Comparative Study of Process Integration and Retrofit Design of a

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Comparative Study of Process Integration and Retrofit Design of a Liquefied Natural Gas (LNG) Regasification Process Based on Exergy Analyses: A Case Study of an LNG Regasification Process in South Korea Seongho Park, Chansaem Park, Ung Lee, Ikhwan Jung, Jonggeol Na, Krishnadash S. Kshetrimayum, and Chonghun Han* School of Chemical and Biological Engineering, Seoul National University, Gwanak-ro 599, Gwanak-gu, Seoul 151-742, South Korea S Supporting Information *

ABSTRACT: Exergy analysis of the retrofit design scheme of a conventional liquefied natural gas (LNG) regasification process in South Korea was considered in this study. A new exergy evaluation method called exergy decomposition is introduced, in which the exergy is decomposed into thermal and chemical exergies. In studying the conventional LNG regasification process, we found that a large portion of chemical exergy is lost by boil-off gas flaring. Of 17 MW of thermal exergy transferred from cold LNG to seawater in the regasification unit, a fraction as large as 16 MW (close to 95%) is wasted because of heat-transfer irreversibility, limiting the rational exergetic efficiency of the overall process to merely 0.847. Previously reported design schemes, namely, the dual Brayton cycle and the organic Rankine cycle, with low-grade heat sources were also evaluated using the new method and were found to limit the overall rational exergetic efficiencies to 0.890 and 0.849, respectively. A new integrated, retrofitted scheme for LNG regasification with a gas-to-liquid (GTL) process is proposed as an alternative to minimize thermal and chemical exergy losses. The integrated LNG regasification−GTL process improves the overall rational exergetic efficiency to 0.868.

1. INTRODUCTION Liquefied natural gas (LNG) has been identified as a major future energy source, because of its cleanness, flexibility, high energy density, and abundance.1−6 However, it is well-known that, over the entire LNG value chain, an enormous amount of energy is consumed in the liquefaction process; specifically, an immense energy of about 850 kWh/ton is required to liquefy the NG extracted from the well.7 The liquefaction process cools NG to −160 °C, and the volume is reduced by a factor of 600. The energy consumed in the liquefaction process is stored in the form of LNG cold exergy, which can be recovered by configuration of the thermodynamic cycle. In the literature, it has been reported that exergy recovery of approximately 240 kWh/ton can be achieved.7,8 Much effort has been made to improve the efficiency of the LNG regasification process. Kim et al. established a model of the Rankine power cycle using the cold energy of LNG.9 Qiang et al. analyzed the combined power cycle of the LNG regasification process.8 They used propane as the working fluid to configure the organic Rankine cycle (ORC). A cold exergy recovery scheme using the Brayton cycle has also been proposed.10,11 However, all of these examples evaluate the process using different criteria, for example, thermal efficiency, second-law (exergetic) efficiency, or economic standpoints. Of these methods, exergy-based analysis has recently been attracting attention, because it enables the irreversibility of the overall process to be calculated and minimized.12 Exergy analysis can therefore ensure the best use of fuel resources and minimization of overall thermodynamic losses; this is not possible from a pure economic analysis. © 2014 American Chemical Society

Conventional exergy-based evaluation methods use the lumped exergy concept and consider only the irreversibility of the entire process. In this context, lumped exergy means that the LNG exergy entity is not considered separately, but calculated simultaneously. However, using this concept, it is difficult to analyze the exergy losses in the subprocesses or unit operations or the exergy transferred by each stream. Moreover, it is very hard to determine which type of exergy is dominant quantitatively, and as a result, it is difficult to determine the priorities for the exergy recovery subprocess. In this study, exergy analyses were performed on several LNG regasification processes. The exergy decomposition method was used to analyze the entire process with fragmented exergy entities. First, the base LNG regasification process was modeled using a commercial process simulator, and the exergy analysis was performed. Two retrofit concepts reported in the literature were then modeled and evaluated in the same way. With the knowledge gained, through the application of equivalent criteria, a new scheme integrating LNG regasification and gas-to-liquid (GTL) processes is proposed.

2. EXERGY DECOMPOSITION ANALYSIS 2.1. Conventional Thermodynamic Efficiency. Exergy is the maximum useful work potential of a system in a specified Received: Revised: Accepted: Published: 14366

April 17, 2014 August 7, 2014 August 27, 2014 August 27, 2014 dx.doi.org/10.1021/ie501583m | Ind. Eng. Chem. Res. 2014, 53, 14366−14376

Industrial & Engineering Chemistry Research

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related to effluent streams in which the exergy is not recovered (e.g., exergy in exhaust/flue gases, exergy lost to cooling water, streams emitted to the environment). This quantity is more specific than the conventional exergetic efficiency and can be used to evaluate subprocess efficiency and diagnose the main exergy loss in each unit operation. 2.2. Exergy Decomposition Analysis. We propose a new type of exergy analysis method, called exergy decomposition, in which the exergy is decomposed into thermal, chemical, kinetic, and potential exergies. In exergy calculations, for every stream entering and leaving the control volume, the exergy of each stream is subdivided into several types, so the flow of the main exergy component can be easily understood. From eq 1

state; specifically, it measures how much work can be pulled from a certain entity (e.g., mass, flow, or heat) in a specific state that is different from the reference state; for example, the exergy of a hot fluid can be defined as the expansion work of the working fluid that circulates in a reversible Carnot heat engine, if the reference state is taken as a heat sink that receives the heat isothermally. The specific exergy can be calculated using the equation u2 + gz (1) 2 where e is the specific exergy of a certain flow, h is the enthalpy, T is the temperature, s is the entropy, u is the velocity, g is the gravitational constant, and z is the height. The subscript 0 indicates the reference state. The first two terms on the righthand side are thermodynamic-state-related, so these are defined hereafter as the thermal exergy; ech is the chemical exergy; and the remaining terms are the kinetic and potential exergies, respectively. As in the cases of mass and energy, an exergy balance equation can be built for an arbitrary control volume e = (h − h0) − T0(s − s0) + ech +

E i − Eo + EQ − Wsh − I = 0

W Qh

Eout E in

(3)

(4)

This is the ratio of the output exergy to the input exergy. This efficiency can be used to evaluate how efficiently the process is exploiting resources. This concept can be applied to a wider range of processes. However, because it uses lumped exergies, it is not possible to evaluate the subprocesses or unit operations, only the overall process. Another limitation of using the lumped exergy method is that, because it does not provide knowledge of the component exergies, one does not know how much of each component exergy is involved in the process. The rational exergetic efficiency has also been suggested.13 The definition is ψ=

Euseful I+L =1− Eused Eused

ech = ech

(7)

u2 2 ep = gz

(8) (9)

where ech, the chemical exergy, is obtained by calculating the maximum work produced when the considered substance is brought from the environmental state to the reference substances by a reversible reaction. In other words, the sources and sinks of certain types of exergy can be identified, as well as the unit in which the exergy is transferred. Moreover, when different types of exergy are involved in a process, for example, a heat pump or heat engine, the consumption, recovery, or transfer of the exergy in each subprocess can be identified in detail. The source of exergy loss inside the process can therefore be diagnosed component-bycomponent and retrofitted to give an improved version of the process. Finally, using the exergy decomposition method, the relative sizes of all exergy values can be compared; therefore, the recovery priority for the retrofit design can be determined directly. 2.3. LNG Exergy. The LNG exergy is decomposed into two terms, depending on the reference state: thermal exergy and chemical exergy. The kinetic and potential energies are neglected, because they are very small compared with the others. First, when the reference state is regarded as NG at ambient temperature and atmospheric pressure, LNG in the cold state can do work by configuring a reversible heat engine. Next, based on the original purpose of the LNG, the fuel exergy should be taken into account; specifically, the LNG chemical exergy is related to the heating value that NG essentially holds

(2)

This quantity tells how much work (W) one can obtain from the heat (QH) put into the system. Although it provides some intuition, it cannot be directly applied to non-heat-engine processes such as LNG regasification. In LNG regasification, our main focus is not use of the heat source, but exploitation of the cold LNG as a high-quality heat sink. Next, the conventional exergetic efficiency, also known as the second-law efficiency, can be defined as ηII =

(6)

ek =

Equation 2 is the steady-state exergy balance equation, where Ei is the inflow exergy; Eo is the outflow exergy; EQ is the exergy transferred by heat; Wsh is the shaft work done by the system; and I is the exergy loss, the irreversibility due to entropy generation. There are several criteria for the evaluation of a thermodynamic process. First, the simplest and most widely used efficiency is the thermal efficiency, which can be defined as ηth =

eth = (h − h0) − T0(s − s0)

ech = φ × LHV

(10)

where ϕ is an empirical parameter with a typical value of approximately 1.04 for LNG, and LHV is the lower heating value, also known as the net combustion value. The reference state of the chemical exergy is the products of the complete combustion process, CO2 and H2O. The nature of the LNG chemical exergy is the energy stored in the form of atomic bonds between carbon and hydrogen. The physical meaning of this is that the stored chemical exergy can be extracted by combustion of the fuel. The exergy values of all of the streams involved in the LNG regasification process were obtained using the exergy decomposition method. In particular, by focusing on the LNG in the cold state, which is the main source of thermal

(5)

where Euseful is the exergy of the desired output, Eused is the exergy consumed, I is the internal exergy losses due to the process irreversibilities, and L is the external exergy losses 14367

dx.doi.org/10.1021/ie501583m | Ind. Eng. Chem. Res. 2014, 53, 14366−14376


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Figure 1. Schematic diagram of the conventional LNG regasification process (based on a figure in ref 13).

exergy, and the boil-off gas (BOG), which holds a large fraction of the chemical exergy, every streamflow and unit was evaluated. Equation 2 was used to solve the exergy balance equation in every unit and subprocess in the LNG regasification processes to assess them in terms of exergetic efficiency.

Table 1. Exergy Values of Representative Streams in a Conventional LNG Regasification Process

3. BASE-CASE MODELING 3.1. Model Basis. First, a conventional LNG regasification process model was constructed, based on the work of Park et al.14 The regasification concept used in their study was based on the existing Pyeongtaek LNG receiving terminal in South Korea, as shown in Figure 1.15−17 The model developed in this work, however, is slightly different from that presented in the literature, in that an additional BOG flaring system and a detailed seawater vaporizer design were considered. In addition, different feed stream conditions and product specifications (i.e., temperature, pressure, and composition) were used. Consequently, the steady-state simulation results differ from those of the reference model. The assumption of a minimum LNG discharge flow rate and the corresponding amount of BOG generation was retained. The amount of flaring BOG was determined by referring to the data for the Incheon LNG-receiving terminal, reported in the literature,18 and was calculated from the product of the ratio of the LNG discharge flow rate for the Pyeongtaek plant and that for the Incheon plant and the amount of flaring BOG generated from the storage tank in the Incheon plant. The operating conditions of major units such as pumps, compressors, and the recondenser were taken from the literature. Moreover, the isentropic efficiencies of each pump and compressor were assumed to be 0.85 and 0.75, respectively. Through the entire simulation, the commercial process simulator Aspen Plus V7.3 was used, and the Peng−Robinson equation of state19 was used to describe the thermodynamic behavior and phase equilibrium of the NG. The simulation results for particular streams are presented in Table S1 of the Supporting Information. 3.2. Exergy Analyses. Based on the simulation results, stream-by-stream and unit-by-unit exergy analyses were conducted. In particular, the reference state of the thermal exergy was determined as 25 °C and 1 bar, and that of the chemical exergy was taken from Kotas’ work.13 The main exergy flows in the process are presented in Table 1. The chemical exergies of LNG streams 14 and 15 are the actual LNG fuel values that will be used in the downstream consumption region, which means that they have virtually nothing to do with exergy use in the retrofit design. The exergy

a

stream numbera

thermal exergy (MW)

chemical exergy (MW)

2 14 15 18 19

0.73 55.01 38.29 0.09 1.11

544.74 3068.53 3068.53 − −

Stream numbers correspond to those in Figure 1.

calculation showed that the chemical exergy of the BOG discarded to the flaring stack dominated the thermal exergy held by the LNG. In other words, the fuel value of NG is 10− 100 times higher than the LNG cold exergy. After regasification, the amount of thermal exergy held by the NG (stream 15) is still high, as the downstream pressure demand is quite high. In other words, for a case such as the South Korean NG distribution system, where the distance between the storage tank and the consumer is large, high kinetic exergy is needed for the NG to flow through the pipeline network, which, in turn, means that a high NG pressure is needed after regasification. This analysis suggests that one can extract at most ∼17 MW of LNG cold exergy, which is about 1/30th of the chemical exergy of flaring BOG. Finally, the absolute amount of exergy transferred from the LNG to seawater in the open rack vaporizer should be noted. LNG loses 17 MW of work potential, but the exergy increase of the seawater is only 1 MW. This means that the exergy loss term in the exergy balance equation (I in eq 2] is approximately 16 MW). The reason for this large loss is intrinsic entropy generation in the heat-transfer process. As can be seen in Figure 2, of all of the unit operations, LNG regasification has the largest irreversibility. To prevent such internal heat-exchange irreversibility, the two composite curves must meet as closely as possible, which can be achieved by selecting a suitable working fluid. The conclusions from the analyses can be summarized as follows: • The dominant exergy loss in the conventional LNG regasification process is caused by BOG flaring. Therefore, the recovery of the chemical (fuel) exergy of flaring BOG takes precedence over all other types of exergy recovery. 14368



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(dotted red box), and at the same time, the LNG cold exergy is recovered through a secondary closed Brayton cycle working with helium (dash-dotted blue box). The conditions for the rest of the unit, such as the BOG compressors and recondenser, were set equal to those of the base-case model. The steady-state simulation results are presented in Table S2 of the Supporting Information. 4.1. Model Basis. 4.1.1. Open Brayton Cycle. In this subprocess (see Figure 4, where the numbers in the graph

Figure 2. Exergy loss for each unit operation in the conventional LNG regasification process.

• Maximum use of the LNG thermal (cold) exergy depends on the downstream NG pressure demands. The higher the requirement is, the less the exergy gain is. • Of all the unit processes, the exergy loss in the open-rack heat exchanger is the largest. However, through appropriate working fluid selection, entropy generation can be minimized; thereby, thermal exergy use can be maximized.

Figure 4. Temperature−entropy diagram for the primary open Brayton cycle.

4. ALTERNATIVE 1: DUAL BRAYTON CYCLE An alternative process for the recovery of the chemical exergy of flaring BOG was modeled based on literature reports (Figure 3).10,11 In this process, the flaring BOG originally discarded to the stack is introduced into the primary open Brayton cycle

correspond to the stream numbers in Figure 3), air is first compressed (1 → 3) and then cooled by an intercooler (3 → 4), after which it is subjected to further pressurization (4 → 5). It is then combusted together with compressed flaring BOG in the combustor, resulting in a high-temperature, high-pressure

Figure 3. Schematic diagram of an alternative process for LNG regasification using a dual Brayton cycle (based on a figure in ref 10). 14369



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gas (5 → 8). This gas subsequently expands through the gas turbine (GT1), producing a certain amount of electricity (8 → 9). Low-pressure exhaust gas from the turbine is then discarded to the environment after rejecting part of its thermal exergy to the helium circulating in the closed Brayton cycle (9 → 10). Al-Doori’s work20 was consulted to determine the operating parameters in the cycle. For GT1, it was assumed that the total pressure ratio was 20, the air-to-fuel ratio was 35 (mass-based), and the intercooling effectiveness was 0.94. The pressure of the exhaust gas outlet stream was determined to be 1 bar, and no pressure drop was assumed throughout the heat exchanger. The isentropic efficiencies of all of the compressors and the turbine were assumed to be 0.75. 4.1.2. Closed Brayton Cycle Working with Helium. The purpose of this secondary Brayton cycle is to use both the hot thermal exergy from the exhaust gas and the cold exergy from the pressurized LNG. The cycle was configured and the operating parameters were determined based on Dispenza et al.’s work.10 First (see Figure 5, where numbers in the graph correspond to the stream numbers in Figure 3), the pressurized working

crossover occurred. Over the entire feasible range, the maximum flow rate of the working fluid was chosen to recover as much as possible of the corresponding exergy. A step size of 1 (t/h) was used for convenience, and the number of manual adjustments did not exceed 20. 4.2. Exergy Analyses. Exergy analyses were performed on each of the two cycles, and representative results are presented in Tables 2 and 3. Table 2. Results of the Exergy Analysis for the Primary Open Brayton Cycle stream or unita

thermal exergy (MW)


2 LP compressor HP compressor NG compressor GT1

0.73 −

544.74 −

− 81.80

−

−

85.25

−

−

5.81

−

−

−302.02b

a b

pure exergy (electricity) (MW)

Stream number and unit names correspond to those in Figure 3. Negative sign indicates that this unit recovers exergy.

Table 3. Results of the Exergy Analysis for the Secondary Closed Brayton Cycle stream or unita 9 10 helium compressor 14 15 GT2 a b

Figure 5. Temperature−entropy diagram for the secondary closed Brayton cycle, working with helium.

thermal exergy (MW)



− − −

156.24 126.41 −

− − 36.91

53.34 37.12 −

2974.84 2974.84 −

− − −44.21b

Stream numbers and unit names correspond to those in Figure 3. Negative sign indicates that this unit recovers exergy.

First, the results of analysis of the primary Brayton cycle show that a total of 718.33 MW of exergy is used, and 302.02 MW of it is recovered by GT1, resulting in a rational exergetic efficiency (ψ) of 42%. It should be noted that the chemical exergy of stream number 2 (544.74 MW) accounts for approximately 75.8% of the total consumed exergy. Next, as shown by the data in Table 3, through the secondary closed Brayton cycle, a total of 82.96 MW of the exergy is transferred to the helium, and 44.21 MW is recovered. The rational exergetic efficiency of this subprocess (ψ) is therefore 53%. The working fluid receives 29.83 MW from the exhaust gas in the upper HX, 16.22 MW from LNG in the LNG regasifier, and 36.91 MW in the helium compressor. Note that the exergy delivered by the cold LNG (16.22 MW) accounts for only 19.6% of the total exergy used. There are two problems in the secondary helium cycle. The first is high energy consumption for working fluid compression. The helium compressor, in which the helium pressure increases from 3 to 28.2 bar, consumes approximately 37 MW, which is 83% of the shaft work produced in GT2. Because of this large amount of compression work, the net exergy recovery is poor, resulting in a low rational exergy efficiency (ψ) for the cycle. However, this can be solved by using the ORC instead of the Brayton cycle. In the ORC, the working fluid undergoes a phase change from vapor to liquid state through the condenser

fluid, namely, helium, absorbs heat from the exhaust gas (11 → 12) and expands through GT2, where electricity is produced (12 → 13). Subsequently, it rejects heat to the LNG through the regasifier, so that cold exergy is recovered from the LNG and transferred to the helium (13 → 16). Low-pressure, lowtemperature helium is then pressurized by the compressor (16 → 11), and that completes the heat engine cycle. As stated in the literature,10 we set the outlet pressure of GT2 at 3 bar, the pressure ratio of the helium compressor at 9.4, and the minimum temperature differences at the pinch point of the upper HX and the LNG regasifier at 34 and 31 K, respectively. The NG outlet stream temperature was set at 5 °C. This was achieved by varying the LNG inlet stream’s mass flow rate. A numerical solution of 218.4 t/h was obtained using the secant root-finding algorithm. Based on the total 230 t/h extraction of LNG in the base-case model, 95% of it was regasified in the secondary closed Brayton cycle. After the determination of all other operating conditions, the mass flow rate of helium (72 t/h) was determined. This was done manually, by checking the composite curves of each heat exchanger. In every step-by-step adjustment of the variable, we observed whether or physical violations such as temperature 14370



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The pressurized LNG is overpressurized using an LNG pump, and it then gives its thermal exergy to the working fluid circulating in the ORC loop. Vaporized NG then obtains heat from an external heat source, and then high-pressure, hightemperature NG releases its exergy through a gas turbine (GT2). In the ORC subprocess (see Figure 7, where numbers in the graph correspond to the stream numbers in Figure 6), the working fluid changes its state from gas phase to liquid phase in the LNG regasifier (1 → 2). The liquefied fluid is pressurized using a propane pump (2 → 3) and then absorbs heat from an external heat source (3 → 4). The high-pressure, hightemperature working fluid then expands through the gas turbine (GT1), producing electricity (4 → 1). As described in the literature,8 pure propane was used as the working fluid in the ORC, and the temperature of the external heat source was assumed to be 90 °C. A low-grade heat source of infinite capacity and a minimum-temperature of 10 °C were also assumed. We arbitrarily set the discharge pressure of the LNG pump at 90 bar; this will have to be verified in terms of process optimization (not done in this study). The pressure of the directly expanded NG (stream 15) was set at 74.53 bar, the same value as the demand pressure in the base-case model. Next, the ORC operating conditions were determined as follows. Three variables, namely, the inlet/outlet pressures of GT1 and the propane flow rate, were determined so as to maximize exergy recovery through GT1. Second, the condition of the cold stream flowing to the LNG regasifier (stream 5) was calculated by LNG pump simulation (−120.26 °C, 90 bar), so the lower bound of the discharge pressure of GT1 (p1) was specified as 0.0113 bar, which is the vapor pressure at −110.26 °C. Finally, a feasible value of p1 was sought, together with the flow rate of propane. The composite curve of the LNG regasifier was checked by manually adjusting both variables, to determine whether a physically infeasible situation occurs (see Figure 8). If p1 is too low, the hot composite curve decreases downward, violating the pinch point; if the propane flow rate is too high, both the sensible-heat region of the vapor phase and the latent heat of the vaporization region (horizontal line of the hot curve) overwhelm the liquid-phase sensible-heat region, resulting in infeasible operation of the propane pump. In contrast, when the flow rate is too low, the latent-heat region shortens horizontally, and at the same time, the two sensibleheat regions become steeper vertically, resulting in temperature crossover. With several steps of manual adjustment, the feasible values of p1 and the flow rate were determined to be 1.5 bar and 161.46 t/h, respectively. The steady-state simulation results are presented in Table S3 of the Supporting Information. 5.2. Exergy Analyses. Exergy analyses were performed on the ORC subprocess, and the results are presented in Table 5. A total of 62.70 MW of exergy entered the entire subprocess. However, as mentioned before, because an infinite-capacity, free, low-grade heat source is given, 46.75 MW of exergy delivered by the heat source can be neglected. Note that the main source of the remaining 15.95 MW of exergy is the thermal exergy transferred from the cold LNG to the ORC loop. Approximately 37% of the thermal exergy introduced into the system was recovered from the two turbines. In other words, the rational exergetic efficiency (ψ) of this subprocess is 0.37. From GT1, 4.59 MW of exergy was recovered by expansion work of the propane working fluid, and from GT2, 1.37 MW was obtained by direct expansion of overpressurized

(corresponding to the LNG regasifier unit in Figure 3), whereas throughout the Brayton cycle, it remains in the gas phase. Compared with the fluid of the gas phase, less work is required to increase the pressure of the liquid flow

δW = V (βT − 1) dP

(11)

where β is the volume expansivity and is defined as β=

1 ⎛⎜ ∂V ⎞⎟ V ⎝ ∂T ⎠ p

A fluid with a high specific volume and high volume expansivity (e.g., a gas phase) requires more work, and vice versa. Therefore, if the ORC is used, the pressure increase of the liquid fluid is usually achieved using a pump rather than a compressor, resulting in a huge energy saving. The second problem is that not all the thermal exergy content of the exhaust gas is being used. As mentioned before, the exhaust gas (streams 9 and 10 in Figure 3) gives its exergy to the working fluid, i.e. helium (streams 11 and 12 in Figure 3). The data in Table 3 show that the amount of transferred exergy is 29.83 MW. The thermal exergy initially held by the exhaust gas is five times more than this, so the rest of the 126.41 MW is discarded to the environment. This unnecessary exergy waste is due to the ill-balanced energy quantities in the heat source and the heat sink. In other words, the maximum amount of usable LNG cold exergy (16.22 MW) is far less than that of the hot exergy of the exhaust gas (156.24 MW). An additional heat sink such as cooling water should be considered to balance the exergy inflow and outflow of the heat engine, thereby increasing the overall exergy efficiency of the cycle. The overall process efficiency is summarized in Table 4, together with those of alternative processes. The rational Table 4. Rational Exergetic Efficiencies of the Alternative Processes for LNG Regasification process

ψ

base case dual Brayton cycle organic Rankine cycle gas-to-liquid process

0.847 0.890 0.849 0.868

exergetic efficiency of this dual Brayton cycle scheme is 0.890, which is a 4.3% point improvement on the base-case efficiency. This is mainly the result of NG fuel value recovery through the primary open Brayton cycle. The thermal exergy of the LNG, although the amount is small compared to that of the chemical exergy of the BOG, is also recovered throughout the secondary helium Brayton cycle.

5. ALTERNATIVE 2: ORC WITH LOW-GRADE HEAT SOURCE 5.1. Model Basis. An alternative process for the recovery of the thermal exergy of cold LNG was modeled based on literature reports (Figure 6).8 The stream numbers in parentheses correspond to the stream numbers in Figure 1. In this alternative model, the LNG regasification part was modified, but all other parts of the base-case model were the same. In the base-case model, the thermal exergy transferred from the LNG was being discarded to seawater through a heat exchanger, whereas in this retrofit model, it is used in the ORC configuration. 14371



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Figure 6. Schematic diagram of an alternative process for LNG regasification using an organic Rankine cycle (based on a diagram in ref 8).

most of the exergy loss occurs in the LNG regasifier. In other words, a vast amount of exergy is consumed there, which is 47% of the LNG thermal exergy transferred. This is because of the intrinsic irreversibility of the heat-transfer process, which can be minimized by selection of an appropriate working fluid. Pure propane might therefore not be the best candidate for the working fluid in this system. By mixing different types of refrigerant in an optimized composition, the composite curves of both the refrigerant and LNG can be adjusted to minimize the irreversibility in the heat exchanger.

6. ALTERNATIVE 3: PROCESS INTEGRATION WITH A GTL PROCESS 6.1. Design Philosophy. Recently, the GTL process has been highlighted as a promising process, because it produces environmentally friendly diesel fuel from abundant NG.21−23 In this process, NG is first converted to syngas, using the reforming reaction, and the syngas is converted to the final product, namely, Fischer−Tropsch (FT) liquid fuel. Many commercial on-shore GTL facilities have been constructed and operated. The design philosophy behind the integration of the conventional LNG regasification process with GTL is to save the large quantity of chemical exergy lost through BOG flaring by excess LNG extraction and, at the same time, to convert it to a high-value-added, environmentally benign product. Moreover, the thermal exergy of the cold LNG can be recovered together with high-quality heat from the reformer and FT reactor, resulting in high thermal efficiency of the thermodynamic cycle (Figure 10). The details of the idea are as follows: First, as can be seen in Figure 11, monthly LNG demand varies greatly. Therefore, for a certain period of time, when LNG demand is low, surplus LNG is stored in a storage tank. The amount of BOG is proportional to the mass of LNG in the storage tank,24 so a large quantity of BOG is generated during

Figure 7. Temperature−entropy diagram for the organic Rankine cycle, working with propane.

LNG. There is a trade-off between the two shaft works, because if the degree of overpressurization of LNG by the LNG pump increases, then the amount of exergy recovered from GT2 will increase. However, conversely, that from GT1 will decrease because, as the boiling point of the LNG increases, the discharge pressure of GT1 should be adjusted to a higher value. The rational exergetic efficiency (ψ) of the entire process is 0.849, which is a 0.2% point improvement on that of the basecase model (Table 4). This is because the absolute amount of usable thermal exergy of the LNG is very low compared to the chemical exergy loss from BOG flaring. In this alternative process, only the former was recovered, and the latter was discarded in the same way as in the base-case model. Finally, it should be noted that a large portion of valuable exergy is lost in the form of entropy generation. Figure 9 shows results from the exergy analysis of every unit. It can be seen that 14372



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Figure 8. Composite curves of an LNG regasifier for various scenarios: cold stream, pressurized LNG heated to the gaseous state; base case, finally determined operating conditions of working fluid (p1 = 1.5 bar, flow rate = 161.46 t/h); scenario 1, low discharge pressure of working fluid from gas turbine (p1 = 0.1 bar, flow rate = 161.46 t/h); scenario 2, high mass flow rate of working fluid (p1 = 1.5 bar, flow rate = 216.00 t/h); scenario 3, low mass flow rate of working fluid (p1 = 1.5 bar, flow rate = 144.00 t/h). (This graph corresponds to the analytical results reported in the last paragraph of section 5.1 in this article.)

Table 5. Results of the Exergy Analysis for the Organic Rankine Cycle stream or unita

thermal exergy (MW)

heat source 1 5 6 LNG pump propane pump GT1 GT2 heat source 2

31.64

−

−

55.17 39.76 − −

3068.53 3068.53 − −

− − 0.30 0.24

− − 15.11

− − −

−4.59b −1.37b −

a



Stream numbers and unit names correspond to those in Figure 6. Negative sign indicates that this unit recovers exergy.

Figure 9. Exergy loss of each unit operation in the alternative process for LNG regasification process using an organic Rankine cycle.

this period. The aim is to reduce the quantity directly by discharging the excess LNG, thereby preventing the waste of chemical exergy through BOG flaring. Furthermore, the exergy content of the excess LNG is recovered through the GTL process. So, because GTL can be seen as a process of carbon chain formation from an NG feedstock, it is theoretically possible that chemical exergy held by the LNG can be preserved through this process. The only difference between NG and the GTL product (FT fuel) is in their specific heating values, and the difference is not large.25 Finally, the conventional GTL process loses exergy mainly from the reformer and FT reactor, because they are highly endothermal or exothermal units. In general, the exergy is recovered using heat-exchange networks or combined heat and power generation. Therefore, in this alternative process model,

both schemes are used to improve the exergetic efficiency. Our design, however, might have greater efficiency, because we have an additional high-quality heat sink, namely, cold LNG. In other words, LNG in the cold state is used as a heat sink, and at the same time, it delivers its thermal exergy through the regasification process. 6.2. Model Basis. The study by Panahi et al.22 was used as the basis for determining the subprocess structure and operating conditions for each unit of the GTL process. As in that study, an autothermal reformer (ATR) and prereformer were used for syngas production. Pure oxygen was assumed to be provided from an external air-separation unit (ASU) and fed directly into the ATR. It is also assumed that superheated steam is produced by heat exchange with the hot outlet stream of the ATR and fed into the prereformer together with the NG feed

b

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Figure 10. Schematic diagram of the process integrating LNG regasification with a gas-to-liquid process. Stream numbers in parentheses in the lower part of this figure correspond to stream numbers in the upper part of this figure.

the target syngas composition (i.e., a H2/CO ratio of 2), but the 3% purge split ratio was retained. Adiabatic operation of the prereformer and ATR and isothermal operation of the FT reactor at 210 °C were assumed. The pressures of the reformer and FT reactor were 30 and 20 bar, respectively, as suggested in the reference model.22 In the first Rankine cycle, for hot exergy recovery from the ATR, water was used as the working fluid, whereas in the second ORC, for use of the hot exergy of the FT tail gas and cold exergy of the LNG, propane was used. The operating conditions for both cycles were determined step-by-step, following the same order as used in alternative model 2. First, the maximum evaporation pressure was determined from the temperature of the heat source, and then the lower bound of the minimum condensing pressure was obtained from the temperature of the heat sink. For maximum expansion work recovery from the turbine unit, both the discharge pressure and working fluid flow rate were calculated by checking the composite curves of the heat exchangers. 6.3. Exergy Analyses. Exergy analyses were performed on the GTL subprocess; the results are presented in Table 6. A total 836.90 MW of exergy entered the GTL subprocess (neglecting the chemical exergy of the LNG), and 618.72 MW was recovered. The main sources of input thermal exergy were excess LNG from the GTL feed, cold LNG going into the regasification unit, and heat duty for the heaters; those of the input chemical exergy account for the fuel value of the excess LNG. Some exergy was recovered by direct expansion of pressurized NG at GT1. A large portion of the exergy was recovered in the form of synthetic FT fuel. The exergy was also additionally recovered through GT2 and GT3 in the ORC subprocesses, unless otherwise wasted as exothermal heat of reaction; their thermal efficiencies (ηth) are 60% and 32%, respectively. The overall rational exergetic efficiency (ψ) increased by 2.1% for two reasons. First, the amount of BOG generated decreased, because of excess LNG extraction, which reduced

Figure 11. Monthly variations (2009) in the natural gas demand in South Korea.

and recycled gas. A fired heater was simulated using four independent heater models, and component separator models were used to simulate CO2 and H2O separations, with 95% separation efficiencies. Two parallel Rankine cycles were configured for the recovery of the thermal exergies of the hot syngas product from the ATR, hot FT tail gas from the FT reactor, and cold LNG from the LNG regasification process. The molar flow rate of excess LNG was assumed to be 50 t/ h, one-quarter of the original discharge flow rate of the basecase model. Based on the boil-off rate calculation from a model in the literature,24 total BOG generation was reduced by approximately 27.4%. The flow rates of the other feed streams, namely, pure oxygen and superheated steam, were simply determined using the same ratio of these and the NG feed stated in the reference model. The recycled gas flow rate to the prereformer is different from that in the reference model, because the split ratio (SP2 in Figure 10) was adjusted to meet 14374



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processes using dual Brayton cycles (alternative 1), the ORC (alternative 2), and integration with GTL (alternative 3), are 0.847, 0.890, 0.849, and 0.868, respectively. Other factors such as capital costs, layout constraints, and controllability cannot be evaluated using the proposed model. Because some literature results were directly used in the models, further analyses such as optimization of subprocesses and sensitivity responses of several important design variables need to be conducted in the future. It is expected, however, that using the efficiency criteria and procedures described in this article, more precise, accurate, and specific evaluation methods will be developed not only for LNG regasification but also for other processes involving thermal or chemical exergy use.

Table 6. Results of the Exergy Analysis for the Gas-to-Liquid Subprocess stream or unita

thermal exergy (MW)

66 20 29 CMP1 FH1 FH2 FH3 FH4 HT4 HT3

54.77 12.21 24.23 − 19.75 16.00 2.99 2.12 1.34 21.97

GT1 GT2 GT3 67 47 49

− − − 38.12 0.05 0.34

chemical exergy (MW) Exergy In 3054.68 680.93 − − − − − − − − Exergy Out − − − 3054.68 497.31 39.30

pure exergy (electricity) (MW) − − − 0.59 − − − − − −

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ASSOCIATED CONTENT

S Supporting Information *

Additional information as noted in text. This material is available free of charge via the Internet at http://pubs.acs.org.

−1.02b −34.20b −8.38b − − −

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AUTHOR INFORMATION

Corresponding Author

*Tel.: 82-2-880-1887. E-mail: [email protected].

a

Stream numbers and unit names correspond to those in Figure 10. b Negative sign indicates that this unit recovers exergy.

Notes

the exergy loss through BOG flaring. Second, as mentioned before, additional exergy recovery is achieved by the direct expansion work of the NG and the shaft work of the working fluid in the two Rankine cycles. There is therefore a trade-off between BOG loss and LNG exploitation. If more LNG is converted to FT fuel, less BOG is lost, but more exergy is lost, because of process irreversibility. Although GTL integration increases the overall rational exergetic efficiency (ψ), the effect is smaller than in the dual Brayton cycle case. This is because, despite the reduction in flaring BOG losses, the absolute amount of chemical exergy is still large; this exergy should therefore be recovered using an additional Brayton cycle to improve the overall efficiency, which is beyond the scope of this study.

ACKNOWLEDGMENTS This research was supported by the second phase of the Brain Korea 21 Program in 2014; the Institute of Chemical Processes of Seoul National University; an MKE; and a grant from the LNG Plant R&D Center funded by the Ministry of Land, Transportation and Maritime Affairs (MLTM) of the Korean government. The Energy Efficiency & Resources Core Technology Program of the Korea Institute of Energy Technology Evaluation and Planning (KETEP) granted financial resources from the Ministry of Trade, Industry & Energy, Republic of Korea (Nos. 2010201020006D, 20132010201760, 20132010500050, and 2012T100201687).

The authors declare no competing financial interest.

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Abbreviations

7. CONCLUDING REMARKS In this study, exergy analyses were performed to evaluate several LNG regasification processes. The exergy decomposition method was used and applied to every stream and unit in the entire process. The proposed method makes it possible to see how efficiently a process uses the thermal and chemical exergies. If conventional lumped exergy is used, meaningful information such as the priorities for retrofit design and quantitative comparisons of many exergy sources and sinks cannot be obtained. The exergy analyses (Tables 4 and 7) show that the rational exergetic efficiencies of the four processes, namely, the conventional LNG regasification process (base case) and the

ASU = air-separation unit ATR = autothermal reformer BOG = boil-off gas FT = Fischer−Tropsch GTL = gas-to-liquid (process) LHV = lower heating value LNG = liquefied natural gas ORC = organic Rankine cycle NG = natural gas Parameters

ech = specific chemical exergy, MJ/(kg s) I = exergy loss in a system due to irreversibility, MJ/s L = external exergy losses related to effluent streams QH = heat absorbed by a system, MJ/s W = shaft work produced by a system, MJ/s β = volume expansivity of a fluid, K−1 ηII = second-law efficiency ηth = thermal efficiency ϕ = empirical parameter for calculation of chemical exergy of a fuel from its low heating value ψ = overall rational exergetic efficiency ψch = rational chemical exergetic efficiency

Table 7. Summary of the Exergy Balance Calculations for the Alternative LNG Regasification Processes process

exergy in (MW)

exergy out (MW)

base case dual Brayton cycle organic Rankine cycle gas-to-liquid process

3742 3952 3743 4276

3171 3518 3178 3712

NOMENCLATURE

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ψth = rational thermal exergetic efficiency

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(22) Panahi, M.; Rafiee, A.; Skogestad, S.; Hillestad, M. A Natural Gas to Liquids Process Model for Optimal Operation. Ind. Eng. Chem. Res. 2011, 51 (1), 425−433. (23) Panahi, M.; Skogestad, S. Selection of Controlled Variables for a Natural Gas to Liquids Process. Ind. Eng. Chem. Res. 2012, 51 (30), 10179−10190. (24) Shin, M. W.; Shin, D.; Choi, S. H.; Yoon, E. S. Optimal operation of the boil-off gas compression process using a boil-off rate model for LNG storage tanks. Korean J. Chem. Eng. 2008, 25 (1), 7− 12. (25) Boundy, R. G.; Diegel, S. W.; Wright, L. L.; Davis, S. C. Biomass Energy Data Book; Oak Ridge National Laboratory: Oak Ridge, TN, 2011.

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Comparative Study of Process Integration and Retrofit Design of a

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