Comparison between Different Configurations of Internally and

Article pubs.acs.org/IECR

Comparison between Different Configurations of Internally and Externally Heat-Integrated Distillation by Numerical Simulation Dawei Chen, Xigang Yuan,* Lianghua Xu, and K. T. Yu State Key Laboratory of Chemical Engineering and School of Chemical Engineering and Technology, Tianjin University, Tianjin 300072, China ABSTRACT: The internally heat-integrated distillation column (HIDiC) is often compared with heat-pump-assisted distillation columns and claimed to be more energy efficient. Recently, the idea of simplifying the HIDiC by several external heat exchangers has come into favor and been thought to be a more practical configuration with greatly reduced capital costs. To investigate the energy and economic performances of heat-pump-assisted distillation along with the HIDiC and simplified HIDiC, comparisons based on numerical simulation were made in the present work. A propylene−propane splitter was chosen for the comparison. Compression ratios and total numbers of stages were optimized to achieve an optimal configuration in terms of total annualized costs for each scheme before comparison with others. For the HIDiC and its simplifications, the impacts of different heat distribution patterns were also considered. It was found that the results depend strongly on the basis selected for comparison, and some discussions and remarks are also presented.

1. INTRODUCTION Distillation is the most well-established and widely used separation technique in process industries, but at the same time, it is known to be a large energy consumer with low thermodynamic efficiency. Therefore, improving the energy efficiency of distillation would bring substantial benefits and has been accorded high priority. In a conventional distillation column, heat is supplied at the bottom reboiler and rejected at the top condenser, which means that the distillation process is actually operated by the degradation of energy from bottom to top. A straightforward approach to energy saving would involve transferring the heat rejected from the rectifying section to the stripping section of the same column. Such an approach, however, requires the heat rejected to be upgraded to a higher temperature level by compression. Many distillation schemes based on this conception have been proposed, among which the heat pump approach1 might be the simplest external heat-integrated configuration. In a typical heat pump system, called direct vapor recompression (VRC, Figure 1a), the vapor exiting the top of a distillation column is compressed to a desired pressure high enough to form a positive temperature difference with respect to the bottom of the same column, so that the upgraded vapor could be used as a hot utility to drive the bottom reboiler before being condensed and circulating back to the column. Another alternative scheme, called the internally heatintegrated distillation column (HIDiC, Figure 1b), combines the concept of a heat pump and the advantage of diabatic distillation, and its research has drawn much attention over the past 20 years.2,3 Rather than transferring heat from the topmost to the bottommost location in a distillation column as in the VRC configuration, heat transfer in a HIDiC takes place gradually along the column between the stripping and rectifying sections and causes continuous evaporation and condensation, respectively, in the two sections. To maintain a positive temperature difference for the heat transfer, the rectifying section has to be operated at a higher pressure than the © 2013 American Chemical Society

stripping section, but with a lower compression ratio than that for VRC, as compression across the whole temperature span of the column is no longer necessary. The concept of internal heat integration was thoroughly evaluated by Mah et al.4 under the name of secondary reflux and vaporization (SRV) as early as 1977, and even earlier conceptual origins can be found in the literature.5,6 Persistent studies of the HIDiC scheme have been made since 1990s, and reviews can be found in the works of Nakaiwa et al.2 and Olujic et al.3 As a result of such studies, the superiority of the HIDiC over conventional distillation columns has been demonstrated by some successful pilot-plant experiments.7,8 As the simplest application of heat-pump-assisted distillation, VRC has been applied in the separations of close-boiling mixtures and found to bring substantial energy saving.9 In contrast, the HIDiC, although it has been confirmed to be thermodynamically more efficient than VRC,10 is still on the way toward industrialization. One of the difficulties with HIDiCs is the complexity of implementing efficient heat exchange between the two column sections, regardless of the use of concentric, multitube, split,11 or compact12 configurations. Additional difficulties lie in the need to improve process controllability,13,14 especially when process irreversibility is further reduced by heat integration between the vapor distillate and feed.15,16 To avoid the major difficulty of accommodating enough heat-transfer devices within the limited space inside a HIDiC,11 the idea of a simplified HIDiC was investigated by Suphanit17,18 and also independently developed by Chen et al.19 under the name SIHIDiC. The core idea of the simplified HIDiC (or SIHIDiC as denoted by Chen et al.) is to approximate its Received: Revised: Accepted: Published: 5781

January 10, 2013 March 15, 2013 March 19, 2013 March 19, 2013 dx.doi.org/10.1021/ie400112k | Ind. Eng. Chem. Res. 2013, 52, 5781−5790

Industrial & Engineering Chemistry Research

Article

design or analysis of a HIDiC or VRC. A review of previous works on the analysis of HIDiC and/or VRC configurations3,22−26 reveals that comparisons between the VRC and HIDiC schemes under the same minimum temperature driven force for heat transfer are significantly different from those under different temperature driven forces (equivalently, pressure differences) and that the impact of the number of trays on the comparison has not been sufficiently considered. In the present work, a number of simplified HIDiC schemes are proposed and analyzed, and comparisons are made among VRC and the various HIDiC schemes on the basis of the same temperature and total number of stages. The study was performed for a propylene−propane mixture, which is a representative close-boiling system that is extensively used in the literature for the same purposes. A series of base cases were first simulated to compare the energy consumptions of all schemes (VRC, HIDiC, and simplified HIDiC) under the same basis. Then, both variables were allowed to change, so that performances under different numbers of theoretical stages and compression ratios could be thoroughly compared using the total annualized cost (TAC) as the criterion.

2. CONFIGURATIONS OF VARIOUS SCHEMES TO BE CONSIDERED The separation task of producing polymer-grade propylene from a mixture mainly composed of propane and propylene was selected for the case study, as it is widely accepted to be among the most promising applications of both VRC and HIDiC schemes because of its extremely low relative volatility. The feed composition and conditions are the same as those used by Olujic et al.,23 which is actually taken from a state-of-the-art VRC propylene−propane splitter in commercial use. For comparison, distillates in all configurations are drawn as vapor. A pressure drop of 6.2 mbar per stage was universally employed. Simulations were conducted using the software package Aspen Plus, version 11.1, in combination with a spreadsheet program. The Peng−Robinson model27 was used as the thermodynamic method. A conventional distillation column was first simulated to determine the optimum feed stage for each total number of theoretical stages using a sensitivity study block provided in the package, and the feed positions of VRC, HIDiC, and all simplified HIDiC configurations were determined accordingly. In the VRC scheme, a cooler was placed right after the throttling valve to partially condense the flashing recycling stream before it enters the column. The vapor fraction of its outlet stream is thereby controlled to make sure that the state of the heat pump working stream exiting the reboiler is just over the dew point, which means a minimum vapor flow through the compressor and the lowest horse power needed. The HIDiC (Figure 2a) was simulated as two interconnected distillation columns using the RADFRAC28 block. Heat transfers between corresponding stages in the two sections were realized by assigning side duties. The heat distribution pattern along the adiabatic sections was confirmed to have an impact on the overall energy performance.18 Two typical heat distribution schemes, namely, uniform heat-transfer area and uniform heat distribution, were usually used in simulation studies.17 The configuration in which the stripping section stages are thermally connected with the same number of stages in the upper part of the rectifying section with uniform heat distribution was employed in the present work, as this was confirmed by previous studies17,25,29,30 to be the most energy-

Figure 1. Schematics of (a) VRC and (b) HIDiC.

quasireversible diabatic operation by restricting all heat transfers to several discrete locations. Such an approach is known to be able to reduce capital costs significantly while retaining the original reversible advantages to the greatest extent possible.18,19 Moreover, compared with the complicated HIDiC structure, it is much easier and cheaper to have the rectifying and stripping sections built as two individual columns and the heat-transfer devices as several external heat exchangers between the two sections. It was also found that, by optimizing heat-transfer allocation among the external heat exchangers, a simplified configuration can yield even better energy saving performance than the original HIDiC.18−20 Optimal decision making from among the options of heatintegrated configurations and parameters in terms of an economic criterion is challenging. Although a systematic approach to the decision has very recently become available,21 deeper understanding of the behaviors of heat integration and its impact on the economic performance of a distillation column based on detailed analysis remains important for the 5782

dx.doi.org/10.1021/ie400112k | Ind. Eng. Chem. Res. 2013, 52, 5781−5790


Article

Figure 2. Schematics of a HIDiC and its various simplifications for propylene−propane separation: (a) HIDiC, (b) 3HX, (c) 2HX, (d) 1HXa, (e) 1HXb. 5783



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operating hours per year and costs for the utilities used as follows: 0.1 $/kWh for electricity, 0.03 $/t for cooling water, and 13 $/t for low-pressure steam. The total annualized cost (TAC), which takes into account the contributions of both operating costs and capital investments, is generally used as an indicator of the optimum configuration in terms of economics and was thus also employed in our study for the evaluation of various schemes. TAC was calculated by adding utility costs to capital investments divided by a plant lifetime of 10 years.

efficient scheme for a propylene−propane splitter. When altering the compression ratio in HIDiC cases (including the simplified HIDiC), the operating pressure of the rectifying section is kept unchanged, whereas those of the stripping section and feed are adjusted, with the feed thermal condition remaining constant. This ensures that the temperature level demand for the cold utility does not change, which, in this case, can be satisfied simply by using cooling water. Because the total amount of heat to be rejected from the rectifying section is larger than the amount required in the stripping section for the system studied, an external reboiler is no longer needed. The total amount of heat transferred between the two sections was adjusted to reduce the reboiler duty to nearly zero (less than 10 kW), while leaving behind a small portion of extra vapor to be condensed using an external condenser. For simplified HIDiCs, three different numbers (three, two, and one) of external heat exchangers were considered, abbreviated as 3HX, 2HX, and 1HX, respectively. It was found that avoiding heat transfers between some of the top stages can lead to better energy performance than a uniformly distributed scheme.18 In the 3HX configuration (Figure 2b), the upper one-third of all heat-integrated stages in the HIDiC scheme were simulated adiabatically for simplicity as the total number of stages was varied, and three external heat exchangers were evenly placed among the remaining stages. For the 2HX scheme (Figure 2c), a straightforward middle-bottom placement within all stages was confirmed to be sufficient. In the 1HX scheme (Figure 2d,e), the impact of the heat-transfer location on the rectifying side was intensively studied, and to guarantee the requirement of no external reboiling, the single heat exchanger was always connected to the bottom stage of the stripping section. The distribution of heat loads was also considered in all simplified HIDiCs. An optimization28 block was used to allocate heat duties among several heat exchangers to minimize the horse power of the compressor, and a design specification was set to reduce the reboiler duty to a very low value (10 kW).

4. COMPARISON UNDER SAME TOTAL NUMBER OF STAGES AND TEMPERATURE APPROACH In this section, a series of base cases (Table 1) are compared under the same total number of stages and minimum temperature difference for heat transfer. The results are reported in Table 2. Table 1. Separation Task and Feed Conditions of the Base Case parameter feed composition (mol %) propylene propane i-butane feed flow(kg/h) feed vapor fraction number of stages in the stripping section number of stages in the rectifying section pressure drop per stage (mbar) feed pressure/pressure at the top of the stripping section (bar) pressure at the top of the rectifying section for HIDiC schemes (bar) minimum temperature difference for heat transfer (°C) distillate specification (mol %) propylene bottom specification (mol %) propylene

3. COLUMN SIZING AND COST ESTIMATION Column size was estimated inside the RADFRAC block with an 80% flooding limit based on Fair’s correlation. Parameters such as tray design and spacing were taken to be the same as in ref 23 and were used for all configurations. Column height was calculated based on the number of theoretical stages, which is thought to be an acceptable approximation because the overall efficiency of the base commercialized VRC actually reaches 91%.23 Correlations given by Douglas31 were updated by the Marshall & Swift equipment cost index used in the estimation of equipment costs taking into account the influence of the current price level; the value of M&S index is 1463.2 for the year of 2008.32 The principles of choosing coefficients and parameters in the correlations used by Olujic et al.23 were employed. The simplified approach for calculating the installed cost of the HIDiC proposed by Olujic et al.23 was also employed. Each stage-to-stage heat transfer was treated as implemented by an individual heat exchanger, and the cost was calculated from the required heat-transfer area. When calculating the heat-transfer area, constant values of overall heat-transfer coefficients were used for different types of devices, namely, 1000 W/(m2 K) for HIDiC heat-transfer panels and thermosiphon reboilers and 800 W/(m2 K) for condensers. The utility cost calculations assumed 8000

value 52 47 1 112000 37% 62 171 6.2 11.2 14.6 ∼5.5 99.6 1.1

In Table 2, only a slight difference is found between the required compression ratios of the VRC and HIDiC schemes to reach the same minimum temperature difference, which is inconsistent with the common belief that a HIDiC requires a lower compression ratio than VRC.3 Moreover, because the vapor flow through the compressor in a HIDiC is indicated by simulation results to exceed that in VRC by as much as 23.7%, leading to an increase of 15% in the compressor-power requirement; the utility cost of a HIDiC, which is primarily the electricity cost of the compressor, is also higher than that of VRC by a similar proportion. In terms of capital costs, VRC is obviously the cheapest scheme, whereas the HIDiC scheme is the most expensive. As can be seen from Table 2, a large number of heat-transfer devices are actually responsible for more than 28.3% (1.41 million $/year) of the whole capital investment (4.98 million $/year) for a HIDiC, which explains why the HIDiC demands higher capital costs. VRC is shown to be the optimum configuration in terms of TAC. The results of the comparisons performed in our work differ from those in some previous studies,3,23−26 and this is actually due to the different basis employed for comparison. According to the simulation results in Table 2, a reduction of up to 26.3% (from 4.98 million $/year for the HIDiC 5784



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Table 2. Simulation Results of the Base Case configuration

a

parameter

VRC

HIDiC

3HX

2HX

1HXa

1HXb

compression ratio minimum temperature difference (°C) compressor power (kW) vapor flow through the compressor (kmol/h) condenser/cooler duty (kW) total heat integration (kW) total heat-exchanger area (m2) utility cost (106$/year) heat-exchanger costa (106$/year) capital cost (106$/year) TAC (106$/year)

1.56 5.49 6745 19806 −5233 72851 11048 5.52 0.40 3.15 8.67

1.50 5.52 7754 24508 −5989 81445 11882 6.27 1.41 4.98 11.25

1.51 5.52 7679 24218 −5933 80464 12550 6.21 0.62 3.67 9.88

1.51 5.50 7644 24111 −5888 80147 12629 6.18 0.55 3.57 9.74

1.51 5.52 8204 25762 −6444 85898 15555 6.63 0.50 3.58 10.21

1.53 5.52 7473 22699 −5704 75166 13626 6.04 0.46 3.42 9.46

Excludes cost of external condensers.

configuration to 3.67 million $/year for 3HX) could be achieved for the capital cost of the HIDiC by reducing the number of heat-transfer locations to only three. However, further reductions in the number of heat exchangers merely brought slight benefits. For energy consumption, it was shown that even better results than for the original HIDiC could be obtained by optimizing the heat duties for two or three external heat exchangers. As mentioned previously for the single-heat-exchanger scheme, there are many possible locations along the entire upper part of the rectifying section (see Figure 2d,e) to which the single heat exchanger can be connected, whereas the heat rejected is always supplied to the bottom of the stripping section to ensure zero reboiling requirement. Two representative configurations (1HXa and 1HXb) are listed in Table 2, in which the heat-transfer locations are at the bottom and the very middle, respectively, of the upper part of the rectifying section. A more detailed analysis of the influence of heat-transfer locations along with a further comparison was also conducted and is illustrated in Figures 3−5.

Figure 3 depicts a generalized schematic to represent heat transfers in all configurations studied in this work, including VRC, HIDiC, and simplified HIDiC. Only one heat exchanger is installed between the two sections for simplicity, which is adequate for conceptual illustration. Nevertheless, conclusions drawn from the simple analysis of a single heat exchanger can be easily extended to multiple-heat-exchanger schemes. In Figure 3, a heat exchanger is located somewhere in the middle part along the column length, connecting points C and D in the stripping and rectifying sections, respectively, which we call heat transfer from D to C for short. In the case of external cooling or reboiling demands, a condenser and a reboiler might still be needed. If the location where heat is rejected from the rectifying side or supplied to the stripping side is shifted, both the compressor-power and heat-transferarea requirements will change accordingly to reach the specified product purities. For instance, if the heat-rejection point is moved upward from D to D′, the consequent impact is twofold. On one hand, the internal reflux is provided at a higher position, leading to an increase in liquid and vapor flows between D and D′, which causes a greater degree of separation and, hence, higher product purities. Because comparisons are made for uniform product purities, the total heat integration between the two sections can then be reduced, causing a decrease in the liquid and vapor flows from point C all the way through the compressor to point D. A saving of compressor power is thereby achieved with reduced passing vapor flow. On the other hand, the temperature difference for heat transfer also changes as the heat-integrating location varies. For a typical distillation column, the temperature profile shows a monotonic decrease from bottom to top. Therefore, because D′ is lower in temperature than D, the temperature difference between D′ and C is not as large as that in the original heat transfer from D to C. The loss in driving force must be compensated by either an increased heat-transfer area, which means an increase in capital costs, or a higher compression ratio of the compressor (higher utility costs). Obviously, completely opposite tendencies hold for the case in which D is moved downward. A similar analysis of the effect of altering the heat supply location on the stripping side can also be easily carried out. For example, if the heat-exchanging location is shifted from C to C′, less compressor power is needed because of the enhanced separation taken place between C and C′, whereas additional efforts should be made to compensate for the decrease in temperature driving force. In summary, as the heat-integrated location on either side is altered, the overall energy perform-

Figure 3. Generalized schematic of heat transfer between the rectifying and stripping sections. 5785



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ance depends on two distinctive factors, namely, the vapor flow through the compressor and the temperature difference for heat transfer, and they are always contradictory. Therefore, it is the result of the conflict between the two factors that determines whether an alternative configuration results in further energy savings. On the basis of the preceding analysis, a further comparison among various alternatives of the single-heat-exchanger scheme was conducted. The results are shown in Figures 4 and 5. First,

amount of total heat integration required also increases significantly. To exclude the influence of unequal capital costs in the comparisons, the compression ratio for each case was then adjusted to reach a uniform heat-exchanger-area requirement, so that the most energy efficient configuration could be obtained simply by a direct comparison of compressor power. The resultant curve is shown in Figure 5. Heat transfer from the

Figure 5. Comparison of compressor power under the same heatexchanger-area requirement (but with varying compression ratio).

top of the rectifying section to the bottom of the stripping section yields the best performance, which is thermodynamically equivalent to the VRC scheme. Analysis showed that the factor of vapor flow through the compressor overwhelms the influence of the temperature driving force and dominates the overall energy performance. The close-boiling characteristic of the propylene−propane system should be responsible for this result. Specifically, because the temperature span of the whole column is limited, a shift in heat-exchanger location might not bring about a marked change in the temperature driving force, so that a minimal impact on the compressor-power requirement is made by the factor of temperature difference. A similar analysis could also be made to explain why VRC is superior to the HIDiC in this round of comparisons. According to the generalized illustration in Figure 3, VRC is thermodynamically equivalent to a heat transfer from B to E, whereas a HIDiC involves a series of parallel heat transfers such as that from D to C. Again, analysis in terms of both factors should be carried out. In the VRC scheme, heat is rejected from the very same location as in a conventional distillation column and supplied to the bottom reboiler, so that minimum vapor flow through the compressor could be achieved, because B is the highest possible location for D′ and E is the lowest possible location for C′. Simulation results also support the conclusion that VRC requires lower vapor flow through the compressor than any other configuration (Table 2). The advantage of the HIDiC lies in its superiority in terms of compression ratio over VRC, which was, however, found to be merely marginal in this case. In fact, this also has to do with the close-boiling characteristic of the propylene−propane system. Taking location B in Figure 3, for example, heat rejected from that point in the HIDiC is no longer supplied to E as in the VRC scheme but is transferred directly to A instead, which results in an increase in the temperature difference for heat transfer. A reduction in compression ratio is therefore possible to reach the

Figure 4. Effects of heat-transfer location on the rectifying side with a constant compression ratio on (a) vapor flow rate through the compressor and (b) minimum temperature difference for heat transfer.

the compression ratio was fixed as the heat-integrated stage in the rectifying section was altered, so that the impact of the heatintegrated location on the two mentioned factors (i.e., vapor flow through the compressor and temperature difference for heat transfer) could be observed. In regard to vapor flow through the compressor (Figure 4a), a lower heat-rejection point farther from the main condenser corresponds to greater vapor flow through the compressor and, hence, greater compressor power. For the factor of temperature difference (Figure 4b), heat-transfer driving force declines linearly as expected from the lowest heat-rejection position to the topmost alternative. When heat integration takes place between the corresponding stages (stage 61) of the two sections, which we call parallel heat transfer, the smallest heat-exchanger area is required. As the heat exchanger is connected to a higher location in the rectifying section, the heat-exchanger area requirement rises abruptly because of the decreased temperature driving force. However, moving the heat-rejection point downward does not lead to opposite reductions, because the 5786



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Figure 6. TAC curves of various schemes under different total numbers of stages and various minimum temperature differences: (a) VRC, (b) HIDiC, (c) 3HX, (d) 2HX, (e) 1HXa, (f) 1HXb.

5. COMPARISON UNDER VARIABLE TOTAL NUMBER OF STAGES AND TEMPERATURE DIFFERENCE All comparisons in the previous section were made by assuming the same total number of stages for all schemes, which, however, gave results with limitations. With an enhanced degree of reversibility, a HIDiC actually requires more theoretical stages to compensate for the loss in mass-transfer driving force essential in accomplishing a prescribed separating task. It is necessary to determine whether a HIDiC could show better performance with different numbers of theoretical stages and also to find the optimum total number of stages for each

same minimum temperature driving force. The extent to which the savings in compression ratio could be achieved depends completely on how much lower the temperature of A is than that of E, that is, the temperature span of the whole stripping section. Nevertheless, this value of temperature span reaches only 5.85 °C for the close-boiling system studied in the present work, and hence, a limited impact on the compression ratio is expected. 5787



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scheme in terms of TAC. Moreover, the compression ratio remains another crucial parameter to be optimized. A high compression ratio certainly leads to high energy consumption but requires a relatively small heat-exchanger area with a large temperature difference for heat transfer, whereas a lower compression ratio reduces the compressor power at the expense of an increased heat-exchanger-area requirement owing to a corresponding decrease in the temperature driving force. For a thorough comparison among the optimum states of all schemes, simulation studies for at most five different numbers of theoretical stages were carried out to find the optimum number of stages for each scheme. The influence of compression ratio was also examined by setting five minimum temperature difference levels for each total number of stages. The value for the base commercial VRC propylene−propane splitter (around 7 °C) was selected as the upper bound for the temperature difference, and the lower bound was set to 1 °C as a practical limit. The obtained TAC values were plotted against temperature difference, and the performances at different total numbers of stages are represented by the separate curves shown in Figure 6. As can be seen, all TAC curves show similar ascending trends for a wide range of temperature differences, but with flat regions around lower temperature differences. A lowest point of around 2.5 °C was found for the HIDiC scheme, whereas the optimum values for the other schemes seemed not as obvious. Minimum TAC values appearing at such low temperature differences indicate a relatively weak influence of the heatexchanger costs on the TAC that is overwhelmed by the compressor cost nearly proportional to the compressor power. Therefore, generally speaking, a low compression ratio would be preferable for a heat-integrated distillation column in terms of TAC. For the HIDiC and its simplifications, adding 40 stages from 191 to 231 did bring substantial reductions to the TAC, but the effects were less pronounced when the number of stages was further increased to 271, especially in the case of the HIDiC and 1HXb configurations, where hardly any benefits were observed. A total number of 271 stages was found to be the optimum value for most HIDiC-type schemes, with 1HXa as an exception, in which the best performance was seen for 311 stages. Not surprisingly, in the VRC scheme, simulations with the largest number of stages actually led to the highest TAC values, because VRC exhibits the same irreversibility as the conventional distillation column. According to the simulation results, a total of 231 stages seemed to be adequate. The industrial choice of a total of 191 stages23 is also reasonable, because very little difference in TAC between the two choices was found. The optimum number of stages for VRC is obviously smaller than those for HIDiC-type schemes, in which the decreased mass-transfer driving force caused by the increase in reversibility has to be compensated by increased theoretical stages. The TAC curves obtained at the optimal total number of stages for each scheme were then plotted in a single graph as shown in Figure 7 for an overall comparison. It can be seen from the graph that the HIDiC and VRC schemes gave the highest and lowest costs, respectively. Crowded between these extremes are various simplified HIDiC schemes. Obviously, the TAC values of the HIDiC could be significantly reduced by minimizing the number of heat exchangers. Every time a heat exchanger is removed, some economic benefits could be

Figure 7. Overall comparison of TAC among all schemes (optimal number of stages selected).

obtained, and if well arranged, the single-heat-exchanger scheme should be the best among all simplifications. A more detailed comparison between the VRC and HIDiC schemes in terms of both utility and capital costs was also made to obtain a better understanding of their performance for variable total numbers of stages and temperature differences. Figure 8a primarily illustrates the influence of total number of stages on the TAC of both schemes. As can be clearly seen, the utility costs of the HIDiC could be significantly reduced by using more theoretical stages, which confirms that the

Figure 8. Comparison between VRC and HIDiC: (a) utility cost, (b) capital cost. 5788



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should favor the HIDiC and the simplified HIDiCs as well, and as a result, using a more rigorous cost model for heat exchangers will not conflict with the conclusions drawn from Figure 7. Because the heat-exchanger costs constitute a large proportion of the TAC in the HIDiC scheme and, therefore, have a significant impact on its economic performance, great efforts must be made to develop novel internal heat-transfer devices of higher efficiency as well as lower costs.

advantage of reversibility demands more stages to be fully presented, whereas that of VRC is not sensitive to an increase in the number of theoretical stages. However, even in the best case, the HIDiC still requires higher utility costs than VRC. For the comparison of capital costs (Figure 8b), a significant difference is observed between the two schemes, which indicates the superiority of VRC over the HIDiC in such a case.

6. CONCLUDING REMARKS In this work, several configurations of internally and externally heat-integrated distillation were evaluated and compared. The influences of the total number of stages and the temperature difference for heat transfer on the energy and economic performances of all configurations were investigated by sensitivity studies. Generally speaking, a low temperature difference for heat transfer was favored for all schemes. It was shown that the HIDiC and simplified HIDiC configurations obviously required more theoretical stages than VRC and needed to be optimized in terms of the number of stages to fully exhibit their advantages with increased reversibility. From the comparison results, it was found that the standard upon which the comparison was based actually had a great influence on the results obtained. We showed that, for a typical propylene−propane splitter of industrial scale, the VRC scheme was superior to the HIDiC and simplified HIDiC configurations in terms of both energy consumption and economics on the basis of the uniform minimum temperature difference for heat transfer. Only marginal differences between the required compression ratios of the HIDiC and VRC schemes to reach the same minimum temperature driven force difference was observed for the close-boiling system studied. Moreover, because the vapor flow rate through the compressor in the HIDiC was much larger than that in the VRC scheme because of the increased internal reflux, the HIDiC actually consumed greater compressor power. It should be noted that lower energy consumption than obtained in the VRC scheme might still be possible as long as the HIDiC is operated at a lower minimum temperature difference; nevertheless, in that case, accommodating a larger heat-transfer area could be more costly. Considerable reductions in capital costs could be achieved by simplifying the internal heat integration to several external heat exchangers, and similar (if not better) energy-saving performance could be retained. For the system studied, the option with only two external heat exchangers was shown to yield satisfactory results. Moreover, if operated at a much lower temperature difference (e.g., 1HXb at 2.5 °C), the simplified HIDiC could surely compete with, or even exceed, the VRC scheme at a higher compression ratio (e.g., at a temperature difference of 7 °C) in terms of total annualized cost (see Figure 7). Although indicated by simulation results, a very low temperature difference seems to be economically most preferred for all heat-integration configurations, such close temperature approach might be impractical because of restrictions in heat-exchanger manufacture and operation. The optimal temperature difference for heat exchange appearing at such low values might come from an underestimation of the costs of heat exchangers, as the installed costs of the HIDiC and simplified HIDiC configurations in the present work were all estimated using a simplified approach.23 This can be improved by employing a more reliable cost model for heat exchangers. However, lower estimations of heat-exchanger costs

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AUTHOR INFORMATION

Corresponding Author

*E-mail: [email protected]. Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS The authors acknowledge the National Key Fundamental Research Program of China for supporting this research under Grant 2012CB720500.

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REFERENCES

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