ARTICLE pubs.acs.org/EF
Recovery of Flue Gas Energy in Heat-Integrated Gasification Combined Cycle (IGCC) Power Plants Using the Contact Economizer System Vhutshilo A. Madzivhandila,† Thokozani Majozi,*,†,‡ and Toshko K. Zhelev§ †
Department of Chemical Engineering, University of Pretoria, Pretoria 0002, South Africa Modelling and Digital Science, Council for Scientific and Industrial Research, Pretoria 0002, South Africa § Department of Chemical and Environmental Science, University of Limerick, Limerick, Ireland ‡
ABSTRACT: Recovery of low potential energy of flue gases, mainly from industrial boilers, has become one of the problems of interest in research. In this work, the contact economizer system is used to recover low potential heat from the gas turbine exhaust (flue gas) stream of a heat-integrated gasification combined cycle (IGCC) design of the Elcogas plant adopted from previous studies. The underlying support for this idea was the direct relationship between efficiency of the IGCC and the boiler feedwater temperature. Recovery of the flue gas heat to preheat the boiler feedwater was demonstrated to be capable of further increasing the thermal efficiency of the plant. The methodology developed is divided into two parts, i.e., determining the maximum boiler feedwater temperature attainable and applying Mickley’s graphical technique for dehumidification, following a slightly different procedure that allows for the calculation of the exact ratio between the liquid-phase heat-transfer coefficient and the gas-phase masstransfer coefficient, to demonstrate how the aforementioned temperature is achieved. The grand composite curve is used to check whether the determined boiler feedwater temperature is feasible. A case study on the Elcogas plant illustrated that the developed method is capable of increasing the gross efficiency from 54 to 55%. This increase in efficiency, however, has a penalty of operating at higher boiler and heat recovery steam generator (HRSG) pressures.
1. INTRODUCTION The integrated gasification combined cycle (IGCC) is one of the innovative electrical power generation systems that promises to provide a large share of the future world’s energy needs. This system combines two primary technologies: gasification and the combined cycle, made up of the gas turbine and the steam turbine. The combination of these two technologies divides the IGCC into two subsystems: the gas-side subsystem and the steam-side subsystem. These two subsystems are integrated by a gas cooler (often referred to as a boiler) and a heat recovery steam generator (HRSG) to form one system. The IGCC has proven to have several advantages over conventional coal power generation plants. The IGCC has higher thermal efficiencies and low emissions and uses a considerably low amount of water, as compared to conventional coal power generation plants. A lot of research on the IGCC is currently on the way, and already, certain improvements on various aspects of the process have managed to increase the thermal efficiency of the system. Most of the research conducted in this field has focused on improving or optimizing the performance and integration of the components of the gas-side subsystem, i.e., the coal gasification unit, the gas turbine, and the air separation unit, and their operating parameters.14 Only a few researchers have focused on improving the steam-side subsystem and optimizing the use of energy available within the system.58 If we consider the amount of energy available for use in this system, this should be one the areas of focus. The objective of this contribution is to recover low potential heat from the gas turbine exhaust stream of a heat-IGCC plant using a contact economizer system (CES) to further improve the overall efficiency of the IGCC. r 2011 American Chemical Society
The basis of this contribution was an interesting direct relationship between the boiler feedwater temperature and the thermal efficiency of the IGCC.
2. NECESSARY BACKGROUND 2.1. IGCC. It is assumed at this stage that the reader is familiar with the fundamental structure and concepts of the IGCC shown in Figure 1. The basic structure of the IGCC can viewed as a system made up of two major subsystem as previously mentioned. The gas-side subsystem is made up of the air separation unit (ASU), the gasifier, the gas cleanup section, and the gas turbine. The steam-side subsystem on the other hand in made up the boiler, the HRSG, and the steam turbine. Integration between the two subsystems can be seen at the boiler and the HRSG. An interesting integration within the gas-side subsystem exists. The gas turbine provides a portion of its compressed air to the ASU, which in turn supplies the oxygen to the coal gasification unit (gasifier). The nitrogen produced in the ASU is redirected back to the gas turbine to reduce NOx formation and increase the gas turbine output. A portion of the nitrogen is also used to pressurize the coal in the coal gasification unit. Proper integration of these units was therefore a convenient way to improve the efficiency of the IGCC. Received: January 9, 2011 Revised: March 10, 2011 Published: March 11, 2011 1529
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Figure 1. Basic flow sheet of an IGCC plant.
Figure 2. Heat-integrated design of the Elcogas plant.
The stream of concern for this study is the gas turbine (GT) exhaust stream leaving the HRSG en route to stack, as shown in Figure 1. More details on this core stream will be given in the following sections. 2.2. Pinch Analysis. Pinch analysis techniques were recently adopted in a study to maximize the use of energy within IGCC plants.9 This study involved determining the maximum energy available in an IGCC plant (Elcogas plant) using the GCC, targeting the amount of steam that can be produced using the maximum available energy under the plant conditions, and last, constructing a design that meets the aforementioned target. The final design for this study is given in Figure 2. 2.3. CES. The CES is a low potential heat recovery system allotted to explore the simultaneous management of heat and mass transfer between a gas stream and a liquid desiccant stream. Figure 3 represents a typical CES for the recovery of heat from a gas stream using water as a desiccant. This process involves direct heat transfer between the hot gas stream and a cold circulating water stream accompanied by dehumidification of the gas, in a packed-bed column.
The heated desiccant that leaves at the bottom of the packed-bed column can then be used as a source of heat for other operations. The literature has shown that the circulating water can only be heated to a point where its temperature equals the wet-bulb temperature of the gas, provided that an infinite heat-exchange area is available. At this point, fog formation is a possibility.10 Consequently, heat transfer within the column should be maintained in such a way that the circulating water leaves the column at a certain temperature “ΔT” degrees lower than the wet-bulb temperature of the gas at all times. This allows for thermal driving forces and also ascertains that fogging conditions will not be attained in the tower. “ΔT” values as low as 2.5 °C are feasible for packed-bed columns and have been previously demonstrated in the literature.11 A brief derivation of the design equations for such dehumidification systems is given below, with a large portion of their development presented elsewhere.12 Consider a differential height “dZ” of a forced draft, counter-current, adiabatic, constant cross-section dehumidification tower, in which a gas and a liquid desiccant are directly contacted, as shown in Figure 4. 1530
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Figure 3. Basic structure of the CES.
Heat transfer during such a dehumidification process can be viewed as a two-path process, with the first path being from the bulk gas phase to the interface (referred to as the gas-phase path) and the second path being from the interface to the bulk liquid phase (referred to as the liquid-phase path). Division of the heat transfer into two separate paths is necessary because of the fact that, for the liquid-phase path, heat transfer is entirely a result of temperature potential, whereas two mechanisms are involved in the gas-phase path heat transfer. For the gas-phase path, heat is transferred as a result of both the temperature potential and the mass transfer (latent heat). The enthalpy of the humid gas is defined by eq 1. HG ¼ cs ðTG T0 Þ þ λ0 Y
ð1Þ
The expression for the change in the gas-phase path enthalpy in terms of the temperature is given by eq 2. GdHG ¼ Gd½cs ðTG T0 Þ þ λ0 Y ¼ Gcs dTG þ Gλ0 dY ð2Þ The two terms on the right-hand side of eq 2 represent sensible heat (temperature potential) transfer and the latent heat transfer, respectively. These two terms, i.e., sensible heat transfer and the latent heat transfer, can be further represented by eqs 3 and 4, respectively.
Figure 4. Schematic diagram of a forced draft dehumidification tower.
The gas enters the differential section of the column at a mass rate of “G”, a bulk temperature of “TG þ dTG”, enthalpy of “HG þ dHG”, and humidity of “Y þ dY”. The liquid desiccant on the other hand enters the differential section at a mass rate of “L” and a bulk temperature of “TL”. In the differential section, the gas stream and the liquid stream exchange heat and mass and then come out slightly changed. It is assumed that, at the gasliquid interface, the gas is saturated at the interface temperature “Ti”.
Gcs dTG ¼ hG aH ðTG Tt ÞdZ
ð3Þ
Gλ0 dY ¼ λo kG aM ðYG Yt ÞdZ
ð4Þ
For the liquid-phase path, eq 5 is the expression for the change in enthalpy in terms of the temperature. LcL dTL ¼ hL aH ðTL Ti ÞdZ = GdHG
ð5Þ
Upon integration, eq 5 yields eq 6, one of the basic equations used in determining the tower height, as will be explained later. Z HG2 dHG h L aH Z ð6Þ ¼ G HG1 ðTL Tt Þ 1531
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The design equations can now be derived by combining the equations above. A combination of eqs 24 gives eq 7, the basis of the first design equation. GdHG ¼ hG aH ðTG Tt ÞdZ λ0 kG aM ðY Yt ÞdZ
ð7Þ
The Lewis relation, given by eq 8, can be introduced into eq 7 assuming that the area for mass transfer (aM) is equal to the area for heat transfer (aH). This approximate relation, i.e., eq 8, has proven to be accurate for airwater systems. Substituting eq 8 into eq 7 leads to eq 9. hG ¼ cs kG
ð8Þ
GdHG ¼ kG aM ½cs ðTG Tt Þ þ λ0 ðY Yt ÞdZ
ð9Þ
Following the definition of a humid gas given by eq 1, eq 9 can be reduced to eq 10. GdHG ¼ kG aM ðHG Ht ÞdZ
ð10Þ
Equation 10 can then be combined with eq 5, resulting in our first design equation, i.e., the “tie line”, given by eq 11. This tie line represents the ratio of the relative rates of enthalpy transfer through the gas phase and the liquid desiccant phase. HG Ht h L aH ¼ TL Tt kG aM
ð11Þ
An enthalpy balance applied to the two combined phases yields eq 12. dðGHG Þ ¼ d½LcL ðTL T0 Þ
ð12Þ
Normally, the change in the liquid rate because of condensation or evaporation would be negligible. This simplifies eq 12 to eq 13. GdHG ¼ LcL dTL
ð13Þ
Upon integration, eq 13 yields our second design equation, i.e., the “operating line”, given by eq 14. This operating line is a straight line connecting the gas enthalpy and the desiccant temperature. ðHG2 HG1 Þ ¼
LcL ðTL2 TL1 Þ G
ð14Þ
A summary of assumptions under which these design equations operate is as follows: (i) At the watergas interface, the gas is saturated at the interface temperature (Ti). (ii) The change in the water flow rate because of evaporation or condensation is negligible. (iii) The heat-transfer area (aH) is equal to the mass-transfer area (aM). This assumption holds if and only if the interfacial area of packing inside the tower is fully wetted. A set back in the design of such dehumidification or humidification systems is the lack of information on the values of the masstransfer coefficient (kGaM) and heat-transfer coefficient (hLaH). Published values of such coefficients are extremely insufficient to cover a range of design problems. The majority of the available rate coefficient data or correlations is in the form of overall transfer coefficients. These overall transfer data or correlations would be exact if the liquid-film heat-transfer coefficient is infinitely large or if the equilibrium curve of enthalpy versus temperature is linear. These conditions are rarely met in practice, and although the use of an overall enthalpy coefficient to determine the tower/column volume is usually satisfactory, application of an overall driving force could lead to erroneous results. This is mainly the case of situations
Figure 5. Enthalpytemperature diagram indicating the dehumidification process.
significantly different from the experimental conditions used to obtain the coefficients. 2.4. Mickley’s Graphical Technique. Mickley presented a simple and improved graphical method for the design of forced draft air-conditioning equipment.10 This method is an extension of the “enthalpy potential” method proposed by Merkel.13 Mickley’s graphical method is still recognized as the most convenient method for determining the size of the equipment for direct contact systems. All operating conditions of the equipment can be quickly determined by this method, and the danger of fog formation can be ascertained. Figure 5 represents typical results of Mickley’s graphical technique applied to a dehumidification process to help explain the method. Curve “EF” in this figure represents the equilibrium line or saturation curve constructed under the assumption that the gas at the liquidgas interface is saturated. Curve “AB” is the operating line constructed from eq 14, with point “A” corresponding to the entering bulk-gas enthalpy and the leaving bulk-water temperature. The slope of this line is LcL/G, as indicated in eq 3. The construction of the dehumidification path begins at point “C”, which represents the gaswater vapor mixture at the bottom of the tower. A tie line with slope hL/kG is drawn from point “A” to intersect the equilibrium line at point “G”. The coordinates of this intersection point are Ti and Hi. Equation 15 is an equal counterpart of eq 3, obtained by substituting eq 8 into the latter equation. The ratio of eq 10/eq 15 is the basis of Mickley’s graphical technique. This ratio is given by eq 16. GdTG ¼ kG aH ðTG Tt ÞdZ
ð15Þ
dHG HG Ht ¼ dTG TG Tt
ð16Þ
Integration of eq 15 yields eq 17, the second basic equation used in the calculation of the tower height. Z HG2 dHG kG aH Z ð17Þ ¼ G ðT T Þ L t HG1 A straight line drawn from point “C” to point “G” then gives the direction of the initial tangent to the gas path. The slope of this 1532
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line is (HG Ht)/(TG Tt), the ratio of the enthalpy driving force to the temperature driving force. By virtue of eq 16, the slope of curve “CG” also represents dHG/dTG, the rate of change of bulk-gas enthalpy with bulk-gas temperature. Assuming that this slope is constant over a small interval, point “H” represents the bulk-gas enthalpy and temperature at a short distance above the bottom of the tower. The construction is extended with a new tie line and a new direction of the path line tangential to point “H”. This exercise is repeated until a complete gas path (curve “CD”) is achieved.
Table 1. Fuel Properties property
3. BASIS OF THE CONTRIBUTION Equation 18, a representation of how the amount of energy available (Qavailable) within the IGCC system could be used up by the steam path to improve the thermal efficiency, was adopted from previous studies and leads us to the basis of this contribution.9 Qavailable ¼ mðH _ sl HBFW Þ þ mλ _ v þ mðH _ sp Hss Þ
coal
coke
mix
humidity
11.80
7.00
9.40
ash (%)
41.00
0.26
20.68
carbon (%)
36.27
82.21
59.21
hydrogen (%)
2.48
3.11
2.80
nitrogen (%)
0.81
1.90
1.36
oxygen (%)
6.62
0.02
3.32
sulfur (%)
0.93
5.50
3.21
LHV (MJ/kG) HHV (MJ/kG)
13.10 13.58
31.99 32.65
22.55 23.12
ð18Þ
Manipulation of eq 18 by replacing λv by Hss Hsl as per the definition yields eq 19. _ sp mH _ BFW Qavailable ¼ mH
ð19Þ
The rearrangement of eq 19 by simply writing mH _ sp as the subject of the formula yields eq 20, the overall energy balance of the plant. Equation 20 states that the energy carried by the superheated steam to the steam turbine (Qsp) is equal to the sum of the energy available within the IGCC system (Qavailable) and the energy enthalpy boiler feedwater carried into the system. _ BFW mH _ sp Qsp Qavailable þ mH
ð20Þ
If we consider or assume that Qavailable is constant, it is clear from eq 20 that an increase in HBFW will result in an increase in Qsp. According to eq 21, an increase in Qsp will in turn result in an increase in the overall thermal efficiency (ηIGCC), where WST, the steam turbine power output, is given by eq 22. ηIGCC ¼
WGT þ WST Qcoal
WST ηST Qsp
ð21Þ ð22Þ
Substituting eqs 20 and 22 into eq 21, results in eq 23, i.e., the basis of this contribution. ηIGCC ¼
mη _ ST WGT þ ηST Qavailable ðHBFW Þ þ Qcoal Qcoal
ð23Þ
If we consider keeping m_ constant, for example, in the case of a heatIGCC plant, where a maximum m_ possible to satisfy the cooling was obtained, eq 23 is then a straight line equation, where ηIGCC is only a function of TBFW.9 It is evident from eq 23 that an increase in TBFW will result in an increase in ηIGCC. This important result is the basis of this contribution and led to the methodology discussed in the following section.
4. METHODOLOGY The method developed is divided into two parts: i.e., determining the maximum boiler feedwater temperature attainable and applying Mickley’s graphical technique for dehumidification to demonstrate how the aforementioned temperature is achieved. The grand composite curve is used to check whether the determined boiler feedwater temperature is feasible.
Figure 6. Grand composite curve with steam targeting. Application of this method with reference to Figure 3 is as follows, given the conditions of the GT exhaust (G, HG2, TG2, and TW). Part 1: (1) Specify a “ΔTmin_column” for the packed column and a “ΔTmin_HX” for the heat exchanger. (2) Specify TL2 to be at least “ΔTmin_column” lower than the wet-bulb temperature (TW) of the gas to allow for thermal driving forces. (3) Determine the maximum attainable TBFW2 that does not violate “ΔTmin_HX” in the heat exchanger and then calculate the corresponding ηIGCC from eq 23. Validation of TBFW2 is performed on the grand composite curve. Part 2: (4) Draw the equilibrium curve on the “HG” versus “TL” graph. (5) Specify the circulating water flow rate (L) and then calculate the L/G ratio. (6) Draw the operating line using eq 14 and apply Mickley’s graphical technique to determine the gas dehumidification path as follows: (6.1) Specify TG1 and HG1 such that the gas is still above its condensation point. (6.2) Choose a starting hLaH/kGaM and apply Mickley’s graphical technique to determine the dehumidification path. (6.3) Check if the dehumidification path intersects the point (TG, HG) = (TG1, HG1) in step 6.1 and proceed as follows: (i) If the dehumidification path intersects the point (TG, HG) = (TG1, HG1), then the design can be carried out with the specified hLaH/kGaM ratio. (ii) Otherwise repeat steps 6.2 and 6.3 while changing hLaH/kGaM in step 6.2 until the dehumidification path intersects the aforementioned point. (7) Repeat steps 5 and 6 with a new L/G ratio to optimize the cooling process. (8) Determine the tower height (Z), as discussed below. Determining the tower height after obtaining the hLaH/kGaM ratio becomes a matter of solving three simple simultaneous equations. The integrals in the left-hand side of eqs 6 and 17 can be solved graphically after determining the dehumidification path. In addition to the determined hLaH/kGaM ratio, this leaves us with three equations and three unknowns, of which the tower height (Z) is one of them. The other two unknowns are hLaH and kGaM. The three equations can then be solved 1533
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Table 2. Data for the Heat-Integrated Design property
magnitude
m_ (kg/s)
134.2
ηST
0.36
ηIGCC
0.54
WGT (MW)
200
Qavailable (MW)
434.8
Qcoal (MW)
669.3
Figure 9. Results of Mickley’s graphical technique at a L/G ratio of 1.5.
Figure 7. Results of Mickley’s graphical technique at a L/G ratio of 1.1.
Figure 10. Results of Mickley’s graphical technique at a L/G ratio of 1.5.
Figure 8. Grand composite curve with a new steam target. for these three unknowns, hence determining the exact transfer (heat and mass) coefficients required for the process.
5. CASE STUDY The heat-integrated design of the Elcogas plant, represented by Figure 2, was used as a case study for this contribution. This design is based on the fuel (mixture of coal and coke) composition given in Table 1. The stream of concern, i.e., the gas turbine exhaust, is indicated by the heavily bolded line. It is worth emphasizing that the gas desulfurization process of this plant removes about 99.91% of the sulfur present in the gas after gasification. This results in minor sulfur dioxide emissions, which is often excluded in the gas turbine exhaust composition. Figure 6 shows the GCC and the steam targeting used in constructing Figure 2. The green line in Figure 6 represents the
Figure 11. Mickley’s graphical technique at a L/G ratio of 0.8.
amount of energy available within the system after process process exchange. The blue line in Figure 6 represents how the maximum amount of steam, used in obtaining the plant design given in Figure 2, was targeted. For this case, boiler feedwater (BFW) enters the boiler at 25 °C, as shown by the lowest point of the middle curve, and leaves the HRSG en route to the steam turbine at a temperature of 506 °C. Figure 2, therefore, is a design that satisfies the steam target obtained using up the energy available within the system. 1534
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Energy & Fuels The CES was adopted for application in this design to recover the low potential heat and increase the overall efficiency of the design. Application of the CES is on the gas turbine exhaust stream of the design that enters the stack at a temperature of 90 °C. The rest of the design data is given in Table 2. The composition of the gas turbine exhaust stream was assumed to be the design composition of a typical IGCC. This stream was assumed to behave similar to air under these conditions. Consequently, enthalpy data were obtained from the air water psychometric chart. The eight-step procedure given in section 3 was followed in the application of the CES to increase ηIGCC.
6. RESULTS AND DISCUSSION Figure 7 shows the results of the case study after application of the eight-step graphical methodology. These results were obtained at a L/G ratio of 1.1. The maximum attainable TBFW2 was found to be 57.3 °C. Figure 8 shows the GCC of the process
Figure 12. Mickley’s graphical technique at a L/G ratio of 0.8.
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discussed in the case study with a new steam target. The yellow line in this case represents the effect of increasing the BFW temperature to the aforementioned TBFW2 value, on the system. In other words, the yellow dotted curve represents the effect of the CES on the steam target discussed in Figure 8. It is evident from the first slope of the yellow line, which is the same as the first slope of the initial steam target represented by the blue solid line, that the heat capacity flow rate (C) is kept constant. Because C is given by the product of specific heat _ it follows that, at a constant capacity (cp) and BFW flow rate (m), m, _ cp also remains constant to maintain the same C used in the initial steam targeting. Consequently, for a higher BFW temperature, i.e., the aforementioned TBFW2, a higher boiler pressure is required to maintain the same C. In this case, a boiler and HRSG pressure of 189 bar was required, as opposed to the initial boiler pressure of 120 bar for the initial steam target. This resulted in an increase in ηIGCC from 0.54 to 0.55. A TL2 of 59 °C at a ΔTmin_column of 3 °C satisfied the required conditions for step 3 of the aforementioned methodology, while a ΔTmin_HX of 1.7 °C was maintained for the heat exchanger. The outlet gas temperature and enthalpy (TG1 and HG1) where specified to be 64 °C and 450 kJ/kg, as indicated by the small blue circle on the gas curve of Figure 7. The corresponding hLa/kGa that satisfied the conditions of step 5 of the methodology was 7.04 kJ kg1 K1. Further improvements could be performed on these results to improve the cooling process by altering L/G, as explained in step 7 of the methodology. Figure 9 shows the results of Mickley’s graphical technique at a L/G ratio of 1.5 and the same hLa/kGa ratio used in constructing Figure 7. It is evident from Figure 9 that temperature driving forces, i.e., the space between the gas curve and the equilibrium curve, are larger at the bottom of the column and decrease rapidly as the gas approaches outlet conditions (TG1, HG1) at the top of the
Figure 13. Integration between the CES and IGCC. 1535
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Energy & Fuels column. In fact, cooling is so rapid at the bottom of the column in such a way that the gas does not meet the outlet conditions; i.e., the gas curve does not intersect the point (TG1, HG1). To meet the outlet conditions at a higher L/G ratio, a larger hLa/kGa ratio is required. For the case considered in Figure 9, a hLa/kGa ratio of 8.992 kJ kg1 K1 meets the gas outlet conditions, as shown in Figure 10. Figure 11 shows the results of Mickley’s graphical technique at a L/G ratio of 0.8 and the same hLa/kGa ratio used in constructing Figure 7. It is evident from Figure 11 that temperature driving forces decrease gradually as the gas approaches outlet conditions (TG1, HG1) at the top of the column. In this case, the outlet gas exits at a temperature higher than the specified temperature (TG1). To meet the outlet conditions at a lower L/G ratio, a smaller hLa/kGa ratio is required. For the case considered in Figure 11, a hLa/kGa ratio of 5.761 kJ kg1 K1 meets the gas outlet conditions, as shown in Figure 12. This optimization of the cooling or dehumidification process is entirely influenced by the design requirements of the tower; hence, it was not included in this paper, but rather the application of the proposed methodology is emphasized. Integration of the CES with IGCC. Figure 13 shows the integration between the CES and IGCC design discussed in section 2. Instead of feeding BFW at 25 °C, as was the case in the aforementioned design, BFW is fed at 57.3 °C after preheating in the CES. This allows for an increase in ηIGCC as discussed in section 3 and demonstrated in the first few paragraphs of this section. The heated BFW stream is represented by the blue dotted line in Figure 13. The gas turbine exhaust gas is represented by the solid red line in Figure 13. The final superheated steam stream to the steam turbine, which is at a higher temperature than in the initial design, is responsible for the increased ηIGCC. This stream is represented by the orange dotted line in Figure 13.
7. CONCLUSION A new method for recovering low potential heat in a direct CES was presented. Application of this method in a CES proved effective in recovering the low potential heat from the GT exhaust stream to improve the efficiency of the IGCC plant. The method used avoids the use of correlations to estimate the ratio between the liquid-phase heat transfer and the gas-phase mass transfer but rather gives a way to calculate the exact ratio for the cooling task at hand. A case study on the heat-integrated design of the Elcogas plant illustrated that the developed method is capable of increasing the gross efficiency from 54 to 55%. ’ AUTHOR INFORMATION Corresponding Author
*E-mail:
[email protected].
’ NOMENCLATURE aH = heat-transfer area aM = mass-transfer area ASU = air separation unit BFW = boiler feedwater CES = contact economizer system cs = humid heat G = mass flow rate of the gas
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HBFW = enthalpy of the boiler feedwater HG = enthalpy of the gas Hi = enthalpy at the gasliquid interface HL = enthalpy of the liquid hL = heat-transfer coefficient of the liquid phase HRSG = heat recovery steam generator Hsl = enthalpy of the saturated liquid in the boiler Hsp = enthalpy of the superheated steam leaving the HRSG Hss = enthalpy of the saturated steam leaving the boiler IGCC = integrated gasification combined cycle kG = gas-phase mass-transfer coefficient L = mass flow rate of the liquid Qcoal = calorific value of coal Ti = temperature at the liquidgas interface TL = liquid-phase temperature T0 = reference temperature 1 = top of the packed-bed column 2 = bottom of the column m_ = maximum boiler feed flow rate WGT = gas turbine power output WST = steam turbine power output Y = humidity of the gas Z = tower height Greek Symbols
ηST = thermal efficiency of the steam turbine λ0 = latent heat of vaporization of water at T0 λV = latent heat of vaporization of water
’ REFERENCES (1) Bi, X. T.; Liu, X. High density and high solids flux CFB risers for steam gasification of solid fuels. Fuel Process. Technol. 2010, 91, 915–920. (2) Kim, Y. S.; Lee, J. J.; Kim, T. S.; Sohn, J. L.; Joo, Y. L. Performance analysis of syngas-fed gas turbine considering the operating limitations of its components. Appl. Energy 2010, 87, 1602–1611. (3) Corella, J.; Toledo, J. M.; Molina, G. A review on dual fluidizedbed biomass gasifier. Ind. Eng. Chem. Res. 2007, 46, 6831–6839. (4) Minchener, A. J. Coal gasification for advanced power generation. Fuel 2005, 84, 2222–2235. (5) Cormos, C. C. Evaluation of energy integration aspects for IGCC-based hydrogen and electricity co-production with carbon capture and storage. Int. J. Hydrogen Energy 2010, 35, 7485–7497. (6) Vlaswinkels, E. E. Energetic analysis and optimization of an integrated coal gasification-combined cycle power plant. Fuel Process. Technol. 1992, 32, 47–67. (7) Jiang, L.; Lin, R.; Lin, H.; Cai, R.; Liu, Z. Study on thermodynamic characteristic and optimization of steam cycle steam in IGCC. Energy Convers. Manage. 2002, 43, 1339–1348. (8) Emun, F.; Gadalla, M.; Majozi, T.; Boer, D. Integrated gasification combined cycle (IGCC) process simulation and optimization. Comput. Chem. Eng. 2010, 34, 331–334. (9) Madzivhandila, V.; Majozi, T.; Zhelev, T. Process integration as an optimization tool in clean coal technology: A focus on IGCC. Chem. Eng. Trans. 2009, 18, 941–946. (10) Mickley, H. S. Design of forced draft air conditioning equipment. Chem. Eng. Prog. 1949, 45, 739–749. (11) Zhelev, T. K.; Semkov, K. A. Cleaner flue gas and energy recovery through pinch analysis. J. Cleaner Prod. 2004, 12, 165–170. (12) Venkataraman, C. CL 332/334-UG LAB, Cooling Tower-MT 305; http://www.che.iitb.ac.in/courses/uglab/manuals/coollabmanual. pdf (accessed on Jan 2010). (13) Kloppers, J. C.; Kr€oger, D. G. Cooling tower performance: A critical evaluation of Merkel assumptions. Res. Dev. J. 2004, 20, 24–29. 1536
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