Energy Consumption and Exergy Analysis of MEA-Based and Hydrate

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Energy Consumption and Exergy Analysis of MEA-Based and Hydrate-Based CO2 Separation Nan Xie,† Bin Chen,‡ Chenghua Tan,† and Zhiqiang Liu*,†,§ †

School of Energy Science and Engineering, Central South University, Changsha 410083, China AECC Hunan Aviation Powerplant Research Institute, Zhuzhou 412002, China § Collaborative Innovation Center of Building Energy Conservation & Environmental Control, Zhuzhou 412007, China ‡

S Supporting Information *

ABSTRACT: Carbon capture and storage (CCS) is regarded as the most efficient approach in handling the global warming crisis. MEA-based CO2 capture is a well-developed chemical absorption method with a long period of industrial application. A novel hydrate-based gas separation (HBGS) method, with a wide range of advantages, has recently received special attention from researchers. In this study, two different CO2 separation processes were simulated utilizing Aspen Plus software. The feasibility of both processes was validated, and the process energy consumption and exergy loss were also compared at the same flue gas condition. Some efforts have also been made to investigate the effects of different operation parameters on the process energy efficiency. Results show that the first law efficiency of the MEA-based CO2 separation system is 88.19% and the second law efficiency of the system is 38.32%, while the corresponding values of the hydrate-based separation system are 74.15% and 38.85%, respectively. Cooling of the lean amine solution and the regeneration process occupy the largest portion of exergy loss in the MEA separation system. In the hydration separation, the flue gas compression and cooling are the major causes for exergy loss.

1. INTRODUCTION Rising CO2 emission has led to various environmental issues, and global warming, most of all, remains a major concern among governments and scientists. It is common sense to see that the anthropogenic emission of CO2 is mainly due to the burning of fossil fuels, especially the combustion process in large power plants. Carbon capture and storage (CCS) technology is now seen as the most valid solution to reduce CO2 emission, essentially known as precombustion, oxy-combustion, and postcombustion methods. Physical absorption, chemical absorption, pressure swing adsorption, cryogenic condensation, and membrane separation are common techniques in capturing carbon dioxide, particularly the MEA-based absorption.1,2 As a well-developed amine-based absorption technique, it has a long history of industrialization. However, it is still hindered by its own weakness, such as the energy intensive process of solvent regeneration, high corrosion on equipment, etc.3,4 Hydrate-based gas separation (HBGS) is now considered as a novel way to avoid the massive atmospheric buildup of CO2. Hydrate-based CO2 capture is realized by the selective partition of CO2 between the hydrate phase and the gaseous phase. Gas hydrates, also known as clathrate hydrates, are nonstoichiometric, crystal compounds which have cagelike structures formed by water and gas molecules.5 Structure identification and investigation are carried out by researchers to better understand the nature of gas hydrates.6−8 Generally, there are three different kinds of crystalline structures (sI, sII, sH) depending on the © XXXX American Chemical Society

sizes of the gas molecules. Most research is focused on the formation/dissociation conditions, enthalpy, additives, and phase equilibrium.9−19 Kumar et al.10 investigated the effects of thermodynamic and kinetic promoters, in which the promoters used in this study were sodium dodecyl sulfate (SDS) and tetrahydrofuran (THF). Their results showed the gas uptake was increased by almost three-fold in the presence of kinetic promoter SDS. The hydrate formation pressure was obviously reduced with the addition of thermodynamic promoter THF. Three different ternary systems of CO2 + water + (TBAB/TBAC/TBAF) were reported by Li et al.11 An isochoric pressure-search method was utilized to measure the equilibrium data of these semiclathrate hydrates. The lowest formative pressure at the same temperature was found in the system of tetra-n-butyl ammonium fluoride (TBAF). The synergic effect of TBAB in the presence of both cyclopentane and dodecyl trimethylammonium chloride (DTAC) was carefully examined by Li et al.12,13 Gas uptakes, CO2 selectivity, CO2 separation efficiencies, induced time, and hydrate formation rate were also investigated by them.20 They found that the CO2 concentration reaches approximately 93% by separation involving only one stage. Similar investigations were made by many Received: Revised: Accepted: Published: A

September 10, 2017 November 20, 2017 November 29, 2017 November 29, 2017 DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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Industrial & Engineering Chemistry Research other researchers,13,21−23 who discovered that the captured CO2 concentration in all these studies can be enriched to a relatively high level in only one or two separation stages. A comprehensive review was given to the properties of CO2 hydrates and the hydrate-based CCS by Ma et al.24 This technology is still at its early stage because of the complicated and stochastic process of hydrate formation and diverse experimental conditions. Hence, they suggested that large-scale system design and economic and environmental impact analysis should be further conducted. Xu and Li made original comments and suggestions on further development of the hydrate-based CCS technology, including equilibrium conditions, additives, molecular-level measurement methods, process optimization, new apparatuses and cost, etc.25 Energy consumption estimation of the HBGS process was carried out by Tajima et al.26 Hydration separation processes of three greenhouse gases were studied and compared with liquefaction separation processes in terms of the energy consumed. They found that pure CO2 hydrate separation consumed a huge amount of energy due to the extremely high formation pressure. This high pressure can be dramatically alleviated by adding appropriate additives. After reviewing the literature, we found no research on the comparison between the MEA-based and the hydrate-based separation processes. Thus, it is essential to make a comprehensive comparison in terms of energy efficiency. As a commonly used tool for simulating various CO2 capture processes in power plants,27−29 Aspen Plus is utilized in this study as well. Based on the simulated results, the energy consumption estimation and the exergy analysis of a two-stage hydrated-based CO2 separation process are carried out. The results are then compared in detail with those of a MEA-based separation process. Furthermore, sensitivity studies of various operational factors which influence the process energy efficiencies are performed as well. Our research presents a thorough comparison between a well-developed and a novel CO2 separation technology, providing results with a deeper understanding of CCS technology. This work will also be useful in the energy conservation and the process optimization of CCS technology in large power plants.

absorbent on equipment is neglected; (2) compression, drying, and transportation processes of products are also ignored; (3) there is no heat dissipation in the heat transfer process; (4) the main reactions of the present separation process are given in Table 2.2,30 Gaseous CO2 forms a semiclathrate hydrate with the promoter tetra-n-butyl ammonium fluoride (TBAF). In this crystal structure, a part of the cage structure is broken to encage the large ion TBA+. And the fluoride anion can form a polyhedral hydrogen-bonded cage structure with water molecules and change the structure of this crystal inclusion compound, moderating the formation condition as a result. Basic assumptions of the hydrate-based CO2 capture process are given as follows: (1) The pressure of flue gas after the first compression is assumed to be 0.5 MPa, and 2.5 MPa after the second compression, and the compression is considered as an adiabatic process with the efficiency of 0.8; (2) the minimum temperature difference in exchangers is specified at 10 °C and the COP (coefficient of performance) of the refrigeration unit is equal to 3; (3) the hydrate dissociation temperature is set to be 25 °C; (4) compression work imposed on the feed gas is recovered by the expansion of the off-gas stream in a turbine of which the efficiency is 0.8; (5) the molar fraction of the additive TBAF is assumed to be 0.29%; thus, the dissociation enthalpy of the CO2 + TBAF hydrate is estimated as 144 kJ/mol using the Clausius−Clapeyron equation based on the available equilibrium data.

3. PROCESS SIMULATION 3.1. System description. Simulation diagrams of two separation processes are shown in Figure 1 and Figure 2, respectively. In Figure 1, the feed gas is cooled to the absorption temperature at first and then introduced into the ABSORBER, in which the CO2 is absorbed by a lean solution to form a CO2rich solution. The clean flue gas leaves the vessel from the upper outlet. The carbon-rich solvent is then pumped into the exchanger (EXCHANGE) and heated up by the lean solution issuing from the STRIPPER to an appropriate temperature for regeneration. After leaving the exchanger, the rich CO2 solvent is heated by the superheated steam from the reboiler to release its burden of CO2. The CO2-rich gas exits the stripper after condensation as the product of the current system. The regenerated solution (CO2-lean solution) is cooled (in COOLER2) to the absorbing temperature and enters the absorber again to conduct further work. The MEA solution loss is supplemented by using a MIXER. In Figure 2, a two-stage separation process is described to produce gaseous CO2 with a concentration higher than 90% by volume. The high temperature flue gas is precooled by an ambient fluid before it enters the compressor. After the first compression in an adiabatic compressor (SCOMP1), the pressurized gas mixture is cooled again to a given temperature by the lean CO2 gas emitted from the hydrate formation reactor (FORMAT1). And then, the feed gas is adiabatically pressurized for the second time to reach the formation pressure. The feed gas is cooled to the formation temperature by using a refrigeration unit (REFRIG1) and is then introduced into the hydrate reactor to form a CO2 hydrate in the TBAF solution. In the dissociation reactor (DISSOC1), the hydrate is heated up by an ambient fluid until it breaks down to the CO2-rich gas and the CO2-lean solvent. The latter is piped back to the formation unit to maintain the cycle operation. The former is pressurized once again to undergo another similar process to a

2. BASIS PARAMETERS OF FLUE GAS AND PROCESS ASSUMPTIONS 2.1. Flue gas parameters. Based on the flue gas emitted from a typical 1000 MW thermal power plant after desulfurization and denitration, the temperature and the pressure of the flue gas are fixed at 140 °C and 0.1 MPa, respectively. The concentration of CO2 in the flue gas is assumed to be 16%, which is within the reasonable range of 14% to 18%. The composition and the simplified parameters of the flue gas are shown in Table 1. 2.2. Process assumptions. Some reasonable assumptions are currently made in the simulation of two different CO2 separation processes. The basic assumptions of the MEA-based CO2 capture process are given as (1) corrosion behavior of the Table 1. Basis Parameters of the Flue Gas in Present Study Component

Mole fraction

Molar mass

Feed rate (kmol/h)

Feed rate (t/h)

CO2 N2 O2 H2O

0.16 0.71 0.05 0.08

44 28 32 18

7143 31697 2232 3571

314 888 71 64 B

DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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Industrial & Engineering Chemistry Research Table 2. Reactions of the MEA-Based CO2 Capture Process2,30 No.

Type

Reaction equations

1 2 3 4 5

Equilibrium Equilibrium Equilibrium Equilibrium Equilibrium

MEA+ + H2O ↔ C2H7NO + H3O+ CO2 + 2H2O ↔ H3O+ + HCO3− HCO3− + H2O ↔ H3O+ + CO3−2 2H2O ↔ H3O+ + OH− MEACOO− + H2O ↔ C2H7NO + HCO3−

Coefficients for calculating the equilibrium constants (Keq) in the following equations: ln(Keq) = A + B/T + C* ln(T) + D*T A A A A A

= = = = =

−3.038, B = −7008.36, C = 0, D = −0.0031 231.47, B = −12092.1, C = −36.78, D = 0 216.05, B = −12431.7, C = −35.48, D = 0 132.899, B = −13445.9, C = −22.48, D = 0 −0.52, B = −2545.53, C = 0, D = 0

Figure 1. Process flowsheet of the MEA-based CO2 separation.

Figure 2. Process flowsheet of the hydrate-based CO2 separation.

higher concentration. In the current system, part of the compression work imposed on the feed gas is recovered from the off-gas by using a TURBINE. 3.2. Selection of model. Thermodynamics and kinetics of the MEA-based CO2 capture system are modeled using the “MEA Property Insert” of Aspen Plus software. Thus, the electrolyte−NRTL model is currently utilized to describe the MEA−water−CO2 system. Besides, Elec Wizard in Aspen Plus is employed to describe the complex chemistry and thermodynamics of the MEA property. GMELCC-1, GMELCD-1, GMELCE-1, and GMELCN-1 (electrolyte−molecules/electrolyte−electrolyte binary energy parameters C, D, E, and nonrandomness parameter alpha in the electrolyte NRTL activity coefficient model) are modified for the present research. The ELECNTRL model is utilized to handle various problems involving aqueous electrolyte solution for different concentrations. Detailed configurations of the absorber and the stripper are listed in Table 3. The block types and the parameters of other individual equipment are shown in Table 4. The properties of the streams in the hydrate-based separation system of interest are simulated by using the PRMHV2 model. Based on the Peng−Robinson equation of state and the modified Huron−Vidal mixing rule, this model has an excellent

Table 3. Detailed Configurations of the Absorber and the Stripper Model type Calculation type Number of stages Condenser type Reboiler type Reflux Entry position Stage 1/Condenser pressure (kPa) Column pressure drop (kPa)

absorber

stripper

RadFrac Equilibrium 10 None None

RadFrac Equilibrium 10 Partial-Vapor Kettle 1.2 On-Stage 69 10

On-Stage 220

Table 4. Block Types and Parameters of Individual Device in the MEA-Based Process Name

Type

Specification

PUMP COOLER1 EXCHANGE COOLER2

Pump HeatX HeatX HeatX

Discharge pressure, 0.26 MPa Hot stream outlet temperature, 40 °C Cold stream outlet temperature, 92 °C Hot stream outlet temperature, 40 °C

degree of predictive accuracy for high pressure cases (>150 bar), like the streams in the current process. Table 5 shows the C

DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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Industrial & Engineering Chemistry Research Ex in − (Exout + Ir) = W

Table 5. Block Types and Parameters of Individual Device in the Hydrate-Based Process Name

Type

SCOMP1

Compr

SCOMP2

Compr

SCOMP3

Compr

COOLER EXCHANGE1 EXCHANGE2 TURBINE REFRIG1 REFRIG2

HeatX HeatX HeatX Compr HeatX HeatX

where Exin represents the exergy that enters the system, kJ; and Exout means the exergy that leaves the system, kJ; Ir refers to the exergy loss and also represents the degradation of energy caused by process dissipation, kJ; W is the net work output of a system or device, kJ. Exergy efficiency, also known as the second law efficiency, can be calculated by

Specification Isentropic; Discharge pressure, 0.5 MPa; Mechanical efficiency 0.8 Isentropic; Discharge pressure, 2.5 MPa; Mechanical efficiency 0.8 Isentropic; Discharge pressure, 0.9 MPa; Mechanical efficiency 0.8 Hot stream outlet temperature, 50 °C Hot stream outlet temperature, 90 °C Hot stream outlet temperature, 60 °C Isentropic; Mechanical efficiency, 0.8 Hot stream outlet temperature, 4.5 °C Hot stream outlet temperature, 4.5 °C

ηex = W /Extotal = 1 − ηex,loss = 1 − Ir /Extotal

4. ENERGY CONSUMPTION ESTIMATION AND EXERGY ANALYSIS METHOD 4.1. Energy balance algorithm. It is a fundamental principle that energy cannot be created or destroyed, but it can be transferred or transformed into different forms with its total amount unchanged. Based on the first law of thermodynamics, the energy balance can be described as Q in − Q out = ΔQ (1)

5. RESULTS AND DISCUSSION 5.1. MEA-based CO2 separation process. Energy balance and distribution of the MEA-based CO2 separation system (MEA system) are carefully calculated (details are shown in Table S1 in the Supporting Information), at the reference state T0 = 298.15 K, P0 = 1 atm. The major energy input is the superheated steam (Stream 12) from the reboiler, which occupies 78.29% of the total energy input. Only 0.18% of the total energy is taken away by the regenerating gas (Stream 16), while 36.59% of the total energy output is caused by the cooling water (Stream 20) leaving the cooler. And the overall first-law efficiency of the MEA system is 88.19%. Exergy balance and distribution are also calculated based on the simulation results (details in Table S2). It can be found that 78.93% of the total exergy input is also achieved by the superheated steam. The maximum exergy loss exists in the processes of the MEA solvent regeneration and the regenerating gas cooling in the stripper (37.14% of the total). The exergy loss of all devices is 61.68% of the total; hence, the exergy efficiency of the current system is only 38.32%. According to our energy consumption estimation and exergy analysis, the main consumption in this process is the heating of the CO2-rich solution (Stream 6) to release its burden of target gas. So, energy saving in the stripper is the key problem. High cost of the heating process is determined by the essential characteristics of the absorbents. Thus, searching for new absorbents and new additives seems to be the only way, which is the main constraint of this technique. 5.2. Hydrate-based CO2 separation process. Compositions of several typical streams in this process are shown in Figure 3. The concentration of gaseous CO2 is increased from 16% (mole fraction) in the flue gas (a) to 60% in the first-stage separated gas (c), and eventually to 90% in the second-stage separated gas (e). The CO2 concentrations in the off-gas (b, d) are very low, which proves the validity and feasibility of the CO2 separation by using hydrate. Energy consumption as well as exergy analysis are carried out (as shown in Table S3 and Table S4 of the Supporting Information). The electric power for pressurizing and cooling the streams is the major energy input of the hydrate system. The chemical reaction heat needed in the dissociation units can be provided by an environmental fluid because the endothermal CO2 hydrate dissociation can happen at ambient temperature. But the heat removal in the formation units can only be realized by consuming the refrigeration

where Qin represents the energy entering the system, kJ; and Qout means the energy that leaves the system, kJ; ΔQ is the energy change in the system, kJ. The thermal efficiency can be defined as (2)

where W means the useful work, kJ; and Qtotal is the energy consumed in the system, kJ. Thus, the thermal efficiency is currently adopted as a major evaluation index of the present study. Simulation results of both processes for energy consumption calculations are illustrated in Figure S1 and Figure S2 in the Supporting Information. 4.2. Exergy analysis method. On the basis of the second law of thermodynamics, the exergy efficiency is also taken as a major criterion to evaluate the quality of the energy, defined as the maximum potential of the energy that can be converted into useful work. In the processes of our interest, the total exergy consists of physical exergy and chemical exergy. Physical exergy can be given as follows: Exphys = − (H0 − H1) + T0(S0 − S1)

(3)

where the H0 means the enthalpy of the stream at reference state (T0, P0), kJ; S0 means the entropy at reference state, kJ/K; H1 is the enthalpy of the stream, kJ; S1 is the entropy of the stream, kJ/K. And chemical exergy can be given as Exchem =

∑ xiEx0,i + ∫ RT0∑ xi ln xi i

i

(6)

where Extotal represents the total exergy input into the system, kJ; and ηex,loss means the exergy loss efficiency, which shows the distribution of exergy loss in a certain process. Therefore, the exergy analysis for both CO2 separation processes is conducted based on the results obtained in the simulation and then compared and discussed in detail. Simulation results of both processes for specific calculations of exergy loss are illustrated in Figure S1 and Figure S2 in the Supporting Information.

block types and the simulation parameters of the hydrate-based separation.

ηt = W /Q total

(5)

(4)

where xi represents the mole fraction of component i; Ex0,i is the standard chemical exergy of component i, kJ; R represents the universal gas constant, 8.314J/mol·K; T0 is the temperature of the reference state, K. All processes in nature are irreversible and, hence, lead to exergy loss, which represents the irreversibility of the system or process; thus, the exergy balance equation can be described as D

DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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the flue gas to the formation pressure reaches more than onehalf of the total input. The energy consumption per unit weight of CO2 is currently estimated as well. The energy consumption per unit in the hydrate system is 1.117 kWh/kg, being much lower than 2.848 kWh/kg of the MEA system. This indicates that the separation process by hydration has a significant advantage. These results are also compared with previous literature, in which the energy consumption per unit is 0.853 kWh/kg in the research of Tajima et al.26 Our results are higher because the energy consumed for cooling the flue gas before its compression is considered in our research, and the separated CO2 flow rate in our study is 284 t/h higher than the value (186 t/h) in the literature. Moreover, the enthalpy in their study was 73 kJ/mol while the value in our work is 144 kJ/mol because of the addition of TBAF. Figure 5 shows the exergy loss in each device. It can be observed that the exergy loss has a concentrative distribution for both processes. In the MEA system, a considerable amount of exergy loss which can barely be recovered was taken away by the cooling process in the condenser. Hence the regeneration and the cooling in the stripper led to the major part of exergy loss. For this reason, further energy saving in the MEA system might be hard to achieve. In the hydrate-based system, nearly one-third of the total exergy loss was found in the refrigeration units (23.62%) as well as the adiabatic compressors (8.62%). Thus, to improve the energy efficiency of both systems, sensitivity studies focusing on the stripper, compressors, and refrigeration units are quite necessary. 5.4. Sensitivity analysis. In the hydrate-based separation system, a large amount of exergy loss was caused by the cooling process in refrigeration units. There is a value to study the effect of refrigeration performance on the process energy efficiency. Thus, the effect of COP on the energy consumption per unit weight of CO2 was investigated as shown in Figure 6. With the increase of COP of refrigeration units, the energy consumption per unit goes down apparently, but this trend slows down with its continuous increase. Therefore, improving COP value is favorable in the hydrate-based system, but its effect is less marked for COP larger than 2.5. This means pursuing an excessive COP value will be both ineffective and impractical. Figure 7 shows the variation of the energy consumption per unit by improving the efficiency of compressors. The variation curve also presents a descending tendency with the rising efficiency of compressors. Thus, energy conservation in this system can also be achieved by improving the efficiency of compressors.

Figure 3. Compositions of different streams: (a) flue gas; (b) off-gas after the first-stage separation; (c) first-stage separated gas; (d) off-gas after the second-stage separation; (e) second-stage separated gas.

power 3 and 4 (in Table S3). Only 0.09% of the energy input is taken away by the gaseous mixtures (Stream 20) after the final separation process. A turbine is utilized to recover the energy in the first-stage off-gas (Stream 8), but still, around 3% leaves the system with the off-gas (Stream 9). The overall first-law efficiency of this system is only 74.15%. In the exergy estimation, the electric power input for the compression occupies more than 55% of the total input, but the maximum exergy loss is in the refrigerators. A part of exergy loss is caused by the emission of the off-gas (Stream 9 and 17). The exergy efficiency of the hydrate separation system is 38.85%. Except for the heat-exchange network optimization aiming at a certain power plant, or the improvements on compressors and refrigerators, effective energy saving can also be realized by moderating the hydrate formation condition using suitable additives, which declares a remarkable potential of energy conservation. 5.3. Comparison of both separation processes. By contrast, the energy efficiency of the hydrated-based separation system is lower than that of the MEA absorption system, but with a slightly higher exergy efficiency. Higher exergy efficiency means a slower energy degradation. Figure 4 shows the distributions of the energy input in these two systems. The energy distribution in the MEA system is quite uneven, and the thermal energy of the reboiler steam plays the dominant role in the whole process, determined by the solution regeneration of the absorbent. In the hydrate system, the distribution also seems to be an uneven split. The electric power for pressurizing

Figure 4. Input energy distribution of both CO2 separation processes: the MEA-based (left) and the hydrate-based (right). E

DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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Figure 5. Exergy loss in different devices.

Figure 6. Effect of COP on the process energy consumption per unit. Figure 8. Effect of solvent flow rate on the heat duty of the reboiler and the COOLER2.

heat duty in the reboiler and the lean solvent cooler. An obvious linear relation between the heat duty and the mass flow rate was observed for both devices when increasing the mass flow rate from 6100 t/h to 7000 t/h. The thermal energy from the reboiler was partly absorbed by the rich CO2 solution as the reaction heat was needed for maintaining the CO2 desorption process. The rest was taken out of the system by the lean solvent and the regenerating gas. In the lean solvent cooler, the heat received from the reboiler was eventually taken away by an environmental fluid. In the regenerating gas, a large amount of water vapor and some volatile gases (in practical operation) must be condensed and removed before leaving the system. This also helps acquire a higher concentration of the target gas. Figure 9 shows the relation between the heat duty of the condenser and the mass flow rate of solvent. This positive function also shows a linear variation but with a small difference. An increasing slope was spotted on the curve in the range of 6100 t/h to 6300 t/h. It can be explained as when the flow rate of flue gas remains unchanged, increasing the solvent flow rate will certainly augment the CO2 absorption, which leads to a nonlinear increase in the heat duty of the condenser because of the rising amount of the regenerating gas, while continuously

Figure 7. Effect of compressor efficiency on the process energy consumption per unit.

In the MEA-based separation process, nearly a half of the total exergy was expended on the regeneration of the MEA solvent and the cooling of the lean solvent and the regenerating gas. This involves three major devices: the lean solvent cooler (COOLER2), the reboiler, and the condenser (both in STRIPPER). Hence the effect of the mass flow rate of MEA solvent on the thermal load of these three devices was investigated carefully. Figure 8 presents the variation curves of the F

DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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between the literature and the present work. Furthermore, the energy efficiency of the removal process can also be enhanced by implementing various modifications, such as solvent properties variation, device structural adjustment, and operation optimization.

6. CONCLUSIONS High energy consumption is seen as the major weakness of CO2 capture and separation technology. Much research was conducted to improve the energy performance. In this work, energy consumption and exergy loss of two different CO2 separation systems were estimated from the simulation results, and then compared in detail. The main conclusions are as follows: (1) Carbon dioxide can be efficiently separated from the flue gas emitted from a thermal plant in both systems. At the same flue gas condition, the first law efficiency of the MEA-based CO2 capture system is 88.19%, higher than the value 74.15% of the hydrate-based system. The second law efficiency is 38.32%, being slightly lower than the value 38.85% of the latter. (2) In the process of MEA absorption, most energy is consumed by the solvent regeneration, during which the energy conservation should be considered carefully. The heat loads of the reboiler, the condenser, and the cooler are significantly influenced by the solvent flow rate. The separation rate of CO2 can obviously affect the energy performance as well. (3) In the hydrate-based process, the refrigeration process causes the largest portion of the total consumption. Raising the COP of refrigerators helps achieve energysaving, but this effect disappears with its continuous increase. Besides, improving the compressor efficiency can also dramatically enhance the energy efficiency of the removal process. (4) Searching for new absorbents and additives might be the only way to reduce the regeneration energy consumption in the MEA-based CO2 separation. Therefore, the hydratebased separation can be an alternative method with a promising prospect, although it has a lot of requirements of devices and limits for now. In summary, with some improvements to the hydrate-based CO2 separation process, this technique will show a distinct advantage of energy savings. Future research focusing on the energy efficiency analysis, technology optimization, as well as structural design of devices is greatly encouraged before its industrial application.

Figure 9. Effect of solvent flow rate on the heat duty of the condenser.

increasing the MEA solvent to an excessive degree will not augment the absorbed amount anymore because the CO2 in the flue gas is limited. Thus, the heat duty of the condenser transforms into a single value linear function of the solvent flow rate. In addition, it can also be found in Figure 8 and Figure 9 that the thermal load in the cooler was one order higher than that in the condenser. This indicates that the thermal energy in the regeneration process was mostly taken away by the lean solvent, not the regenerating gas. Thus, the lean solvent cooler is also considered as a key point of energy saving, where a considerable thermal energy might be recovered somehow. Most energy was carried into the MEA system with the high temperature steam, providing the regenerating energy for CO2 desorption. As presented in Figure 10, the regenerating energy

■

Figure 10. Effect of the CO2 removal rate on the regenerating energy.

ASSOCIATED CONTENT

S Supporting Information *

The Supporting Information is available free of charge on the ACS Publications website at DOI: 10.1021/acs.iecr.7b03729. Total energy and exergy input/output of both CO2 separation systems (Table S1, Table S2, Table S3, and Table S4) and simulation results of the streams in both CO2 separation processes (Figure S1 and Figure S2) (PDF)

in the MEA-based separation system obviously augments with the rising removal rate. At removal rate around 90%, the regenerating energy in our study was 5.3424 MJ/kg, being slightly higher than 5.13 MJ/kg in previous research.31 This proves the reliability of the results in the present work. The suitable range for industrial application was recommended by the authors as 3.2 MJ/kg to 4.2 MJ/kg. Hence, a lot of efforts were made by them in order to optimize the operation, and preferable results were obtained. The average value of the regenerating energy in the literature32 was 4.12 MJ/kg, within the reasonable range mentioned above. This difference might be due to the discrepancy of the amount and the composition of flue gas

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AUTHOR INFORMATION

Corresponding Author

*Tel: +86-731-88879863. E-mail: [email protected]. G

DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX

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Industrial & Engineering Chemistry Research ORCID

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Nan Xie: 0000-0002-9502-4501 Zhiqiang Liu: 0000-0002-6504-0543 Notes

The authors declare no competing financial interest.

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ACKNOWLEDGMENTS This research was financially supported by the National Natural Science Foundation of China (Grant No. 51376198 and 51676209), and Fundamental Research Funds for the Central Universities of Central South University (No. 2017zzts129).

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DOI: 10.1021/acs.iecr.7b03729 Ind. Eng. Chem. Res. XXXX, XXX, XXX−XXX