rate of entrainer, kmol/h HA = operating hours per year, hr/year HP = horsepower, bhp LC = height of a distillation column, m LD = height of a decanter, m NA = location of aqueous phase reflux NADW = number of stages above the dividing wall NDW,ADC = number of stages of the ADC along the dividing wall NDW,ERDC = number of stages of the ERDC along the dividing wall NF = feed location NIHE,ADC = location of the intermediate heat exchanger of the ADC NIHE,ERDC = location of the intermediate heat exchanger of the ERDC OC = operating cost, $/year Q = vapor flow rate into a compressor, kmol/h QCOMP = compressor power, kW QCON = condenser heat duty, kW QD = flow rate into a decanter, kmol/h QUCONS = total utility consumption, kW QIHE,ADC = heat duty of the intermediate heat exchanger of the ADC, kW QIHE,ERDC = heat duty of the intermediate heat exchanger of the ERDC, kW QPRE = preheater heat duty, kW QREB = reboiler heat duty, kW QREB,ADC = reboiler heat duty of the ADC, kW QREB,ERDC = reboiler heat duty of the ERDC, kW QT‑CON = trim-condenser heat duty, kW TIHE,ADC = temperature in the intermediate heat exchanger of the ADC, K TIHE,ERDC = temperature in the intermediate heat exchanger of the ERDC, K ΔTADC = temperature span of the ADC, K ΔTCON = log-mean temperature driving force in a condenser, K ΔTD = temperature difference between inlet and outlet cooling water, K ΔTERDC = temperature span of the ERDC, K ΔTPRE = log-mean temperature driving force in a preheater, K ΔTREB = temperature driving force in a reboiler, K TAC = total annual cost, $/year TUC = total utility consumption, kW UCON = overall heat-transfer coefficient for a condenser, Btu/(h·ft2·°F) UPRE = overall heat-transfer coefficient for a preheater, Btu/(h·ft2·°F) UREB = overall heat-transfer coefficient for a rebolier, Btu/(h·ft2·°F) VV = volume of a decanter, m3

ADC = azeotropic distillation column CHADS = conventional heterogeneous azeotropic distillation system CR = compression ratio CYH = cyclohexane ERDC = entrainer recovery distillation column HADWDC = heterogeneous azeotropic dividing-wall distillation column HADWDC-VRHP(ADC) = process design with the VRHP connected to the ADC HADWDC-VRHP(ADC-ERDC) = process design with the VRHP connected to the ADC and the ERDC HADWDC-VRHP(ERDC) = process design with the VRHP connected to the ERDC IPA = isopropyl alcohol PIE = pyridine TUE = toluene VRHP = vapor recompression heat pump I-HADWDC-VRHP(ADC-ERDC) = a typical design with the VRHP connected to the ADC and the ERDC for example I II-HADWDC-VRHP(ADC-ERDC) = a typical design with the VRHP connected to the ADC and the ERDC for example II II-HADWDC-VRHP(ERDC) = a typical design with the VRHP connected to the ERDC for example II Greek Letters

■

βL = liquid split ratio βpbt = payback time, year ε = motor efficiency εL = characteristic value of the HADWDC εU = characteristic value of the HADWDC μ = isentropic coefficient λV = steam latent heat, Btu/lb

AUTHOR INFORMATION

Corresponding Authors

*Phone: +86-10-64437805. Fax: +86-10-64437805. E-mail: [email protected] (K. Huang). *Phone: +886-3-5715131 ext. 33624. Fax: +886-3-5721694. E-mail: [email protected] (S.-J. Wang). 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) and Ministry of Science and Technology of Taiwan under Grant No. MOST 103-2221-E-007-138.

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