Energy-Saving Dividing-Wall Column Design and Control for

Article pubs.acs.org/IECR

Energy-Saving Dividing-Wall Column Design and Control for Heterogeneous Azeotropic Distillation Systems Yi Chang Wu,† Hao-Yeh Lee,‡ Hsiao-Ping Huang,† and I-Lung Chien*,† †

Department of Chemical Engineering, National Taiwan University, Taipei 10617, Taiwan Department of Chemical Engineering , National Taiwan University of Science and Technology, Taipei 10607, Taiwan

‡

ABSTRACT: In this paper, the energy-saving potential of a heterogeneous azeotropic dividing-wall column is investigated by demonstrating an example for the separation of pyridine and water using toluene as entrainer. The original two-column system includes a heterogeneous azeotropic column with top decanter and another column served as preconcentrator column for the fresh feed and also served as recovery column for aqueous outlet stream from decanter. It is demonstrated that this complex twocolumn system can be thermally coupled into a dividing-wall column with top decanter. By comparing the optimized design of this dividing-wall column with the original design, a significant reduction (29.48%) in steam cost can be obtained. Furthermore, because important control degree-of-freedoms (two reboiler duties) are still preserved in this dividing-wall column, control performance of this proposed design is found to be comparable to that of the original two-column system.

1. INTRODUCTION The use of a dividing-wall column is a very promising process intensification technology, allowing significant energy reduction for the separation of mixtures with three or more components. Up to 30% energy savings for various case studies were found in open literature by utilizing 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 advanced controllers such as LQG/LQR, H∞, and model predictive control (MPC) have been proposed in open literature.8−15 A good complete reference for the subject of dividing-wall columns can be found in Yildirim et al.16 They gave 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 further energy-saving possibility of distillation systems for separating azeotropes. In a book discussing the design and control of such topic (cf. Luyben and Chien17), various separation methods were illustrated with real industrial examples. Further energy-saving for pressure-swing distillation systems can be obtained via a heat-integration method by combining the condenser of highpressure column with the reboiler of low-pressure column.18 For extractive distillation systems, operating either an extractive distillation column or the entrainer recovery column at high pressure will have an adverse effect on the separation. Instead, a dividing-wall column by thermally coupled the two columns can be designed to save reboiler duty. There are a number of recent papers in open literature discussed the design and control of extractive dividing wall column system.19−23 Wu et al.24 gave a critical assessment of the energy-saving potential of extractive dividing-wall column, using three industrial examples. It was found that, although the savings of the overall reboiler duty can be determined by using this © 2014 American Chemical Society

dividing-wall design, often the actual steam cost is adversely increased. The reason is because the dividing-wall design for the extractive distillation system needs to combine two reboilers into one. With the fact that a heavy entrainer to be typically used for extractive distillation system, a higher steam grade is often needed for the overall reboiler. Another drawback on the control performance was also demonstrated, because of the loss of one important control degree-of-freedom. This paper intends to closely investigate the energy-saving potential of dividing-wall columns for another completely different azeotropic separation method commonly used in industry. It is called heterogeneous azeotropic distillation system. Arifin and Chien25 studied a heterogeneous azeotropic system for separating isopropanol and water using cyclohexane as the light entrainer. However, this separation method resulted in significantly higher energy requirement than the competing separation method via extractive distillation.26 Kiss and Suszwalak27 also demonstrated similar results for the ethanol dehydration system comparing extractive distillation system vs heterogeneous azeotropic system. Thus, studying of dividingwall column for either isopropanol or ethanol dehydration system28 via heterogeneous azeotropic distillation is not worthwhile. In this paper, a pyridine and water separation system that was studied by Wu and Chien29 will be used as a demonstrating example to investigate the energy-saving potential of the dividing-wall column for heterogeneous azeotropic distillation systems. This separation system has difference residue curve maps (RCM) types, compared to the alcohol dehydration systems. To the best of our knowledge, there is no paper in the open literature that has studied the detailed design and control of dividing-wall column, which is for this type of heterogeneous Received: Revised: Accepted: Published: 1537

September 22, 2013 November 30, 2013 January 1, 2014 January 1, 2014 dx.doi.org/10.1021/ie403136m | Ind. Eng. Chem. Res. 2014, 53, 1537−1552

Industrial & Engineering Chemistry Research

Article

Figure 1. Conceptual design flowsheet and material balance lines of the original two-column system.

the NRTL thermodynamic model parameters were taken from Pommier et al.30 and are shown in Table 1. As shown in Figure 1, the pyridine−water mixture has an azeotrope with composition of ∼77 mol % H2O and an azeotropic temperature of 94.89 °C. By adding toluene into the system, two additional azeotropes are formed. One desirable azeotrope (toluene-water) is heterogeneous, with an azeotropic temperature of 84.53 °C, which is the minimum temperature for the entire ternary system. Another azeotrope between pyridine and toluene is also formed with a higher azeotropic temperature of 110.15 °C. As shown in Figure 1, the distillation boundaries divide the residue curve maps (RCMs) of the ternary system into three distillation regions. The idea of the conceptual design is to let column C-2 serve as a preconcentrator column for the fresh feed; it also serves as a recovery column for the aqueous outlet stream. With this

azeotropic distillation system with this RCM type and also with one column served as the preconcentrator column and also as the recovery column. Besides the issue on the energy-saving, dynamic controllability of the dividing-wall column will also be closely examined. A comparison of the control performance will be made to that of the original heterogeneous azeotropic column system. Conventional tray-temperature control strategy will be used for wider industrial applications.

2. ORIGINAL TWO-COLUMN SYSTEM Wu and Chien29 studied a pyridine and water separation system via heterogeneous azeotropic distillation, using toluene as an entrainer. The conceptual design flowsheet and material balance lines are shown in Figure 1. In their simulation study, 1538

dx.doi.org/10.1021/ie403136m | Ind. Eng. Chem. Res. 2014, 53, 1537−1552


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Table 1. NRTL Model Parametersa for the Pyridine Dehydration System Comp. i

Comp. j

Comp. i

Comp. j

Comp. i

Comp. j

water

toluene

water

pyridine

toluene

pyridine

source aij aji bij (K) bji (K) cij eij eji a

Aspen LLE 627.0528 −247.8792 −27269.3555 14759.7598 0.2 −92.7182 35.582

Pommier et al. 0 0 895.315 209.4203 0.6932 0 0

Pommier et al. 0 0 133.173 −30.3653 0.2992 0 0

Aspen Plus NRTL is described as follows: ln γi =

∑j xjτjiGji ∑k xkGki

⎡

+ ∑j

xjGij ⎢τ ∑k xkGkj ⎢ ij

⎣

−

∑m xmτmjGmj ⎤ ⎥ ∑j xkGkj ⎥⎦

where Gij = exp(−αijτij), with τij = aij + bij/T + eij ln T, αij = cij, τii = 0, and Gii = 1.

Figure 2. Optimal design of the original two-column design.

3. ENERGY-SAVING DESIGNS 3.1. Stripping-Column Design. An alternative design to further save energy in the heterogeneous azeotropic distillation system was proposed in Chang et al.32 by incorporating a stripping column in their three-column system for isopropyl alcohol dehydration. Their idea, from the perspective of energy conservation, of a stripping column is a tower without a condenser, in which energy carried by the overhead vapor is conserved instead of being removed in a condenser, and is considered to be more energy-efficient than the conventional recovery column. This design concept is adapted here for this pyridine and water system. The resulting flowsheet with the same feed

design, composition of the V1 stream does not need to be close to the toluene−water binary heterogeneous azeotrope. Note that it is very difficult to reach this binary azeotrope by going through a very narrow funnel. As shown in the material balance lines of Figure 1, pyridine loss through the aqueous outlet stream can be prevented by purifying this stream further at the C-2 column. The optimized design flowsheet by minimizing total annual cost (TAC) can be seen in Figure 2 using Aspen Plus31 simulation. The two product purity specifications (pyridine and water) are all set at 99.9 mol %. In the next section, energysaving designs will be developed with the same feed conditions and product specifications as in Figure 2 for direct comparison. 1539



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Figure 3. Stripping-column design.

the system be reduced, the number of columns can also be reduced to a single column shell to save space at the plant site. 3.2. Dividing-Wall Column Design. The design concept can be seen in Figure 4 with the upper flowsheet showing the thermally coupled design and the lower flowsheet showing the equivalent dividing-wall column design. The way to thermally couple the two columns is to eliminate the condenser at top of the recovery column by putting the top vapor directly back to the heterogeneous azeotropic column. In order to provide the liquid flow into the recovery column (as reflux in the original two-column design), a liquid sidedraw is designed to drawoff at the same stage of heterogeneous azeotropic column to the recovery column. This way, the fresh feed, as well as the aqueous outlet stream, do not need to be restricted to entering the C-2 column at top stage. In order to have the additional benefit of combining two columns into only a single column shell to save space requirement, the number of stages for the heterogeneous azeotropic column below the liquid sidedraw stage should be the same as the total number of stages of column C-2. In this way, the conceptual design flowsheet can be seen as in the lower flowsheet of Figure 4 with a dividing wall at lower part of a single column. For simplicity purposes, we assume that the optimized total numbers of stages of the original two-column system are preserved in this study. That is, the NT of the heterogeneous azeotropic column is kept at 14 and that of the recovery column is kept at 10. Since the lower part of the dividing-wall column should have same number of stages on the rightside and leftside of the dividing wall, the liquid draw location is set

condition and product purity specifications is shown in Figure 3. Note that the feed location for this stripping column must be at the first stage. In the original two-column system, the distillate composition of the combined column was optimized to be at 21 mol % pyridine. However, for the stripping column configuration, the design degree-of-freedom at the top of the conventional recovery column is lost. Thus, the top vapor composition for this stripping column is actually an output variable (at 18.96 mol % pyridine). By comparing the overall reboiler duty of the original twocolumn system with this stripping-column design, it is found that significant savings can be realized (17180.9 kW vs 12740.4 kW). However, one potential deficiency of this design is for the vapor recycle from the stripping column to the heterogeneous azeotropic column. In order to ensure that the vapor flow would travel as designed is to add a small compressor (as in the simulation flowsheet) to raise the top vapor pressure of the stripping column to 1.09 atm, so that this pressure is the same as the pressure at stage 6 of column C-1. Of course, an alternative way is to raise the operating pressure at the stripping column to ensure that the top vapor pressure is always higher than the pressure at stage 6 of column C-1. However, raising operating pressure will make the separation in column C-2 more difficult. We did not investigate the effect of raising the operating pressure of column C-2 for the purpose of eliminating the use of an expensive compressor. Instead, a thermally coupled design will be proposed in the next section. This thermally coupled design can further be thought as a dividing-wall column. In this way, not only can the energy requirements of 1540



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Figure 4. Illustration of the thermally coupled design and the equivalent dividing-wall column.

operating cost and capital cost by the new design. Conversely, a design with less total number of stages in the dividing-wall column, will result in reducing the capital cost; however, the reboiler duty will be increased, comparing to the case by keeping the total number of stages the same. This tradeoff between the capital cost and the operating cost will complicate the benefit of the new design versus the original design. Figure 5 displays the simulation results by varying both liquid sidedraw flow rate (L1) and the feed location (NF3). The results show that the total reboiler duty is minimized when L1 is set to be 65 kmol/h and the NF3 is set to be 5. With L1 set to be greater than 65 kmol/h, the two product purities cannot be maintained at their specifications. The resulting optimized flowsheet for the dividing-wall column design is shown in Figure 6. Note that the overall reboiler duty is further reduced to 12116.3 kW. This represents a significant 29.48% reduction

to be the fourth stage (from the top). The only other two variables needed to be determined for the proposed design are the liquid sidedraw flow rate and the feed location of the fresh feed (and the aqueous outlet stream). The left reboiler duty is used to maintain pyridine product purity at 99.9 mol % and the right reboiler duty is used to maintain water product purity at 99.9 mol %. The toluene makeup flow is set to balance out the tiny entrainer loss through the two product streams. The resulting total reboiler duty will be used as the objective function to determine the optimal liquid sidedraw flow rate and the feed location of the fresh feed. The reason for keeping the total numbers of stages the same is because any savings of the energy requirements by the dividing-wall column design will also directly translate into decreasing of the column diameter, because of a lesser vapor rate. This way, we can clearly demonstrate the savings of both 1541



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Figure 5. Plots for determining values L1 and NF3 of the dividing wall column.

in the reboiler duty, compared with the original two-column system. Because it is assumed that only one column shell is required for this dividing-wall column, the equivalent diameter (De) of the column shell at the lower part of the dividing-wall column is calculated as the illustration in Figure 7. It is assumed that the De value is back-calculated, so that a total cross-sectional area from both sides of the dividing wall can be provided. For this dividing-wall column, there are two parts with different diameters. The upper part has a diameter of 2.30 m, and the lower part has an equivalent diameter of 2.42 m. These values will be used to calculate the capital cost for the upper and lower parts of column shell and trays, for comparison purposes with the original two-column system and the stripping-column design. For practical consideration, by keeping the diameter of this dividing-wall column the same for both the upper and

Figure 7. Illustration of how to calculate the equivalent diameter (De) on the lower part of dividing-wall column.

lower parts, a unified diameter between 2.30 and 2.42 m can alternatively be designed. The liquid and vapor composition profiles for column C-1 of the original two-column system are shown in Figure 8 and compare with the dividing-wall column system (C-1 plus C-2 in Figure 6). It is shown that both designs serve the purpose of carrying large amounts of water to the column top (see vapor composition profiles). Note that there is no observation of remixing effect at upper part of column C-1 in the original twocolumn system. Thus, the energy saving is not due to the elimination of remixing effect for the column sequence, but rather by a close thermal coupling of the original two columns.

Figure 6. Proposed dividing-wall column design. 1542



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Figure 8. Liquid and vapor composition profiles of the heterogeneous azeotropic systems: (a) original two-column design and (b) dividing-wall column design.

Table 2. Results Comparison of the Three Designs for the Pyridine Dehydration System Original Two-Column System

Stripping-Column Design

Dividing-Wall Column Design

configuration

C1

C2

C1

C2

C1

C2 + C3

annualized capital cost for column shell (× 103 $/yr) annualized capital cost for column tray (× 103 $/yr) annualized capital cost for reboiler (× 103 $/yr) annualized capital cost for condenser (× 103 $/yr) annualized capital cost for compressor (× 103 $/yr) steam cost (× 103 $/yr) cooling water cost (× 103 $/yr) electricity cost (× 103 $/yr) makeup flow (× 103 $/yr) total reboiler duty (kW) (× 103 difference) total steam cost (× 103/yr) (% difference) TAC (× 103 $/yr) (% difference)

270.57 35.91 217.83 168.46

147.07 15.83 234.54 130.85

254.64 32.88 139.93 191.99

141.66 14.99 227.10

57.59 5.10 112.32 188.01

185.04 22.11 236.44

742.61 27.45

832.03 22.78

375.87 35.47

268.04 33.66

842.42

3.71 17180.9 (0%) 1574.64 (0%) 2849.64 (0%)

3.3. Comparison. In this section, TAC, reboiler duty, and steam cost of the above two cases will be compared with those of the original two-column system. TAC calculations includes annual total operating cost, total installed capital cost divided by a payback period, and also a stream cost for small makeup flow. Operating cost includes the costs of steam and cooling water, and capital cost includes the costs of columns, trays, reboilers, and condensers. The capital cost for the internal wall is assumed to be negligible, in comparison with the other capital costs. The payback period is assumed to be three years, and the wastewater treatment cost is also assumed to be negligible in the overall TAC. For the stripping-column design,

56.52 791.80

8.93 3.71 12740.4 (−25.85%) 1167.67 (−25.85%) 2275.49 (−20.15%)

3.71 12116.3 (−29.48%) 1110.46 (−29.48%) 1954.44 (−31.41%)

additional costs associated with compressor and electricity are also included. The formulas for the installed capital costs for all the process equipments can be found on pages 572−575 in Appendix D of Douglas’ book, Conceptual Design of Chemical Processes.33 In calculating the capital costs, an M&S index of 1536.5 was used. For the heterogeneous azeotropic distillation system, low-pressure steam (50 psig at 147 °C) is commonly used, because a light entrainer is often used for this type of separation system. The unit prices of the steam at $6.60/1000 kg and cooling water at $0.02/m3 were adopted from Table 23.1 of Seider et al.’s book, Product and Process Design Principles Synthesis, Analysis, and Evaluation.34 1543



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Figure 9. Overall control strategy of original two-column system.

sacrifice needs to be made. In the control study, only conventional control structure by using multiloop traytemperature control will be implemented for wider industrial applications. Note that, in the extractive dividing-wall column system of Wu et al.,24 the control performance was hampered because of losing one important degree-of-freedom (a reboiler duty). It is not the case here, because both reboiler duties are still preserved in this energy-saving design. 4.1. Dynamic Results of the Original Two-Column System. Wu and Chein29 studied the overall control strategy of the two-column system in Figure 2. The overall control strategy is summarized in Figure 9. The inventory control loops were designed as follows. For column C-1 with the decanter, the aqueous phase level was controlled by manipulating the aqueous outlet flow; the column bottom level was controlled by manipulating the bottom pyridine product flow; the column top pressure was controlled by manipulating the top vapor flow; and the organic phase level was controlled by entrainer makeup flow to avoid having possible snowballing effect in the recycle system. For column C-2, the reflux drum level was controlled by the distillate flow (feed to heterogeneous azeotropic column); the column bottom level was controlled by the bottom water product flow; and the column pressure was controlled by the condenser duty. With this inventory control strategy, the remaining manipulated variables for column C-1 are the organic reflux flow and the reboiler duty, and the remaining manipulated variables for column C-2 are the reflux flow and the reboiler duty. The suggested temperature control strategy for columns C-1 and C-2 was to hold the ninth and eighth stage

Table 2 shows the results of the original two-column system in Figure 2, the stripping-column design in Figure 3 and the dividing-wall column design in Figure 6. It is shown that the overall reboiler duty of the proposed dividing-wall column design is significantly less than that of the original two-column design (29.48% reduction). This significant reduction can directly be reflected as the reduction of the steam cost. Note that for the study of extractive dividing-wall column in Wu et al.,19 it was found that the savings on the reboiler duty cannot be directly reflected to the savings on the steam cost. Conversely, the steam cost of the extractive dividing-wall column was usually higher than the original two-column system. As for the TAC comparison, the proposed dividing-wall column design also shows significant reduction than the original two-column system. The large reduction of 31.41% in TAC is mainly contributed by the savings in the steam cost, as well as the column shell. Note that the TAC of the original two-column system in Table 2 is slightly higher than a table in Wu and Chien29 The reason is because more updated M&S index is used here in the calculation. By comparing the results of the proposed dividing-wall column design with strippingcolumn design, TAC and the steam cost are all lower for the proposed design. An additional benefit involves the space requirement at the plant site, because only a single-column shell needs to be used for the proposed design.

4. CONTROL OF DIVIDING-WALL COLUMN SYSTEMS In this section, dynamics and control of the proposed design flowsheet in Figure 6 will be examined, to determine if any 1544



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Figure 10. Closed-loop responses of original two-column system with ±20% feed composition changes (solid trace: +20% pyridine composition; dashed trace: −20% pyridine composition). 1545



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Figure 11. Original two-column system with ±20% feed flow rate changes.

temperatures, respectively, by manipulating their reboiler duties. For the other two free manipulated variables, the ratio

control scheme was designed so that these two free manipulated variables can be adjusted in the face of 1546



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Figure 12. Open-loop sensitivity analyses.

disturbances. The organic reflux flow was set to maintain a constant ratio to the feed flow of column C-1 and the reflux ratio of column C-2 was also maintained. The control performances for the feed composition and feed flow rate disturbances are shown in Figures 10 and 11. These closed-loop results were done by using Aspen Dynamics via pressure-driven simulation. It was shown that both products can be maintained at high purity, despite large ±20% disturbance changes. These control performances will be compared with the closed-loop results of the dividing-wall column system under same disturbances. Note for the feed flow rate changes as considered to be measured disturbances, a feedforward scheduling scheme by adjusting the tray-temperature set points at different feed rates can be implemented to eliminate small deviations of the final product compositions. The authors29 did not show the improved results for the feed rate disturbances. 4.2. Dynamic Results of the Dividing-Wall Column System. Similar inventory control strategy is adapted for the dividing-wall column system except for the organic phase level. In the original two-column system, the organic phase level is controlled by manipulating a small entrainer makeup flow. Another important manipulated variable affecting the organic phase level is the organic reflux flow. This flow was set to maintain a constant ratio to the feed flow of column C-1 (an internal flow of the overall system). Using this ratio scheme, the organic reflux flow rate can be adjusted in the face of unmeasured feed composition disturbance. However, for the dividing-wall column system, this internal flow is realistically considered as not measurable, so that the above ratio scheme cannot be implemented. Therefore, the organic phase level in the dividing-wall column system is set to be controlled by the organic reflux flow. The small entrainer makeup flow is set to maintain at a constant ratio with the fresh feed. The original

worry of using organic reflux flow to control the organic phase level may cause snowball effect will be examined in the closedloop simulations. For the dividing-wall column system, there are only two remaining manipulated variables (two reboiler duties). Conventional tray-temperature control strategy will be implemented to select two control points to pair with these two manipulated variables. Another variable which was considered by other dividing-wall column papers as possible manipulated variable (liquid split ratio) is more realistically considered here to be a disturbance assuming that the actual liquid split ratio may not be the same as the designed ratio. The open-loop sensitivity analyses are performed to determine the temperature control points. When perturbing one manipulated variable, the other manipulated variable is kept at base case values. The results of the open-loop sensitivity analyses are shown in Figure 12. The stage number is counting from top to bottom with stage 1 as the top tray of section C-1 in Figure 6, L1 as the top tray of section C-2 on the left side of the dividing wall and R1 as the top tray of section C-3 on the right-hand side of the dividing wall. The most sensitive and near-linear stage on the left-hand side of the dividing wall is at stage L8 and the most sensitive and near-linear stage on the right-hand side of the dividing wall is at stage R12. Since, dynamically, the left reboiler duty should affect the temperature on the left-hand side of the dividing wall first, the pairing is easily determined by using left reboiler duty to hold the L8 tray temperature and using the right reboiler duty to hold the R12 tray temperature. The PI tuning constants for the two traytemperature loops are determined by using the relay feedback test provided in Aspen Dynamics with Tyreus and Luyben tuning rules.35 The overall control strategy for the dividing-wall column is shown in Figure 13. Note that there is a fictitious compressor at 1547



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Figure 13. Overall control strategy of dividing-wall column.

are maintained at high purity with corresponding changes of the production rates. Note again with these disturbances, the snowballing effect is again avoided. The organic reflux flow is correspondingly increased/decreased with the feed flow rate changes. The third disturbance considered in this paper is the liquid split ratio. Let us assume that the actual ratio is different than the designed ratio by either +10% or −10%. The closed-loop dynamic simulations with these two unmeasured disturbances are shown in Figure 16. Again, the control performances are all very satisfactory with both products still maintained at high purity.

the top of section C-3 to ensure vapor traffic can correctly be flowed from section C-3 (which is the right-hand side of the dividing wall) to section C-1. This type of arrangement was commonly used in the open literature (e.g., Ling and Luyben9) for dynamic study of the dividing-wall column systems. In reality, the vapor flow will naturally move from the right-side of the dividing wall (section C-3) to the upper section (section C1), because of the pressure drop. The closed-loop dynamic simulations with ±20% changes in the feed composition are shown in Figure 14. Both temperature control points are able to quickly return back to their setpoints. The two product flow rates (B2 and B3) are increased/ decreased according to the unmeasured feed composition changes. Both products are maintained at high purity, despite these unmeasured disturbances. The only minor problem is on the dynamic transient response for the pyridine product purity. The pyridine purity dipped down to 0.991 during −20% feed pyridine composition change but quickly return to higher purity of over 0.998. From Figure 14, it is also demonstrated that snowballing effect can still be avoided with this proposed control strategy because the organic reflux flow is steadily maintained at new value under a particular disturbance change. The second disturbance considered is the throughput changes. This can be achieved by changing the feed flow rate. Figure 15 displays the closed-loop dynamic simulations with ±20% changes in the feed flow rate. Note that both products

5. CONCLUSIONS In this paper, dividing-wall column design for heterogeneous azeotropic distillation system is proposed. It is demonstrated by a pyridine dehydration example that the original two-column system can be thermally coupled into a dividing-wall column with a single-column shell. It is found that significant savings (29.48%) on the reboiler duty can be realized by the proposed design. Unlike extractive dividing-wall column systems, the savings on the reboiler duty can directly be reflected as the reduction of the steam cost. Significant reduction on the total annual cost (31.41%) can also be realized with the additional benefit of having only one column shell. The result is general 1548



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Figure 14. Closed-loop responses of a dividing-wall column with ±20% feed composition changes (solid trace: +20% pyridine composition; dashed trace: −20% pyridine composition). 1549



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Figure 15. Dividing-wall column with ±20% feed flow rate changes.

and should be able to apply to other heterogeneous azeotropic distillation systems.

As for the control performance, the original two important control degree-of-freedoms (two reboiler duties) are still 1550



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Figure 16. Dividing-wall column with ±10% liquid split ratio changes.

preserved in the proposed design. A conventional traytemperature control strategy can still be implemented to

manipulate these two reboiler duties. It is demonstrated that the two products can still be maintained at high purity, despite 1551



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feed composition, feed flow rate, and liquid split ratio disturbances.

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(20) Bravo-Bravo, C.; Segovia-Hernández, J. G.; Gutiérrez-Antonio, C.; Durán, A. L.; Bonilla-Petriciolet, A.; Briones-Ramirez, A. Extractive Dividing Wall Column: Design and Optimization. Ind. Eng. Chem. Res. 2010, 49, 3672. (21) Ibarra−Sánchez, J. J.; Segovia-Hernández, J. G. Reducing Energy Consumption and CO2 Emissions in Extractive Cistillation, Part II: Dynamic Behavior. Chem. Eng. Res. Des. 2010, 88, 135. (22) Murrieta-Dueñas, R.; Gutiérrez-Guerra, R.; Segovia-Hernández, J. G.; Hernández, S. Analysis of Control Properties of Intensified Distillation Sequences: Reactive and Extractive Cases. Chem. Eng. Res. Des. 2011, 89, 2215. (23) Kiss, A. A.; Ignat, R. M. Innovative Single Step Bioethanol Dehydration in an Extractive Dividing-Wall Column. Sep. Purif. Technol. 2012, 98, 290. (24) Wu, Y. C.; Hsu, P. H. C.; Chien, I. L. Critical Assessment of the Energy-Saving Potential of an Extractive Dividing-Wall Column. Ind. Eng. Chem. Res. 2013, 52, 5384. (25) Arifin, S.; Chien, I. L. Combined Preconcentrator/Recovery Column Design for Isopropyl Alcohol Dehydration Process. Ind. Eng. Chem. Res. 2007, 46, 2535. (26) Arifin, S.; Chien, I. L. Design and Control of an Isopropy Alcohol Dehydration Process via Extractive Distillation Using Dimethyl Sulfoxide as an Entrainer. Ind. Eng. Chem. Res. 2008, 47, 790. (27) Kiss, A. A.; Suszwalak, D. J-P.C. Enhanced Bioethanol Dehydration by Extractive and Azeotropic Distillation in DividingWall Columns. Sep. Purif. Technol. 2012, 86, 70. (28) Sun, L. Y.; Chang, X. W.; Qi, C. X.; Li, Q. S. 2011, Implementation of Ethanol Dehydration Using Dividing-Wall Heterogeneous Azeotropic Distillation Column. Sep. Sci. Technol. 2011, 46, 1365. (29) Wu, Y. C.; Chien, I. L. Design and Control of Heterogeneous Azeotropic Column System for the Separation of Pyridine and Water. Ind. Eng. Chem. Res. 2009, 48, 10564. (30) Pommier, S.; Massebeuf, S.; Kotai, B.; Lang, P.; Baudouin, O.; Floquet, P.; Gerbaud, V. Heterogeneous Batch Distillation Processes: Real System Optimisation. Chem. Eng. Process. 2008, 47, 408. (31) Aspen Plus, Release: AspenONE; Aspen Technology, Inc.: Cambridge, MA, 2009. (32) Chang, W. T.; Huang, C. T.; Cheng, S. H. Design and Control of a Complete Azeotroopic Distillation System Incorporating Stripping Columns for Isopropyl Alcohol Dehydration. Ind. Eng. Chem. Res. 2012, 51, 2997. (33) Douglas, J. M. Conceptual Design of Chemical Processes; McGraw−Hill: New York, 1988. (34) Seider, W. D.; Seader, J. D.; Lewin, D. R.; Widagdo, S. Product and Process Design Principles Synthesis, Analysis, and Evaluation; Wiley: Hoboken, NJ, 2010. (35) Tyreus, B. D.; Luyben, W. L.; Tuning, P. I. Controllers for Integrator/Deadtime Processes. Ind. Eng. Chem. Res. 1992, 31, 2625.

AUTHOR INFORMATION

Corresponding Author

*Tel.: +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 R.O.C. is greatly appreciated. REFERENCES

(1) Glinos, K. A.; Malone, M. F. Optimality Regions for Complex Column Alternatives in Distillation Systems. Chem. Eng. Res. Des. 1988, 66, 229. (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 ThreeProduct (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. (9) Ling, H.; Luyben, W. L. New Control Structure for Divided-Wall Columns. Ind. Eng. Chem. Res. 2009, 48, 6034. (10) Ven Diggelen, R. C.; Kiss, A. A.; Heemink, A. W. Comparison of Control Strategies for Dividing-Wall Columns. Ind. Eng. Chem. Res. 2010, 49, 288. (11) Buck, C.; Hiller, C.; Fieg, G. Applying Model Predictive Control to Dividing Wall Columns. Chem. Eng. Technol. 2011, 34, 663. (12) Kiss, A. A.; Bildea, C. S. A Control Perspective on Process Intensification in Dividing-Wall Columns. Chem. Eng. Process. 2011, 50, 281. (13) Kiss, A. A.; Rewagad, R. R. Energy Efficient Control of a BTX Dividing-Wall Column. Comput. Chem. Eng. 2011, 35, 2896. (14) Ling, H.; Cai, Z.; Wu, H.; Wang, J.; Shen, B. Remixing Control of Divided-Wall Columns. Ind. Eng. Chem. Res. 2011, 50, 12694. (15) Rewagad, R. R.; Kiss, A. A. Dynamic Optimization of a DividingWall Column Using Model Predictive Control. Chem. Eng. Sci. 2012, 68, 132. (16) Yildirim, Ö .; Kiss, A. A.; Kenig, E. Y. Dividing Wall Columns in Chemical Process Industry: A Review on Current Activities. Sep. Purif. Technol. 2011, 80, 403. (17) Luyben, W. L.; Chien, I. L. Design and Control of Distillation Systems for Separating Azeotropes; Wiley: Hoboken, NJ, 2010. (18) Luyben, W. L. Design and Control of a Fully Heat-integrated Pressure-swing Azeotropic Distillation System. Ind. Eng. Chem. Res. 2008, 47, 2681. (19) Gutiérrez-Guerrz, R.; Segovia-Hernández, J. G.; Hernández, S. Reducing Energy Consumption and CO2 Emissions in Extractive Cistillation. Chem. Eng. Res. Des. 2009, 87, 145. 1552


Energy-Saving Dividing-Wall Column Design and Control for

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