Critical Assessment of the Energy-Saving Potential of an Extractive

Article pubs.acs.org/IECR

Critical Assessment of the Energy-Saving Potential of an Extractive Dividing-Wall Column Yi Chang Wu, Paul Hen-Chia Hsu, and I-Lung Chien* Department of Chemical Engineering, National Taiwan University, Taipei 10617, Taiwan ABSTRACT: In this paper, the energy-saving potential of an extractive dividing-wall column (EDWC) is investigated. One potential drawback on the EDWC design is that two reboilers in the original design need to be combined into one reboiler. Since a heavy entrainer is used in the extractive distillation system, often cases show that the total reboiler duty is reduced but with adversely increasing the total steam cost. Three industrial separation systems of isopropyl alcohol-water, dimethyl carbonatemethanol, and acetone-methanol have been used as demonstrating examples for critical assessment of the energy-saving potential of the EDWC design. It is found that although the savings of the overall reboiler duty can be made by using the EDWC design, only the acetone-methanol system using water as an entrainer actually saves on the steam cost. The control performance of the EDWC design is also hampered because of losing one important control degree-of-freedom.

1. INTRODUCTION Dividing-wall (thermally coupled) column is a process intensification technology to obtain energy-savings for the separation of ternary mixtures. By way of eliminating the remixing effect in the conventional column systems, up to 30% in energy savings can be realized using this technology.1−7 Because the design structure of the dividing-wall column is more complex and more interactive than the conventional column system, various control structures including conventional control strategies as well as advance controllers such as LQG/LQR, H∞, and model predictive control have been proposed in open literature.8−15 A good complete reference for the subject of dividing-wall columns is in Yildirim et al.16 They gave a comprehensive review of current industrial applications of dividing-wall columns and related research activities including column configurations, design, modeling, and control issues. It is estimated about 350 industrial applications could be expected by 2015. This paper focuses on investigating the energy-saving possibility of extractive distillation systems for separating azeotropes. Gutiérrez-Guerra et al.17 studied the thermally coupled extractive distillation sequence of three separation systems. Significant energy savings in the range between 20 and 30% were found in contrast to the conventional arrangement. Moreover, they use theoretical control properties such as the condition number and the minimum singular value to predict the dynamic behavior of EDWC could be even better than that in the conventional sequence. Bravo-Bravo et al.18 used the constraint stochastic multiobjective optimization technique to design the extractive dividing-wall column of four systems including nheptane/toluene/aniline, tetrahydrofuran/water/1,2-propanediol, isopropyl alcohol (IPA)/water/DMSO, and acetone/ water/octanoic acid. In the study of the reactive distillation production of dimethyl carbonate (DMC), Wang et al.19 studied a thermally coupled extractive distillation system for the purpose of separating DMC and methanol. Using phenol as an entrainer, they found 17.6% reboiler duty can be saved as compared to the two-column sequence without thermal coupling. Kiss and Suszwalak20 studied bioethanol dehydration by extractive and © 2013 American Chemical Society

azeotropic distillation in dividing-wall columns. The separation system via extractive distillation leads to a lower energy requirement as compared to the azeotropic distillation case. By further considering EDWC configuration, an additional 9.4% in the reboiler duty can be saved. Kiss and Ignat21 developed an innovative EDWC column to further combine the preconcentrator column to the two-column extractive distillation system for diluted fresh feed. The above-mentioned literature showed that the dividing-wall column is a promising configuration for industrial applications. This paper intends to closely investigate the potential energysaving benefit of EDWC as compared to the convention twocolumn system. One potential drawback of the EDWC configuration is that two reboilers in the original design need to be combined into one reboiler. This may cause a deficiency in the steady-state economics because of the high boiling-point temperature of the heavy entrainer. On the dynamic controllability issue, there may also be a problem with losing of one control degree-of-freedom by combining two reboilers into one. Three industrial examples will be used as demonstrating examples for critical assessment of the energy-saving potential of the EDWC design.

2. CASE STUDY: IPA DEHYDRATION Arifin and Chien22 studied an isopropyl alcohol (IPA) dehydration process via extractive distillation using DMSO as an entrainer. The optimal design flowsheet of this process has been established showing that the total annual cost (TAC) and the energy requirement are significantly less than a competing design flowsheet via heterogeneous azeotropic distillation.23 It is shown that the TAC is reduced by as much as 32.7%, and the required steam cost is also cut by 30.3%. To further assess the energy-savings of this extractive distillation system by using a Received: Revised: Accepted: Published: 5384

December March 18, March 20, March 20,

25, 2012 2013 2013 2013

dx.doi.org/10.1021/ie3035898 | Ind. Eng. Chem. Res. 2013, 52, 5384−5399

Industrial & Engineering Chemistry Research

Article

Figure 1. Optimal design flowsheet of the conventional two-column system for IPA dehydration.

dividing-wall configuration, this industrial example is chosen for a detail investigation of the design flowsheet via EDWC. 2.1. Conventional Two-Column System. The physical property information of this separation system can be found in the original paper. The optimal design flowsheet established in Arifin and Chien22 is reproduced in Figure 1 using Aspen Plus24 simulation with only a slight difference. The main difference is due to the setting of the IPA composition in the F2 stream. In Figure 1, the IPA loss through the D2 stream is less than that of the original flowsheet. This will make the production rate of D1 a little higher but requiring a little higher reboiler duty in C1. We will use Figure 1 as a base case for investigating the potential for further energy-savings via EDWC. The design flowsheet includes an extractive distillation column and an entrainer recovery column. The IPA-water mixture and the entrainer (DMSO) are fed into the extractive distillation column with the stages above the entrainer feed as the rectifying section, the stages between the two feeds as the extractive section, and the stages below the feed stage of the IPA-water mixture as the stripping section. The presence of DMSO alters the relative volatility between IPA and water causing IPA to move toward the top part and water to move toward the bottom part of this column. In the rectifying section, there is essentially no water, thus simple separation between IPA and DMSO is performed with pure IPA going to the distillate and DMSO returning to the extractive section. In the stripping section, IPA as the lightest component is stripping toward the extractive section of the column resulting in only negligible IPA in the column bottoms. This column bottom stream is fed into the entrainer recovery column to produce water in the distillate and DMSO in the column bottoms. DMSO as a heavy entrainer is recycled back to the extractive distillation column. To balance tiny entrainer losses in both distillate streams, the small makeup stream of an entrainer is added. A cooler is also designed to cool the returned entrainer stream to a temperature close to that of the entrainer feed stage. The liquid composition profiles of the two columns are shown in Figure 2. For the entrainer recovery column with negligible IPA, the composition profile is normal with a light component

Figure 2. Liquid composition profiles of the two-column system for IPA dehydration.

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Figure 3. Design configuration of the thermally coupled two-column system. (a) Simulated configuration. (b) Configuration with main column and side rectifier.

effect to save energy is to design a thermally coupled two-column system detailed in the next subsection. 2.2. Thermally Coupled Two-Column System. The configuration of a thermally coupled two-column system is shown in Figure 3. In this configuration (Figure 3a), the vapor traffic of the new extractive distillation column (without reboiler) is provided by a sidedraw from the entrainer recovery column. The liquid outlet from the bottoms of this extractive distillation column serves as the feed to the entrainer recovery column. An equivalent design configuration of a thermally coupled twocolumn system is shown in Figure 3b with a main column with

(water) enriched toward the top of the column and a heavy component (DMSO) to the bottoms. As for the composition profile of the extractive distillation column, the extractive section is rather large. The reason is because of very tight IPA purity specification (99.9999 mol %) for the semiconductor usage, thus the water has to be diminished in this section. Further examining the composition profile of the extractive distillation column, it is found that a prominent remixing effect is exhibited in the stripping section with one composition (water) going through a maximum in the middle of this section and then decreasing toward the column bottoms. A way to eliminate this remixing 5386



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Figure 4. Design flowsheet of the thermally coupled two-column system for IPA dehydration.

reboiler and another side rectifier. In this configuration, the stripping section of the entrainer recovery column with reboiler is lumped into the new extractive distillation column. We used the design configuration of Figure 3a in the Aspen Plus24 simulation. We assumed that the optimized conventional two-column system is now undergoing a retrofit to change to a thermally coupled design. The natural choice is to keep the original optimal configuration the same in this new system. Thus the rectifying section, the extractive section, and the stripping section of the extractive distillation column and also the total stage of the second columns are all kept the same. The only optimized variables are the vapor sidedraw flow rate and the location of the sidedraw. It was found that varying the sidedraw location differing from the original feed location of the original twocolumn system had a negligible effect on the operating cost. In the following, for simplicity reason only the vapor sidedraw flow rate will be varied until reaching a minimum value where the product specifications can still be met. The resulting flowsheet for the thermally coupled design is shown in Figure 4. It is observed that the two product purities are all maintained at same high purity. The overall reboiler duty is reduced from 2670.52 KW (= 1643.92 + 1026.60) in Figure 1 to 2487.79 KW. The reduction in the reboiler duty (6.84%) is not that significant as in Gutiérrez-Guerra et al.17 but more in line with the result in Kiss and Suszwalak.20 The liquid composition profile in the new extractive distillation column without reboiler is shown in Figure 5. It is evident that the remixing effect is eliminated. An important observation in Figure 4 is that the bottom temperature of the new entrainer recovery column is at 196.45 °C. With this high temperature, high-pressure steam at 450 psig (237 °C) has to be used for providing the heat duty in the reboiler. However for the conventional two-column system as in Figure 1, the bottom temperature of the C1 column is lower at 152.03 °C. Thus, the reboiler duty for the C1 column can instead be provided by medium-pressure steam at 150 psig (185 °C). The decision of using what kind of steams is to provide at least 10 °C temperature difference between the steam and the process fluid. Since the price of medium-pressure steam is cheaper than

Figure 5. Liquid composition profile of the new extractive distillation column without reboiler for IPA dehydration.

the high-pressure steam, the total steam cost of the conventional two-column system is actually less than the thermally coupled system. A complete comparison of the Total Annual Cost (TAC), the reboiler duty, and the steam cost will be shown in Section 2.4. 2.3. Extractive Dividing-Wall Column. One of the other major benefits of an extractive dividing-wall column is the savings in the space requirement from a two-column system to a dividing-wall column with a single shell. By changing the design flowsheet in Figure 4 to increase the number of stages in the rectifying section of the entrainer recovery column, the modified design flowsheet is shown in Figure 6. This modification is made so that uneven column heights on the left and the right sides of the dividing-wall can be avoided. Notice that although the rectifying section is increased in the entrainer recovery column, the reboiler duty is actually slightly increased from 2487.79 KW in Figure 4 to 2535.74 KW in Figure 6. The main reason is 5387



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Figure 6. Design flowsheet of the extractive dividing-wall column for IPA dehydration.

reboiler, and condenser. The capital cost for the internal wall is assumed to be negligible in comparison with the other capital costs. The formula for the installed costs can be found in Appendix D of Douglas’ book.25 The utility cost includes the steam and the cooling water costs. The unit price of the steam and cooling costs can be found in Table 23.1 of Seider et al. book.26 The pricing of the high-pressure steam (450 psig) is at $14.50/1000 kg and that of medium-pressure steam (150 psig) is at $10.50/1000 kg. Table 1 shows the results of the conventional two-column system in Figure 1, the thermally coupled design in Figure 4, and the dividing-wall design in Figure 6. It is shown that the TACs of the two new design configurations are all greater than that of the conventional design. The reboiler duty of the thermally coupled design is less than that of the conventional design. However, the steam cost of this design is actually 22.44% more than that of the conventional design. The reason is because of the unit price difference of the medium-pressure steam vs the high-pressure steam. We used Seider et al. book26 for the steam pricing. This information may vary in time or depend on the plant location. However, the pricing of the high-pressure steam can reasonably be assumed to always be higher than the medium- and lowpressure steam.

because the feed location of the entrainer recovery column is not at an optimized location anymore. Because it is assumed that only one column shell is required with this dividing wall configuration, the equivalent diameter (De) of the column shell on the top part of the dividing wall is calculated as the illustration in Figure 7. It is assumed that the De

3. CASE STUDY: DMC AND METHANOL SEPARATION From the results of the above example, it is possible to assess the energy-saving potential of an extractive dividing-wall column before doing rigorous simulation. In the following two sections, we will use two other industrial examples to illustrate this point. 3.1. Conventional Two-Column System. Hsu et al.27 developed an extractive distillation system for the separation of dimethyl carbonate (DMC) and methanol using aniline as the entraniner. The optimal design flowsheet established in Hsu et al.27 is reproduced in Figure 8. The remixing phenomenon at the stripping section of the extractive distillation column is again observed as in Figure 9. Note that the remixing effect for this system is even more severe than that of the IPA dehydration system in Figure 2. Thus, it is possible to develop the thermally

Figure 7. Illustration of how to calculate the equivalent diameter on the top part of the dividing-wall column.

is back-calculated so that total cross-section area from both sides of the dividing wall can be provided. For the dividing-wall column, there are two parts with different diameters. The top part has the equivalent diameter of 0.988 m, and the bottom part has the diameter of 1.206 m. 2.4. Comparison. In this section, TAC, reboiler duty, and steam cost of the above three cases will be compared. The TAC includes the annualized capital cost and the utility cost with a payback period for the capital cost assuming to be three years. The itemized capital costs include column shell, column tray, 5388



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Table 1. Results Comparison of Three Designs for the IPA Dehydration System conventional two-column configurations annualized capital cost for column shell (1000 $/y) annualized capital cost for column tray (1000 $/y) annualized capital cost for reboiler (1000 $/y) annualized capital cost for condenser (1000 $/y) steam cost (1000 $/y) cooling water cost (1000 $/y) annualized cooler capital cost (1000 $/y) cooling water cost for cooler (1000 $/y) entrainer makeup (1000 $/y) total reboiler duty (KW) (% difference) total steam cost (1000 $/y) (% difference) TAC (1000 $/y) (% difference) a

C1 207.64 20.09 64.52 47.60 236.99 3.49 31.04 2.01 0.19 2670.52 (0%) 479.03 (0%) 1067.80 (0%)

C2 125.94 10.68 41.53 31.35 242.04 2.68

thermally coupled design C1

C2

228.24 177.47 22.96 16.02 0 73.83 52.88 26.70 0 586.54 4.11 2.10 30.06 2.01 0.21 2487.79 (−6.84%) 586.54 (22.44%) 1223.14 (14.55%)

dividing-wall column C1

C2 a

259.02 145.79b 27.59a 14.08b 0 84.87 52.46 26.77 0 597.84 4.06 2.10 31.29 2.14 0.22 2535.74 (−5.05%) 597.84 (24.80%) 1248.25 (16.90%)

Calculations based on upper section of the EDWC. bCalculations based on lower section of the EDWC.

Figure 8. Optimal design flowsheet of the conventional two-column system for DMC and methanol separation.

coupled configuration or the dividing-wall configuration to reduce the required reboiler duty of this separation system. However from Figure 8, it is observed that the bottom temperature at the C1 column is at 163.77 °C while that of the C2 column is at 188.06 °C. With the design guideline of minimum temperature difference at 10 °C, medium-pressure steam (150 psig, 185 °C) is required for the C1 column and highpressure steam (450 psig, 237 °C) is required for the C2 column. From the experience in Section 2, it is conjecture that no benefit can be gained by replacing the column sequence to a more complicated thermally coupled or dividing-wall configurations. 3.2. Comparison of the Conventional Two-Column System to Energy-Saving Designs. The results from rigorous Aspen Plus simulations of the three designs are summarized in Table 2 with the flowsheets of the thermally coupled design and the dividing-wall column shown in Figures 10 and 11, respectively. From Figure 10, more reduction of the reboiler duty (16.18%) from thermally coupled design can be made as compared to the conventional design. This is mainly due to the elimination of a more severe remixing effect in previous Figure 9. However from the calculations in Table 2, the steam cost is

Figure 9. Liquid composition profile of the extractive distillation column for DMC and methanol separation.

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Table 2. Results Comparison of Three Designs for the DMC and Methanol Separation conventional two-column configurations annualized capital cost for column shell (1000 $/y) annualized capital cost for column tray (1000 $/y) annualized capital cost for reboiler (1000 $/y) annualized capital cost for condenser (1000 $/y) steam cost (1000 $/y) cooling water cost (1000 $/y) annualized cooler capital cost (1000 $/y) cooling water cost for cooler (1000 $/y) entrainer makeup (1000 $/y) total reboiler duty (KW) (% difference) total steam cost (1000 $/y) (% difference) TAC (1000 $/y) (% difference) a

C1 291.46 34.36 131.51 109.94 456.39 7.08 44.74 4.02 4.72 4109.02 (0%) 678.77 (0%) 1488.26 (0%)

C2 100.75 8.14 34.79 35.30 222.39 2.68


C2

302.70 176.95 36.11 17.47 0 80.72 106.12 17.45 0 812.01 6.70 0.90 45.31 4.02 2.57 3444.12 (−16.18%) 812.01 (19.63%) 1609.04 (8.12%)


C2 a

318.29 158.25b 38.84a 16.46b 0 89.65 105.93 17.58 0 837.74 6.69 0.91 46.78 4.22 2.74 3553.27 (−13.53%) 837.74 (23.42%) 1644.09 (10.47%)


Figure 10. Design flowsheet of the thermally coupled two-column system for DMC and methanol separation.

0.994 and methanol purity at 0.995. The remixing phenomenon at the stripping section of this extractive distillation column is again observed as in Figure 13. Thus, it is possible to develop the thermally coupled configuration or the dividing-wall configuration to reduce the required reboiler duty of this separation system. From the design flowsheet in Figure 12, it is observed that the bottom temperature at the C1 column is at 90.48 °C while that of the C2 column is at 104.77 °C. Note that for both columns lowpressure steam (50 psig, 147 °C) can be used as the heat source for both reboilers. Thus, reducing reboiler duty corresponds also to the reduction of the actual steam cost. The results from rigorous simulations of the three designs are summarized in Table 3 with the flowsheets of the thermally coupled design and the dividing-wall column shown in Figures 14 and 15, respectively. The pricing of the low-pressure steam (50 psig) in Seider et al.26 book is $6.60/1000 kg. From Figure 14, reduction of the reboiler duty (7.57% less) from the thermally coupled design can be made as compared to the conventional design. From the calculations in Table 3, the steam cost is also at 7.57% less than that of the conventional design. The dividing-

19.63% more than that of the conventional design. The dividingwall configuration, although it can save the space requirement, the reboiler duty, and the steam cost are slightly greater than that of the thermally coupled design.

4. CASE STUDY: ACETONE AND METHANOL SEPARATION 4.1. Using Water As an Entrainer. From the previous two industrial examples, it is concluded that the energy-savings can only be realized if the two columns in the extractive distillation system use the same steam grade. However, a heavy entrainer is often required in the extractive distillation system. Thus, the above situation can rarely be found in the literature. The only industrial example we can find is the acetone-methanol separation using water as the entrainer. Luyben28 developed a design flowsheet of this separation system via extractive distillation. This flowsheet is reproduced in Figure 12 with only minor differences in the reboiler duties of the two columns. The purity specifications of the two products are kept the same as the flowsheet in Luyben28 with acetone purity at 5390



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Figure 11. Design flowsheet of the extractive dividing-wall column for DMC and methanol separation.

Figure 12. Optimal design flowsheet of the conventional two-column system for acetone and methanol separation using water as an entrainer.

Figure 12. From the two design flowsheets, it appears that using DMSO as an entrainer is more economical for this separation system. One drawback of this design flowsheet in Figure 16 is that the DMSO is a much heavier entrainer. From the flowsheet, the bottom temperature at the C1 column is at 124.50 °C while that of the C2 column is at 194.72 °C. This reveals that low-pressure steam can be use in C1, but much more expensive high-pressure steam is required for C2. We also repeated the study by obtaining the flowsheets of the thermally coupled design and the dividing-wall column. The results are as predicted and summarized in Table 4. There is no benefit by further considering energy-saving options as opposed to the conventional design. One interesting observation by comparing this table to Table 3 is that the energy requirement of the dividing-wall column by using water as an entrainer is

wall configuration not only can save the space requirement but also can reduce the reboiler duty and the steam cost even more (9.72%) than that of the thermally coupled design. In this case, the benefit from increasing the rectifying section of the entrainer recovery column outperforms the pitfall from not optimizing the feed location. Note also that the TAC of the dividing-wall column is also a little less than that of the conventional design. 4.2. Using DMSO As an Entrainer. Luyben29 proposed to use a more effective entrainer (DMSO) for the acetonemethanol separation. We also reproduced this design flowsheet in Figure 16. The only difference is that the two product specifications are kept the same as in Figure 12 so that a direct comparison can be made. The total reboiler duty of 12140.75 KW using DMSO as an entrainer is much less than that of using water as an entrainer (18031.08 KW). The total numbers of stages of the two columns in Figure 16 are also less than that in 5391



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5.1. Dynamic Results of the Dividing-Wall Column. The difficult to reject unmeasured disturbance of feed composition changes will be considered in this study. One additional disturbance that needs to be considered for this design configuration is the vapor split ratio from the lower part of the column to the upper right or upper left part of the dividing wall. The mismatch of this vapor split ratio is inevitable because no regulation can be made inside of the column. The goal of the control strategy is to keep the two product purities as close as to the base case condition (acetone purity at 0.994 and methanol purity at 0.995) despite various disturbances. It is assumed that online composition measurement is either unavailable or troublesome with the maintenance issue, thus a tray temperature control strategy will be used to indirectly control the product purities. There are two pressure control loops and three level control loops in this system. Both condenser pressures on the top rightand left-side of the dividing-wall column are controlled by manipulating the respective condenser duty. Both reflux drum levels are controlled by manipulating the two product flow rates. An important inventory control loop is the bottom level of the dividing-wall column. The control of this level was adapted from the original extractive distillation system as suggested by Luyben28 and Grassi30 to be held by the entrainer makeup flow. However, because this flow is very small, thus the bottom level essentially floats indicating if the entrainer inventory in the system is balanced or not. With this control pairing, the entrainer feed flow to the dividing-wall column is ratio-controlled to the fresh feed flow rate. The production rate can be made to increase/decrease by changing the set point of the fresh feed flow loop. Another ratio control scheme is implemented to hold the vapor split ratio between the internal vapor flow up the column to the vapor sidedraw flow. Since in a real situation holding this ratio at a constant value may not be feasible, disturbance tests by varying this ratio value will be introduced in the dynamic simulation. There are three manipulated variables left which can be used in the tray temperature loop(s). They are the following: the reflux flow rates on the top right- and left-side of the dividing-wall column and the reboiler duty. A closed-loop sensitivity analysis is done to see which manipulated variable(s) can be fixed during unmeasured feed composition changes. The way to do the analysis is to introduce these disturbances into the system while keeping the two product purities and bottom composition the

Figure 13. Liquid composition profile of the extractive distillation column for acetone and methanol separation using water as an entrainer.

significantly less that the conventional design by using DMSO as an entrainer ($1466.80 × 103 vs $2240.14 × 103) even though the TAC is 10% higher. The design flowsheet in Figure 15 will be used in Section 5 as a base case to investigate the dynamics and control issues of EDWC in the face of various disturbances.

5. CONTROL OF AN EXTRACTIVE DIVIDING-WALL COLUMN Note from the above investigation, the only separation system that can benefit from the dividing-wall design is the acetonemethanol system using water as an entrainer. The energy requirement for the design flowsheet via Figure 15 is 9.72% less than that of the conventional design. This energy requirement is also significantly lower (34.5% less) than that of the conventional design using DMSO as an entrainer. The additional benefit of this design configuration is the space requirement of needing only one column. However, one important control degree-offreedom of this design is lost due to the integration of two reboilers into one. In the following, we will critically examine the dynamics and control of this process in the face of various disturbances. In the study, conventional control structure using multiloop PID control is assumed for wider industrial applications.

Table 3. Results Comparison of Three Designs for the Acetone and Methanol Separation Using Water As an Entrainer conventional two-column configurations annualized capital cost for column shell (1000 $/y) annualized capital cost for column tray (1000 $/y) annualized capital cost for reboiler (1000 $/y) annualized capital cost for condenser (1000 $/y) steam cost (1000 $/y) cooling water cost (1000 $/y) annualized cooler capital cost (1000 $/y) cooling water cost for cooler (1000 $/y) entrainer makeup (1000 $/y) total reboiler duty (KW) (% difference) total steam cost (1000 $/y) (% difference) TAC (1000 $/y) (% difference) a

C1

C2

988.57 183.48 156.05 403.90 984.47 32.07 80.18 4.51 18038.67 (0%) 1624.64 (0%) 4252.96 (0%)

335.15 43.69 141.42 236.33 640.17 22.98


C2

1005.01 470.28 187.37 59.87 0 246.21 404.85 210.40 0 1501.74 32.19 19.30 80.13 4.50 16673.56 (−7.57%) 1501.74 (−7.57%) 4222.43 (−0.72%)


C2

1165.69a 269.01b 232.47a 35.61b 0 270.96 403.94 191.39 0 1466.80 32.08 16.62 86.50 5.07 16284.60 (−9.72%) 1466.80 (−9.72%) 4176.77 (−1.79%)

Calculations based on upper section of the EDWC. bCalculations based on lower section of the EDWC. 5392



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Figure 14. Design flowsheet of the thermally coupled two-column system for acetone and methanol separation using water as an entrainer.

Figure 15. Design flowsheet of the extractive dividing-wall column for acetone and methanol separation using water as an entrainer.

same as in the base case by varying the above three manipulated variables. If one or more of the three manipulated variables can be fixed at a constant value or at a constant ratio to another process variable, this manipulated variable needs not to be used in the tray temperature control loop. The results of this analysis are summarized in Table 5. It is observed that all three manipulated variables (or some ratio) need to be changed in the face of feed composition disturbances. Thus, the control structure of three tray temperature loops will be designed next. Another observation foreseeing the problem with the threepoint temperature control strategy is that no matter there is +5% or −5% feed acetone composition changes the correct actions for the three manipulated variables are always to the increasing side.

This fact will actually require a sign of gain reversal for the tray temperature control loops which is impossible to achieve using conventional PID control. There will always be some product purity deviations by using a three-point temperature control strategy. The open-loop sensitivity analyses are performed to determine the temperature control points. When perturbing one manipulated variable, the other two manipulated variables are kept at base case values. The simulation is done using the flowsheet as in Figure 15 with an extractive distillation column with left-side condenser and without reboiler and a second column with rightside condenser and a reboiler. The stage number is counting from top to bottom with stage L1 as the left condenser and stage R1 as 5393



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Figure 16. Optimal design flowsheet of the conventional two-column system for acetone and methanol separation using DMSO as an entrainer.

Table 4. Results Comparison of Three Designs for the Acetone and Methanol Separation Using DMSO As an Entrainer conventional two-column configurations

C1

annualized capital cost for column shell (1000 $/y) annualized capital cost for column tray (1000 $/y) annualized capital cost for reboiler (1000 $/y) annualized capital cost for condenser (1000 $/y) steam cost (1000 $/y) cooling water cost (1000 $/y) annualized cooler capital cost (1000 $/y) cooling water cost for cooler (1000 $/y) entrainer makeup (1000 $/y) total reboiler duty (KW) (% difference) total steam cost (1000 $/y) (% difference) TAC (1000 $/y) (% difference) a

C2

481.55 69.51 109.36 216.50 979.28 12.25 124.50 17.04 3.33 12140.75 (0%) 2240.14 (0%) 3795.86 (0%)

242.73 29.55 118.17 120.63 1260.86 10.70


C2

494.95 308.32 72.00 37.35 0 196.04 218.50 107.79 0 2747.19 12.43 9.31 120.56 17.03 4.69 11652.17 (−4.02%) 2747.19 (22.63%) 4346.15 (14.50%)


C2

574.61a 218.84b 89.45a 28.22b 0 223.72 215.46 114.56 0 2826.22 12.16 9.40 124.64 17.93 2.42 11987.38 (−1.26%) 2826.22 (26.16%) 4457.62 (17.43%)


other two stages are selected to be stage L54 and stage R23. The logical manipulated variable for stage 66 is the reboiler duty. For stage L54 and R23, the left reflux flow and the right reflux flow are respectively selected. The overall control strategy for the dividing-wall column is shown in Figure 18. As for the controller tunings, all level loops are P-only controlled with Kc = 2.0. All pressure loops are tightly tuned with Kc = 20 and τI = 12 min. The controller tuning for the three crucial temperature loops are determined using iterative relayfeedback tests with Tyreus and Luyben tuning rules.31 The relayfeedback tests are performed iteratively one-at-a-time until the tuning parameters converged. The tuning constants for the temperature loop in the left-upper section are Kc = 5.5 and τI = 22 min; the tuning constants for the temperature loop in the rightupper section are Kc = 2.2 and τI = 21 min; and the tuning constants for the temperature loop in the bottom section are Kc = 9.1 and τI = 2.6 min. The closed-loop dynamic simulations with ±20% changes in the feed composition are shown in Figure 19. All three temperature control points are able to return back to their set points, although it takes about 25 h until the dynamic transients die-out. The two product flow rates are increased/decreased

Table 5. Results from Closed-Loop Sensitivity Analysis for the Dividing-Wall Column feed acetone composition +5% change

feed acetone composition −5% change

cases

base case

R1 (kmol/h) (% difference) R2 (kmol/h) (% difference) Qr (KW) (% difference) R1/D1 (% difference) R2/D2 (% difference) Qr/B (KW/kmol/ h) (% difference)

901.19 (0%) 238.56 (0%) 16284.60 (0%) 3.324 (0%)

989.108 (+9.76%)

970.062 (+7.64%)

280.727 (+17.68%)

246.562 (+3.34%)

17396.00 (+6.82%)

16866.80 (+3.58%)

3.484 (+4.81%)

3.759 (+13.09%)

0.881 (0%)

1.092 (+23.95%)

0.868 (−1.48%)

14.83 (0%)

15.830 (+6.74%)

15.365 (+3.61%)

the right condenser. The numbers without the prefix L or R are the stages in the lower part of the column without a dividing wall. The results of the open-loop sensitivity analyses are shown in Figure 17. The most sensitive and near linear stage in the stripping section of the dividing-wall column is at stage 66. The 5394



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Figure 17. Open-loop sensitivity analyses.

Figure 18. Overall control strategy of the dividing-wall column.

according to the unmeasured feed composition changes (see two bottom-right plots in Figure 19). The major problem is that the high-purities of the two products (see two bottom-left plots in Figure 19) are not able to maintain. During +20% feed acetone

composition, the acetone product purity is dropped from 0.994 to 0.985. The dynamic transient responses of the two product purities are also not favorable. The acetone purity dipped down to 0.957 during −20% feed acetone composition change and the 5395



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Figure 19. Closed-loop responses of the dividing-wall column with ±20% feed composition changes (solid: +20% acetone composition; dashed: −20% acetone composition).

entrainer makeup flow. The entrainer feed flow is also ratioed to the fresh feed flow as in the dividing-wall column. There are four manipulated variables which can be used in tray temperature loops. They are the following: reflux flow and reboiler duty of the extractive distillation column as well as the reflux flow and reboiler duty of the entrainer recovery column. Closed-loop sensitivity analysis is also done to see which manipulated variable(s) can be fixed during unmeasured feed composition changes. The results are summarized in Table 7. For the two manipulated variables in the extractive distillation column, it is found the least change is the reflux ratio (R1/D1). Thus, this ratio is fixed with the reboiler duty (Qr1) used by a tray temperature control loop. The same observation is found for the entrainer recovery column with fixing of the reflux ratio (R2/D2) and using the reboiler duty (Qr2) in another tray temperature control loop. Another important observation for Table 6 is that there is no problem now with the corresponding direction of the reboiler

methanol purity dipped down to 0.857 during +20% feed acetone composition change. Another unmeasured disturbance of vapor split ratio is also introduced in the dynamic simulation to test the control strategy. A summarized table with several different magnitudes of the above two unmeasured disturbances are shown in Table 6. The table shows the final purities of the two products with holding the three temperature control points at their set points. It is shown that the acetone purity can be dropped to 0.991 with −20% vapor split ratio change. 5.2. Dynamic Results of the Convention Two-Column System. For comparison purposes, the base case of the conventional two-column system in Figure 12 is also converted to Aspen Plus Dynamics. The control strategy used in the dynamic simulation is the same as in Luyben.28 An important inventory control loop is the bottom level of the entrainer recovery column which is controlled by manipulating the 5396



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The closed-loop dynamic simulations with ±20% changes in the feed composition are shown in Figure 20. The dynamic results are similar to the results in Luyben28 with the two temperature control points quickly returning to their set points in 3 h. The deviation of the acetone purity during +20% feed acetone composition change is also not as severe as in the results of dividing-wall column. The comparisons of the final steadystate results in the face of different magnitudes of the feed composition disturbances to that of the dividing-wall column are summarized in Table 6. The control performance particularly on the acetone product purity is much better than that of the dividing-wall column.

Table 6. Comparison of Control Performances of a DividingWall Column vs a Conventional Two-Column System configurations dividing-wall column disturbance feed acetone composition +5% feed acetone composition −5% feed acetone composition +10% feed acetone composition −10% feed acetone composition +20% feed acetone composition −20% vapor split ratio +10% vapor split ratio −10% vapor split ratio +20% vapor split ratio −20%

conventional two-column

acetone molar purity

methanol molar purity

acetone molar purity

methanol molar purity

0.992830

0.995284

0.993609

0.994779

0.994449

0.99486

0.994308

0.995129

0.991173

0.995579

0.993116

0.994437

0.995294

0.994477

0.994548

0.99518

0.985317

0.996201

0.991728

0.993015

0.995595

0.994025

0.994876

0.995047

0.994267

0.994881

-

-

0.992821

0.995250

-

-

0.994267

0.994881

-

-

0.990696

0.995348

-

-

6. CONCLUSIONS Although dividing-wall column is a promising technology to save energy and cut down the space requirement of a distillation system for separating ternary azeotrope, the energy-saving potential for the extractive dividing-wall column is limited. The main reason is because a heavy entrainer is required in the extractive distillation to effectively enhance the relative volatility of the original components. This fact most often results in cheaper steam grade being used in the extractive distillation column, but a more expensive steam grade is needed for the entrainer recovery column. Using three industrial systems as examples, important observations of this paper are summarized as below: 1. The overall reboiler duty can be saved by using the EDWC configuration. In our study, the savings of the overall reboiler duty are ranging from 1.26% to 13.53% for the four studied systems with larger savings to be expected for a system with more severe remixing effects. 2. It is found that only the acetone-methanol-water system can benefit from the more complicated EDWC configuration with both savings of the total reboiler duty and the steam cost. The evaluation of the energy-saving potential for a particular extractive distillation system can be done by just checking the two bottom temperatures in the conventional extractive distillation system. If the same steam grade can be used for both columns, savings in overall reboiler duty of EDWC can directly be translated to the savings in steam cost. 3. The above findings were using the unit price of steam in Seider et al. book26 for the calculations. The authors realize that the unit price of steam depends largely on the plant location, and it also varies in time. As such, they cannot be representative for the whole world. With information of the unit price of steam of the plant, it can easily be calculated that how much reboiler duty savings are needed to make the steam cost of the EDWC design break-even to that of the conventional design. Together with expected savings of the reboiler duty of less than 20% in these four systems, the decision on the economics of EDWC can be made. 4. Combining of two reboilers in EDWC loses one important control degree-of-freedom in the overall control strategy. It is illustrated that the control performance is hampered by this more complicated design. The reason is because coordinated control actions of the two manipulated variables (reboiler duties) for the conventional two-column system are needed to maintain product purity during feed composition disturbances. For example in the dynamic simulation of section 5, Qr1 has to be increased with the corresponding decrease of Qr2 during +5% feed acetone composition changes. This kind of coordinated actions of the two reboiler duties cannot be done by the dividingwall column with only one overall reboiler.

Table 7. Results from Closed-Loop Sensitivity Analysis for the Conventional Two-Column System cases R1 (kmol/h) (% difference) R2 (kmol/h) (% difference) Qr1 (KW) (% difference) Qr2 (KW) (% difference) R1/D1 (% difference) R2/D2 (% difference) R2/F2 (% difference) Qr1/B1 (KW/ kmol/h) (% difference) Qr2/B2 (KW/ kmol/h) (% difference)

feed acetone composition +5% change

feed acetone composition −5% change

899.29 (0%) 434.69 (0%) 10914.5 (0%) 7113.3 (0%) 3.319 (0%) 1.605 (0%) 0.318 (0%) 7.972 (0%)

983.48 (+9.36%)

825.57 (−8.2%)

416.51 (−4.18%)

450.89 (+3.73%)

11727.3 (+7.45%)

10177.4 (−6.75%)

6756.4 (−5.02%)

7427.9 (+4.42%)

3.456 (+4.13%)

3.208 (−3.34%)

1.605 (+0.92%)

1.585 (−1.26%)

0.307 (−3.22%)

0.326 (+2.7%)

8.652 (+8.53%)

7.361 (−7.67%)

6.477 (0%)

6.152 (−5.02%)

6.764 (+4.43%)

base case

duty to +5% or −5% changes in the feed acetone composition. By looking into the coordinated correct directions for the two reboiler duties, it is noted that Qr1 has to be increased with the corresponding decrease of Qr2 during +5% feed acetone composition changes. This kind of coordinated responses of the two reboiler duties cannot be done by the dividing-wall column with only one overall reboiler duty. The control points are determined again by the open-loop sensitivity analysis. The purpose for each column is to find a tray temperature control point so that the sensitivity to the reboiler duty is large and linear enough. The two temperature control points are stage 51 for the extractive distillation column and stage 21 for the entrainer recovery column. 5397



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Figure 20. Closed-loop responses of the conventional two-column system with ±20% feed composition changes (solid: +20% acetone composition; dashed: −20% acetone composition).

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(2) Triantafyllou, C.; Smith, R. The Design and Operation of Fully Thermally Coupled Distillation Columns. Trans. Inst. Chem. Eng. 1992, 70A, 118. (3) Wolff, E. A.; Skogestad, S. Operation of Integrated Three-Product (Petlyuk) Distillation Columns. Ind. Eng. Chem. Res. 1995, 34, 2094. (4) Annakou, O.; Mizsey, P. Rigorous Comparative Study of EnergyIntegrated Distillation Schemes. Ind. Eng. Chem. Res. 1996, 35, 1877. (5) Emtir, M.; Rev, E.; Fonyo, Z. Rigorous Simulation of Energy Integrated and Thermally Coupled Distillation Schemes for Ternary Mixture. Appl. Therm. Eng. 2001, 21, 1299. (6) Schultz, M. A.; Stewart, D. G.; Harris, J. M.; Rosenblum, S. P.; Shakur, M. S.; O’Brien, D. E. Reduce Costs with Dividing-Wall Columns. Chem. Eng. Prog. 2002, 98, 64. (7) Hernández, S.; Pereira-Pech, S.; Jiménez, A.; Rico-Ramírez, V. Energy Efficiency of an Indirect Thermally Coupled Distillation Sequence. Can. J. Chem. Eng. 2003, 81, 1087. (8) Gómez-Castro, F. I.; Segovia-Hernández, J. G.; Hernández, S.; Gutiérrez-Antonio, C.; Briones-Ramirez, A. Dividing Wall Distillation Columns: Optimization and Control Properties. Chem. Eng. Technol. 2008, 31, 1246.

AUTHOR INFORMATION

Corresponding Author

*Phone: +886-3-3366-3063. Fax: +886-2-2362-3040. E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The research funding from the National Science Council of the R. O. C. under grant no. NSC 100-2221-E-002-115-MY3 is greatly appreciated.

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REFERENCES

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Critical Assessment of the Energy-Saving Potential of an Extractive

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