Unique Design Considerations for Maximum-Boiling Azeotropic

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Unique Design Considerations for Maximum-Boiling Azeotropic Systems via Extractive Distillation: Acetone/Chloroform Separation Yen-Hsiang Wang, and I-Lung Chien Ind. Eng. Chem. Res., Just Accepted Manuscript • DOI: 10.1021/acs.iecr.8b03125 • Publication Date (Web): 29 Aug 2018 Downloaded from http://pubs.acs.org on August 30, 2018

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Industrial & Engineering Chemistry Research

Paper submitted for publication in Ind. Eng. Chem. Res.

Unique Design Considerations for Maximum-Boiling Azeotropic Systems via Extractive Distillation: Acetone/Chloroform Separation

Yen-Hsiang Wang and I-Lung Chien*

Department of Chemical Engineering National Taiwan University Taipei 10617, Taiwan

Revised: August 11, 2018

*

Corresponding author. I-Lung Chien, Tel: +886-3-3366-3063; Fax: +886-2-2362-3040; E-mail: [email protected]

1 Environment ACS Paragon Plus

Industrial & Engineering Chemistry Research 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60

ABSTRACT If two dissimilar molecules exhibit attraction behavior, a maximum-boiling azeotrope can be formed. There are very few papers in open literature studying the design of extractive distillation system for separating of such systems. The design flowsheet of the two-column system is exactly the same as the ones for minimum-boiling azeotrope systems. However, because of different topology of ternary diagram for the maximum-boiling system, unique design consideration should be taken to properly design such separation system. A demonstrating example of separating acetone and chloroform via extractive distillation is presented in this paper. A distillation boundary will be formed by adding a heavy entrainer into the system. Depending on the degree of curvature of distillation boundary, feasible minimum entranier-to-feed ratio can be estimated with given product purity specifications. This minimum ratio should be considered as another screening tool for selecting suitable entrainer of maximum-boiling azeotrope systems. An optimal design flowsheet by adding a newly proposed N-methyl-2-pyrrolidone as suitable heavy entrainer is developed for this separation system. Significant savings of total annual cost and energy requirement can be obtained with this newly proposed extractive distillation system as compared to that of the two other published systems in open literatures using either dimethyl sulfoxide or ethylene glycol as entrainer. A simple heat-integration scheme is also proposed to further save energy of this proposed system. Keywords: maximum-boiling azeotrope, extractive distillation, acetone/chloroform, heavy entrainer, N-methyl-2-pyrrolidone, process design.


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1. Introduction When there is repulsion between two dissimilar molecules, a minimum-boiling azeotrope may occur with the azeotrope boils at a temperature that is lower than the boiling point of the lighter pure component. For fewer occurrences in industry, there are still a number of systems that exhibit molecular attractions, and these systems can result in maximum-boiling azeotropes. Both above azeotropic mixtures create difficulty on the feasibility for distillation-based separations. A recent review by Mahdi et al.1 listed the distillation-based separation methods including: pressure-swing distillation, extractive distillation, and azeotropic distillation. There are over fifty chemical separation systems listed in the three tables of this review paper with cited references. However, the references for the design of maximum-boiling azeotrope are relatively much fewer. Only two separation systems were mentioned in this review paper including one to separate ethyl-acetate/chloroform2 and another one to separate ethylenediamine/water3. Separation systems in both of these two references are operated in either batch or semi-batch mode. The separation systems for maximum-boiling azeotropic mixtures that operated in continuous mode were also very scarce in open literature. Luyben4 developed a design flowsheet of a hypothetical maximum-boiling azeotropic mixture via pressure-swing distillation. In a later paper, the design flowsheet of methanol/trimethoxysilane separation using pressure-swing distillation was proposed by Luyben5. Another system of ethylenediamine/water via pressure-swing distillation was also developed in Li et al.6 and Fulgueras et al.7. Wang et al.8 developed a pressure-swing distillation system for the separation of N-heptane and isobutanol which exhibited minimum- and maximum-boiling azeotropes at different operating pressures.



As for using extractive distillation, Hostrup et al.9 studied acetone-chloroform separation using either benzene, 1-hexanol or methyl-n-pentyl ether as extrainer. Luyben10 proposed to add dimethyl sulfoxide (DMSO) as heavy entrainer to enhance the relative volatility of acetone over chloroform. Comparison of extractive distillation and pressure-swing distillation for the above system was given in Luyben11 and found out that pressure-swing distillation was much more expensive for this separation than extractive distillation in terms of both capital and energy. An alternative heavy entrainer of ethylene glycol (EG) was proposed for the acetone/chloroform separation system by Shen et al.12. The same research group also showed the feasibility analysis of acetone/chloroform separation by either heavy, light, or water as entrainer13. Another separation system of Phenol/cyclohexanone was studied by Li et al.14 by adding a heavy entrainer of acetophenone into the system. The entrainer recovery column was also operated at different pressure to avoid phenol-acetophenone azeotrope. Shen et al.15 extended thermodynamic insights on batch extractive distillation to continuous operation. In their study, a feasible extractive distillation of chloroform and vinyl acetate using n-butyl acetate as entrainer was illustrated. Anokhina and Timoshenko16 studied methyl acetate and chloroform separation using ethylene glycol, DMSO or dimethylformamide as entrainer. The final system we can find in open literature is to compare extractive distillation and heterogeneous azeotropic distillation for the ethylenediamine dehydration system17. In this paper, we use the system having more appearances in open literatures (acetone/chloroform separation) as an illustrative example to explain the unique design considerations for this kind of separation systems as opposed to the minimum-boiling extractive distillation system, although having the same two-column design flowsheet.


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Another contribution of this paper is to propose an alternative N-methyl-2-pyrrolidone (NMP) as a more suitable heavy entrainer for this separation system. Comparisons of the design result using our proposed entrainer versus the results from using the other two entrainers in literatures (DMSO or EG) will be shown in the following sections.

2. Extractive Distillation System For Acetone/Chloroform System 2.1. Thermodynamic model and conceptual design To conduct rigorous investigation of this acetone/chloroform separation system, UNIQUAC model is used to describe the non-ideal behavior in liquid phase and the behavior in gas phase is assumed to be ideal (IG). All binary thermodynamic model parameters used in this study can be found in Table S1 of the supporting information. Figure 1 shows the Txy plot of this system at 1.1 atm. With an equal molar feed as was studied in Luyben10, a regular column can obtain pure acetone at distillate stream while bottom stream can only obtain a mixture near maximum-boiling azeotropic composition of 33.90 mol % acetone at azeotropic temperature of 340.66 K. This means that some of the light component acetone will still go to the column bottoms carrying by the maximum-boiling azeotrope. One thinking to circumvent this problem is to feed a suitable heavy entrainer into the same column to greater enhance the relative volatility of acetone over chloroform to prevent acetone from going to the column bottoms. A second entrainer recovery column can be added to recycle the heavy entrainer back to the first column. The conceptual design flowsheet of this two-column extractive distillation system as shown in Figure 2 is exactly the same as the ones for minimum-boiling systems. For the extractive distillation column, it is customary to call the top section from top tray to the feed tray of heavy entrainer as rectifying section. The middle section



from the feed tray of heavy entrainer to feed tray of fresh feed is called extractive section while the bottom section from feed tray of fresh feed to column bottoms as stripping section. For minimum-boiling systems with the help of heavy entrainer, the purpose of extractive section is to prevent one component (usually the heavy boiler) from going up to the column top. Therefore, usually the extractive section contains quite a few stages to fulfill this purpose. As for maximum-boiling system, the difficulty of separation is not on the top part of this column but on the bottom part. With the help of heavy entrainer adding into this system, now this column is able to prevent acetone from going down to the column bottoms originally carried by the maximum-boiling azeotrope in a regular column. Thus straightly speaking, now the extractive section should be from the feed tray of heavy entrainer to the column bottoms and this section serves to diminish acetone toward bottom of this column. 2.2. Entrainer selection To analyze three components with heavy entrainer adding into the system, a ternary diagram with residue curve maps (RCM) can be drawn. For maximum-boiling system with no additional azeotrope formed with heavy entrainer, the topology of RCM is classified as Serafimov’s class 1.0-218-20. The unique feature of this RCM class is that a distillation boundary is formed to divide the ternary diagram into two distillation regions. As for minimum-boiling system with heavy entrainer, no distillation boundary is formed. Distillation boundary constrains of obtaining two pure components at opposite side of distillation boundary as products from an entire sequence of distillation columns. According to the heuristic rule in Doherty and Malone21, it is possible for the conceptual design flowsheet as in Figure 2 to become feasible only if the distillation


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boundary is curved. In this paper, it is illustrated that the curvature of distillation boundary together with the topology of RCM and material balance lines for the candidate ternary diagram can be used as a screening tool for entrainer selection. The feed stream information is chosen to be exactly the same as the one in Luyben10 with feed flow rate at 100 kmol/h, feed temperature at 320 K, and equal molar of acetone and chloroform. The two product purities need to be at 99.5 mol% or over. Figure 3 shows the ternary diagram, RCM, and material balance lines of the acetone/chloroform separation system with three alternative entrainers of NMP (proposed), DMSO (in Luyben10), and EG (in Shen et al.12). Let’s use NMP in Figure 3(a) as an example to demonstrate the finding for minimum feasible entrainer-to-feed ratio from this figure. In Figure 3(a), the fresh feed location is labeled as FF. By adding entrainer into the extractive distillation column, the merge point of FF and NMP can be estimated at the intersection point of two material balance lines. The reason is that the location of the distillate product should be located at acetone (lower-right corner) while the location of B1 point needs to be estimated at near the intersection of the distillation boundary with the chloroform-NMP side. In this way, the material balance line for the second entrainer recovery column can be drawn as a straight line satisfying B2 at NMP (lower-left corner) and D2 at chloroform (top corner). Thus, all stage compositions of the second entrainer recovery column are actually at the different distillation region as compared with that of extractive distillation column. With the strong curvature of this distillation boundary, it becomes possible to have feed location of second column at different distillation region as opposed to the locations of B2 and D2. From lever rule, the minimum feasible entrainer-to-feed ratio can be estimated to be 0.56. This is a hard constraint for the entrainer flow rate to be always needed higher



than 56 kmol/h, otherwise, no feasible separation can be achieved for the second column. Using the same thinking, the minimum feasible entrainer-to-feed ratios of EG and DMSO can be estimated as in Figs. 3(b) and 3(c) to be 1.74 and 1.26, respectively. This shows that the loading of extractive distillation system (and also entrainer recovery column) can be much relieved by using NMP as entrainer as compared to that of EG or DMSO. In Figure 3(a), the isovolatility curve can also be easily drawn using Aspen Plus simulation package. It is shown by a dash-dotted line from the azeotropic point of acetone-chloroform to the binary side of chloroform-NMP. This means that at this isovolatility curve, the relative volatility of acetone to chloroform is 1.0. Thus, the ternary diagram is divided into two regions with different relative volatility order of acetone vs. chloroform. In the much larger lower region, the relative volatility of acetone to chloroform is greater than 1.0. This can easily be predicted by the binary side of acetone-chloroform in this region. At this lower binary side, the acetone at 332.10 K would volatile more toward distillate than the maximum-boiling azeotrope at 340.66 K. The above explanations can serve as another evidence that acetone should be the distillate product of extractive distillation column. With the curvature nature (concave or convex) of distillation boundary for a maximum-boiling system, the distillate product of extractive distillation column as well as the feasibility of the entire extractive distillation system having high-purity distillate products at two column tops can be determined. For this acetone-chloroform system with three different entrainers as in Figure 3, it is not possible to have a feasible two-column sequence with chloroform as the distillate product of extractive distillation column. However for the minimum-boiling system, it is possible to have alternative


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distillate product for the extractive distillation column with different extrainers. The ways of using isovolatility curve to determine the feasible distillate product of extractive distillation column for minimum-boiling system was explained in Hsu et al.22 There will not have a similar argument for maximum-boiling systems mainly because of the presence of distillation boundary. Note again for minimum-boiling system, no distillation boundary is present. The ability of candidate entrainers to enhance the relative volatility of acetone over chloroform can be evaluated by FLASH2 module of Aspen Plus. With a fixed value of entrainer-to-feed ratio and setting the flasher to bubble-point condition, the relative volatility of acetone over chloroform can be calculated by the liquid and vapor compositions of the two flasher outlet streams. Figure 4 displays the enhancement of relative volatility for the three candidate entrainers. With equal molar flow rate of entrainer vs. fresh feed, it is demonstrated that NMP can increase the relative volatility to 4.0 while DMSO or EG can only increase it to less than 3.0. The main goal of the rectifying section of extractive distillation column is to separate acetone vs. entrainer and that of the entrainer recovery column is to separate chloroform vs. entrainer. From the Txy plots of these two pairs in Figure 5, it reveals that the separation should be quite easy. Since NMP is a commonly used solvent in industry, there is no carcinogenic problem of using it as entrainer for the separation of acetone and chloroform.

3. Comparative Study of Different Extractive Distillation Systems for the Separation of Acetone and Chloroform 3.1. Compare with DMSO as entrainer



The design flowsheet using DMSO as entrainer for this separation was developed in Luyben9. In this section, rigorous simulations are conducted to compare the economics of design flowsheets using our proposed entrainer versus the one in Luyben10. For a fair comparison, the total stages of two columns and the purity specifications of four column outlet streams were set to be exactly the same as that in Luyben9. In our proposed design flowsheets, iterative sequential search to minimize the total reboiler duty is performed to find the optimal entrainer-to-feed ratio and also the suitable feed locations of two columns. The operating pressures of two columns were set to be the same as that in Luyben10 with negligible tray pressure drops. The results of rigorous simulation study are displayed in Figure 6. Note that the four purity specifications were: chloroform mole fraction of 0.005 and acetone mole fraction of 0.001 at distillate and bottoms of extractive distillation column; and entrainer mole fraction of 0.0001 and chloroform mole fraction of 0.0001 at distillate and bottoms of second column. It is observed that the optimal entrainer-to-feed ratio of NMP system is only at 0.74, while that of DMSO system was set at 1.64. The total reboiler duty of our proposed system is significantly lower than that of DMSO system in Luyben10. To compare the economics of two design flowsheets, total annual cost (TAC) is calculated based mainly by the formulae in Luyben23. The TAC includes annual total operating cost (TOC, $/y) plus total capital cost (TCC, $) divided by a payback period of three years. The energy cost includes steam cost in reboilers and cooling water cost in condensers and cooler. The cost of makeup entrainer is negligible as compared to other costs. The capital cost includes column shells, reboilers, condensers and cooler. The formulae used in TAC calculations are summarized in Table S2 of the supporting information. The economic comparison of two design flowsheets in Figure 6 can be


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found in Table 1. By changing entrainer to NMP, significant savings of 23.2% and 21.0% in total operating cost and total annual cost, respectively, can be realized. 3.2. Compare with EG as entrainer In Shen et al.12, a design flowsheet using EG as entrainer was developed. Again for a fair comparison in the following simulation study, the total numbers of stages for two columns of using NMP as entrainer were set to be the same as that in Shen et al.12. The product specifications in Shen et al.12 were mostly set for the major components of that outlet streams, e.g. 0.9951 mole fraction for acetone product stream; 0.9944 mole fraction for chloroform product stream; and 0.9999 for EG outlet stream. The remaining one for the bottoms of extractive distillation column, it was set to have only 0.001 mole fraction of acetone in that stream. The two flowsheets with exactly the same feed and product conditions are shown in Figure 7. It is observed that the entrainer-to-feed ratio in the flowsheet of Shen et al.12 is much larger than that in our proposed system. This larger entrainer-to-feed ratio for EG case as compared to the case of DMSO in Figure 6 also agrees well with the prediction by using the topology of distillation boundary and material balance lines as explained in Section 2 and Figure 3. Table 2 compares the economics of the two systems in Figure 7. With detail calculations from rigorous simulation results, the economic advantage of NMP flowsheet as compared to that of EG was clearly demonstrated. The total operating cost and TAC can be significantly reduced by 30.1% and 26.2%, respectively, by changing the entrainer from EG to NMP.

4. Proposed Optimized Extractive Distillation System for the



Separation of Acetone and Chloroform 4.1. Design flowsheet with entrainer cooler In previous comparisons to two design flowsheets in open literatures, the total numbers of stages for two columns were fixed as the same as that in the previous design. In this section, the design flowsheet using NMP as entrainer is further improved to include the total numbers of stages of two columns in the iterative sequential optimization procedure to minimize total annual cost. Before doing the optimization study, several other design variables in the flowsheet was set as follows. The operating pressures of two columns are again set at 1.1 atm to allow for using cooling water in two condensers. A cooler is placed at the entrainer recycle stream to cool down the entrainer temperature before going back to extractive distillation column. Since cooling water is used in this cooler, a sub-cooled entrainer temperature of 320 K is assumed in the following simulation study. The optimization procedure is outlined in Figure 8. There are six design variables in this two-column system, including: entrainer-to-feed ratio (FE/FF); total stages of extractive distillation column (NT1); fresh feed location (NFF); entrainer feed location (NFE); total stages of entrainer recovery column (NT2); feed location of second column (NFF2). The optimization procedure can be divided into two steps. In the first step with a fixed FE/FF ratio, the minimized total annual cost of extractive distillation column (TAC1) can be independently obtained by iteratively search for the best NT1, NFF, and NFE. In the second step, the total TAC of the overall system will be calculated for a specific value of FE/FF ratio by iteratively search for the best NT2 and NFF2. Since there will be a trade-off effect of FE/FF ratio on the two columns, all best results under a specific FE/FF ratio will be collected to determine the final optimized


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design flowsheet. In all the simulation studies, the three product purity specifications are: 0.995 mole fraction for acetone product stream; 0.995 mole fraction for chloroform product stream; and 0.9999 for NMP outlet stream. As for the bottoms of extractive distillation column, it is set so that the acetone contents divided by the total of acetone plus chloroform contents to be at 0.0049. The setting will ensure that the major impurity at chloroform product stream is acetone, not NMP. The detail results from the optimization procedure can be found in Figure 9. At FE/FF=0.6 and NT1=32, the optimized NFF=13 and NFE=3 can be viewed from the upper left plot of Figure 9. With NFE above the 3rd stage, the purity specifications cannot be achieved. The upper right plot summarized each best result by varying NT1 but with a fixed FE/FF=0.6. It is observed the optimized NT1=32. The lower left plot further search for the best NT2 and NFF2 at the case of FE/FF=0.6, NT1=32, NFE=3, and NFF=13. The results showed the optimized NT2=15 and NFF2=5. The lower right plot of Figure 9 summarized all the results using the optimization procedure in Figure 8 and then picked the lowest TAC at each FF/FF ratio. It is found that the optimized FE/FF is at 0.6. The optimized design flowsheet can be found in Figure 10. As compared with the two previous flowsheets in Figs. 6(b) and 7(b), the total reboiler duty is further reduced by means of taller columns. An observation worth mentioning is that the total stages between NFE and NFF are less than the cases in open literatures for minimum-boiling azeotropic separation. Usually these total stages (named extractive section for minimum-boiling system) were needed to be large in order to circumvent the separation difficulty at upper part of this column by diminishing the heavy component



at top stage of this so called extractive section. However, for the maximum-boiling system as in this case, the separation difficulty is actually shifted at the lower part of this column by avoiding the light component from going down the column. Thus, it is important to have enough column stages from NFE to column bottoms so that to achieve this separation goal. The economics of this optimal design flowsheet is shown in Table 3. It is demonstrated that the TAC is further reduced as compares to the ones in Tables 1 and 2. In the following section, possible further energy-saving schemes will be discussed. The purpose is to develop an industrial easily implementable design schemes to further save energy of this separation system. 4.2. Proposed energy-saving design flowsheet Before developing the final proposed design flowsheet of this separation system, two other commonly used energy-saving design schemes are investigated to explain the reasons for declining any further investigation. The two energy-saving schemes are: Dividing-wall column (thermally-coupled design) and double-effect distillation by operating two columns at different pressures. By observing the design flowsheet in Figure 10, it is possible to thermally-coupled this two-column sequence into a single column with upper dividing-wall. The way is to combine two reboilers into one at the bottoms of entrainer recovery column. To still provide vapor traffic to extractive distillation column, a vapor side stream can be designed to draw-off some vapor from a certain stage of entrainer recovery column to feed into the bottom stage of extractive distillation column. By considering this design configuration as one dividing-wall column to save capital cost of column shell, the stripping section of entrainer recovery column is the bottom part of this column, the


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rectifying section of entrainer recovery column is the right-side of upper dividing-wall, and the whole extractive distillation column becomes the left-side of upper dividing-wall. Although this energy-saving design configuration can eliminate the re-mixing effect at bottom part of extractive distillation column to make the duty of combined reboiler lower than the sum of duties of the original two reboilers. The saving of actual steam cost needs to be carefully investigated. Wu et al.24 investigated four extractive distillation systems and found that only the acetone-methanol-water system can benefit from changing the original two-column design into a dividing-wall column. The main reason for saving of reboiler duty but not the actual steam cost is because usually different steam grades are required for the original two columns. This is exactly the same situation in this acetone/chloroform separation system. Notice from Figure 10, the bottom temperature of first column is at 391.3 K requiring only low-pressure steam, while that of second column is at 481.1 K requiring more expensive high-pressure steam. Because of this reason, the detail design configuration of dividing-wall column was not investigated any further. Another way to save energy is to operate two columns at different operating pressure so that the energy removal from the condenser of high-pressure column can be used (at least partially) by the reboiler of low-pressure column. By observing four temperatures at condensers and reboilers in Figure 10, the only possibility is to let the heat removal of second-column condenser to be used by first-column reboiler. There are two ways to make the temperature of second-column condenser to be higher than that of the first-column reboiler (preferably with a temperature difference of 10 K). One way is to operate the first column at vacuum pressure. However, it can be



expected that the condenser temperature of this first-column would become lower to prohibit the using of cheap cooling water as cooling medium, not to mention the cost of operating at vacuum. The other way is to operate the second column at even higher pressure to provide this temperature difference. However, it is expected that the bottom temperature of second column would become much higher to prohibit the using of even high-pressure steam as heating medium of this reboiler. By just checking the optimal design flowsheet in Figure 10, it can reveal that the above two energy-saving design schemes are all infeasible. The other design scheme which can easily be implemented in industry is to exchange heat of a hot stream at bottoms of entrainer recovery column to a cold stream of fresh feed. Figure 11 displays this design flowsheet with such heat exchangers. Note that all other design variables considered in previous section were fixed at original values. There are two heat-exchangers in Figure 11 in order to divide the exchanged-heat into two different situations of liquid-to-liquid and liquid-to-vaporizing liquid. This way to more accurately estimate heat transfer area was suggested in Luyben25. In Figure 11, HX1 was used to heat-up fresh feed into saturated liquid and HX2 was used to further partially vaporize the feed stream into a stream with vapor fraction of 0.51. The minimum approaching temperature difference was set at 10 K in the simulation. Comparing the design flowsheets in Figs. 10 and 11, it can be seen that the reboiler duty of first column is significantly reduced from 1003.6 kW to 721.7 kW because of warmer fresh feed into this column. The economics of these two design flowsheets are itemized in Table 3. With this simple heat-integration scheme, the total operating cost and TAC can further be reduced by 15.4% and 9.3%, respectively. The liquid composition profiles of two columns are displayed in Figure 12. The


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separation of entrainer recovery column should be easy (as can be expected from Figure 5(b)) and exhibits similar behavior as in the minimum-boiling system. However, the behavior in the composition profile of extractive distillation column is not the same as that in a minimum-boiling system. In this case, the commonly called “extractive section” is quite short with only eleven stages out of total of 32 stages. The separation difficulty of the maximum-boiling system is actually on the bottom part of this column. With the presence of high composition of NMP at stages starting from the entrainer feed location to column bottoms, the avoidance of acetone toward the column bottoms can be satisfied. It is also observed that at stages below fresh feed location NMP is diluted because of feed stream into this column. From the optimization procedure, the number of stages below fresh feed location is quite high in order to allow acetone to become diminishing at column bottoms.

5. Conclusions Extractive distillation of maximum-boiling system is characteristically different than that of minimum-boiling system. By adding a heavy entrainer into the system, a distillation boundary will be formed for maximum-boiling system. If the curvature of this distillation boundary is flat, it is not possible to develop a feasible two-column system with high-purity products. Using the acetone/chloroform separation system as an

illustrative

example,

a

method

for

estimating

the

minimum

feasible

entrainer-to-feed ratio is developed by just using the topology of ternary diagram with distillation boundary and material balance lines. Besides the ability to enhance relative volatility, this minimum ratio can be considered as an important screening tool for finding the most suitable entrainer of a particular extractive distillation system. A new entrainer of NMP is also proposed in this paper for the separation of



acetone and chloroform. By comparing the performance of this new entrainer with that of DMSO or EG used in open literature, it is confirmed that significant total operating cost and total annual cost can be gained by the proposed two-column extractive distillation system. Further energy-saving design schemes such as dividing-wall column or multi-effect distillation were also discussed in this paper. An industrial easily implementable heat-integration scheme with feed-effluent-heat-exchanger is proposed to significantly provide 33% further savings of reboiler duty at extractive distillation column.

Acknowledgements The research funding from the Ministry of Science and Technology of the R. O. C. under grant no. MOST 107-2221-E-002-102 is greatly appreciated.

Supporting Information This part includes some detailed information in tables, and it is available free of charge via the Internet at http://pubs.acs.org/.

References (1)

Mahdi, T.; Ahmad, A.; Nasef, M. M.; Ripin, A. State-of-the-art technologies for separation of azeotropic mixtures. Separation & Purification Reviews 2015, 44, 308-330.

(2)

Stéger, C.; Rév, E.; Horváth, L.; Fonyo, Z.; Meyer, M.; Lelkes, Z. New extractive


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Page 22 of 38

Table 1. Economic comparison of proposed NMP system vs. DMSO system DMSO

NMP

Configurations

EDC

ERC

EDC

Capital cost for column (1000 $) Capital cost for reboiler (1000 $ ) Capital cost for condenser (1000 $) Capital cost for cooler (1000 $) Capital cost for FEHE (1000 $) TCC (1000 $) Steam cost (1000 $/y) Cooling water cost (1000 $/y) Cooling water cost for cooler (1000 $/y) Entrainer makeup (1000 $/y) Total reboiler duty (kW) Total steam cost (1000 $/y) TOC (1000 $/y) TAC (1000 $/y)

158.7 148.6 93.7 64.3 -

76.9 71 79.1 -

137 73.5 108.8 71.9 83.7 67.4 48.1 590.4(-14.7%) 249.4 251.9 6.3 5.9 7.2 1998.5 501.3 520.7(-23.2%) 717.5(-21%)

692.3 333.8 318.8 7.5 7.5 10.7 2610.4 652.6 678.3 909.1


ERC

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Table 2. Economic comparison of proposed NMP system vs. EG system EG

NMP

Configurations

EDC

ERC

EDC

Capital cost for column (1000 $) Capital cost for reboiler (1000 $ ) Capital cost for condenser (1000 $) Capital cost for cooler (1000 $) Capital cost for FEHE (1000 $) TCC (1000 $) Steam cost (1000 $/y) Cooling water cost (1000 $/y) Cooling water cost for cooler (1000 $/y) Entrainer makeup (1000 $/y) Total reboiler duty (kW) Total steam cost (1000 $/y) TOC (1000 $/y) TAC (1000 $/y)

202.7 118.1 74.9 83 -

59.6 82.2 63.9 -

168.8 62.5 88.3 72.1 85.7 69.9 41.5 588.8(-13.9%) 229.8 252.7 6.5 6.4 5.8 1914 482.5 501.2(-30.1%) 697.5(-26.2%)

684.4 329.2 5.3 16.2 2738.2 690.2 717.1 945.2

361 5.4 -


ERC


Page 24 of 38

Table 3. Economic comparison of proposed system with and without heat-integration NMP

NMP FEHE

Configurations

EDC

ERC

EDC

Capital cost for column (1000 $) Capital cost for reboiler (1000 $ ) Capital cost for condenser (1000 $) Capital cost for cooler (1000 $) Capital cost for FEHE (1000 $) TCC (1000 $) Steam cost (1000 $/y) Cooling water cost (1000 $/y) Cooling water cost for cooler (1000 $/y) Entrainer makeup (1000 $/y) Total reboiler duty (kW) Total steam cost (1000 $/y)

178.1 83.4 86.2 39.9 -

82 65.1 60.7 -

224.8 216.3 6.6 5 5.5 1763.6 441.1

182.1 83.5 68 66.5 104 61.5 59.4 625(4.9%) 150.5 223 8.7 5.1 1455.7 373.2

TOC (1000 $/y)

458.2

387.3(-15.4%)

TAC (1000 $/y)

657.2

595.8(-9.3%)

595.4


ERC

Page 25 of 38 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


Figure Captions Figure 1. Txy plot of acetone/chloroform system at 1.1 atm Figure 2. Conceptual design flowsheet with important design variables Figure 3. Ternary diagram with distillation boundary, isovolatility curve, and material balance lines for estimating minimum feasible entrainer-to-feed ratio (a) NMP system; (b) DMSO system; (c) EG system Figure 4. Enhancement of relative volatility by three candidate entrainers Figure 5. Txy plots at 1.1 atm (a) NMP/acetone system; (b) NMP/chloroform system Figure 6. Comparison of design flowsheets with same feed and product specifications (a) DMSO as entrainer; (b) NMP as entrainer Figure 7. Comparison of design flowsheets with same feed and product specifications (a) EG as entrainer; (b) NMP as entrainer Figure 8. Sequential iterative optimization procedure for this study Figure 9. Optimization results for the determination of important design variables Figure 10. Design flowsheet of the acetone/chloroform separation system via extractive distillation. Figure 11. Proposed design flowsheet with heat-integration Figure 12. Liquid composition profiles of EDC and ERC.



Page 26 of 38

Figure 1


Page 27 of 38 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


Figure 2

Entrainer recycle

Qc 1

Qc 2

FE/FF =.? Entrainer makeup

D1 Entrainer feed (FE) NFE NT1 =.? NFE =.? NFF =.? Acetone-Chloroform feed (FF)

D2 NT2 =.? NFF2 =.?

Extractive distillation column

Entrainer recovery column B1

NFF2

NFF

Qr1

Qr2

B2



Page 28 of 38

Figure 3 (a)

(b)

(c)


Page 29 of 38

Figure 4

Relative volatility (Acetone/Chloroform)

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


Ethylene glycol N-methyl-2-pyrrolidone DMSO

Entrainer/Feed ratio(mole basis)



Page 30 of 38

Figure 5


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Figure 6

(a) Fresh Entrainer 320 K 1.1 atm 164.4 kmol/hr Chloroform 0.0001 DMSO 0.9999

320 K, -1046.65 kW

Qc = -737.8 kW

D1 332.2 K 1.1 atm 50.03 kmol/hr Acetone 0.995 Chloroform 0.005

1.1atm

4 MAKEUP 0.0057 kmol/hr

10 Fresh Feed 320 K 1.1 atm 100 kmol/hr Acetone 0.5 Chloroform 0.5

21

Qc = -736.6 kW

1.1atm

EDC NT1 22 NFE 4 NFF 10 RR 0.8 D1m

6 B1 407.6 K 1.1 atm 214.36 kmol/hr Acetone 0.001 Chloroform 0.232 DMSO 0.766

ERC NT2 12 NF2 6 RR 0.8 D 0.83 m

11

D2 337.3 K 1.1 atm 49.96 kmol/hr Acetone 0.0043 Chloroform 0.9955 DMSO 0.0001

B2 467.6 K 1.1 atm 164.4 kmol/hr Chloroform 0.0001 DMSO 0.9999

Qr = 1120.5 kW

Qr = 1489.9 kW

(b) Fresh Entrainer 320 K 1.1 atm 79.99 kmol/hr Chloroform 0.0001 NMP 0.9999

320 K, -714 kW

Qc = -619.2 kW


1.1atm



21

Qc = -575.6 kW

1.1atm

EDC NT1 22 NFE 4 NFF 10 RR 0.5 D 0.87 m

Qr = 1113.2 kW

4 B1 402.3 K 1.1 atm 129.8 kmol/hr Acetone 0.001 Chloroform 0.383 NMP 0.615

ERC NT2 12 NF2 4 RR 0.4 D 0.79 m

11

Qr = 885.3 kW


D2 337.2 K 1.1 atm 49.88 kmol/hr Acetone 0.0026 Chloroform 0.997 NMP 0.0001

B2 481.1 K 1.1 atm 79.99 kmol/hr Chloroform 0.0001 NMP 0.9999


Page 32 of 38

Figure 7

(a) Fresh Entrainer 320 K 1.1 atm 221.48 kmol/hr Chloroform 0.0001 EG 0.9999

320 K, -1594.7 kW

Qc = -522.6 kW


1.1atm

3 MAKEUP 0.021 kmol/hr Fresh Feed 320 K 1.1 atm 100 kmol/hr Acetone 0.5 Chloroform 0.5

1.1atm


9

29

Qc = -531.5 kW


ERC NT2 10 NF2 4 RR 0.3 D 0.76 m

4 B1 397.3 K 1.1 atm 271.53 kmol/hr Acetone 0.001 Chloroform 0.183 EG 0.815

9

B2 473.38 K 1.1 atm 221.48 kmol/hr Chloroform 0.0001 EG 0.9999

Qr = 1268.8 kW

Qr = 1469.4 kW

(b) Fresh Entrainer 320 K 1.1 atm 63.82 kmol/hr Chloroform 0.0001 NMP 0.9999

320 K, -569.74 kW

Qc = -642.7 kW


1.1atm


13 Fresh Feed 320 K 1. 1atm 100 kmol/hr Acetone 0.5 Chloroform 0.5

29

Qc = -612.3 kW

D2 337.3 K 1.1 atm 50.04 kmol/hr Acetone 0.0022 Chloroform 0.9944 NMP 0.0033

1.1atm


3 B1 394 K 1.1 atm 113.86 kmol/hr Acetone 0.001 Chloroform 0.437 NMP 0.562

Qr = 1025.8 kW


ERC NT2 10 NF2 3 RR 0.4 D 0.794 m

9

Qr = 888.2 kW


Page 33 of 38 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


Figure 8



Page 34 of 38

Figure 9


Page 35 of 38 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


Figure 10

Fresh Entrainer 320 K 1.1 atm 59.98 kmol/hr Chloroform 0.0001 NMP 0.9999

320 K, -535.48 kW

Qc = -648.2 kW

1.1atm



31


D1 332.2 K 1.1 atm 50 kmol/hr Acetone 0.995 Chloroform 0.0048 NMP 0.00016

Qc = -490.5 kW

D2 337.3 K 1.1 atm 50 kmol/hr Acetone 0.0049 Chloroform 0.995

1.1atm

5 B1 391.3 K 1.1 atm 109.995 kmol/hr Acetone 0.0022 Chloroform 0.452 NMP 0.545

Qr = 1003.6 kW


ERC NT2 15 NF2 5 RR 0.19 D 0.73 m

14

Qr = 760 kW



Page 36 of 38

Figure 11

HX1 Qex = 78 kW 330.4 K

Fresh Feed 320 K 1.1 atm 100 kmol/hr Acetone 0.5 Chloroform 0.5 354.4 K

339.8 K

Fresh Entrainer 330.49 K 1.1 atm 65.9917 kmol/hr Chloroform 0.0001 NMP 0.9999

Qc = -865 kW

HX2 Qex = 478.1 kW 481.1 K

340.09K Vapor fraction 0.57

D1 332.2 K 1.1 atm 50 kmol/hr Acetone 0.995 Chloroform 0.0049

Qc = -500.6 kW

D2 337.3K 1.1atm 50 kmol/hr Acetone 0.0049 Chloroform 0.995

1.1atm

1.1atm



13

B1 394.8 K 1.1 atm 115.99 kmol/hr Acetone 0.0021 Chloroform 0.429 NMP 0.568

31

Qr = 671.9 kW


5

ERC NT2 15 NFF2 5 RR 0.21 D 0.74 m

14

Qr = 783.8 kW


Page 37 of 38 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29 30 31 32 33 34 35 36 37 38 39 40 41 42 43 44 45 46 47 48 49 50 51 52 53 54 55 56 57 58 59 60


Figure 12



Table of Contents (TOC) Graphic


Page 38 of 38

Unique Design Considerations for Maximum-Boiling Azeotropic

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