Performance Enhancement of Reactive Dividing Wall Column Based

Jun 14, 2019 - It is also a high energy intensive process and the thermodynamic ..... The condenser pressures of the both columns are fixed at 100 kPa...
0 downloads 0 Views 2MB Size
Download PDF
Article Cite This: Ind. Eng. Chem. Res. 2019, 58, 12179−12191

pubs.acs.org/IECR

Performance Enhancement of Reactive Dividing Wall Column Based on Self-Heat Recuperation Technology Jie Li, Fengjiao Zhang, Qi Pan, Yilei Yang, and Lanyi Sun* State Key Laboratory of Heavy Oil Processing, College of Chemical Engineering, China University of Petroleum (East China), Qingdao,Shandong 266580, China

Downloaded via CHALMERS UNIV OF TECHNOLOGY on August 5, 2019 at 03:05:06 (UTC). See https://pubs.acs.org/sharingguidelines for options on how to legitimately share published articles.

S Supporting Information *

ABSTRACT: Reactive dividing wall column (RDWC) is a highly integrated configuration where reaction and separation take place in a single equipment simultaneously. In this study, three self-heat recuperative reactive dividing wall column processes are developed to reduce the energy consumption of RDWC. Latent heat and sensible heat in the processes are recovered by selfheat recuperation technology. The heat exchanger network using pinch analysis is applied to optimize the energy utilization. Total energy consumption, total annual cost, CO2 emissions, and thermodynamic efficiency are evaluated. Compared with the RDWC, the results demonstrate that the optimal design of the self-heat recuperative processes reduces the energy consumption by 58.1% and decreases the total annual cost by 17.8%. Apart from that, the CO2 emissions of the optimal design are 87.9% lower than the RDWC and the thermodynamic efficiency can be increased to 31.4%.

1. INTRODUCTION Distillation is the most commonly used technology for the separation of multicomponent mixtures. It is also a high energy intensive process and the thermodynamic efficiency of conventional distillation is about 5−20%.1 Therefore, many researchers focus on the process intensification technology of energy-saving distillation.2 As a representative of energy-efficient distillation, the dividing wall column (DWC) has great potential in reducing energy consumption and capital cost.3−5 A study showed that DWC achieved 10−60% energy savings and 10−50% capital cost reduction for the separation of different systems.6 Reactive distillation (RD) is an intensified configuration combining reaction and separation into a single shell, which eliminates the limitation of conversion and phase equilibrium.7 The reactive dividing wall column (RDWC) takes the advantages of both RD and DWC to further reduce energy consumption and capital cost of the process.5,8 Li et al.9 made a comparison between the conventional RD and RDWC processes for the hydrolysis of methyl acetate. It was shown that 20.1% energy savings can be attained by RDWC. Zheng et al.10 designed an RDWC to synthesize diethyl carbonate, and the results showed that approximately 18.7% of the energy consumption and © 2019 American Chemical Society

13.9% of the TAC can be saved by the RDWC compared with the conventional RD process. An et al.11 proposed an RDWC for the esterification of acetic acid with methanol. The RDWC achieved 7.7% saving in energy consumption, 8.3% reduction in operating cost, and 15.5% decrease in capital cost in comparison with the conventional RD process. In the conventional distillation column, high-temperaturelevel energy is needed in the reboiler while low-temperaturelevel energy is removed from the condenser. The vapor recompression heat pump (VRHP) can significantly improve the temperature level of the overhead vapor stream through a compressor, such that the energy can be transferred from a heat source to a heat sink.12−14 The application of VRHP in distillation has been studied extensively, demonstrating that VRHP has a great potential in saving energy.15−17 In consideration of the remarkable energy-saving contributions of VRHP and RDWC, the integration of VRHP and RDWC has been investigated by a few researchers. Liu et al.18 Received: Revised: Accepted: Published: 12179

April 30, 2019 June 6, 2019 June 14, 2019 June 14, 2019 DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

thermodynamic efficiency, economic performance, and environmental performance were improved notably as well. Based on these investigations, the HEN using pinch analysis is an effective tool for the establishment of self-heat recuperative processes, which possesses the advantage of energy saving. In the work of An et al.,11 an RDWC configuration was proposed for the production of methyl acetate and favorable energy-saving results were achieved compared with the conventional RD. However, no further energy-saving research has been made for this RDWC process. As mentioned above, the self-heat recuperation technology and HEN using pinch analysis are effective and feasible measures or tools in the reduction of energy consumption; both of them can be extended to the design of a more energy-saving RDWC process. Although the superiority of RDWC has been demonstrated by plenty of literature reports, the integration of the HEN using pinch analysis and self-heat recuperation technology into the RDWC process has not been studied yet. In this paper, the RDWC for the production of methyl acetate is extended in order to reduce energy consumption as much as possible. Therefore, three self-heat recuperative RDWC (SHR-RDWC) processes are proposed to enhance the thermodynamic efficiency. The HEN using pinch analysis is applied to optimize the energy utilization. The proposed SHR-RDWC processes are compared with the RDWC in terms of energy consumption, economics, environmental performance, and thermodynamic efficiency.

proposed a heat pump assisted RDWC for the transesterification of methyl acetate and n-butanol, in which compressed overhead vapor exchanges heat with the intermediate reboiler. Compared with the RDWC, the total annual cost was reduced by 9.6% with a payback period of 8 years and CO2 emissions were reduced by 14.76%. Feng et al.19,20 proposed intensification schemes combining VRHP with an upper partitioned reactive dividing wall column and a lower partitioned reactive dividing wall column, respectively. With the intensified processes, significant reductions in total annual cost, total utility consumption, and CO2 emissions were obtained in contrast to the RDWC. Feng et al.21 developed and assessed a vapor recompression heat pump assisted RDWC for the production of n-propyl acetate. The results revealed that the VRHP-RDWC gave the most energy-saving and economical design. The total annual cost of the intensified process was reduced by 25.53% and the thermodynamic efficiency was increased significantly in contrast to the RDWC. However, the integration of VRHP and RDWC still remains a problem of exergy loss due to the insufficient utilization of sensible heat in the process, which is expected to be further integrated for sustainable development. For further energy savings, self-heat recuperation technology proposed by Kansha et al.22 has been developed based on minimizing the exergy loss to reduce energy consumption in chemical processes. This technology promotes the reuse of both latent heat and sensible heat of the process.23 The wasted heat was recuperated thoroughly by compressors and exchanged with the cooling load, leading to no need for external heat or minor energy requirements.24−26 Self-heat recuperation technology has been applied to many distillation processes, which have attained a large reduction in energy consumption. Christopher et al.27 proposed a self-heat recuperative process for the separation of propylene and propane, in which the overhead vapor of the column was compressed and recovered heat was used to heat the reboiler, and then the remaining sensible heat was supplied to preheat the feed stream and the compressor inlet stream. The energy consumption has been reduced by 66 and 45% compared with the conventional process and simple vapor recompression. Matsuda et al.24 applied the self-heat recuperation technology to the distillation process. Energy consumption has been reduced significantly in the modified process through self-heat recuperation technology. A heat exchanger network (HEN) using pinch analysis is an effective tool for the optimization of energy utilization that can reduce energy consumption in a huge range of processes, large and small plants.28 Recently, the HEN using pinch analysis has been extended to the design of self-heat recuperative distillation processes. Xia et al.29 used the HEN to design and optimize the self-heat recuperative pressure-swing distillation process based on the pinch analysis. The improved process resulted in 72.39% saving of energy and 36.55% reduction of TAC in comparison with the conventional pressure-swing distillation process. Chen et al.30 improved the economic and environmental performance of heterogeneous azeotropic distillation by the self-heat recuperation technology and made full use of all the heat in the process by HEN based on the pinch analysis. Xia et al.31 applied the HEN to the design of self-heat recuperative dividing wall column (SHR-DWC) for the production of n-butanol based on the pinch analysis. The SHR-DWC worked out 71.13% savings of energy consumption in contrast to the basic process. Besides,

2. PERFORMANCE EVALUATION METHODS In this work, four criteria including total energy consumption (TEC), total annual cost (TAC), CO2 emissions, and thermodynamic efficiency are used to evaluate the proposed configurations. 2.1. Total Energy Consumption (TEC). The TEC is used to evaluate the total energy consumption of the process; it specifically refers to the consumption of external energy including steam and electricity. The TEC is defined as follows: TEC =

∑ Q REB + ∑ Q PRE + f ∑ WCOMP

(1)

where QREB (kW) represents the reboiler duty; QPRE (kW) denotes the feed preheater duty; WCOMP (kW) is the compressor duty; the multiplication factor (f = 3) is employed to convert the compression work into thermal energy which can produce an equivalent amount of electrical work.32 2.2. Total Annual Cost (TAC). As an important indicator to measure the economic performance of a process, TAC considers the operating cost (OC) and capital cost (CC) together. It is calculated by the following equation:33 TAC = operating cost +

capital cost paypack period

(2)

The OC contains the costs of steam, cooling water, catalyst, and electricity. The CC is the installed costs of major equipment including column shells, column trays, heat exchangers, and compressors. Since construction and installation of the dividing wall column are much more difficult than for the conventional column, a penalty of 20%34,35 is assumed in this paper. A payback period of 8 years is used, with an 8000 h/year operating time.19,20 The detailed calculation formulas for OC and CC are presented in Table S1 of the Supporting Information. Douglas’s33 cost formulas are transferred into the 12180

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research CEPCI inflation index based on the equation presented in the literature36 and adopted for the calculation of capital cost. The minimum temperature difference of 5 °C is assumed in all the heat exchangers.19−21 Compressor efficiency is assumed to be 80% in all the configurations.37 2.3. CO2 Emissions. CO2 emissions can be used to assess the environmental benefits of a distillation process. A large amount of CO2 will be discharged to the environment with the burning of fossil fuels (such as coal, heavy fuel oil, and natural gas). The CO2 emissions (kg/h) reported by Gadalla et al.38 are shown as follows: iQ yi C% yz zzα CO2 emissions = jjjj Fuel zzzzjjj k NHV {k 100 {

thermodynamic efficiency =

Q Proc λProc

(hProc − 419)

thermodynamic efficiency =

kf

V CH3COOCH3(MeAc) + H 2O(water) kr

(4)

r = mcat(k f aHAcaMeOH − k raMeAca H2O)

i −49190 zy zz k f = 2.961 × 104 expjjj k RT { i −69230 yz zz k r = 1.348 × 106 expjjj k RT {

(6)

where B (kJ/kmol) denotes the exergy; H (kJ/kmol) is the enthalpy; T0 (=298 K) is the ambient temperature; S [kJ/ (kmol·K)] is the entropy; m (kmol/h) is the stream flow rate. The lost work is calculated by eq 7: LW =

i

∑ Wi − Δ(mB) + ∑ jjjjj1 − i

i

k

T0 yzz zQ Ti zz{ i

(12)

(13)

where R [=8.314 kJ/(kmol·K)] is the gas constant; T (K) is the temperature. The simulation of the esterification reaction is implemented using the RadFrac module in Aspen Plus V10. The universal quasi-chemical activity coefficient (UNIQUAC) model is used to calculate the liquid activity coefficients. The Hayden− O’Connell second virial coefficient model is adopted to predict the solvation and dimerization of acetic acid in the vapor phase. The binary interaction parameters for the models are provided in An’s work.11 The catalyst with a density of 770 kg/ m3 occupied half of the tray holdup volume.7 It is assumed that the weir height is 10 cm while the liquid holdup is calculated based on the weir height and the column diameter.7 3.2. Conventional Reactive Dividing Wall Column. The flow sheet of the conventional RDWC proposed by An et al.11 is shown in Figure 1. Two coolers are added to the outlet streams since the products have to be cooled for the sake of storage or wastewater treatment. The reactants’ feed ratio is set to 1:1.2 with 20% excess methanol. The RDWC is composed of a reactive distillation column (REA) and a rectifying column (REC) linked with two interconnecting streams. The light product methyl acetate with a purity of 98 wt % is obtained from the top of the RDWC. The heavy product water with a purity of 99 wt % is obtained from the bottom of the RDWC. A vapor side stream containing the unreacted methanol is

(5)

in

(11)

where r (kmol/s) is the reaction rate; mcat (kg) is the mass of the catalyst; ai is the activity of component i; kf and kr [kmol/ (kg·s)] are rate constants of the forward reaction and reverse reaction, respectively, which are temperature dependent given as follows:

∑ mB − ∑ mB out

(10)

The reaction takes place in the liquid phase. The acid ionexchange resin Amberlyst 15 is used as catalyst. The catalyst weight based kinetic model reported by Tang et al.42 is described as follows:

where λProc (=2085 kJ/kg) and hProc (=2755.5 kJ/kg) are the latent heat and enthalpy of steam delivered to the process, respectively. The TFTB (=1800 °C) is the flame temperature of the boiler flue gases, and the TStack (=160 °C) represents the stack temperature. The boiler feedwater is assumed to be 100 °C with an enthalpy of 419 kJ/kg. The T0 (=25 °C) is the ambient temperature. Natural gas is chosen as the fuel in this study. The NHV and C% are 51 600 kJ/kg and 75.4, respectively. CO2 emissions caused by the generation of electricity are also involved in this work. The CO2 emissions are 0.0565 g/kJ when the electricity is generated from the combustion of natural gas.40 2.4. Thermodynamic Efficiency. Thermodynamic efficiency can be used to assess the improvement and effectiveness of the configurations. Thermodynamic efficiency depends on the minimum separation work (Wmin) and lost work (LW) of the process. The exergy energy (B) and Wmin are calculated by the following equations:41

Wmin =

(9)

CH3COOH(HAc) + CH3OH(MeOH)

39

B = H − T0S

Wmin Wmin − LW

3. PROCESS DESCRIPTION 3.1. Kinetic and Thermodynamic Model. This work considers the production of methyl acetate (MeAc) by the esterification of methanol (MeOH) and acetic acid (HAc). The reversible reaction can be described by the following expression:

(3)

TFTB − T0 TFTB − TStack

(8)

If the value of Wmin is negative, the thermodynamic efficiency becomes

where QFuel (kW) represents the amount of fuel burnt; NHV (kJ/kg) is the net heating value of a fuel; C% is the carbon content; α (=3.67) denotes the ratio of the molar masses of CO2 and C. QFuel is calculated by the following equation: Q Fuel =

Wmin − LW Wmin

(7)

where LW (kJ/h) represents the lost work; Wi (kJ/h) is positive for the work done by the surroundings on the system; Ti (K) is the temperature of heat source or heat sink; Qi (kJ/h) is positive for heat transfer from the surroundings. Wmin depends on the energy stored in the inlet and outlet streams. When the value of Wmin is positive, the thermodynamic efficiency is given by 12181

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

limiting reaction rate above stage 12 is mainly attributed to the content of methanol, while that for the reaction rate below stage 12 is ascribed to the content of acetic acid. The generation or consumption amounts of the other components gradually decrease on the other reactive stages. Based on the criteria mentioned in section 2, the TEC and TAC for the basic RDWC are 1221.3 kW and 449.4 × 103 $/year, respectively. Besides, the CO2 emissions and thermodynamic efficiency are 286.1 kg/h and 22.4%, respectively.

4. PROCESS IMPROVEMENT 4.1. Reactive Dividing Wall Column with Feed Preheating (RDWC-FP). The thermal condition of the feed can be adjusted to improve the thermodynamic efficiency of the distillation column.43 The reactive dividing wall column with feed preheating is conducted before the development of self-heat recuperative processes. On the one hand, reboiler duty can be reduced with feed preheating. On the other hand, the hot streams in the process will provide the required heat for the cold feed stream through heat integration or self-heat recuperation technology so that using existing heat sources available in the plant is possible. In the conventional reactive dividing wall column, acetic acid is fed to the top of the reactive zone while methanol is fed to the stage near the bottom of the reactive zone since acetic acid is much heavier than methanol. The feed temperatures of HAc and MeOH are all set to 30 °C, which is much lower than the stage temperatures. Exergy loss in the vicinity of the feed location is substantial because of the mixing of different state streams. In addition, a large amount of heat is required to vaporize these liquid streams. Thus, the reboiler duty can be reduced by preheating the feed streams. Figure 2 demonstrates

Figure 1. Flow sheet diagram and process parameters of RDWC.

rectified in the REC with high purity (99 wt %) methanol obtained from the top. The recovered methanol is mixed with the fresh MeOH and then fed into the REA. The condenser pressures of the both columns are fixed at 100 kPa in order to use cheap cooling water as cold utility. A pressure drop of 0.7 kPa is assumed for each stage.10,21 Pressure drops in shells and tubes of heat exchangers are neglected. In the RDWC, the reflux ratios of the REA and REC (RR1 and RR2) are adjusted to meet the purities of methyl acetate and methanol while the vapor side stream flow rate (FVR) is varied to meet the specified purities of water. The design variables including the number of reactive stages (NRX), feed locations of acetic acid and methanol (NHAc and NMeOH), withdraw location of the vapor side stream (NVR), number of stages in the rectifying and stripping sections (NR and NS), and number of stages of REC (NREC) are varied to obtain the optimum process. We further optimize the RDWC process with TAC as the optimization objective based on the sequential iterative optimization procedure. Procedure and optimization results are shown in Figures S1 and S2. Detailed optimal process parameters are given in Figure 1. The composition profiles of the RDWC are shown in Figure S3. From the feed stage to the bottom in the stripping section, the composition of the intermediate component (methanol) increases first due to the reduction of light components and withdrawn by the vapor side stream at a high content, which avoids the remixing of internal streams. Then the vapor side stream is further rectified by the REC column, in which the excess methanol is obtained from the overhead stream and the bottom liquid stream returns to the REA column. The temperature profiles of the RDWC are shown in Figure S4. The feed arrangement which drives the heavy reactant (acetic acid) to the top and the light reactant (methanol) to the bottom results in the nonmonotonic distribution of temperature profiles. In the stripping section, the temperature increases dramatically stage by stage because of the enrichment of water. Figure S5 displays the reactive profiles of the RDWC. The component generation rates in the column reveal that the reactive mainly proceeds between stage 5 and stage 16. The consumption rates of reactants reach a maximum value at stage 12 because of the maximum concentration of reactants. The

Figure 2. Effect of feed thermal condition of HAc and MeOH on reboiler duty.

the effect of feed thermal condition (ψ) of HAc and MeOH on reboiler duty with reflux ratios and vapor side stream flow rate adjusted to satisfy the product purities. ψ is equal to the vapor fraction of the stream. It can be seen from Figure 2 that preheating the HAc has a minor effect on the reduction of reboiler duty. However, the reboiler duty decreases significantly when the ψ of MeOH increases and saturated vapor corresponds to the optimal thermal condition. The reboiler duty is reduced because of the vaporization of feed and less heat is required to separate the intermediate component in the stripping section. The feed thermal condition of MeOH is fixed at 1 for the reboiler duty has the minimum value. 12182

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

Figure 3. Flow sheet diagram and process parameters of RDWC-FP.

Figure 4. (a) T−H diagram and (b) GCC of the RDWC-FP.

in the stripping section result in the higher temperatures. The reactive profiles of RDWC-FP are shown in Figure S8. The generation or consumption amounts of the components reach a maximum value on stage 12, which is the same as the conventional RDWC. The total energy consumption of RDWC-FP is reduced by 9.3% although reboiler duty is decreased sharply. The feed preheater still needs external heat, which is not acceptable in the present economic and environmental scenario. To address the problem, SHR technology which can recover the latent heat and sensible heat in the process will be applied to the RDWC and investigated in section 4.2. 4.2. Self-Heat Recuperative Reactive Dividing Wall Column. 4.2.1. Energy Analysis. Ahead of the design of selfheat recuperative processes, the stream data of RDWC-FP are extracted and listed in Table S2. As shown in Figure 4, the temperature−enthalpy (T−H) diagram and grand composite curve (GCC) are plotted based on ΔTmin = 5 °C. The T−H diagram is a helpful method for the visualization of stream temperature and heat content28 as demonstrated in Figure 4a. The y-axis represents the temperature, whereas the x-axis

It can be found that the bottom stream has a large amount of sensible heat which can be recuperated to heat the cold stream MeOH. A feed-effluent heat exchanger (FEHE) can be implemented between the bottom product and cold stream MeOH. The flow sheet and process parameters of RDWC-FP are shown in Figure 3. Two feed preheaters are added to the stream MeOH. The cold stream MeOH exchanges heat with the hot stream water in the FEHE first and then is heated by external heat in the feed preheater (PHE1) to reach the target temperature. The recovered methanol from the overhead of REC is recycled directly to the feed without further condensing and cooling since the vapor fraction of the recycle stream is 1. The composition profiles of the RDWC-FP are shown in Figure S6. As compared with the conventional RDWC without feed preheating, the composition of methanol in the stripping section decreases significantly leading to 48.2% reduction in reboiler duty. The temperature profiles of RDWC-FP are given in Figure S7. The trends of the profiles are similar to those of the conventional RDWC. However, the temperatures in the stripping section are slightly higher than those in the conventional RDWC. The lower concentrations of methanol 12183

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

Figure 5. P−T curves for (a) REA_OVHD and (b) REC_OVHD.

represents the heat content which is usually called the enthalpy of streams. The hot composite curve (HCC) represents the overall energy supplements, while the cold composite curve (CCC) represents the total energy requirements of the process. The overlap (shaded area) of the HCC and CCC in Figure 4a is the heat that can be recovered. The displacements of QCW and QH represent the demands of cold utility and hot utility, respectively. In the RDWC-FP, the QCW is for cold utility required in CON1, CON2, and C1 while QH is for hot utility required in REB and PHE1. It is obvious that a great deal of hot utility and cold utility are required in the process. Thus, the energy-saving potential of the process is large and measures can be taken to further reduce the requirements of utilities. Figure 4b gives the GCC of the streams which represents the minimum amount of heating or cooling demands at any specific temperature.28 The residuals at the top and bottom ends tell us the minimum hot and cold utilities (QH and QCW). Besides, pinch points can be easily detected from the GCC. The GCC is divided into several segments as shown in Figure 4b. The oblique lines represent the sensible heat, while the horizontal lines represent the latent heat in the process. What is worth mentioning is that the horizontal lines contain large quantities of heat in a small temperature range. Segments BC and DE below the pinch point represent the enthalpies of the overhead vapor streams of REA and REC (REA_OVHD and REC_OVHD), respectively. Segments FG and HI above the pinch point represent the enthalpies of the fresh MeOH stream and the bottom stream of REA (MeOH and BTM), respectively. The region above the pinch point is the net heat sink, while that below the pinch point is the net heat source. As can be seen from Figure 4b, the usage of hot utilities and cold utilities is nearly equal, whereas minor heat can be recovered. Though hot streams release a large amount of heat, the low temperature of hot streams restricts them to match with the cold streams. As a heat recovery system to pump heat backward, heat can be transferred from the lowtemperature level to the high-temperature level by a heat pump. Thus, hot and cold utilities can be cut down and the thermodynamic efficiency of the process can be further improved. Although the implementation of a heat pump system can cut down the steam consumption to some extent, the economics should also be evaluated since the compressor consumes plenty of electricity at the same time. Before designing the SHRRDWC, the coefficient of performance (COP) is used to pre-

estimate the feasibility of the implementation of a heat pump system. The COP is calculated by the following equation:44 COP =

QC W

=

TC TE − TC

(14)

where QC is the heat load; W is the compressor duty; TE (K) is the evaporating temperature; TC (K) is the condensing temperature. If the COP is higher than 10, the heat pump is clearly recommended. If it is between 5 and 10, it should be evaluated in more detail. For a COP lower than 5, it is not recommended. Townsend et al.45 pointed out that energy is saved overall only when the heat pump operates across the pinch. That is, heat must be removed below the pinch point and supplied above the pinch point. The T−H diagram and GCC can be used to analyze the appropriate placement of the heat pump. As demonstrated in Figure 4b, the hot stream REA_OVHD and REC_OVHD contain large amount of condensing heat that can be recovered by the heat pump. For the process stream REA_OVHD, its temperature can be lifted to heat the cold stream BTM or MeOH. The COP between REA_OVHD and BTM is 7.4, indicating that the implementation of a heat pump system should be considered in detail due to the large temperature gap between the two streams. The COP between REA_OVHD and MeOH is 33.3, which is an ideal value for the implementation of a heat pump system. For another hot stream REC_OVHD, although its latent heat is much lower than that of REA_OVHD, its potential energy-saving capacity is also worth investigating. The stream REC_OVHD can be reused to heat BTM or MeOH after elevating its temperature level. The COP between REC_OVHD and BTM is 9.36, and the COP between REC_OVHD and MeOH is 241.5. Therefore, the matching between these streams maybe also feasible and economical. In the heat pump system, the temperature level of hot stream is lifted through compressing. Therefore, the discharge pressure is a direct factor affecting the temperature difference in heat exchangers. The discharge pressure of the compressor should at least ensure that the temperature difference in the heat exchanger is greater than ΔTmin. The compressed hot stream will change from superheating, condensing, and maybe subcooling in the heat exchanger.30 Assuming that countercurrent heat transfer takes place in the heat exchanger, since the majority of the heat recovered from the hot stream is latent heat and phase change occurs in the heat exchanger, the dew point of the hot stream should at least larger than that of the 12184

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

Figure 6. T−H diagram and GCC of SHR-RDWC-I.

Figure 7. Grid diagram of SHR-RDWC-I.

utility requirement is 281.5 kW, and the hot utility is no longer required in the process. In this case, the maximum heat recovery is 1174.1 kW. Due to the work done by the compressor to the hot stream, the total enthalpy of the hot streams is increased. The increment is equal to the work done by the compressor. It is clear that hot utility and cold utility can be cut down significantly. The wasted heat can be recuperated and recycled effectively through the heat pump system. After determining the energy target, SHR-RDWC-I can be designed and optimized in Aspen Energy Analyzer based on the basic HEN design principles.28 The grid diagram is a helpful representation for the design of a heat exchanger network, in which streams are plotted as horizontal lines. High temperatures are on the left while cold temperatures are on the right which describe the countercurrent flow pattern in the heat exchangers.28 The matches between the hot and cold streams are represented by two linked circles. Thus, it is convenient to check the feasibility of exchanger temperatures. The grid diagram of SHR-RDWC-I is illustrated in Figure 7. In SHR-RDWC-I, the hottest hot stream REA_OVHD is matched against the hottest cold stream BTM which gives the best temperature driving forces and ensures feasibility. The MeOH is first matched with water, which causes the maximum load on this match. The residual energy requirement on MeOH is fully supplied by REA_OVHD. Both of the cold streams are heated to the target temperatures, so the remaining heat of the hot streams is removed by three condensers. Three heat exchangers and three

cold stream, and the bubble point of the hot stream is at least larger than that of the cold stream. The supply temperature of BTM is 101.0 °C, and the target temperature is 103.2 °C. Thus, the dew point of the compressed hot stream is at least 108.2 °C and the bubble point is at least 106.0 °C. The boiling point of MeOH is 66.3 °C since the stream is nearly pure component. The bubble point of the compressed hot stream is at least 71.3 °C. After determining the bubble or dew points of the compressed hot streams, the discharge pressures of compressors in the heat pump systems can be determined from the pressure−temperature (P−T) curves of the hot streams as shown in Figure 5. 4.2.2. SHR-RDWC-I. Four possible locations for the heat pump placement have been decided and the method for the determination of the discharge pressures of the compressors has been introduced in section 4.2.1. The stream data can be obtained and the HEN design can be carried out in the following sections. As shown in Figure 4, the hot stream REA_OVHD can be matched with the cold stream BTM or MeOH after the temperature level is lifted as a result of the large amount of latent heat. The heat exchange between the REA_OVHD and BTM is considered in SHR-RDWC-I. The discharge pressure of the compressor in the heat pump system is decided to 464.4 kPa according to Figure 5. All the stream data are extracted as shown in Table S3. The T−H diagram and GCC can be plotted based on the minimum temperature difference, as shown in Figure 6. The height of the HCC has been elevated after being compressed. The minimum cold 12185

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

Figure 8. Grid diagram of SHR-RDWC-II.

the other part is rejected to the condenser directly. Because the compressor duty is directly related to the flow rate of the stream, the operating cost and capital cost can be cut down to some extent. Though less latent heat is contained in the hot stream REC_OVHD, the COP values between REC_OVHD and BTM or MeOH are much higher than that of REA_OVHD, as shown in section 4.2.1. Thus, a more economical process can be conducted if we match the REC_OVHD with the two cold streams. In this case, part of the heat duty of REB or PHE1 can be supplied by REC_OVHD and the inlet flow rate of COMP1 can be further reduced. For the matching between REC_OVHD and MeOH, the discharge pressure of the compressor can be easily decided to 130.0 kPa according to Figure 5, while for the heat exchange between REC_OVHD and BTM, the REC_OVHD has to be compressed to 469.0 kPa and the discharge temperature of the compressor is raised to 181.1 °C, which threads the safe operation of the process.34 Thus, the match between REC_OVHD and BTM is not considered in this paper. After extracting the stream data and analyzing the unreasonable heat exchanges in SHR-RDWC-I, further improvements can be carried out in SHR-RDWC-II. The grid diagram of SHR-RDWC-II is illustrated in Figure 8, and the stream data are provided in Table S4. One more compressor (COMP2) is added to the REC_OVHD stream and one more feed preheater (PHE2) can be added to the flow sheet; thus part of the heat in PHE1 can be moved to PHE2. In this case, the heat load of PHE1 can be reduced accordingly. The REA_OVHD is divided into two parts: one part with a split ratio of βV1 is pressurized by COMP1 to elevate the temperature level for the sake of exchanging heat with REB and guaranteeing the target temperature of MeOH; the other stream is rejected to the condenser directly. The βV1 is changed to satisfy the ΔTmin in PHE1, and the optimum βV1 is 0.859. About 12.6% electricity is saved in this case due to the splitting and implementation of COMP2. The corresponding flow sheet diagram is shown in Figure S10. The T−H diagram and GCC of SHR-RDWC-II are shown in Figure 9. The total enthalpy of hot streams has been

condensers are required in SHR-RDWC-I. The corresponding flow sheet diagram is shown in Figure S9. The total heat recovery in SHR-RDWC-I is 1174.1 kW, which satisfies the aim of maximum heat recovery. No external heat is required in SHR-RDWC-I. With the addition of compressors and heat exchangers, latent heat and sensible heat in the process can be reused. The improved process only consumes a small amount of electricity. Therefore, energy consumption has been reduced dramatically. Compared with the RDWC-FP, the SHR-RDWC-I achieves 36.9% reduction in TEC. Besides, the TAC decreases from 424.6 × 103 to 402.9 × 103 $/year, which has been reduced by 5.1%. Although capital cost increases sharply due to the addition of compressor and heat exchanger, the reductions in operating cost can compensate for the increments. 4.2.3. SHR-RDWC-II. The effectiveness of heat exchange can be achieved when the exergy loss is the lowest, corresponding to the temperature differences in the heat exchangers are minimum,46 which can be identified from the T−H diagram and GCC. Although SHR-RDWC-I has achieved the target of maximum heat recovery, further energy-saving potential can be explored with the aim of minimizing the temperature differences in heat exchangers. As can be seen from Figure 6, no pinch point is detected and all the temperature differences in heat exchangers are higher than the ΔTmin. The process can be further optimized with the aim of minimizing the temperature difference. In SHR-RDWC-I, the effluent temperature of hot stream in PHE1 is still high as shown in Figure 7. The hot stream outlet temperature can be reduced to 71.3 °C when the cold end of PHE1 satisfies the ΔTmin. The excess energy is eventually taken out and wasted by the cooling water. The compressor consumes electricity to compress the hot stream; however, if the consumed energy has not been used to the maximum extent, the improvement is not economical. In this case, not only electricity is wasted but also equipment cost is increased. A reasonable solution to this problem is to split the hot stream REA_OVHD into two parts properly. One part is pressurized by the compressor to elevate the temperature level, whereas 12186

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

Figure 9. T−H diagram and GCC of SHR-RDWC-II.

Figure 10. Grid diagram of SHR-RDWC-III.

configuration and capital cost increases slightly, the added COMP2 induces the reduction of inlet flow rate of COMP1, which decreases the operating cost and capital cost accordingly. 4.2.4. SHR-RDWC-III. The HCC is much closer to the CCC in SHR-RDWC-II which means the temperature differences have been further reduced. However, temperature differences are still high which can be identified from Figure 9. Therefore, the exergy loss in SHR-RDWC-II is still large. Operating conditions can be further adjusted to minimize the temperature differences. The hot stream outlet temperature of REB in SHR-RDWCII is 107.8 °C, which is higher than the bubble point of the hot stream. A small amount of latent heat is fed into PHE1 and exchanges heat with the MeOH. In this case, the temperature level of REA_OVHD is elevated by the compressor and the hot stream is used to exchange heat with a much lower temperature level stream. However, the temperature of the hot

decreased by 23.5 kW due to the reduction of the total compressor duty. As a result, the cold utility requirements have been cut down accordingly. In comparison with the HCC of SHR-RDWC-I, the heat content of the high-temperature level has been reduced. The minimum cooling duty for SHRRDWC-II is 258.8 kW, which is equal to the total condenser duty in SHR-RDWC-II. Since the enthalpy of the hot streams has been decreased, and the CCC shifted on the H-axis relative to the hot streams, an intermediate near-pinch can be detected, which can be identified from the HCC and CCC as a region of close temperature approach and from the GCC that the enthalpy is closing to zero. The SHR-RDWC-II is further optimized based on SHRRDWC-I with the aim of minimizing the temperature difference. Compared with SHR-RDWC-I, the TEC has reduced from 698.0 to 632.6 kW and the TAC has been reduced from 402.9 × 103 to 393.1 × 103 $/year. Although one compressor and one heat exchanger are added in this 12187

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research

Figure 11. Flow sheet diagram of SHR-RDWC-III.

Figure 12. T−H diagram and GCC of SHR-RDWC-III.

discharge pressure of COMP3 is chosen as 166.9 kPa, which is determined by the method mentioned in section 4.2.1. The βV1 is changed to satisfy the ΔTmin in REB, and the optimum βV1 is 0.640. The hot stream outlet temperature of PHE1 is fixed at 71.3 °C. The βV2 is varied to satisfy the ΔTmin in PHE3, and the minimum βV2 is 0.246. The corresponding flow sheet diagram is shown in Figure 11. The T−H diagram and GCC of SHR-RDWC-III are shown in Figure 12. The total enthalpy of hot streams has been decreased by 26.2 kW due to the reduction of compressor duty. As a result, the cold utility requirements have been cut down accordingly. The minimum cooling load for SHRRDWC-III is 232.6 kW, which is equal to the total condenser and cooler duty in SHR-RDWC-III. The HCC is much different since the enthalpy of the hot streams has been decreased and part of the vapor is elevated to a relatively lower

stream cannot be further reduced due to the limitation of ΔTmin in REB. To address the problem, we proposed the SHRRDWC-III based on SHR-RDWC-II. The inlet flow rate of COMP1 is redistributed again so that part of the stream is compressed through a compressor with a smaller compression ratio for the purpose of heating the cold stream MeOH. The inlet flow rate of COMP1 can be reduced, and thereby operating cost can be cut down accordingly. The grid diagram of SHR-RDWC-III is illustrated in Figure 10, and the stream data are provided in Table S5. In SHRRDWC-III, the overhead vapor stream REA_OVHD is divided into three parts: one part with a split ratio of βV1 is pressurized by COMP1 to elevate the temperature level for the sake of exchanging heat with REB. Another part with a split ratio of βV2 is pressurized by COMP3 in order to make up for the remaining heat required by the cold stream MeOH. The 12188

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research Table 1. Comparison of Basic and Improved Configurations parameter TEC (kW) thermodyn efficiency (%) CO2 emissions (kg/h) CC (103 $) OC (103 $/year) TAC (103 $/year) 3 years 5 years 8 years 10 years

RDWC

RDWC-FP

SHR-RDWC-I

SHR-RDWC-II

SHR-RDWC-III

1221.3 22.4 286.1 467.5 391.0

1107.2 (−9.3%) 23.7 259.3 (−9.3%) 475.9 365.1

698.0 (−42.8%) 27.6 47.3 (−83.5%) 1403.4 227.5

632.6 (−48.2%) 29.3 42.9 (−85.0%) 1409.7 216.9

511.4 (−58.1%) 31.4 34.7 (−87.9%) 1378.3 197.3

546.8 484.5 449.4 437.8

523.7 460.3 424.6 412.7

695.2 508.1 402.9 367.8

686.8 498.9 393.1 357.9

646.8 473.0 369.6 335.1

(−4.2%) (−5.0%) (−5.5%) (−5.7%)

(+27.1%) (+4.9%) (−10.4%) (−16.0%)

(+25.6%) (+3.0%) (−12.5%) (−18.3%)

(+20.1%) (−2.4%) (−17.8%) (−23.5%)

The SHR-RDWC-III shows the best environmental performance for the elimination of steam requirement and minimum electricity consumption. The thermodynamic efficiencies of the three self-heat recuperative processes are 27.6, 29.3, and 31.4%, respectively, while that of the RDWC is only 22.4%. Owing to the different schemes of recovering the latent heat and sensible heat, striking reductions in exergy loss of these configurations has been attained compared with the RDWC. Finally, the effect of payback period on TAC is demonstrated in Table 1. Although the SHR-RDWC processes have comparable operating costs, all of the SHR-RDWC processes show poor economic performance with shorter payback periods. In this case, the TAC of the SHR-RDWC processes are dominated by the huge capital investment, especially for the implementation of compressors. Since the significant energy savings largely compensate for the capital cost, the SHR-RDWC processes provide favorable performance in terms of TAC compared with the RDWC and the RDWC-FP for 8 or 10 years payback period. Besides, the differences in TAC are greater for increasing payback periods. This is because it requires a longer period to recover the capital cost. From our results, it is evident that the SHR-RDWC-III is superior among all cases as it exhibits the least TAC for a longer payback period.

temperature level compared with the SHR-RDWC-II. The CCC shifted on the H-axis relative to the hot stream. The HCC and CCC are close to each other, and the HCC is almost parallel to the CCC. There are two intermediate near-pinches caused by the bend of the HCC or CCC which mean minimum temperature differences. The temperature differences in other places are slightly higher than the minimum temperature difference. Therefore, exergy loss has be minimized in the process and all the latent heat raised by the heat pump systems and sensible heat have been utilized rationally. The SHR-RDWC-III is optimized based on SHR-RDWC-II with the aim of minimizing the temperature differences. Although one more compressor is added in this configuration, the TEC decreased from 632.6 to 511.4 kW. The total heat exchanger cost is much higher than those for the other two self-heat recuperative processes because of the smaller temperature approaches. However, the TAC reduced from 393.1 × 103 to 369.6 × 103 $/year compared with SHRRDWC-II. The total compressor duty has been cut down and capital cost and operation cost are saved, which offset the increments of heat exchanger investment. 4.3. Comparison. Detailed comparisons of RDWC, RDWC-FP, and SHR-RDWC are listed in Table 1 for the identification of strengths and weaknesses. Performances of the conventional and improved processes are compared from the views of TEC, TAC, CO2 emissions, and thermodynamic efficiency, and the indicators are considered simultaneously except for only reference of TAC in the selection of the best configuration. TAC is calculated and values are compared as shown in Table 1. In addition, detailed costs of equipment and utilities are also offered in Table S6. The results indicate that the basic RDWC together with the RDWC-FP requires the highest operating cost and the lowest capital cost while the SHRRDWC processes are just the opposite. The addition of compressors and heat exchangers results in the higher capital costs of SHR-RDWC processes. However, we can see from the results that the operating cost of the system can be significantly reduced by the addition of the equipment. Therefore, the reductions in operating cost can compensate for the increments of capital cost. In comparison with the basic RDWC, the TAC of the three self-heat recuperative processes decreases by 10.4, 12.5, and 17.8%, respectively, of which the SHR-RDWCIII shows the best economic performance. CO2 emissions are used to evaluate the environmental benefits of these processes. The CO2 emissions of the three self-heat recuperative processes have been reduced by 83.5, 85.0, and 87.9%, respectively, in comparison with the RDWC.

5. CONCLUSIONS Intensified RDWC configurations via self-heat recuperation technology are proposed for the production of methyl acetate in this paper. The heat exchanger network using pinch analysis is applied to optimize the energy utilization. The SHR-RDWC processes are compared with the RDWC in terms of TEC, TAC, CO2 emissions, and thermodynamic efficiency. The results reveal that the intensified processes based on the selfheat recuperation technology are inferior to the conventional processes for a shorter payback period because of the huge capital investment. Nevertheless, the significant energy savings largely compensate for the capital cost. The optimal self-heat recuperative process (SHR-RDWC-III) achieves 17.8% reduction in TAC when the payback period is 8 years, in contrast to the basic RDWC. The intensified processes also exhibit remarkable environmental performance. In the practical application of a chemical process, it should not only consider the steady-state performance but also dynamic controllability aspects. The implementation of compressors and heat exchangers deepen the complexity of the RDWC process, which challenges the dynamic controllability of the proposed process. Future efforts should be geared toward the control 12189

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research REC = rectifying column RD = reactive distillation RDWC = reactive dividing wall column RR = reflux ratio S = entropy TAC = total annual cost TEC = total energy consumption Wmin = minimum separation work, kJ/kmol

study of the self-heat recuperative reactive dividing wall column.

■

ASSOCIATED CONTENT

S Supporting Information *

The Supporting Information is available free of charge on the ACS Publications website at DOI: 10.1021/acs.iecr.9b02363. Optimization procedure for RDWC; effect of design variables on TAC; composition and temperature profiles of RDWC; reactive profiles of the REA column (RDWC); composition and temperature profiles of RDWC-FP; reactive profiles of the REA column (RDWC-FP); flow sheet diagrams of SHR-RDWC-I and SHR-RDWC-II; cost estimating formulas; stream data; capital costs and operating costs (PDF)

■

■

REFERENCES

(1) de Koeijer, G. M.; Kjelstrup, S. Minimizing entropy production rate in binary tray distillation. Int. J. Thermodyn. 2000, 3, 105−110. (2) Kiss, A. A.; Flores Landaeta, S. J.; Infante Ferreira, C. A. Towards energy efficient distillation technologies − Making the right choice. Energy 2012, 47, 531−542. (3) Zhai, J.; Liu, Y.; Li, L.; Zhu, Y.; Zhong, W.; Sun, L. Applications of dividing wall column technology to industrial-scale cumene production. Chem. Eng. Res. Des. 2015, 102, 138−149. (4) Dejanović, I.; Matijašević, L.; Olujić, Ž . Dividing wall column A breakthrough towards sustainable distilling. Chem. Eng. Process. 2010, 49, 559−580. (5) Weinfeld, J. A.; Owens, S. A.; Eldridge, R. B. Reactive dividing wall columns: A comprehensive review. Chem. Eng. Process. 2018, 123, 20−33. (6) Asprion, N.; Kaibel, G. Dividing wall columns: Fundamentals and recent advances. Chem. Eng. Process. 2010, 49, 139−146. (7) Luyben, W. L.; Yu, C. Reactive Distillation Design and Control; John Wiley & Sons: Hoboken, NJ, 2008. (8) Kaur, J.; Sangal, V. K. Reducing energy requirements for ETBE synthesis using reactive dividing wall distillation column. Energy 2017, 126, 671−676. (9) Li, L.; Sun, L.; Yang, D.; Zhong, W.; Zhu, Y.; Tian, Y. Reactive dividing wall column for hydrolysis of methyl acetate: Design and control. Chin. J. Chem. Eng. 2016, 24, 1360−1368. (10) Zheng, L.; Cai, W.; Zhang, X.; Wang, Y. Design and control of reactive dividing-wall column for the synthesis of diethyl carbonate. Chem. Eng. Process. 2017, 111, 127−140. (11) An, D.; Cai, W.; Xia, M.; Zhang, X.; Wang, F. Design and control of reactive dividing-wall column for the production of methyl acetate. Chem. Eng. Process. 2015, 92, 45−60. (12) Jana, A. K. Heat integrated distillation operation. Appl. Energy 2010, 87, 1477−1494. (13) Jana, A. K. Advances in heat pump assisted distillation column: A review. Energy Convers. Manage. 2014, 77, 287−297. (14) Long, N. V. D.; Kwon, Y.; Lee, M. Design and optimization of thermally coupled distillation schemes for the trichlorosilane purification process. Appl. Therm. Eng. 2013, 59, 200−210. (15) Chaniago, Y. D.; Minh, L. Q.; Khan, M. S.; Koo, K.-K.; Bahadori, A.; Lee, M. Optimal design of advanced distillation configuration for enhanced energy efficiency of waste solvent recovery process in semiconductor industry. Energy Convers. Manage. 2015, 102, 92−103. (16) Gao, X.; Chen, J.; Tan, J.; Wang, Y.; Ma, Z.; Yang, L. Application of mechanical vapor recompression heat pump to doubleeffect distillation for separating N, N-dimethylacetamide/water mixture. Ind. Eng. Chem. Res. 2015, 54, 3200−3204. (17) Shahandeh, H.; Jafari, M.; Kasiri, N.; Ivakpour, J. Economic optimization of heat pump-assisted distillation columns in methanolwater separation. Energy 2015, 80, 496−508. (18) Liu, Y.; Zhai, J.; Li, L.; Sun, L.; Zhai, C. Heat pump assisted reactive and azeotropic distillations in dividing wall columns. Chem. Eng. Process. 2015, 95, 289−301. (19) Feng, S.; Lyu, X.; Ye, Q.; Xia, H.; Li, R.; Suo, X. Performance enhancement of reactive dividing-wall column via vapor recompression heat pump. Ind. Eng. Chem. Res. 2016, 55, 11305−11314. (20) Feng, S.; Ye, Q.; Xia, H.; Li, R.; Suo, X. Integrating a vapor recompression heat pump into a lower partitioned reactive dividing-

AUTHOR INFORMATION

Corresponding Author

*Tel.: +86 13854208340. Fax: +86 86981787. E-mail: [email protected]. ORCID

Lanyi Sun: 0000-0002-3158-6388 Notes

The authors declare no competing financial interest.

■

ACKNOWLEDGMENTS This work was supported by the National Natural Science Foundation of China (Grant 21476261) and the Graduate Innovation Project (YCX2018033). Finally, the authors are grateful to the editor and the anonymous reviewers.

■

NOMENCLATURE ai = activity of component i B = exergy, kJ/kmol CC = capital cost, $ COP = coefficient of performance DWC = dividing wall column H = enthalpy, kJ/kmol HAc = acetic acid HEN = heat exchanger network ki = rate constant, kmol/(kg·s) LS = vapor side stream flow rate, kg/h LW = lost work, kJ/h m = stream flow rate, kmol/h mcat = mass of catalyst, kg MeAc = methyl acetate MeOH = methanol NHAc = feed location of acetic acid NMeOH = feed location of methanol NR = number of stages in rectifying section NRX = number of reactive stages NREC = number of stages of REC NS = number of stages in stripping section NVR = withdraw location of vapor side stream OC = operating cost, $/year Qi = duty of equipment i QFuel = amount of fuel burnt, kW r = reaction rate, kmol/s R = gas constant, kJ/(kmol·K) βV = split ratio REA = reactive distillation column 12190

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191

Article

Industrial & Engineering Chemistry Research wall column for better energy-saving performance. Chem. Eng. Res. Des. 2017, 125, 204−213. (21) Feng, Z.; Shen, W.; Rangaiah, G. P.; Lv, L.; Dong, L. Process development, assessment, and control of reactive dividing-wall column with vapor recompression for producing n-propyl acetate. Ind. Eng. Chem. Res. 2019, 58, 276−295. (22) Kansha, Y.; Tsuru, N.; Sato, K.; Fushimi, C.; Tsutsumi, A. Selfheat recuperation technology for energy saving in chemical processes. Ind. Eng. Chem. Res. 2009, 48, 7682−7686. (23) Kansha, Y.; Ishizuka, M.; Song, C.; Tsutsumi, A. Process intensification for dimethyl ether production by self-heat recuperation. Energy 2015, 90, 122−127. (24) Matsuda, K.; Kawazuishi, K.; Kansha, Y.; Fushimi, C.; Nagao, M.; Kunikiyo, H.; Masuda, F.; Tsutsumi, A. Advanced energy saving in distillation process with self-heat recuperation technology. Energy 2011, 36, 4640−4645. (25) van Duc Long, N.; Lee, M. Novel acid gas removal process based on self-heat recuperation technology. Int. J. Greenhouse Gas Control 2017, 64, 34−42. (26) van Duc Long, N.; Lee, M. A novel NGL (natural gas liquid) recovery process based on self-heat recuperation. Energy 2013, 57, 663−670. (27) Christopher, C. C. E.; Dutta, A.; Farooq, S.; Karimi, I. A. Process synthesis and optimization of propylene/propane separation using vapor recompression and self-heat recuperation. Ind. Eng. Chem. Res. 2017, 56, 14557−14564. (28) Kemp, I. C. Pinch Analysis and Process Integration: A User Guide on Process Integration for the Efficient Use of Energy; ButterworthHeinemann: Oxford, 2007. (29) Xia, H.; Ye, Q.; Feng, S.; Li, R.; Suo, X. A novel energy-saving pressure swing distillation process based on self-heat recuperation technology. Energy 2017, 141, 770−781. (30) Chen, J.; Ye, Q.; Liu, T.; Xia, H.; Feng, S. Improving the performance of heterogeneous azeotropic distillation via self-heat recuperation technology. Chem. Eng. Res. Des. 2019, 141, 516−528. (31) Xia, H.; Ye, Q.; Feng, S.; Chen, J.; Liu, T. Energy-efficient design of downstream separation to produce n-butanol by several heat-integrated technologies. Ind. Eng. Chem. Res. 2018, 57, 13205− 13216. (32) Iwakabe, K.; Nakaiwa, M.; Huang, K.; Nakanishi, T.; Røsjorde, A.; Ohmori, T.; Endo, A.; Yamamoto, T. Energy saving in multicomponent separation using an internally heat-integrated distillation column (HIDiC). Appl. Therm. Eng. 2006, 26, 1362−1368. (33) Douglas, J. M. Conceptual Design of Chemical Processes; McGraw-Hill: New York, 1988. (34) Luo, H.; Bildea, C. S.; Kiss, A. A. Novel heat-pump-assisted extractive distillation for bioethanol purification. Ind. Eng. Chem. Res. 2015, 54, 2208−2213. (35) Yu, H.; Ye, Q.; Xu, H.; Zhang, H.; Dai, X. Design and control of dividing-wall column for tert-butanol dehydration system via heterogeneous azeotropic distillation. Ind. Eng. Chem. Res. 2015, 54, 3384−3397. (36) Dimian, A. C. Integrated Design and Simulation of Chemical Processes; Elsevier: Amsterdam, 2003. (37) Shi, L.; Huang, K.; Wang, S. J.; Yu, J.; Yuan, Y.; Chen, H.; Wong, D. S. H. Application of vapor recompression to heterogeneous azeotropic dividing-wall distillation columns. Ind. Eng. Chem. Res. 2015, 54, 11592−11609. (38) Gadalla, M. A.; Olujić, Ž .; Jansens, P. J.; Jobson, M.; Smith, R. Reducing CO2 emissions and energy consumption of heat-integrated distillation systems. Environ. Sci. Technol. 2005, 39, 6860−6870. (39) Haar, L.; Gallagher, J. S.; Kell, G. S. NBS/NRC Steam Tables: Thermodynamic and Transport Properties and Computer Programs for Vapor and Liquid States of Water in SI Units; Hemisphere Publishing: Washington, DC, 1984. (40) Flórez-Orrego, D.; Silva, J. A. M.; Oliveira, S. de., Jr. Renewable and non-renewable exergy cost and specific CO2 emission of electricity generation: The Brazilian case. Energy Convers. Manage. 2014, 85, 619−629.

(41) Seider, W. D.; Lewin, D. R.; Seader, J. D.; Widagdo, S.; Gani, R.; Ng, K. M. Product and Process Design Principles: Synthesis, Analysis and Evaluation; John Wiley & Sons: New York, 2017. (42) Tang, Y.; Chen, Y.; Huang, H.; Yu, C.; Hung, S. B.; Lee, M. J. Design of reactive distillations for acetic acid esterification. AIChE J. 2005, 51, 1683−1699. (43) Bandyopadhyay, S. Effect of feed on optimal thermodynamic performance of a distillation column. Chem. Eng. J. 2002, 88, 175− 186. (44) Pleşu, V.; Bonet Ruiz, A. E.; Bonet, J.; Llorens, J. Simple equation for suitability of heat pump use in distillation. In 24th European Symposium on Computer Aided Process Engineering; Klemeš, J. J., Varbanov, P. S., Liew, P. Y., Eds.; Elsevier: 2014; Vol. 33, pp 1327−1332. (45) Townsend, D. W.; Linnhoff, B. Heat and power networks in process design. Part I: Criteria for placement of heat engines and heat pumps in process networks. AIChE J. 1983, 29, 742−748. (46) Song, C.; Liu, Q.; Ji, N.; Kansha, Y.; Tsutsumi, A. Optimization of steam methane reforming coupled with pressure swing adsorption hydrogen production process by heat integration. Appl. Energy 2015, 154, 392−401.

12191

DOI: 10.1021/acs.iecr.9b02363 Ind. Eng. Chem. Res. 2019, 58, 12179−12191