Thermally Coupled Reactive Distillation System for the Separations of

Article pubs.acs.org/IECR

Thermally Coupled Reactive Distillation System for the Separations of Cyclohexene/Cyclohexane Mixtures Jieping Yu, Li Shi, Yang Yuan, Haisheng Chen, Shaofeng Wang, and Kejin Huang* College of Information Science and Technology, Beijing University of Chemical Technology, Beijing 100029, People’s Republic China ABSTRACT: Because of the very close boiling points of cyclohexene and cyclohexane, their separation is extremely difficult with conventional distillation systems and reactive separation with water as reactive entrainer offers great economic incentives. In the current work, the synthesis and design of a flow-sheet with two reactive distillation columns (i.e., the hydration and dehydration reactive distillation columns) in series (FSTRDC) is first conducted subject to the minimization of total annual cost (TAC), and this leads to a process design with an excessive use of water. Although the process design facilitates the hydration of cyclohexene into cyclohexanol, it gives rise to a serious remixing effect in the hydration reactive distillation column and poses an unfavorable effect to the decomposition of cyclohexanol into cyclohexene and water in the dehydration reactive distillation column. For the suppression of these deficiencies, the hydration reactive distillation column was then modified to be heated directly with a vapor flow from the reboiler of the dehydration reactive distillation column, and this gives rise to a novel thermally coupled reactive distillation system (TCRDS). The mass and thermal coupling between these two reactive distillation columns involved eliminates completely the remixing effect and reduces greatly the excessive use of water and enables the TCRDS to require substantially less capital investment (CI), operating cost (OC), and TAC than the FSTRDC. The implementation of the TCRDS in terms of a dividing-wall distillation column is finally highlighted and found to lead to further reductions in the CI and TAC. These striking outcomes demonstrate the great significance of including effective mass and thermal coupling to the two reactive distillation columns in series, reactively separating close boiling binary mixtures.

1. INTRODUCTION Cyclohexene is an important raw material for synthesizing polymers, intermediates of pesticides and pharmaceuticals, and cyclohexanol, an important precursor for the synthesis of Nylon-6. It is estimated that more than one million tons of cyclohexene are produced each year worldwide.1 Under industrial operating conditions, cyclohexene is usually produced with the existence of a certain amount of cyclohexane, and this poses a serious limitation on its use as an important raw material. It is, therefore, necessary to remove cyclohexane from cyclohexene, and this actually constitutes a major purification step toward the production of cyclohexene. Because of the very close boiling points between cyclohexene and cyclohexane, their separation is extremely difficult with conventional distillation systems, and the derivation of energy-efficient processes thus is highly desirable and challengeable. So far, only a few studies have been reported on the separations of cyclohexene/cyclohexane mixtures. In 2002, Steyer et al. advocated reactive separations of cyclohexene/cyclohexane mixtures with water as reactive entrainer and proposed a conceptual design of a flow-sheet with two reactive distillation columns in series (FSTRDC) as shown in Figure 1a.2 Because water reacts with cyclohexene to form cyclohexanol and cyclohexanol has a much higher boiling point than cyclohexene and cyclohexane, the FSTRDC offers a great potential of improvement in steady-state economics despite the fact that two reactive distillation columns were involved. The first reactive distillation column worked to directly hydrate cyclohexene into cyclohexanol and was termed the hydration reactive distillation column (HRDC; to be more specific, it is, in most cases, referred to as the FSTRDC−HRDC in the © XXXX American Chemical Society

current work). Cyclohexane was extracted from the top and cyclohexanol compounded with the unconverted cyclohexene from the bottom. The second reactive distillation column worked to dehydrate cyclohexanol into cyclohexene and water and was termed the dehydration reactive distillation column (DRDC; to be more specific, it is, in most cases, referred to as the FSTRDC−DRDC in the current work). Cyclohexene and water were extracted from the top and separated further in the decanter. (Because cyclohexanol escapes easily with water from the bottom of the FSTRDC−DRDC, no water was considered to be withdrawn from there in the current work.) Recently, Qiu et al. proposed a reactive distillation column for the direct hydration of cyclohexene into cyclohexanol by adding further a cosolvent of 1,4-dioxane.3 Although the reaction conversion could be greatly enhanced, the prevention of cyclohexane accumulation within the process made the scheme unfeasible to work even with the addition of a purge system. Chien and his co-workers noticed the great impact of feed ratio between cyclohexene and water to reaction conversion and worked out an efficient system including one reactive distillation column plus two decanters and one follow-up stripper.4 Though these two studies focused merely on the direct hydration of cyclohexene into cyclohexanol, they highlighted the great importance of maintaining high reaction conversion in process synthesis and design and had strong implications to the Received: September 20, 2015 Revised: November 30, 2015 Accepted: December 17, 2015

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DOI: 10.1021/acs.iecr.5b03526 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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Industrial & Engineering Chemistry Research

Figure 1. FSTRDC versus TCRDS: (a) FSTRDC and (b) TCRDS.

In the current work, the synthesis and design of the FSTRDC is first conducted subjected to the minimization of total annual cost (TAC), a summation of operating cost (OC) and annualized capital investment by a given payback period. The resultant process flowsheet is analyzed and its drawbacks are pointed out. A novel thermally coupled reactive distillation system (TCRDS) is then proposed and designed by effectively including mass and thermal coupling between the hydration and dehydration reactive distillation columns involved. Indepth comparisons are made between the FSTRDC and TCRDS and a dividing-wall distillation-column-based implementation of the TCRDS is highlighted. Finally, some insightful remarks on the reactive separations of close boiling binary mixtures are given in the last section.

development of the FSTRDC. Because the amount of water may also affect the dehydration of cyclohexanol into cyclohexene and water, a complicated interplay is quite likely to occur between the FSTRDC−HRDC and FSTRDC−DRDC involved in the FSTRDC. Note that the FSTRDC is actually a train of the hydration and dehydration reactive distillation columns, the drawbacks of conventional distillation systems, for instance, the high degree of irreversibility in mass and heat transfer and remixing effect, are also likely to happen and present consequently negative effect to its steady-state economics. Therefore, it is imperative to study in great detail the synthesis and design of the FSTRDC and on which some energy-efficient schemes can be derived. B


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2. DESIGN AND ANALYSIS OF THE FSTRDC 2.1. Problem Description. The kinetics of the hydration of cyclohexene into cyclohexanol and the dehydration of cyclohexanol into cyclohexene and water is shown

(ANE), respectively. There are four components involved in the mixtures separated, that is, cyclohexene, water, cyclohexanol, and cyclohexane. The commercial software ASPEN PLUS is employed to predict system performance and the NRTL (nonrandom two-liquid) thermodynamic model to estimate the vapor−liquid and vapor−liquid−liquid equilibrium relationships. All the relevant model parameters have been taken from Steyer and Sundmacher and reproduced in Table 2 for quick reference.7 Because the vapor−liquid phase transfers in both forward and backward reactions are accompanied possibly by liquid phase splitting in the FSTRDC−HRDC and FSTRDC−DRDC, phase stability tests need to be conducted stage-by-stage in process simulation. In the current study, the minimization of the TAC (with a payback period of 3 years) is taken as the objective function for process synthesis and design. While the operating cost includes steam cost, cooling water cost, and catalyst cost, the annualized capital investment consists of column shells cost, column stages cost, condensers cost, and reboilers cost. The prices of steam and cooling water are taken from Seider et al. and the formulas for cost estimations are adopted from Douglas (c.f., Appendix A).8,9 In particular, water makeup cost and decanter (whose heat duty is zero) cost are assumed to be negligible as compared with the other operating and equipment costs. The Marshall and Swift Equipment Cost Index (M&S) is assumed to be a value of 1533.3 in the current work. 2.2. Synthesis and Design of the FSTRDC. Because the catalyst Amberlyst 15 can only operate below 393.15 K, a hightemperature constraint must be imposed simultaneously onto the two reactive sections of the FSTRDC−HRDC and FSTRDC−DRDC during process synthesis and design. Both the FSTRDC−HRDC and FSTRDC−DRDC are operated under atmospheric pressure with a pressure drop of 689.01 N/ m2 per stage. The normal boiling points of cyclohexane, cyclohexene, water, and cyclohexanol are 353.93, 356.03, 373.17, and 433.99 K, respectively, and these render an appropriate flexibility to the arrangement of two reactive sections in the FSTRDC−HRDC and FSTRDC−DRDC, respectively. For the FSTRDC−HRDC, its top product is set near to the water-cyclohexane azeotrope (at 342.56 K with a composition of 30.0 mol % of water and 70.0 mol % of cyclohexane) so that phase splitting can occur naturally in its decanter. Similarly, for the FSTRDC−DRDC its top product is set close to water−cyclohexene azeotrope (at 343.86 K with a composition of 31.7 mol % of water and 68.3 mol % of cyclohexene) so that phase splitting can occur naturally in its decanter. The on-spec products of cyclohexane and cyclohexene can thus be obtained from the FSTRDC−HRDC and FSTRDC−DRDC, respectively. Figure 2 shows a systematic procedure derived for the synthesis and design of the FSTRDC. It employs the commonly used grid-search method to sequentially determine

The forward and backward reactions are exothermic and endothermic, respectively. They are heterogeneously catalyzed by solid catalyst Amberlyst 15 with a density of 770 kg/m3.5 The kinetic expression in Langmuir−Hinshelwood form is written as6 ENE H 2O ⎛ K ads K ads r = ⎜⎜mcatk het H 2O ENE NOL 2 (1 + αENEK ads + αH2OK ads ) + αNOLK ads ⎝

⎞⎛ ⎞ ⎟⎜⎜αENEαH O − αNOL ⎟⎟ ⎟ 2 Keq ⎠ ⎠⎝

(2)

where α is the liquid activity, and Kads, the adsorption constant, which are 19.989, 0.056839, and 0.77324, respectively, for water, cyclohexene (ENE), and cyclohexanol (NOL). The reaction rate constant khet and reaction equilibrium constant Keq are given as ⎛ −93687 ⎞⎛ mol ⎞ ⎟ ⎟⎜ k het = 7.7083 × 1012exp⎜ ⎝ RT ⎠⎜⎝ kg s ⎟⎠ cat

(3)

⎡ ⎛1 1 ⎞⎤ ⎟⎥ K eq =4.2907exp⎢3389.38⎜ − ⎝T ⎣ 298.15 ⎠⎦

(4)

Table 1 lists the operating conditions and design specifications for the FSTRDC to be developed. A cycloTable 1. Operating Conditions and Design Specifications parameter

value

condenser pressure (N/m2) stage pressure drop (N/m2) feed flow rate (mol/s) feed thermal condition feed composition (mole fraction) product specification (mole fraction)

q ANE ENE ANE ENE

101325 689.01 27.78 1.0 (saturated liquid) 0.200 0.800 0.990 0.990

hexene/cyclohexane mixture with a flow rate of 27.78 mol/s and a composition of 0.8/0.2 is to be separated with the product specifications of 0.99 for cyclohexene and cyclohexane Table 2. Thermodynamic Model Parameters component i

component j

aij

aji

bij

bji

cij

ENE ENE EBE NOL NOL ANE

NOL H2O ANE H2O ANE H2O

0 0 0 0 0 0

0 0 0 0 0 0

429.305 1705.00 5.10961 160.782 2.39765 2122.93

0.115809 2609.45 7331.87 1318.19 489.733 3012.81

0.802522 0.267206 0.831053 0.359706 0.993301 0.258799

C


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Figure 2. Procedure for the synthesis and design of the FSTRDC and TCRDS.

all the structural and operating variables involved.10 Because the feed flow rate of water represents the most important decision variable that can affect simultaneously the operation of the FSTRDC−HRDC and FSTRDC−DRDC, it is deliberately designated here to be the variable determined in the outmost iteration loop as highlighted with the blocks shadowed. The

resultant optimum process design of the FSTRDC is shown in Figure 3. Although the FSTRDC−HRDC contains 97 stages with the processed mixture of cyclohexene/cyclohexane fed onto stage 80, the FSTRDC−DRDC accommodates 37 stages with the bottom effluent of the FSTRDC−HRDC fed onto stage 18 (here, the top condenser is designated as the first stage D


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Figure 3. Optimum design of the FSTRDC.

and the bottom reboiler as the last stage). The reactive section runs from stage 2 to stage 95 in the former and from stage 6 to stage 36 in the latter. The reboiler heat duties are 5708.02 × 103 W and 3543.31 × 103 W, respectively, for the FSTRDC− HRDC and FSTRDC−DRDC, resulting in a sum of 9251.33 × 103 W. Figure 4a depicts the relationship between the TAC and feed flow rate of water (including here the water makeup) and the optimal feed flow rate of water is found to be 30.56 mol/s, being 37.5% in excess of cyclohexene. The temperature profiles of the FSTRDC−HRDC and FSTRDC−DRDC are shown in Figures 5a and b, respectively. Note that the bottom temperatures of the FSTRDC−HRDC and FSTRDC−DRDC are 385.69 and 379.43 K, respectively (highlighted here with two dashed lines), and both are well below 393.15 K, implying that the catalyst Amberlyst 15 can work normally in the two reactive sections. Because low-pressure steam can maintain a minimum temperature driving force of 10 K in the two reboilers of the FSTRDC−HRDC and FSTRDC−DRDC, it is employed here as hot utility for the FSTRDC. 2.3. Steady-State Analysis of the FSTRDC. In Figures 6a and 7a, the profiles of liquid compositions are delineated, respectively, for the FSTRDC−HRDC and FSTRDC−DRDC of the resultant FSTRDC. As can readily be noted from Figure 6a, there appears to be a severe remixing phenomenon occurring to water at the bottom of the FSTRDC−HRDC (indicated here with a dashed circle), and it is likely to lower considerably the steady-state performance of the FSTRDC. In spite of the fact that the employment of excessive water is favorable to the hydration of cyclohexene into cyclohexanol in the FSTRDC−HRDC, the remixing phenomenon can yield adverse steady-state economics because the more water is employed the greater its detrimental effect becomes. Similar to its conventional counterparts, such a drawback should also be considered to be aroused by the connection of the FSTRDC−

Figure 4. TAC versus water feed flow rate of the FSTRDC and TCRDS.

HRDC and FSTRDC−DRDC in series as indicated by the dashed rectangular in Figure 1a.11,12 In the FSTRDC−DRDC (c.f., Figure 7a), water is found to show a very high presence at E


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Figure 5. Temperature profiles of the FSTRDC: (a) FSTRDC−HRDC and (b) FSTRDC−DRDC.

Figure 6. Liquid composition profiles of the FSTRDC and TCRDS: (a) FSTRDC−HRDC and (b) TCRDS−HRDC.

the top and bottom of the reactive section with its compositions almost near to one. Although the unique

phenomenon stems from the operating characteristics of the FSTRDC−DRDC and the physicochemical properties of the F


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Figure 7. Liquid composition profiles of the FSTRDC and TCRDS: (a) FSTRDC−DRDC and (b) TCRDS−DRDC.

Figure 8. Optimum design of the TCRDS.

reacting mixture separated, it is also closely related to the excessive use of water in the FSTRDC−HRDC. Because the bottom water must move to the top decanter by way of the reactive section and is recycled further to the FSTRDC− HRDC, it poses certainly an unfavorable effect to the

dehydration of cyclohexanol into cyclohexene and water in the FSTRDC−DRDC and this inevitably diminishes the steady-state performance of the FSTRDC. This represents essentially the second most serious drawback next to the remixing phenomenon outlined above. G


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Figure 9. Temperature profiles of the TCRDS: (a) TCRDS−HRDC and (b) TCRDS−DRDC.

3. DERIVATION OF A NOVEL THERMALLY COUPLED REACTIVE DISTILLATION SYSTEM

4. DESIGN AND ANALYSIS OF THE TCRDS 4.1. Synthesis and Design of the TCRDS. In terms of the systematic procedure shown in Figure 2, the synthesis and design of the TCRDS are conducted under the same operating conditions and product specifications as the FSTRDC and result in the optimum process design as shown in Figure 8. Although the TCRDS−HRDC contains 107 stages with its reactive section from stage 2 to stage 105, the TCRDS−DRDC possesses 42 stages with its reactive section from stage 7 to stage 41. The processed mixture of cyclohexene/cyclohexane is now fed onto stage 93 of the TCRDS−HRDC and its bottom effluent fed onto stage 16 of the TCRDS−DRDC. The reboiler heat duty becomes 7836.11 × 103 W and much lower than the sum of two reboiler heat duties of the FSTRDC, implying a great enhancement in thermodynamic efficiency by the inclusion of mass and thermal coupling between the TCRDS−HRDC and TCRDS−DRDC involved. Figure 4b depicts the relationship between the TAC and feed flow rate of water (including here the water makeup). The optimal feed flow rate of water is now reduced to 26.39 mol/s (as compared with 30.56 mol/s of the FSTRDC) and it is certainly due to the inclusion of mass and thermal coupling between the TCRDS− HRDC and TCRDS−DRDC that causes a more effective distribution of water within the TCRDS than within the FSTRDC. The temperature profiles of the TCRDS−HRDC and TCRDS−DRDC involved in the TCRDS are shown in Figure 9a and b, respectively. Although almost no differences can be found between the FSTRDC−HRDC and TCRDS− HRDC, relatively significant differences can be found between the FSTRDC−DRDC and TCRDS−DRDC. The mass and thermal coupling between the TCRDS−HRDC and TCRDS−

As per the above steady-state analysis, modifications are suggested to be made to the FSTRDC, and this gives rise to a novel thermally coupled reactive distillation system (TCRDS) as shown in Figure 1b. For the suppression of the remixing phenomenon at the bottom of the TCRDS−HRDC, mass and thermal coupling has now been introduced between the TCRDS−HRDC and TCRDS−DRDC, and this must be made in good accordance with their inherent operating characteristics. Because water exhibits a very high presence at the bottom of the TCRDS−DRDC, a vapor flow is reasonable to be withdrawn from the reboiler of the TCRDS−DRDC and employed to heat directly the bottom of the TCRDS−HRDC. Such a process modification involves three major advantages. First, a certain amount of water is recycled directly from the bottom of the TCRDS−DRDC to the TCRDS−HRDC and this favors the hydration of cyclohexene into cyclohexanol in the TCRDS−HRDC. It works also as the main thrust to remove the remixing phenomenon at the bottom of the TCRDS−HRDC. Second, with a certain amount of water removed from the bottom of the TCRDS−DRDC, the amount of water that must be moved through the reactive section to the decanter is reduced, thereby lessening the unfavorable effect to the dehydration of cyclohexanol into cyclohexene and water in the TCRDS−DRDC. Third, the employment of the bottom reboiler of the TCRDS−HRDC is no longer necessary, and this helps to lower the capital investment as compared with the FSTRDC. H


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Industrial & Engineering Chemistry Research Table 3. Summary of All Process Designs Studied configurations column cost (103 $) reboiler cost (103 $) condenser cost (103 $) total ci (103 $) steam cost (103 $/year) cooling water cost (103 $/year) catalyst cost (103 $/year) total OC (103 $/year) total reboiler heat duty (103 W) total TAC (103 $/year)

SDC 11552.26 980.59 760.40 13293.24 1213.89 44.08 00.00 1257.96 13280.05 5689.04

(+201.86%)

(+9.20%) (+43.55%) (+117.15%)

FSTRDC 2744.65 698.97 960.12 4403.74 845.63 31.85 274.47 1151.95 9251.33 2619.86

(0.00%)

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

TCRDS 2706.34 554.68 793.18 4054.20 716.27 26.62 246.34 989.23 7836.11 2340.63

TCRDS−DWDC

(−7.94%)

(−14.13%) (−15.30%) (−10.66%)

2602.41 554.68 793.18 3950.27 716.27 26.62 246.34 989.23 7836.11 2305.98

(−10.30%)

(−14.13%) (−15.30%) (−11.98%)

Figure 10. Implementation of the TCRDS in terms of the framework of the DWDC.

4.2. Steady-State Analysis of the TCRDS. The profiles of liquid compositions of the TCRDS−HRDC and TCRDS− DRDC involved in the TCRDS are shown in Figures 6b and 7b, respectively. It can readily be seen that the remixing effect has completely been eliminated at the bottom of the TCRDS− HRDC (highlighted here also with a dashed circle in Figure 6b). Despite the fact that the mass and thermal coupling between the TCRDS−HRDC and TCRDS−DRDC involved leads to great differences in process designs between the FSTRDC and TCRDS, they exhibit actually quite similar liquid composition profiles, implying quite analogous functions toward the separations of cyclohexene/cyclohexane mixtures.

DRDC involved should be responsible for such relatively significant differences. Note that the bottom temperatures of the TCRDS−HRDC and TCRDS−DRDC are 383.93 and 389.14 K, respectively (highlighted here with two dashed lines), and both are below 393.15 K, implying that the catalyst Amberlyst 15 can work normally in the two reactive sections. The bottom temperature of the TCRDS−DRDC still permits the use of low-pressure steam for reboiler heating. In comparison with the FSTRDC, thus no adverse steady-state economics at all has been aroused by the inclusion of mass and thermal coupling between the TCRDS−HRDC and TCRDS− DRDC involved. I


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Industrial & Engineering Chemistry Research 4.3. FSTRDC versus TCRDS. The summary of all the process designs studied are presented in Table 3. The TCRDS appears to be not only much more thermodynamic efficient but also much more cost-effective than the FSTRDC. The former secures, respectively, reductions in the CI, OC, and TAC by 7.94%, 14.13%, and 10.66%, as compared with the latter. The striking outcomes highlight the reasonability and effectiveness of including mass and thermal coupling between the TCRDS− HRDC and TCRDS−DRDC involved.

5. IMPLEMENTATION OF THE TCRDS IN TERMS OF DIVIDING-WALL DISTILLATION COLUMNS With reference to the topological structure of the TCRDS, further process intensification can be attempted and a kind of implementation based on a dividing-wall distillation column (DWDC) is derived as shown in Figure 10 (termed the TCRDS−DWDC, hereinafter, in the current work). Note that the dividing wall runs from the stage 68 to the bottom of the TCRDS−DWDC−HRDC. Unlike the conventional DWDC, the TCRDS−DWDC−DRDC is isolated from the TCRDS− DWDC−HRDC at the top edge of the dividing wall and no liquid splitting is allowed to occur there. At the bottom edge of the dividing wall, the vapor generated at the bottom reboiler splits into two flows that go to the TCRDS−DWDC−HRDC and TCRDS−DWDC−DRDC, respectively. The products of cyclohexane and cyclohexene are extracted from the decanters at the top of the TCRDS−DWDC−HRDC and TCRDS− DWDC−DRDC, respectively. As shown in Table 3, the TCRDS−DWDC leads to further reductions of the CI and TAC by 2.36% and 1.32%, respectively (with reference to the FSTRDC), but it shared the same OC with the TCRDS.

Figure 11. Optimum design of the SDC.

demonstrating the great potential of process intensification in the separations of the cyclohexene/cyclohexane mixtures.

7. CONCLUSIONS Although the FSTRDC permits the separations of cyclohexene/ cyclohexane mixtures with water as the reactive entrainer, it has two inherent drawbacks that considerably worsen its steadystate performance. One is the remixing effect occurred to water at the bottom of the FSTRDC−HRDC. The other is the accumulation of water at the bottom of the FSTRDC−DRDC that must be driven up through the reactive section to the decanter for phase splitting. An unfavorable impact, thus, is given to the dehydration of cyclohexanol into cyclohexene and water. The detrimental impacts of these two drawbacks can be substantially amplified with the excessive use of water to enhance the conversion of cyclohexene into cyclohexanol. By the inclusion of mass and thermal coupling between the TCRDS−HRDC and TCRDS−DRDC (i.e., introducing a certain amount of vapor from the reboiler of the latter to heat directly the bottom of the former), the resultant TCRDS is likely to get rid of completely the remixing effect and reduce, to a great extent, the excessive use of water, thereby securing considerable reductions in capital investment, operating cost, and total annual cost in comparison with the FSTRDC. The unique connection between the TCRDS−HRDC and TCRDS−DRDC allows one to implement the TCRDS in terms of the framework of the DWDC, and this can lead to further reductions in capital investment and total annual cost. The detailed outcomes of the current study have demonstrated the feasibility and effectiveness of the proposed philosophy for process intensification. The proposed philosophy for process synthesis and design is also considered to be of great implications to the reactive separations of other close boiling binary mixtures, including, for example, the separations of C4 mixtures with methanol as reactive entrainer,13 and the separations of m-xylene and p-

6. DISCUSSION For the separations of cyclohexene/cyclohexane mixtures with water as reactive entrainer in the FSTRDC, although the low reaction rates of the hydration and dehydration reactions stand for the main reasons that result in the requirement of very large CI, OC, and TAC, the remixing phenomenon at the bottom of the FSTRDC−HRDC and the excessive employment of water to enhance the conversion of cyclohexene into cyclohexanol are still two great drawbacks that lower substantially the steadystate performance. With the inclusion of mass and thermal coupling between the TCRDS−HRDC and TCRDS−DRDC involved, not only is the remixing phenomenon completely removed at the bottom of the former but also the amount of water used is substantially diminished (and so is the accumulation of water at the bottom of the latter). This is why the TCRDS can lead to substantial reductions in the CI, OC, and TAC as compared with the FSTRDC. The case study has also demonstrated the great importance of following the inherent operation characteristics of the FSTRDC to include mass and thermal coupling between the TCRDS−HRDC and TCRDS−DRDC involved. In order to explore the impacts of employing reactive distillation columns for the separations of cyclohexene/ cyclohexane mixtures, we also conducted the design of a simple distillation column (SDC), and the resultant SDC is shown in Figure 11. Note that it involves totally 300 stages with a reboiler heat duty up to 13280.05 × 103 W. As shown in Table 3, the resultant SDC requires considerably large CI, OC, and TAC as compared with the FSTRDC and the TCRDS, J


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Industrial & Engineering Chemistry Research xylene mixtures with sodium p-xylene as reactive entrainer.14,15 Future studies should also be centered on the examination of the dynamics and operation of the proposed TCRDS and TCRDS−DWDC.

(3) The heat exchanger cost is given by heat exchanger cost[$] ⎛ M&S ⎞ 0.65 ⎟101.3A (2.29 + FC) =⎜ ⎝ 280 ⎠

■

APPENDIX A: TAC ESTIMATIONS FOR ALL PROCESS DESIGNS STUDIED The TAC is given by TAC = OC +

CI β

where FC = (Fd + Fp)Fm = (1.35 + 0) × 3.75 for reboiler, and FC = (Fd + Fp)Fm = (1 + 0) × 3.75 for condenser. (4) The steam cost is given by ⎡ $ ⎤ Cs ⎛ Q R ⎞⎛ hr ⎞ steam cost⎢ ⎥= ⎜ ⎟⎜8150 ⎟ year ⎠ ⎣ year ⎦ 1000lb ⎝ λV ⎠⎝

(A.1)

where OC is the total operating cost, CI, the total installed capital investment, and β, the payback period. The equipment are sized in the following manner. (1) The heat-transfer area for reboiler (AR) is given by AR [ft2] =

(A.9)

where Cs ($) is the saturated steam price, and λV (Btu/lb), the latent heat of steam. (5) The cooling water cost is given by

QR UR ΔTR

(A.2)

⎡ $ ⎤ cooling water cost⎢ ⎥ ⎣ year ⎦

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

A C[ft ] =

=

(A.10)

⎛ ⎡ $ ⎤ $ ⎞⎛ 4 ⎞ catalyst cost⎢ ⎥ = (M kg)⎜price ⎟⎜ ⎟ kg ⎠⎝ year ⎠ ⎣ year ⎦ ⎝

(A.3)

where QC (Btu/h) is the heat duty of condenser, ΔTC (°F), the log-mean temperature driving force, and UC, the overall heattransfer coefficient, which is assumed to be 150 Btu/(hr·ft2·°F) for cooling water. (3) The column length (LC) is given by LC[ft] = 2.0NT

$0.03 ⎛ 1gal ⎞⎛ Q C ⎞⎛ hr ⎞ ⎜ ⎟⎜ ⎟⎜8150 ⎟ 1000gal ⎝ 8.34lb ⎠⎝ 30 ⎠⎝ year ⎠

(6) The catalyst cost is given by

QC UCΔTC

(A.8)

■

(A.11)

AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. Phone: +86 10 64437805. Fax: +86 10 64437805.

(A.4)

Notes

The authors declare no competing financial interest.

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

D12 + D22

■

ACKNOWLEDGMENTS We thank National Natural Science Foundation of China (21076015, 21176015, 21376018, and 21576014) for their financial support.

■

(A.5)

where D1 and D2 are the diameters of the left and right sides along the dividing wall. The capital and operating costs are calculated with the following expressions. (1) The column cost is given by column cost[$] =

⎛ M&S ⎞ 1.066 ⎜ ⎟101.9D C ⎝ 280 ⎠ LC0.802(2.18 + FC)

(A.6)

where M&S indicates the Marshall and Swift Equipment Cost Index, and FC = FmFp = 3.67. (2) The stage cost is given by stage cost[$] =

⎛ M&S ⎞ 1.55 ⎜ ⎟4.7D C LCFC ⎝ 280 ⎠

(A.7)

where FC = Fs + Ft + Fm = 1 + 1.8 + 1.7 = 4.5. K

NOTATION A = heat-transfer area, m2 Cs = saturated steam price, $ D = diameter, m or distillate product, mol/s De = diameter, m F = feed flow rate, mol/s L = liquid flow rate, mol/s LC = column length, m M = mass of catalyst, kg NT = total number of stages QC = condenser heat duty, W Qreb = reboiler heat duty, W QR = reboiler heat duty, W RR = reflux flow rate, mol/s T = temperature, K Tbot = temperature at bottom, K ΔTC = temperature driving force, K ΔTR = temperature driving force, K UC = overall heat-transfer coefficient, Btu/(hr·ft2·°F) DOI: 10.1021/acs.iecr.5b03526 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

Article

Industrial & Engineering Chemistry Research UR = overall heat-transfer coefficient, Btu/(hr·ft2·°F) V = vapor flow rate, mol/s V1 = vapor flow rate into the TCRDS−HRDC or the TCRDS−DWDC−HRDC V2 = vapor flow rate into the TCRDS−DRDC or the TCRDS−DWDC−DRDC

(11) 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. (12) 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. (13) Qi, Z.; Sundmacher, K.; Stein, E.; Kienle, A.; Kolah, A. Reactive Separation of Isobutene from C4 Crack Fractions by Catalytic Distillation Processes. Sep. Purif. Technol. 2002, 26, 147. (14) Terrill, D.; Sylvestre, L.; Doherty, M. F. Separation of Closely Boiling Mixtures by Reactive Distillation. 1. Theory. Ind. Eng. Chem. Process Des. Dev. 1985, 24, 1062. (15) Cleary, W.; Doherty, M. F. Separation of Closely Boiling Mixtures by Reactive Distillation. 2. Experiments. Ind. Eng. Chem. Process Des. Dev. 1985, 24, 1071.

Abbreviations

ANE = cyclohexane CI = total installed capital investment, $ DRDC = dehydration reactive distillation column DWDC = dividing-wall distillation column ENE = cyclohexene FSTRDC = a flow-sheet with two reactive distillation columns in series HRDC = hydration reactive distillation column NOL = cyclohexanol OC = total operating cost, $/year SDC = simple distillation column TAC = total annual cost, $/year TCRDS = thermally coupled reactive distillation system TCRDS−DWDC = thermally coupled reactive distillation system in terms of a dividing-wall distillation column Greek Letters

β = payback period, year βv = vapor split ratio λV = latent heat of steam, Btu/lb Subscripts

bot = bottom C = condenser R = reboiler WM = water makeup

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REFERENCES

(1) Takamatsu, Y.; Kaneshima, T. Process for the Preparation of Cyclohexanol. U.S. Patent 6,552,235 B2, 2003. (2) Steyer, F.; Qi, Z.; Sundmacher, K. Synthesis of Cylohexanol by Three-Phase Reactive Distillation: Influence of Kinetics on Phase Equilibria. Chem. Eng. Sci. 2002, 57, 1511. (3) Qiu, T.; Kuang, C. H.; Li, C. G.; Zhang, X. W.; Wang, X. D. Study of Feasibility of Reactive Distillation Process for the Direct Hydration of Cyclohexene to Cyclohexanol Using a Cosolvent. Ind. Eng. Chem. Res. 2013, 52, 8139. (4) Chen, B. C.; Yu, B. Y.; Lin, Y. L.; Huang, H. P.; Chien, I. L. Reactive Distillation Process for Direct Hydration of Cyclohexene to Produce Cycohexanol. Ind. Eng. Chem. Res. 2014, 53, 7079. (5) Lee, H. Y.; Lee, Y. C.; Chien, I. L.; Huang, H. P. Design and Control of a Heat-Integrated Reactive Distillation System for the Hydrolysis of Methyl Acetate. Ind. Eng. Chem. Res. 2010, 49, 7398. (6) Steyer, F.; Sundmacher, K. Cyclohexanol Production via Esterification of Cyclohexene with Formic Acid and Subsequent Hydration of the Esters Reaction Kinetics. Ind. Eng. Chem. Res. 2007, 46, 1099. (7) Steyer, F.; Sundmacher, K. VLE and LLE Data for the System Cyclohexane + Cyclohexene + Water + Cyclohexanol. J. Chem. Eng. Data 2004, 49, 1675. (8) Seider, W. D.; Seader, J. D.; Lewin, D. R.; Widagdo, S. Product and Process Design: Principles Synthesis, Analysis, and Evaluation; Wiley: Hoboken, NJ, 2010. (9) Douglas, J. M. Conceptual Design of Chemical Processes; McGrawHill: New York, 1988. (10) Yu, J.; Wang, S. J.; Huang, K.; Yuan, Y.; Chen, H.; Shi, L. Improving the Performance of Extractive Dividing-Wall Columns with Intermediate Heating. Ind. Eng. Chem. Res. 2015, 54, 2709. L


Thermally Coupled Reactive Distillation System for the Separations of

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