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Integrating a vapor recompression heat pump into a lower partitioned reactive dividing-wall column for better energy-saving performance
Accepted Manuscript Title: Integrating a vapor recompression heat pump into a lower partitioned reactive dividing-wall column for better energy-saving performance Authors: Shenyao Feng, Qing Ye, Hui Xia, Rui Li, Xiaomeng Suo PII: DOI: Reference:
S0263-8762(17)30384-2 http://dx.doi.org/doi:10.1016/j.cherd.2017.07.017 CHERD 2758
To appear in: Received date: Revised date: Accepted date:
3-1-2017 11-7-2017 12-7-2017
Please cite this article as: Feng, Shenyao, Ye, Qing, Xia, Hui, Li, Rui, Suo, Xiaomeng, Integrating a vapor recompression heat pump into a lower partitioned reactive dividingwall column for better energy-saving performance.Chemical Engineering Research and Design http://dx.doi.org/10.1016/j.cherd.2017.07.017 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Integrating a vapor recompression heat pump into a lower partitioned reactive dividing-wall column for better energy-saving performance
Shenyao Feng, Qing Ye*, Hui Xia, Rui Li, and Xiaomeng Suo
Jiangsu Key Laboratory of Advanced Catalytic Materials and Technology, School of Petrochemical Engineering, Changzhou University, Changzhou, Jiangsu 213164, China *
Corresponding author. Tel.: +86 519 86330355. Fax: +86 519 86330355. E-mail:
[email protected].
Graphical abstract
Highlights: ο·
Integrated lower partitioned reactive dividing-wall column with vapor recompression heat pump models are proposed.
ο·
A remarkable energy-saving performance improvement is achieved by the integrated model.
ο·
CO2 emissions of the process are totally reduced by the integrated model.
Abstract: In this study, four integrated lower partitioned reactive dividing-wall columns (LRDWCs) with vapor recompression heat pump (VRHP) models are developed to reduce the energy consumption of a reactive dividing-wall column. Energy-saving, economic, and environmental performances of the integrated LRDWC with VRHP models are compared with the corresponding LRDWC. The results indicate that the integration of VRHP into the LRDWC significantly reduces the energy consumption and CO2 emissions associated with the process. The application of a preheater further improves the energy-saving and economic performance of the process. With the aid of a preheater, the integrated LRDWC with the VRHP model saves 49.86% energy consumption and 13.76% total annual cost with a payback period corresponding to five years, and reduces all CO2 emissions when compared with the corresponding LRDWC.
Keywords: Vapor recompression heat pump; Reactive dividing-wall column; Energy saving Nomenclature CI
capital investment
COMP
compressor
CON
condenser of the lower partitioned reactive dividing-wall column
DWC
dividing-wall column
IPAC
isopropyl acetate
IPOH
isopropanol
LHHW
Langmuir Hinshelwood Hougen Watson model
LMTD
logarithmic mean temperature difference
LRDWC
lower partitioned reactive dividing-wall column
MEAC
methyl acetate
MEOH
methanol
OC
operating cost
PHE
preheater
RC
recovery
RD
reactive distillation
RDWC
reactive dividing-wall column
REB1
reboiler of the reactive distillation part
REB2
reboiler of the recovery part
TAC
total annual cost
TUC
total utility consumption
UNIQUAC
universal quasi-chemical activity coefficient
URDWC
upper partitioned reactive dividing-wall column
VRHP
vapor recompression heat pump
1.
Introduction
Distillation is the most widely used separation process for multi-component mixtures in chemical industries although it is considerably energy intensive (Jana, 2010; You et al., 2016). Energy consumptions of the distillation process are of great concern due to increasing energy demands (Jana, 2016a, 2016b, 2016c; Liu et al., 2015; Yasuhiko et al., 2012).
A dividing-wall column (DWC) is an attractive design that saves equipment and operating costs of the process and reduces energy consumption when compared with traditional separation sequences (Errico et al., 2009; Novita et al., 2015). Reactive distillation (RD) combines reaction and separation steps in a single unit. The application of RD increases conversion and selectivity, and reduces energy consumption when compared with those of conventional reactor and distillation sequences (Kiss et al., 2007). The reactive dividing-wall column (RDWC) integrates the RD and DWC in a column; thus, it possesses the advantages of both the RD and DWC (Mueller and Kenig, 2007). The upper partitioned reactive dividing-wall column (URDWC, Fig. 1c) and the lower partitioned reactive dividing-wall column (LRDWC, Fig. 1d) correspond to two popular configurations of the RDWC. In general, an URDWC is used in the place of reactive distillation in a direct sequence (Fig. 1a), and an LRDWC is considered as an alternative to reactive distillation in an indirect sequence (Fig. 1b). The selection of an RDWC configuration is based on physical and chemical parameters and the composition of reactants and products. Dai et al. (2015) studied an URDWC for the synthesis of n-propyl propionate, and the results indicated that the URDWC saved 12.4% and 16.4% energy consumption and total annual cost, respectively, when compared with those corresponding to the conventional two-column process. Zheng et al. (2016) developed an LRDWC to synthesize diethyl carbonate. The results showed that the LRDWC saved 18.7% energy consumption and 13.9% total annual cost when compared with those of the conventional RD process. Kiss and Suszwalak (2012a, 2012b) introduced a novel dimethyl ether process based on methanol dehydration in a RDWC. The results indicated that the process achieved up to 58% energy savings, 60% reduced CO2 emissions, and 30% lower capital investment costs when compared with those of the traditional RD process. The findings of the aforementioned studies indicate that a RDWC significantly reduces energy consumption and investment when compared with those of the corresponding RD process.
A vapor recompression heat pump (VRHP) reduces the energy demand of separating close-boiling mixtures (Waheed et al., 2014). A compressor improves low-grade heat of the overhead vapor into a higher-grade heat; thus, compressed vapor provides heat in a heat exchanger. A study by Danziger (1979) suggested that a VRHP can save energy consumption by 80%. Recently, few studies reported on integrating a VRHP into a DWC. Li et al. (2016) integrated a VRHP into an azeotropic DWC system. The total annual cost and CO2 emissions were reduced by 32.22% and 63.79%, respectively, when compared with those of the corresponding two-column azeotropic system. Luyben (2016) developed a three-temperature control configuration with a pressure compensation in a heterogeneous azeotropic dividing-wall system with vapor recompression. The controls solved disturbances in the feed composition and feed rate to a considerable extent. These studies demonstrated that the integration of a VRHP into a DWC reduces the energy consumption of the whole process. The integration of a VRHP into a RDWC can exhibit the same effect. Liu et al. (2015) proposed a heat pump assisted reactive dividing-wall column that reduces 11.5% total energy consumption, 4.3% total annual cost with a capital payback period of 5 years, and 13.4% CO2 emissions when compared with those of the corresponding RDWC. However, studies that integrate a VRHP into a RDWC are still at a nascent stage. There is a lack of studies focusing on the reduction of energy consumption in a RDWC by using a VRHP. An earlier study (Feng et al., 2016) integrated a VRHP into a URDWC model. The integration of a VRHP into a URDWC significantly reduced energy consumption of the complete process, since the results indicated that the URDWC integrated with the VRHP model saved up to 39.22% total energy consumption when compared with that of the corresponding URDWC model. More specifically, URDWC and LRDWC correspond to two different configurations of the RDWC; thus, an integrated LRDWC with the VRHP model should exhibit same effect. However, studies continue to explore the actual energy-saving performance of the integrated LRDWC with the VRHP models.
This study is focused on improving the energy-saving performance of a LRDWC process by integrating a VRHP into a LRDWC model. The transesterification reaction of methyl acetate (MEAC) and isopropanol (IPOH) is selected as the reaction of the LRDWC. In this study, four integrated LRDWCs with the VRHP models are proposed. The VRHP provides heat in the bottom reboilers, and the influence of the preheater on the integrated LRDWC with the VRHP model is investigated. The energy-saving, economic and environmental performances of all models are examined. 2.
Design basis and methodology
2.1 Thermodynamic and kinetic models Isopropyl acetate (IPAC) and methanol (MEOH) are synthesized by a transesterification reaction of MEAC and IPOH. The chemical equation is shown as follows: ππΈπ΄πΆ + πΌπππ» β πΌππ΄πΆ + ππΈππ»
(1)
This reaction is catalyzed by a strongly acidic ion-exchange resin Lewatit MonoPlus S100 H+ (Akyalçın, 2014). The equilibrium constant (Ka) and the rate constant (Kf) of the reaction by using the Langmuir Hinshelwood Hougen Watson (LHHW) model (Akyalçın, 2014) are given as follows:
πΎπ = expβ‘(
702 β 1.81) π
πΎπ (πππ/π β min) = expβ‘(9.03 β
(2) 4999 ) π
(3)
where T denotes absolute temperature in Kelvin. Simulations are performed by using Aspen Plus 8.4 software. The universal quasi-chemical activity coefficient (UNIQUAC) model is selected to depict the vapor-liquid equilibrium of the simulations. The UNIQUAC model parameters of MEOH +IPAC (Resa et al., 2001), MEAC + MEOH (Tu et al., 1997), IPAC + IPOH (Andreatta et al., 2010) and IPAC + MEOH (Resa et al.,
2001) are regressed by using experimental data. As a result, the UNIQUAC model is in agreement with the experimental data. The UNIQUAC model parameters of the system (Suo et al., 2016) are listed in Table 1. 2.2 Total utility consumption (TUC) For the convenience of comparing the energy-saving performance of the processes, the total utility consumption (TUC) is defined as follows: TUC = βππ πΈπ΅ + βπππ πΈ + πβππΆπππ
(4)
where QREB, QPRE, and QCOMP denote the heat duty of the reboiler, heat duty of the preheater and work of the compressor respectively. In addition, f (=3) represents the multiplication factor of the work of the compressor that converts electrical work into thermal power with the same effect (Kumar et al., 2013). 2.3 Total annual cost (TAC) It is necessary to simultaneously consider the capital cost and the operating cost while evaluating the economic performance of a process. Thus, the economic potential of the processes is assessed in terms of total annual cost (TAC)and expressed as follows:
TAC = operatingβ‘cost +
capitalβ‘investment paybackβ‘period
(5)
The capital investment (CI) is calculated by summing up all the investments of the equipment, including column vessels, heat exchangers and compressors. The cost estimating formula (Douglas, 1988; Luyben, 2013) is shown in Table 2. The value of the Marshall and Swift (M&S) index is set as 1293 (Luyben, 2011), and the motor efficiency of the compressor corresponds to 60% (Douglas, 1988). The low-pressure steam produced by the steam boiler is used in reboilers, cooling water at a temperature of 25Β°C is used in the condensers, and electricity produced by the gas turbine is used in the compressors. Heat transfer coefficients of the condensers and the reboilers
correspond to 0.852 and 0.568 kW/(KΒ·m2), respectively (Luyben, 2013). The smaller equipment investment and catalyst costs are neglected, since the catalyst is repeatedly utilized after reproduction. The operating cost (OC) is calculated based on the utility cost for a year with 8000 operating hours. The price of low-pressure steam, cooling water and electricity correspond to 7.78 $/GJ, 0.03 $/kg and 16.8 $/GJ respectively, based on recommendations specified by Luyben (2013, 2011). 2.4 CO2 emission Specifically, CO2 is generated by burning a fuel and air mixture based on a stoichiometric expression (Gadalla et al., 2005) as follows: π¦ π¦ πΆπ₯ π»π¦ + (π₯ + ) π2 β π₯πΆπ2 + π»2 π 4 2
(6)
where x and y represent the number of carbon C and hydrogen H atoms respectively. The CO2 emission ([πΆπ2 ]πΈπππ π ) is described as follows: [πΆπ2 ]πΈπππ π = (
ππΉπ’ππ C% )( )πΌ ππ»π 100
(7)
where Ξ± (=3.67) denotes molar mass proportion of CO2 and C, and NHV (=22000 kJ/kg) represents the net thermal value of fuels. Carbon content (C%) of the coal is set as 0.68 (Seader et al., 2010). The flue gas produced by combusting fuel is used as a heat source in the steam boiler and gas turbine (Jana, 2015). The quantity of the combusted fuel (ππΉπ’ππ ) is calculated as follows:
ππΉπ’ππ =
πππππ ππΉππ΅ β ππ (βππππ β 419) πππππ ππΉππ΅ β ππ π‘πππ
(8)
where πππππ (kJ/kg) and βππππ (kJ/kg) denote latent heat and the enthalpy of steam respectively; and ππ (Β°C), ππΉππ΅ (=1800Β°C), and ππ π‘πππ (=160Β°C) denote the ambient
temperature, flame temperature, and stack temperature respectively (Jantes-Jaramillo et al., 2008). 3.
Lower partitioned reactive dividing-wall column (LRDWC)
The process flow diagram of a LRDWC is shown in Fig. 2. The LRDWC is divided into three parts, namely the top part, reactive distillation (RD) part and recovery (RC) part. The top part, RD part and RC part include 10, 74 and 8 stages, respectively. The RD part is divided into three sections: the rectifying section, reactive section, and stripping section. The esterification reaction of MEAC and IPOH occurs in the reactive section. In addition, IPOH and MEAC are fed into the RD part at the top and bottom of the reactive section respectively with a flowrate of 50 kmol/h. The top stream is condensed in the condenser of the LRDWC (CON) and is then separated into two streams. A stream is refluxed back to the LRDWC, and the other stream is recycled back to the RD part, with the MEAC feed stream. The IPAC product is obtained from the bottom of the RD part with a mole purity corresponding to 99.7% and a flowrate corresponding to 50 kmol/h. Similarly, the MEOH product is obtained from the bottom of the RC part, with mole purity and flowrate corresponding to 99.5% and 50 kmol/h, respectively. The catalyst occupies 50% of the tray holdup volume. The single tray holdup is calculated by using a weir height of 0.05 m that is assumed to be multiplied by the area of the RD part. The top stage pressure of the LRDWC corresponds to 1 atm, and the pressure drop between stages is set as 0.001 atm. The heat duty of the CON, reboiler of the RD part (REB1), and reboiler of the RC part (REB2) correspond to 5127.15 kW, 4895.82 kW and 386.86 kW respectively. The TAC and TUC of the process correspond to 1.66Γ106 $/a and 5282.68 kW, respectively. 4.
Integrating a vapor recompression heat pump into the LRDWC
As shown in Fig. 2, the overhead vapor of the LRDWC is condensed by utilities in the CON, and thus the latent heat of the overhead vapor is wasted. The use of the VRHP is an efficient way to solve this problem. The VRHP recirculates the latent heat after
compressing the overhead vapor into a higher pressure and temperature, and this saves the energy consumption of the entire process. An LRDWC only possesses one overhead vapor stream, while an URDWC possesses two overhead vapor streams (Feng et al., 2016). Hence, all the overhead vapor of the LRDWC is compressed with only one compressor, and thus the overhead vapor is used more effectively in the integrated LRDWC with the VRHP model, when compared with that of the integrated URDWC with the VRHP configuration. 4.1 Model 1 In this section, the VRHP provides heat in the REB2. Fig. 3 shows the flow diagram of the process. Specifically, IPOH and MEAC are fed into the RD part, at the top and bottom of the reactive section respectively. The overhead vapor of the LRDWC is split into two streams. A stream is compressed at a higher pressure and temperature in the compressor (COMP), and it then gives heat to the boiling-up stream of the RC part in the REB2. After releasing the pressure in the valve, the stream in conjunction with the other overhead vapor stream of the LRDWC, is condensed in the CON. The condensed stream is separated into two streams, a stream is refluxed back to the LRDWC and the other stream is recycled back to the RD part along with the MEAC feed stream. The IPAC and the MEOH are obtained from the bottom of the RD part and the RC part respectively. The top stage pressure of the LRDWC corresponds to 1 atm; and the pressure drop between stages is set as 0.001 atm. The mole purity of the IPAC and the MEOH is set above 0.995. The differential temperature driving force of the REB2 is maintained at 5 Β°C (Fonyo and Mizsey, 1994; Waheed et al., 2014). Efficiency of the compressor is set as 0.72, based on an isentropic efficiency of 0.8 and a mechanical efficiency of 0.9 (Fonyo and Benko, 1998; Jana, 2015). In order to satisfy the separation requirement, heat duty of the REB2 should be maintained at 386.86 kW. Fig. 4 shows the effects of the compression ratio of the COMP on the TAC and the TUC. The TUC increases with increases in the compression ratio.
When the compression ratio is less than 2.4, the TAC decreases with increases in the compression ratio. Conversely, the TAC increases with increases in the compression ratio when the compression ratio exceeds 2.4. Therefore, a minimum TAC is obtained at a compression ratio of 2.4. Fig. 5 shows the PβT curve of the overhead vapor composition. As shown in Fig. 3 and Fig. 5, the temperature and pressure of the overhead vapor stream correspond to 53.6 Β°C and 1 atm, respectively (State 1).The temperature of the overhead stream increases to 97.5 Β°C after it is compressed to 2.4 atm (State 2). The temperature of the overhead stream reduces to 69.9 Β°C after heat is provided in the REB2 (State 3). The bottom recycled stream of the RC part is boiled in the REB2 (State 4). The differential temperature driving force of the REB2 is maintained at 5 Β°C, as shown in Fig. 5. The CI of the process increases due to the integration of the VRHP into the LRDWC. Latent heat of the overhead vapor is slightly recycled, and thus the OC of the process decreases slightly. Therefore, the TAC of the process exceeds that of the LRDWC. Fig. 3 indicates the optimal condition of the process. The compression ratio and the output work of the COMP correspond to 2.4 and 40.58 kW respectively. Heat duty of the CON decreases from 5127.15 kW of the LRDWC to 4780.87 kW, due to the integration of the VRHP into the LRDWC. The TAC and the TUC of the process correspond to 1.73Γ106 $/a and 5017.57 kW, respectively. 4.2 Model 2 In section 4.1, the VRHP is integrated into the LRDWC, and the VRHP provides heat in the REB2. However, the reduction in the TUC is slight. The use of the VRHP to provide additional heat is a potential solution to solve this problem, although heat duty of the REB2 is determined to ensure that the separation requirement is satisfied. Hence, it is necessary to examine other heat exchangers. In this section, the VRHP provides heat in the REB1. Fig. 6 shows the flow diagram of the process. This is different from Model 1 because the compressed stream provides heat to the boiling-up stream of the
RD part in the REB1, and the boiling-up stream of the RC part is boiled by utilities in the REB2. The differential temperature driving force of the REB1 is maintained at 5 Β°C. Fig. 7 indicates the effects of the compression ratio on the TAC and TUC. The TAC and TUC increase with increases in the compression ratio. When the compression ratio is below 4.0, the compressed stream is unable to provide sufficient heat to the boiling-up stream of the RD part. The minimum values of the TAC and TUC are obtained when the compression ratio corresponds to 4.0. Fig. 8 shows the PβT curve of the overhead vapor composition. As shown in Fig. 6 and Fig. 8, the temperature and pressure of the overhead vapor stream correspond to 53.6 Β°C and 1 atm respectively (State 1). The temperature of the overhead stream increases to 123.3 Β°C after it is compressed to 4 atm (State 2). The temperature of the overhead stream reduces to 96 Β°C after the provision of heat in the REB1 (State 3). The bottom recycled stream of the RD part is boiled in the REB1 (State 4). The differential temperature driving force of the REB2 is maintained at 5 Β°C, as shown in Fig. 8. Fig. 6 shows the optimal condition of the process. The compression ratio and the output work of the COMP correspond to 4.0 and 884.10 kW respectively. Heat duty of the CON decreases from 5127.15 kW of the LRDWC to 1115.43 kW, due to the integration of the VRHP into the LRDWC. The TAC and TUC of the process correspond to 1.53Γ106 $/a and 3039.15 kW respectively. 4.3 Model 3 In section 4.2, the VRHP provides heat in the REB1. The integration of VRHP into the LRDWC significantly reduces energy consumption. However, the energy-saving performance of the VRHP can be further developed. The temperature of the compressed stream exceeds 30 Β°C after the provision of heat in the REB1 when compared with that of the bottom recycled stream of the RC part, and thus the compressed stream continues to provide heat in the REB2. Fig. 9 indicates the flow diagram of the process. This is different from Model 2 since the provision of heat to the boiling-up stream of the RD
part in the REB1 is followed by provision of heat from the compressed stream to the boiling-up stream of the RC part in the REB2, and the pressure in the valve is subsequently released. In the process, the differential temperature driving force of REB1 is maintained at 5 Β°C. Fig. 10 shows the effects of the compression ratio on the TAC and TUC. In a manner similar to Fig. 7, the TAC and TUC increase when the compression ratio increases; and the TAC and TUC reach minimum values when the compression ratio corresponds to 4.0. Fig. 9 shows the optimum condition of the process. The compression ratio and the output work of the COMP correspond to 4.0 and 884.10 kW respectively. Heat duty of the CON decreases from 1115.43 kW in Model 2 to 728.57 kW. The TAC and TUC of the process correspond to 1.44Γ106 $/a and 2652.29 kW, respectively. 4.4 Model 4 In section 4.3, Model 3 exhibits a better energy-saving performance when compared with that of Model 2. The temperature of the compressed stream exceeds 20 Β°C when compared with that of the overhead vapor of LRDWC after the provision of heat in the REB1 and REB2. Prior to the release of pressure in the valve, the compressed stream provides heat to the inlet stream of the COMP in a preheater (PHE) and the output stream temperature of the COMP increases, and thus the compressed stream in a certain flow rate provides increased heat in the REB1. The heat duty of the REB1 should not be higher, and therefore the flow rate of the compressed stream is reduced. This lowers the output work of the COMP. Fig. 11 shows the flow diagram of the process. In a manner similar to Model 3, the differential temperature driving force of the REB1 is maintained at 5 Β°C. Fig. 12 indicates the effects of the heat duty of the PHE and the compression ratio of the COMP on the TAC and TUC. The TAC and TUC increase when the compression ratio of the COMP increases in a specific heat duty of the PHE. The TUC of the process
reduces when the heat duty of the PHE increases. The TAC reaches a minimum when the heat duty of the PHE corresponds to 150 kW. However, the compressed stream only provides 200 kW of heat in the PHE after providing heat in the REB1 and REB2. The process displays an optimal economic performance when the compression ratio corresponds to 4.0 and the heat duty of the PHE corresponds to 150 kW. However, the process exhibits the smallest TUC when the compression ratio corresponds to 4.0 and the heat duty of the PHE corresponds to 200 kW although the TAC of the process slightly exceeds that of the process when the heat duty of the PHE corresponds to 150 kW. In consideration of the energy-saving performance, the process wherein the compression ratio corresponds to 4.0 and the heat duty of the PHE corresponds to 200 kW is an optimal choice. Fig. 11 indicates this condition. The output work of the COMP and the heat duty of the CON correspond to 882.90 kW and 727.37 kW, respectively. The TAC and TUC of the process correspond to 1.43Γ106 $/a and 2648.69 kW, respectively. 5.
Comparisons
Table 3 shows the energy consumption and CO2 emissions of all models in use., Model 1 and Model 2 reduce 5.02% and 42.47% of the TUC, respectively, when compared with those of the LRDWC. The integration of the VRHP into the LRDWC significantly reduces the TUC of the entire process. An increase in the duty of the heat exchangers reduces the TUC of the entire process. Model 3 and Model 4 reduce 49.79% and 49.86% of the TUC, respectively, when compared with those of the LRDWC. The effective use of the remaining energy reduces the TUC. However, Model 4 displays minor improvements in the TUC when compared with Model 3 due to the limitation of the preheater duty. The integrated LRDWCs with VRHP models exhibit a high environmental performance. The CO2 emissions of Model 1 and Model 2 reduce by 7.32% and 92.68%, respectively, when compared with those of the LRDWC, and Model 3 and Model 4 show no CO2 emissions.
Table 4 shows the costs of all the models. The CI of Model 1 increases by 27.58%, and the OC of Model 1 decreases by 2.29% when compared with those of the LRDWC. In a payback period of 5 years, the TAC of Model 1 increases by 4.39% when compared with those of the LRDWC. The CI of Model 2 increases by 89.72% when compared with those of the LRDWC although the OC of Model 2 decreases by 36.21% when compared with those of the LRDWC. The OC occupies a large part of the TAC in the LRDWC, and thus the TAC of Model 2 reduces by 8.04% when compared with those of the LRDWC for a payback period of 5 years. The CI of Model 3 and Model 4 increase by 90.79% and 89.89%, respectively, when compared with those of the LRDWC. The addition of the heat exchangers increases the CI of the process, and the application of the PHE reduces the costs of heat exchangers and the compressor. Model 3 and Model 4 reduce by 43.55% and 43.62%, respectively, of the OC when compared with those of the LRDWC. The application of the PHE also reduces the cost of cooling water and electricity. The TAC of Model 3 and Model 4 reduce by 13.50% and 13.76%, respectively, when compared with those of the LRDWC for a payback period of 5 years. It is concluded that the integration of a VRHP into an LRDWC increases the CI of the process although it reduces the OC of the process, and thus the TAC of the process reduces in a payback period of 5 years. Furthermore, the reduction of the TAC is better for a longer payback period, as shown in Table 4. 6.
Conclusion
In this study, four integrated LRDWCs with VRHP models are introduced to the transesterification reaction of MEAC and IPOH. The energy-saving, economic and environmental performances are compared before and after the integration of the VRHP into the LRDWC. The integration significantly improves the energy-saving performance of the process; thus, the OC of the process is reduced. Although the CI of the process increases due to the integration of the VRHP into the LRDWC, the TAC of the process reduces under a longer payback period. The integration of VRHP into the LRDWC results in a significantly high environmental performance. The preheater
further improves energy-saving and economic performance of the process. Model 4 shows the best performance when compared with those of all the models introduced in the study. It saves 49.86% of the TUC, 13.76% of the TAC for a payback period of 5 years, and all the CO2 emissions when compared with the LRDWC. The integrated LRDWC with the VRHP model is attractive in terms of industrial applications due to the significant reduction in energy consumption and CO2 emissions.
Acknowledgements We are thankful for the assistance from the staff at the Jiangsu Key Laboratory of Advanced Catalytic Materials and Technology from School of Petrochemical Engineering, Changzhou University.
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Intensif. doi:10.1016/j.cep.2016.09.014
(a)
(b)
(c)
(d)
Fig. 1 Two different structures of RD and the corresponding RDWC configurations: (a) Reactive distillation in a direct sequence; (b) Reactive distillation in an indirect sequence; (c) Upper partitioned reactive dividing-wall column (URDWC); and (d) Lower partitioned reactive dividing-wall column (LRDWC).
CON 5127.15kW
1
REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
MEAC 50kmol/h 60
10
IPOH 50kmol/h 60
30
RD RC 72
84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
REB1 4895.82kW
18 REB2 386.86kW
Fig. 2 Process flow diagram of the LRDWC.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
CON 4780.87kW
53.6 578.81kmol/h
1 42.17 kmol/h
REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
MEAC 50kmol/h 60
10
IPOH 50kmol/h 60
COMP 40.58kW Rc=2.4
30
RD RC 97.5
72
Valve
69.9 84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
REB1 4895.82kW
18 REB2 386.86kW
Fig. 3 Process flow diagram of Model 1.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
1.742
5.03
5.02 1.740 1.739
5.01
1.738 5.00 1.737 1.736
4.99
TUC (MW)
TAC of 5 years (106$/year)
1.741
1.735 4.98 1.734 1.733
4.97 1.8
2.0
2.2
2.4
2.6
Compression ratio
Fig. 4 Effects of the compression ratio on the TAC and the TUC in Model 1.
5 1.0002
4
Pressure (atm)
1.0000
1
0.9998 53.635
3
53.640
53.645
3
2
2
1 50
VF=1 VF=0
4
1 60
70
80
90
100
Temperature (Β°C)
Fig. 5 PβT curve of the overhead vapor composition in Model 1.
CON 1115.43kW
53.6 578.81kmol/h
1 564.62 kmol/h REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
MEAC 50kmol/h 60
10 IPOH 50kmol/h 60
COMP 884.10kW Rc=4.0
30
RD RC 123.3
72
Valve
96
84
18
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
REB1 4895.82kW
REB2 386.86kW
Fig. 6 Process flow diagram of Model 2.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
1.59 3.24 3.22 3.20
1.57
3.18 1.56
3.16
1.55
3.14 3.12
1.54
3.10
TUC (MW)
TAC of 5 years (106$/year)
1.58
3.08
1.53
3.06 1.52
3.04
1.51
3.02 4.0
4.1
4.2
4.3
4.4
Compression ratio
Fig. 7 Effects of the compression ratio on the TAC and TUC in Model 2.
5
VF=1 VF=0
4.1
3 4.0
Pressure (atm)
4
3.9 95
2
3
96
97
98
3 1.0002
1.0000
2
0.9998 53.635
1
53.640
53.645
4
1 50
1
60
70
80
90
100
110
120
Temperature (Β°C)
Fig. 8 PβT curve of the overhead vapor composition in Model 2.
CON 728.57kW
53.6 578.81kmol/h
1 564.62 kmol/h
REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
10 IPOH 50kmol/h 60
RD RC
MEAC 50kmol/h 60
123.3
72
84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
COMP 884.10kW Rc=4.0
30
96
REB1 4895.82kW
18
REB2 386.86kW
Fig. 9 Process flow diagram of Model 3.
Valve
80.1
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
2.84 2.82 2.80
1.48
2.78 1.47 2.76 2.74
1.46
2.72 1.45 2.70 1.44
TUC (MW)
TAC of 5 years (106$/year)
1.49
2.68 2.66
1.43
2.64 4.0
4.1
4.2
4.3
4.4
Compression ratio
Fig. 10 Effects of the compression ratio on the TAC and TUC in Model 3.
CON 727.37kW
53.6 578.81kmol/h 536.64 kmol/h
1
Valve
70.3 REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
10 IPOH 50kmol/h 60
PHE 200kW
30 71.1
RD RC
MEAC 50kmol/h 60
79.3 72
COMP 882.90kW Rc=4.0 141.9
84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
96
REB1 4895.82kW
18
REB2 386.86kW
Fig. 11 Process flow diagram of Model 4.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
(a)
(b) Rc=4.4
1.48
Rc=4.3
2.82 2.80
Rc=4.4
2.78
Rc=4.3
2.76 1.47
TUC (MW)
TAC of 5 years (106$/year)
1.49
Rc=4.2 1.46
Rc=4.1
1.45
Rc=4.2
2.74 2.72
Rc=4.1
2.70 2.68
1.44 2.66
Rc=4.0
Rc=4.0
2.64
1.43 0
50
100
150
Heat duty of the Preheater (kW)
200
0
50
100
150
200
Heat duty of the Preheater (kW)
Fig. 12 Effects of the heat duty of the PHE and the compression ratio on (a) the TAC and (b) the TUC in Model 4.
Table 1 Binary parameters of the UNIQUAC model Component i
MEAC
IPOH
IPAC
MEAC
MEAC
IPOH
Component j
MEOH
MEOH
MEOH
IPOH
IPAC
IPAC
bij
-337.419
-169.757
-401.981
-66.4607
1.96175
67.3403
bji
54.3901
70.9141
68.1792
-40.9002
-20.9966
-163.536
Table 2 Cost estimating formula of the CI Column vessel CI = 17640π·1.066 πΏ0.802, where D denotes column diameter (m), and L denotes column height (m). πΏ = (ππ β 1) Γ 0.61 + 6, where NT denotes actual stage number. Heat exchanger CI = 7296π΄0.65, where A denotes heat transfer area (m2). π΄=
π πβπ
, where Q denotes heat duty of the heat exchanger (kW), and U denotes the heat transfer
coefficient. βT is calculated by using a logarithmic mean temperature difference (LMTD) method. Compressor CI =
π&π 280
Γ 1264.75(βπ)0.82, where hp denotes work of the compressor.
πβπ = βπ/0.8, where bhp denotes brake horsepower of the compressor.
Table 3 Comparisons of energy consumptions and CO2 emissions Parameter Reboiler duty (kW) Condenser duty (kW)
LRDWC
Model 1
Model 2
Model 3
Model 4
4895.82
4895.82
386.86
4780.87
1115.43
728.57
727.37
386.86
4895.82
4895.82
4895.82
386.86
386.86
386.86 5127.15
Heat Exchanger duty (kW) Preheater duty (kW)
200
Compression ratio
2.4
4
4
4
Compressor work (kW)
40.58
884.10
884.10
882.90
5017.57
3039.15
2652.29
2648.69
(-5.02%)
(-42.47%)
(-49.79%)
(-49.86%)
2499.14
197.48
0
0
(-7.32%)
(-92.68%)
TUC (kW)
5282.68
[Savings (%)] CO2 emissions (kg/h) [Savings (%)]
2696.62
Table 4 Comparisons of costs Parameter 3
Column cost (10 $) 3
Reboiler cost (10 $) 3
Condenser cost (10 $)
LRDWC
Model 1
Model 2
Model 3
Model 4
1427.20
1427.20
1427.20
1427.20
1427.20
194.48
168.23
26.25
236.09
225.60
87.60
66.41
66.34
459.02
460.92
528.31
513.23
121.70
1522.56
1522.56
1520.87
2.37
3.52
3.54
3.53
(+27.58%)
(+89.72%)
(+90.79%)
(+89.89%)
1183.66
1096.98
86.68
105.90
98.75
23.04
15.05
15.02
32.73
712.94
712.94
711.97
1.26
0.82
0.73
0.73
(-2.29%)
(-36.21%)
(-43.55%)
(-43.62%)
1.73
1.53
1.44
1.43
(+4.39%)
(-8.04%)
(-13.50%)
(-13.76%)
1.56
1.26
1.17
1.17
(+2.32%)
(-1.70%)
(-23.05%)
(-23.25%)
Heat exchanger (including preheater) cost (103 $) Compressor cost (103 $) 6
CI (10 $)
1.86
[Increase (%)] Steam cost
(103
$/a)
Cooling water cost
(103
$/a) Electricity cost (103 $/a) OC
(106
$/a)
1.29
[Savings (%)] TAC with the payback period of 5 years
(106
1.66
$/a)
[Increase / Savings (%)] TAC with the payback period of 8 years (106 $/a) [Increase / Savings (%)]
1.52
S0263-8762(17)30384-2 http://dx.doi.org/doi:10.1016/j.cherd.2017.07.017 CHERD 2758
To appear in: Received date: Revised date: Accepted date:
3-1-2017 11-7-2017 12-7-2017
Please cite this article as: Feng, Shenyao, Ye, Qing, Xia, Hui, Li, Rui, Suo, Xiaomeng, Integrating a vapor recompression heat pump into a lower partitioned reactive dividingwall column for better energy-saving performance.Chemical Engineering Research and Design http://dx.doi.org/10.1016/j.cherd.2017.07.017 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Integrating a vapor recompression heat pump into a lower partitioned reactive dividing-wall column for better energy-saving performance
Shenyao Feng, Qing Ye*, Hui Xia, Rui Li, and Xiaomeng Suo
Jiangsu Key Laboratory of Advanced Catalytic Materials and Technology, School of Petrochemical Engineering, Changzhou University, Changzhou, Jiangsu 213164, China *
Corresponding author. Tel.: +86 519 86330355. Fax: +86 519 86330355. E-mail:
[email protected].
Graphical abstract
Highlights: ο·
Integrated lower partitioned reactive dividing-wall column with vapor recompression heat pump models are proposed.
ο·
A remarkable energy-saving performance improvement is achieved by the integrated model.
ο·
CO2 emissions of the process are totally reduced by the integrated model.
Abstract: In this study, four integrated lower partitioned reactive dividing-wall columns (LRDWCs) with vapor recompression heat pump (VRHP) models are developed to reduce the energy consumption of a reactive dividing-wall column. Energy-saving, economic, and environmental performances of the integrated LRDWC with VRHP models are compared with the corresponding LRDWC. The results indicate that the integration of VRHP into the LRDWC significantly reduces the energy consumption and CO2 emissions associated with the process. The application of a preheater further improves the energy-saving and economic performance of the process. With the aid of a preheater, the integrated LRDWC with the VRHP model saves 49.86% energy consumption and 13.76% total annual cost with a payback period corresponding to five years, and reduces all CO2 emissions when compared with the corresponding LRDWC.
Keywords: Vapor recompression heat pump; Reactive dividing-wall column; Energy saving Nomenclature CI
capital investment
COMP
compressor
CON
condenser of the lower partitioned reactive dividing-wall column
DWC
dividing-wall column
IPAC
isopropyl acetate
IPOH
isopropanol
LHHW
Langmuir Hinshelwood Hougen Watson model
LMTD
logarithmic mean temperature difference
LRDWC
lower partitioned reactive dividing-wall column
MEAC
methyl acetate
MEOH
methanol
OC
operating cost
PHE
preheater
RC
recovery
RD
reactive distillation
RDWC
reactive dividing-wall column
REB1
reboiler of the reactive distillation part
REB2
reboiler of the recovery part
TAC
total annual cost
TUC
total utility consumption
UNIQUAC
universal quasi-chemical activity coefficient
URDWC
upper partitioned reactive dividing-wall column
VRHP
vapor recompression heat pump
1.
Introduction
Distillation is the most widely used separation process for multi-component mixtures in chemical industries although it is considerably energy intensive (Jana, 2010; You et al., 2016). Energy consumptions of the distillation process are of great concern due to increasing energy demands (Jana, 2016a, 2016b, 2016c; Liu et al., 2015; Yasuhiko et al., 2012).
A dividing-wall column (DWC) is an attractive design that saves equipment and operating costs of the process and reduces energy consumption when compared with traditional separation sequences (Errico et al., 2009; Novita et al., 2015). Reactive distillation (RD) combines reaction and separation steps in a single unit. The application of RD increases conversion and selectivity, and reduces energy consumption when compared with those of conventional reactor and distillation sequences (Kiss et al., 2007). The reactive dividing-wall column (RDWC) integrates the RD and DWC in a column; thus, it possesses the advantages of both the RD and DWC (Mueller and Kenig, 2007). The upper partitioned reactive dividing-wall column (URDWC, Fig. 1c) and the lower partitioned reactive dividing-wall column (LRDWC, Fig. 1d) correspond to two popular configurations of the RDWC. In general, an URDWC is used in the place of reactive distillation in a direct sequence (Fig. 1a), and an LRDWC is considered as an alternative to reactive distillation in an indirect sequence (Fig. 1b). The selection of an RDWC configuration is based on physical and chemical parameters and the composition of reactants and products. Dai et al. (2015) studied an URDWC for the synthesis of n-propyl propionate, and the results indicated that the URDWC saved 12.4% and 16.4% energy consumption and total annual cost, respectively, when compared with those corresponding to the conventional two-column process. Zheng et al. (2016) developed an LRDWC to synthesize diethyl carbonate. The results showed that the LRDWC saved 18.7% energy consumption and 13.9% total annual cost when compared with those of the conventional RD process. Kiss and Suszwalak (2012a, 2012b) introduced a novel dimethyl ether process based on methanol dehydration in a RDWC. The results indicated that the process achieved up to 58% energy savings, 60% reduced CO2 emissions, and 30% lower capital investment costs when compared with those of the traditional RD process. The findings of the aforementioned studies indicate that a RDWC significantly reduces energy consumption and investment when compared with those of the corresponding RD process.
A vapor recompression heat pump (VRHP) reduces the energy demand of separating close-boiling mixtures (Waheed et al., 2014). A compressor improves low-grade heat of the overhead vapor into a higher-grade heat; thus, compressed vapor provides heat in a heat exchanger. A study by Danziger (1979) suggested that a VRHP can save energy consumption by 80%. Recently, few studies reported on integrating a VRHP into a DWC. Li et al. (2016) integrated a VRHP into an azeotropic DWC system. The total annual cost and CO2 emissions were reduced by 32.22% and 63.79%, respectively, when compared with those of the corresponding two-column azeotropic system. Luyben (2016) developed a three-temperature control configuration with a pressure compensation in a heterogeneous azeotropic dividing-wall system with vapor recompression. The controls solved disturbances in the feed composition and feed rate to a considerable extent. These studies demonstrated that the integration of a VRHP into a DWC reduces the energy consumption of the whole process. The integration of a VRHP into a RDWC can exhibit the same effect. Liu et al. (2015) proposed a heat pump assisted reactive dividing-wall column that reduces 11.5% total energy consumption, 4.3% total annual cost with a capital payback period of 5 years, and 13.4% CO2 emissions when compared with those of the corresponding RDWC. However, studies that integrate a VRHP into a RDWC are still at a nascent stage. There is a lack of studies focusing on the reduction of energy consumption in a RDWC by using a VRHP. An earlier study (Feng et al., 2016) integrated a VRHP into a URDWC model. The integration of a VRHP into a URDWC significantly reduced energy consumption of the complete process, since the results indicated that the URDWC integrated with the VRHP model saved up to 39.22% total energy consumption when compared with that of the corresponding URDWC model. More specifically, URDWC and LRDWC correspond to two different configurations of the RDWC; thus, an integrated LRDWC with the VRHP model should exhibit same effect. However, studies continue to explore the actual energy-saving performance of the integrated LRDWC with the VRHP models.
This study is focused on improving the energy-saving performance of a LRDWC process by integrating a VRHP into a LRDWC model. The transesterification reaction of methyl acetate (MEAC) and isopropanol (IPOH) is selected as the reaction of the LRDWC. In this study, four integrated LRDWCs with the VRHP models are proposed. The VRHP provides heat in the bottom reboilers, and the influence of the preheater on the integrated LRDWC with the VRHP model is investigated. The energy-saving, economic and environmental performances of all models are examined. 2.
Design basis and methodology
2.1 Thermodynamic and kinetic models Isopropyl acetate (IPAC) and methanol (MEOH) are synthesized by a transesterification reaction of MEAC and IPOH. The chemical equation is shown as follows: ππΈπ΄πΆ + πΌπππ» β πΌππ΄πΆ + ππΈππ»
(1)
This reaction is catalyzed by a strongly acidic ion-exchange resin Lewatit MonoPlus S100 H+ (Akyalçın, 2014). The equilibrium constant (Ka) and the rate constant (Kf) of the reaction by using the Langmuir Hinshelwood Hougen Watson (LHHW) model (Akyalçın, 2014) are given as follows:
πΎπ = expβ‘(
702 β 1.81) π
πΎπ (πππ/π β min) = expβ‘(9.03 β
(2) 4999 ) π
(3)
where T denotes absolute temperature in Kelvin. Simulations are performed by using Aspen Plus 8.4 software. The universal quasi-chemical activity coefficient (UNIQUAC) model is selected to depict the vapor-liquid equilibrium of the simulations. The UNIQUAC model parameters of MEOH +IPAC (Resa et al., 2001), MEAC + MEOH (Tu et al., 1997), IPAC + IPOH (Andreatta et al., 2010) and IPAC + MEOH (Resa et al.,
2001) are regressed by using experimental data. As a result, the UNIQUAC model is in agreement with the experimental data. The UNIQUAC model parameters of the system (Suo et al., 2016) are listed in Table 1. 2.2 Total utility consumption (TUC) For the convenience of comparing the energy-saving performance of the processes, the total utility consumption (TUC) is defined as follows: TUC = βππ πΈπ΅ + βπππ πΈ + πβππΆπππ
(4)
where QREB, QPRE, and QCOMP denote the heat duty of the reboiler, heat duty of the preheater and work of the compressor respectively. In addition, f (=3) represents the multiplication factor of the work of the compressor that converts electrical work into thermal power with the same effect (Kumar et al., 2013). 2.3 Total annual cost (TAC) It is necessary to simultaneously consider the capital cost and the operating cost while evaluating the economic performance of a process. Thus, the economic potential of the processes is assessed in terms of total annual cost (TAC)and expressed as follows:
TAC = operatingβ‘cost +
capitalβ‘investment paybackβ‘period
(5)
The capital investment (CI) is calculated by summing up all the investments of the equipment, including column vessels, heat exchangers and compressors. The cost estimating formula (Douglas, 1988; Luyben, 2013) is shown in Table 2. The value of the Marshall and Swift (M&S) index is set as 1293 (Luyben, 2011), and the motor efficiency of the compressor corresponds to 60% (Douglas, 1988). The low-pressure steam produced by the steam boiler is used in reboilers, cooling water at a temperature of 25Β°C is used in the condensers, and electricity produced by the gas turbine is used in the compressors. Heat transfer coefficients of the condensers and the reboilers
correspond to 0.852 and 0.568 kW/(KΒ·m2), respectively (Luyben, 2013). The smaller equipment investment and catalyst costs are neglected, since the catalyst is repeatedly utilized after reproduction. The operating cost (OC) is calculated based on the utility cost for a year with 8000 operating hours. The price of low-pressure steam, cooling water and electricity correspond to 7.78 $/GJ, 0.03 $/kg and 16.8 $/GJ respectively, based on recommendations specified by Luyben (2013, 2011). 2.4 CO2 emission Specifically, CO2 is generated by burning a fuel and air mixture based on a stoichiometric expression (Gadalla et al., 2005) as follows: π¦ π¦ πΆπ₯ π»π¦ + (π₯ + ) π2 β π₯πΆπ2 + π»2 π 4 2
(6)
where x and y represent the number of carbon C and hydrogen H atoms respectively. The CO2 emission ([πΆπ2 ]πΈπππ π ) is described as follows: [πΆπ2 ]πΈπππ π = (
ππΉπ’ππ C% )( )πΌ ππ»π 100
(7)
where Ξ± (=3.67) denotes molar mass proportion of CO2 and C, and NHV (=22000 kJ/kg) represents the net thermal value of fuels. Carbon content (C%) of the coal is set as 0.68 (Seader et al., 2010). The flue gas produced by combusting fuel is used as a heat source in the steam boiler and gas turbine (Jana, 2015). The quantity of the combusted fuel (ππΉπ’ππ ) is calculated as follows:
ππΉπ’ππ =
πππππ ππΉππ΅ β ππ (βππππ β 419) πππππ ππΉππ΅ β ππ π‘πππ
(8)
where πππππ (kJ/kg) and βππππ (kJ/kg) denote latent heat and the enthalpy of steam respectively; and ππ (Β°C), ππΉππ΅ (=1800Β°C), and ππ π‘πππ (=160Β°C) denote the ambient
temperature, flame temperature, and stack temperature respectively (Jantes-Jaramillo et al., 2008). 3.
Lower partitioned reactive dividing-wall column (LRDWC)
The process flow diagram of a LRDWC is shown in Fig. 2. The LRDWC is divided into three parts, namely the top part, reactive distillation (RD) part and recovery (RC) part. The top part, RD part and RC part include 10, 74 and 8 stages, respectively. The RD part is divided into three sections: the rectifying section, reactive section, and stripping section. The esterification reaction of MEAC and IPOH occurs in the reactive section. In addition, IPOH and MEAC are fed into the RD part at the top and bottom of the reactive section respectively with a flowrate of 50 kmol/h. The top stream is condensed in the condenser of the LRDWC (CON) and is then separated into two streams. A stream is refluxed back to the LRDWC, and the other stream is recycled back to the RD part, with the MEAC feed stream. The IPAC product is obtained from the bottom of the RD part with a mole purity corresponding to 99.7% and a flowrate corresponding to 50 kmol/h. Similarly, the MEOH product is obtained from the bottom of the RC part, with mole purity and flowrate corresponding to 99.5% and 50 kmol/h, respectively. The catalyst occupies 50% of the tray holdup volume. The single tray holdup is calculated by using a weir height of 0.05 m that is assumed to be multiplied by the area of the RD part. The top stage pressure of the LRDWC corresponds to 1 atm, and the pressure drop between stages is set as 0.001 atm. The heat duty of the CON, reboiler of the RD part (REB1), and reboiler of the RC part (REB2) correspond to 5127.15 kW, 4895.82 kW and 386.86 kW respectively. The TAC and TUC of the process correspond to 1.66Γ106 $/a and 5282.68 kW, respectively. 4.
Integrating a vapor recompression heat pump into the LRDWC
As shown in Fig. 2, the overhead vapor of the LRDWC is condensed by utilities in the CON, and thus the latent heat of the overhead vapor is wasted. The use of the VRHP is an efficient way to solve this problem. The VRHP recirculates the latent heat after
compressing the overhead vapor into a higher pressure and temperature, and this saves the energy consumption of the entire process. An LRDWC only possesses one overhead vapor stream, while an URDWC possesses two overhead vapor streams (Feng et al., 2016). Hence, all the overhead vapor of the LRDWC is compressed with only one compressor, and thus the overhead vapor is used more effectively in the integrated LRDWC with the VRHP model, when compared with that of the integrated URDWC with the VRHP configuration. 4.1 Model 1 In this section, the VRHP provides heat in the REB2. Fig. 3 shows the flow diagram of the process. Specifically, IPOH and MEAC are fed into the RD part, at the top and bottom of the reactive section respectively. The overhead vapor of the LRDWC is split into two streams. A stream is compressed at a higher pressure and temperature in the compressor (COMP), and it then gives heat to the boiling-up stream of the RC part in the REB2. After releasing the pressure in the valve, the stream in conjunction with the other overhead vapor stream of the LRDWC, is condensed in the CON. The condensed stream is separated into two streams, a stream is refluxed back to the LRDWC and the other stream is recycled back to the RD part along with the MEAC feed stream. The IPAC and the MEOH are obtained from the bottom of the RD part and the RC part respectively. The top stage pressure of the LRDWC corresponds to 1 atm; and the pressure drop between stages is set as 0.001 atm. The mole purity of the IPAC and the MEOH is set above 0.995. The differential temperature driving force of the REB2 is maintained at 5 Β°C (Fonyo and Mizsey, 1994; Waheed et al., 2014). Efficiency of the compressor is set as 0.72, based on an isentropic efficiency of 0.8 and a mechanical efficiency of 0.9 (Fonyo and Benko, 1998; Jana, 2015). In order to satisfy the separation requirement, heat duty of the REB2 should be maintained at 386.86 kW. Fig. 4 shows the effects of the compression ratio of the COMP on the TAC and the TUC. The TUC increases with increases in the compression ratio.
When the compression ratio is less than 2.4, the TAC decreases with increases in the compression ratio. Conversely, the TAC increases with increases in the compression ratio when the compression ratio exceeds 2.4. Therefore, a minimum TAC is obtained at a compression ratio of 2.4. Fig. 5 shows the PβT curve of the overhead vapor composition. As shown in Fig. 3 and Fig. 5, the temperature and pressure of the overhead vapor stream correspond to 53.6 Β°C and 1 atm, respectively (State 1).The temperature of the overhead stream increases to 97.5 Β°C after it is compressed to 2.4 atm (State 2). The temperature of the overhead stream reduces to 69.9 Β°C after heat is provided in the REB2 (State 3). The bottom recycled stream of the RC part is boiled in the REB2 (State 4). The differential temperature driving force of the REB2 is maintained at 5 Β°C, as shown in Fig. 5. The CI of the process increases due to the integration of the VRHP into the LRDWC. Latent heat of the overhead vapor is slightly recycled, and thus the OC of the process decreases slightly. Therefore, the TAC of the process exceeds that of the LRDWC. Fig. 3 indicates the optimal condition of the process. The compression ratio and the output work of the COMP correspond to 2.4 and 40.58 kW respectively. Heat duty of the CON decreases from 5127.15 kW of the LRDWC to 4780.87 kW, due to the integration of the VRHP into the LRDWC. The TAC and the TUC of the process correspond to 1.73Γ106 $/a and 5017.57 kW, respectively. 4.2 Model 2 In section 4.1, the VRHP is integrated into the LRDWC, and the VRHP provides heat in the REB2. However, the reduction in the TUC is slight. The use of the VRHP to provide additional heat is a potential solution to solve this problem, although heat duty of the REB2 is determined to ensure that the separation requirement is satisfied. Hence, it is necessary to examine other heat exchangers. In this section, the VRHP provides heat in the REB1. Fig. 6 shows the flow diagram of the process. This is different from Model 1 because the compressed stream provides heat to the boiling-up stream of the
RD part in the REB1, and the boiling-up stream of the RC part is boiled by utilities in the REB2. The differential temperature driving force of the REB1 is maintained at 5 Β°C. Fig. 7 indicates the effects of the compression ratio on the TAC and TUC. The TAC and TUC increase with increases in the compression ratio. When the compression ratio is below 4.0, the compressed stream is unable to provide sufficient heat to the boiling-up stream of the RD part. The minimum values of the TAC and TUC are obtained when the compression ratio corresponds to 4.0. Fig. 8 shows the PβT curve of the overhead vapor composition. As shown in Fig. 6 and Fig. 8, the temperature and pressure of the overhead vapor stream correspond to 53.6 Β°C and 1 atm respectively (State 1). The temperature of the overhead stream increases to 123.3 Β°C after it is compressed to 4 atm (State 2). The temperature of the overhead stream reduces to 96 Β°C after the provision of heat in the REB1 (State 3). The bottom recycled stream of the RD part is boiled in the REB1 (State 4). The differential temperature driving force of the REB2 is maintained at 5 Β°C, as shown in Fig. 8. Fig. 6 shows the optimal condition of the process. The compression ratio and the output work of the COMP correspond to 4.0 and 884.10 kW respectively. Heat duty of the CON decreases from 5127.15 kW of the LRDWC to 1115.43 kW, due to the integration of the VRHP into the LRDWC. The TAC and TUC of the process correspond to 1.53Γ106 $/a and 3039.15 kW respectively. 4.3 Model 3 In section 4.2, the VRHP provides heat in the REB1. The integration of VRHP into the LRDWC significantly reduces energy consumption. However, the energy-saving performance of the VRHP can be further developed. The temperature of the compressed stream exceeds 30 Β°C after the provision of heat in the REB1 when compared with that of the bottom recycled stream of the RC part, and thus the compressed stream continues to provide heat in the REB2. Fig. 9 indicates the flow diagram of the process. This is different from Model 2 since the provision of heat to the boiling-up stream of the RD
part in the REB1 is followed by provision of heat from the compressed stream to the boiling-up stream of the RC part in the REB2, and the pressure in the valve is subsequently released. In the process, the differential temperature driving force of REB1 is maintained at 5 Β°C. Fig. 10 shows the effects of the compression ratio on the TAC and TUC. In a manner similar to Fig. 7, the TAC and TUC increase when the compression ratio increases; and the TAC and TUC reach minimum values when the compression ratio corresponds to 4.0. Fig. 9 shows the optimum condition of the process. The compression ratio and the output work of the COMP correspond to 4.0 and 884.10 kW respectively. Heat duty of the CON decreases from 1115.43 kW in Model 2 to 728.57 kW. The TAC and TUC of the process correspond to 1.44Γ106 $/a and 2652.29 kW, respectively. 4.4 Model 4 In section 4.3, Model 3 exhibits a better energy-saving performance when compared with that of Model 2. The temperature of the compressed stream exceeds 20 Β°C when compared with that of the overhead vapor of LRDWC after the provision of heat in the REB1 and REB2. Prior to the release of pressure in the valve, the compressed stream provides heat to the inlet stream of the COMP in a preheater (PHE) and the output stream temperature of the COMP increases, and thus the compressed stream in a certain flow rate provides increased heat in the REB1. The heat duty of the REB1 should not be higher, and therefore the flow rate of the compressed stream is reduced. This lowers the output work of the COMP. Fig. 11 shows the flow diagram of the process. In a manner similar to Model 3, the differential temperature driving force of the REB1 is maintained at 5 Β°C. Fig. 12 indicates the effects of the heat duty of the PHE and the compression ratio of the COMP on the TAC and TUC. The TAC and TUC increase when the compression ratio of the COMP increases in a specific heat duty of the PHE. The TUC of the process
reduces when the heat duty of the PHE increases. The TAC reaches a minimum when the heat duty of the PHE corresponds to 150 kW. However, the compressed stream only provides 200 kW of heat in the PHE after providing heat in the REB1 and REB2. The process displays an optimal economic performance when the compression ratio corresponds to 4.0 and the heat duty of the PHE corresponds to 150 kW. However, the process exhibits the smallest TUC when the compression ratio corresponds to 4.0 and the heat duty of the PHE corresponds to 200 kW although the TAC of the process slightly exceeds that of the process when the heat duty of the PHE corresponds to 150 kW. In consideration of the energy-saving performance, the process wherein the compression ratio corresponds to 4.0 and the heat duty of the PHE corresponds to 200 kW is an optimal choice. Fig. 11 indicates this condition. The output work of the COMP and the heat duty of the CON correspond to 882.90 kW and 727.37 kW, respectively. The TAC and TUC of the process correspond to 1.43Γ106 $/a and 2648.69 kW, respectively. 5.
Comparisons
Table 3 shows the energy consumption and CO2 emissions of all models in use., Model 1 and Model 2 reduce 5.02% and 42.47% of the TUC, respectively, when compared with those of the LRDWC. The integration of the VRHP into the LRDWC significantly reduces the TUC of the entire process. An increase in the duty of the heat exchangers reduces the TUC of the entire process. Model 3 and Model 4 reduce 49.79% and 49.86% of the TUC, respectively, when compared with those of the LRDWC. The effective use of the remaining energy reduces the TUC. However, Model 4 displays minor improvements in the TUC when compared with Model 3 due to the limitation of the preheater duty. The integrated LRDWCs with VRHP models exhibit a high environmental performance. The CO2 emissions of Model 1 and Model 2 reduce by 7.32% and 92.68%, respectively, when compared with those of the LRDWC, and Model 3 and Model 4 show no CO2 emissions.
Table 4 shows the costs of all the models. The CI of Model 1 increases by 27.58%, and the OC of Model 1 decreases by 2.29% when compared with those of the LRDWC. In a payback period of 5 years, the TAC of Model 1 increases by 4.39% when compared with those of the LRDWC. The CI of Model 2 increases by 89.72% when compared with those of the LRDWC although the OC of Model 2 decreases by 36.21% when compared with those of the LRDWC. The OC occupies a large part of the TAC in the LRDWC, and thus the TAC of Model 2 reduces by 8.04% when compared with those of the LRDWC for a payback period of 5 years. The CI of Model 3 and Model 4 increase by 90.79% and 89.89%, respectively, when compared with those of the LRDWC. The addition of the heat exchangers increases the CI of the process, and the application of the PHE reduces the costs of heat exchangers and the compressor. Model 3 and Model 4 reduce by 43.55% and 43.62%, respectively, of the OC when compared with those of the LRDWC. The application of the PHE also reduces the cost of cooling water and electricity. The TAC of Model 3 and Model 4 reduce by 13.50% and 13.76%, respectively, when compared with those of the LRDWC for a payback period of 5 years. It is concluded that the integration of a VRHP into an LRDWC increases the CI of the process although it reduces the OC of the process, and thus the TAC of the process reduces in a payback period of 5 years. Furthermore, the reduction of the TAC is better for a longer payback period, as shown in Table 4. 6.
Conclusion
In this study, four integrated LRDWCs with VRHP models are introduced to the transesterification reaction of MEAC and IPOH. The energy-saving, economic and environmental performances are compared before and after the integration of the VRHP into the LRDWC. The integration significantly improves the energy-saving performance of the process; thus, the OC of the process is reduced. Although the CI of the process increases due to the integration of the VRHP into the LRDWC, the TAC of the process reduces under a longer payback period. The integration of VRHP into the LRDWC results in a significantly high environmental performance. The preheater
further improves energy-saving and economic performance of the process. Model 4 shows the best performance when compared with those of all the models introduced in the study. It saves 49.86% of the TUC, 13.76% of the TAC for a payback period of 5 years, and all the CO2 emissions when compared with the LRDWC. The integrated LRDWC with the VRHP model is attractive in terms of industrial applications due to the significant reduction in energy consumption and CO2 emissions.
Acknowledgements We are thankful for the assistance from the staff at the Jiangsu Key Laboratory of Advanced Catalytic Materials and Technology from School of Petrochemical Engineering, Changzhou University.
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(a)
(b)
(c)
(d)
Fig. 1 Two different structures of RD and the corresponding RDWC configurations: (a) Reactive distillation in a direct sequence; (b) Reactive distillation in an indirect sequence; (c) Upper partitioned reactive dividing-wall column (URDWC); and (d) Lower partitioned reactive dividing-wall column (LRDWC).
CON 5127.15kW
1
REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
MEAC 50kmol/h 60
10
IPOH 50kmol/h 60
30
RD RC 72
84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
REB1 4895.82kW
18 REB2 386.86kW
Fig. 2 Process flow diagram of the LRDWC.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
CON 4780.87kW
53.6 578.81kmol/h
1 42.17 kmol/h
REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
MEAC 50kmol/h 60
10
IPOH 50kmol/h 60
COMP 40.58kW Rc=2.4
30
RD RC 97.5
72
Valve
69.9 84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
REB1 4895.82kW
18 REB2 386.86kW
Fig. 3 Process flow diagram of Model 1.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
1.742
5.03
5.02 1.740 1.739
5.01
1.738 5.00 1.737 1.736
4.99
TUC (MW)
TAC of 5 years (106$/year)
1.741
1.735 4.98 1.734 1.733
4.97 1.8
2.0
2.2
2.4
2.6
Compression ratio
Fig. 4 Effects of the compression ratio on the TAC and the TUC in Model 1.
5 1.0002
4
Pressure (atm)
1.0000
1
0.9998 53.635
3
53.640
53.645
3
2
2
1 50
VF=1 VF=0
4
1 60
70
80
90
100
Temperature (Β°C)
Fig. 5 PβT curve of the overhead vapor composition in Model 1.
CON 1115.43kW
53.6 578.81kmol/h
1 564.62 kmol/h REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
MEAC 50kmol/h 60
10 IPOH 50kmol/h 60
COMP 884.10kW Rc=4.0
30
RD RC 123.3
72
Valve
96
84
18
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
REB1 4895.82kW
REB2 386.86kW
Fig. 6 Process flow diagram of Model 2.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
1.59 3.24 3.22 3.20
1.57
3.18 1.56
3.16
1.55
3.14 3.12
1.54
3.10
TUC (MW)
TAC of 5 years (106$/year)
1.58
3.08
1.53
3.06 1.52
3.04
1.51
3.02 4.0
4.1
4.2
4.3
4.4
Compression ratio
Fig. 7 Effects of the compression ratio on the TAC and TUC in Model 2.
5
VF=1 VF=0
4.1
3 4.0
Pressure (atm)
4
3.9 95
2
3
96
97
98
3 1.0002
1.0000
2
0.9998 53.635
1
53.640
53.645
4
1 50
1
60
70
80
90
100
110
120
Temperature (Β°C)
Fig. 8 PβT curve of the overhead vapor composition in Model 2.
CON 728.57kW
53.6 578.81kmol/h
1 564.62 kmol/h
REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
10 IPOH 50kmol/h 60
RD RC
MEAC 50kmol/h 60
123.3
72
84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
COMP 884.10kW Rc=4.0
30
96
REB1 4895.82kW
18
REB2 386.86kW
Fig. 9 Process flow diagram of Model 3.
Valve
80.1
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
2.84 2.82 2.80
1.48
2.78 1.47 2.76 2.74
1.46
2.72 1.45 2.70 1.44
TUC (MW)
TAC of 5 years (106$/year)
1.49
2.68 2.66
1.43
2.64 4.0
4.1
4.2
4.3
4.4
Compression ratio
Fig. 10 Effects of the compression ratio on the TAC and TUC in Model 3.
CON 727.37kW
53.6 578.81kmol/h 536.64 kmol/h
1
Valve
70.3 REC 61.54kmol/h 53.6 MEAC=0.662 MEOH=0.338
10 IPOH 50kmol/h 60
PHE 200kW
30 71.1
RD RC
MEAC 50kmol/h 60
79.3 72
COMP 882.90kW Rc=4.0 141.9
84
IPAC 50kmol/h, 91 MEAC=1.18e-05 IPOH=5.50e-04 IPAC=0.997 MEOH=0.002
96
REB1 4895.82kW
18
REB2 386.86kW
Fig. 11 Process flow diagram of Model 4.
MEOH 50kmol/h, 64.8 MEAC=0.003 IPOH=0.002 MEOH=0.995
(a)
(b) Rc=4.4
1.48
Rc=4.3
2.82 2.80
Rc=4.4
2.78
Rc=4.3
2.76 1.47
TUC (MW)
TAC of 5 years (106$/year)
1.49
Rc=4.2 1.46
Rc=4.1
1.45
Rc=4.2
2.74 2.72
Rc=4.1
2.70 2.68
1.44 2.66
Rc=4.0
Rc=4.0
2.64
1.43 0
50
100
150
Heat duty of the Preheater (kW)
200
0
50
100
150
200
Heat duty of the Preheater (kW)
Fig. 12 Effects of the heat duty of the PHE and the compression ratio on (a) the TAC and (b) the TUC in Model 4.
Table 1 Binary parameters of the UNIQUAC model Component i
MEAC
IPOH
IPAC
MEAC
MEAC
IPOH
Component j
MEOH
MEOH
MEOH
IPOH
IPAC
IPAC
bij
-337.419
-169.757
-401.981
-66.4607
1.96175
67.3403
bji
54.3901
70.9141
68.1792
-40.9002
-20.9966
-163.536
Table 2 Cost estimating formula of the CI Column vessel CI = 17640π·1.066 πΏ0.802, where D denotes column diameter (m), and L denotes column height (m). πΏ = (ππ β 1) Γ 0.61 + 6, where NT denotes actual stage number. Heat exchanger CI = 7296π΄0.65, where A denotes heat transfer area (m2). π΄=
π πβπ
, where Q denotes heat duty of the heat exchanger (kW), and U denotes the heat transfer
coefficient. βT is calculated by using a logarithmic mean temperature difference (LMTD) method. Compressor CI =
π&π 280
Γ 1264.75(βπ)0.82, where hp denotes work of the compressor.
πβπ = βπ/0.8, where bhp denotes brake horsepower of the compressor.
Table 3 Comparisons of energy consumptions and CO2 emissions Parameter Reboiler duty (kW) Condenser duty (kW)
LRDWC
Model 1
Model 2
Model 3
Model 4
4895.82
4895.82
386.86
4780.87
1115.43
728.57
727.37
386.86
4895.82
4895.82
4895.82
386.86
386.86
386.86 5127.15
Heat Exchanger duty (kW) Preheater duty (kW)
200
Compression ratio
2.4
4
4
4
Compressor work (kW)
40.58
884.10
884.10
882.90
5017.57
3039.15
2652.29
2648.69
(-5.02%)
(-42.47%)
(-49.79%)
(-49.86%)
2499.14
197.48
0
0
(-7.32%)
(-92.68%)
TUC (kW)
5282.68
[Savings (%)] CO2 emissions (kg/h) [Savings (%)]
2696.62
Table 4 Comparisons of costs Parameter 3
Column cost (10 $) 3
Reboiler cost (10 $) 3
Condenser cost (10 $)
LRDWC
Model 1
Model 2
Model 3
Model 4
1427.20
1427.20
1427.20
1427.20
1427.20
194.48
168.23
26.25
236.09
225.60
87.60
66.41
66.34
459.02
460.92
528.31
513.23
121.70
1522.56
1522.56
1520.87
2.37
3.52
3.54
3.53
(+27.58%)
(+89.72%)
(+90.79%)
(+89.89%)
1183.66
1096.98
86.68
105.90
98.75
23.04
15.05
15.02
32.73
712.94
712.94
711.97
1.26
0.82
0.73
0.73
(-2.29%)
(-36.21%)
(-43.55%)
(-43.62%)
1.73
1.53
1.44
1.43
(+4.39%)
(-8.04%)
(-13.50%)
(-13.76%)
1.56
1.26
1.17
1.17
(+2.32%)
(-1.70%)
(-23.05%)
(-23.25%)
Heat exchanger (including preheater) cost (103 $) Compressor cost (103 $) 6
CI (10 $)
1.86
[Increase (%)] Steam cost
(103
$/a)
Cooling water cost
(103
$/a) Electricity cost (103 $/a) OC
(106
$/a)
1.29
[Savings (%)] TAC with the payback period of 5 years
(106
1.66
$/a)
[Increase / Savings (%)] TAC with the payback period of 8 years (106 $/a) [Increase / Savings (%)]
1.52