performance analysis of combined humidified gas...

Desalination 196 (2006) 84–104

Performance analysis of combined humidified gas turbine powergeneration and multi-effect thermal vapor compression

desalination systems — Part 1: The desalination unit and itscombination with a steam-injected gas turbine power system

Yongqing Wanga*, Noam Liorb

aSchool of Energy Science and Engineering, Harbin Institute of Technology, Harbin, 150001, P.R. ChinaTel. +86 (451) 8641 2178; Fax: +86 (451) 8641 2078; email: [email protected]

bDepartment of Mechanical Engineering and Applied Mechanics, University of Pennsylvania,Philadelphia, PA 19104-6315, USA

Received 27 October 2005; accepted 18 January 2006

Abstract

Humidified gas turbines (HGT) have been identified as a promising way of producing power. The use of the steam-injected gas turbine (STIG) HGT cycle in a combined power and water desalination system was analyzed using energyand exergy performance criteria. A brief description and rationale of the background of HGT cycles and dual-purposepower and water systems is given. A thermal desalination unit was modeled and analyzed, and the results led to theselection of a multi-effect thermal vapor compression (METVC) unit for producing fresh water from seawater for bothgeneral use and humidification; then the performance of a STIG-based combined system was investigated. The analysisperformed improved the understanding of the combined STIG power and water desalination process and of ways toimprove and optimize it. Some specific conclusions are that: (1) a METVC desalination system is preferred to a multi-effect evaporation one when the pressure of the motive steam is high enough, >~3 bar, to run a steam jet ejector; (2) thesteam injection rate in the STIG cycle has a strong effect on water and power production, offering good flexibility fordesign and operation; (3) higher pressure ratios and higher steam injection rates in the STIG cycle increase powergeneration, but decrease water production rates, and higher turbine inlet temperatures increased both power and waterproduction; (4) a distinct water production gain can be obtained by recovering the stack gas energy. The results indicatethat such dual-purpose systems have good synergy, not only in fuel utilization, but also in operation and designflexibility.

Keywords: Integrated power and water system; Humidified gas turbine; Multi-effect thermal vapor compressiondesalination; Energy and exergy analyses

*Corresponding author.

doi:10.1016/j.desal.2006.01.010

0011-9164/06/$– See front matter © 2006 Elsevier B.V. All rights reserved.

Y. Wang, N. Lior / Desalination 196 (2006) 84–104 85

1. IntroductionHumidified gas turbines (HGT) that use gas–

water mixtures as the working fluid have beenidentified as a promising way to generate power,and their incorporation into a dual-purposesystem producing both power and fresh water bydesalination was investigated, especially regard-ing some synergies between the power generationand water desalination processes. Compared withcombined cycles, the main features of HGTinclude high efficiency, high specific power out-put, reduced specific investment cost, reducedNOx emission and improved part-load perfor-mance [1,2]. There are several configurations ofHGT cycles, with the steam-injected gas turbine(STIG) cycle, humid air turbine (HAT) orevaporative gas turbine (EvGT) cycle being themost representative.

In a STIG plant, steam is produced in a heatrecovery steam generator (HRSG) using theexhaust heat of the gas turbine stack, and is theninjected into the gas turbine combustion chamber.In a HAT or EvGT cycle, compressed air fromthe compressor is humidified with hot water in ahumidifier that recovers the low-grade heat in thesystem, and is then heated by the exhaust gasfrom the gas turbine before entering the com-bustor. Different from HRSG in which steam isproduced at a certain saturation temperature, thehumidifier is an air–water direct-contact com-ponent in which water evaporates noniso-thermally. With the potential of reaching anefficiency over 60%, the HAT and EvGT cycleshave been considered as strong future competitorsto the combined cycle.

Large water consumption is a major disadvan-tage of HGT, which restricts the use of the plant,especially in water-short areas. A LM5000STIG™ plant commercialized by General Electric,for instance, consumes about 29 t/d water perMW power output (1450 t/d water with a poweroutput of 50.7 MW) when running under a fullSTIG pattern [3]. Two categories of solutions are

proposed to solve the problem: one is to recoverthe steam in the flue gas for reuse, and another isto produce useful water by water desalination.

Introducing a condenser to recover the watervapor in the flue gas and reuse after treatment hasbeen widely studied [3–6]. Owing to the com-bustion reaction of hydrogen and oxygen in thecombustor, the flue gas contains more water thanconsumed. For instance, for each kg of air enter-ing the compressor and 0.15 kg steam injectedinto the combustor, a CH4-fueled STIG cycle witha pressure ratio of 10 and firing temperature of1300EC contains 0.2176 kg water vapor in theflue gas. Both theoretical and experimental resultsshowed that self-sufficient water recovery waspossible [3–6]; as a further consideration, the heatgained in the process by the condenser coolantcould be used in district heating [7].

As to using desalination for producing thewater for humidification, Cerri [8] proposed amulti-stage flash (MSF) unit driven by exhaustgas heat of turbines for producing demineralizedwater from seawater for a STIG cycle; a smallreverse osmosis (RO) unit was suggested [9] toproduce water for steam injection, and the cal-culations predicted a minor effect of the additionof this system on the final operating cost of thepower plant. The benefit of the integration ofHGT with desalination is not limited to pro-duction of water for the cycle only. Just as inother power plants, the existence of the low-temperature heat in HGT cycles provides a favor-able condition for power and water cogeneration,and the second-law based energy utilizationmodel of cogeneration systems indicates anadvantageous performance over single-purposeunits [10].

In some water-short areas, large quantities oflow-grade thermal energy, of which the tempera-ture is usually lower than 130EC, are needed torun thermal desalination units for producing freshwater from saline water. It is a great waste ofexergy to provide such low-grade heat for

Y. Wang, N. Lior / Desalination 196 (2006) 84–10486

desalination by burning fuel in a boiler. Com-bining power plants with desalination units pro-duces a great synergy between power generationand water production. Such combined power andwater production systems are usually called dual-purpose plants [11–13]. Most of the dual-purposesystems operating in the world are the combina-tion of steam turbines with thermal desalinationunits; but in recent years, there has been anobvious interest in moving to gas turbines orcombined cycle-based systems [14,15] owing totheir increasing installed capacity.

The synergy of the combination of power andwater production has significant energy, economyand environment benefits. For example, couplingMSF with a steam turbine plant showed a 44.4%energy saving of water production, from40 kWh/m3 equivalent work for a water-onlysystem to 22.3 kWh/m3 for water–power cogene-ration [16], and a 44.7% water cost reductionfrom $2.66/m3 to $1.47/m3 [17]. A life-cycleassessment showed that the environmental load ofthermal desalination technologies was reducedabout 75% when operating in a hybrid plantbased on a combined cycle [18].

A combination of HGT (STIG) and desali-nation was studied [19], based on a single-shaftgas turbine of 38.3 MWe, in which 10 kg/s steamproduced in the HRSG was injected into thecombustor, and the balance was used to run twosteam jet ejectors and then two multi-effectdistillation desalination units. A direct-type gas–seawater heat exchanger was introduced torecover further the exhaust heat of the flue.

The objective of this two-part paper was tostudy the energy, exergy, and water productionperformance of integrated power and water sys-tems that use HGT as the prime mover. Tounderstand the performance and parameter selec-tion of water production, a thermal desalinationunit was modeled and analyzed, and a combinedsystem based on a STIG plant was investigated.In a separate paper, Part 2, an EvGT-based

combined system was studied, and the resultsdiscussed to clarify further the performance of thetwo combined systems.

2. Multi-effect thermal vapor compression(METVC) desalination

2.1. Desalination processes

MSF, multi-effect evaporation (MEE), thermalvapor compression (TVC) and RO are fourcommonly used desalination processes. The firstthree are heat-driven, but also consume pumpingwork. Some of the advantages of thermaldesalination are its lower sensitivity to thesalinity and quality of the feedwater, and itsability to produce much higher quality distillatewhich can also be used for applications requiringwater purer than needed for drinking, such as forboiler feed [19]. RO is driven only by mechanicalwork. Compared with current commercial thermaldesalination, RO is much more energy-efficient,more compact, and more flexible in design andoperation because of its modular structure andsimper and quicker start-up/shut-down charac-teristics [13,16]. RO has been increasing itsmarket share in recent years, yet thermal desali-nation still dominates the seawater desalinationmarket. Being driven by low-temperature heat,thermal desalination is suitable to be combinedwith power generation or other industrial pro-cesses to improve the total energy efficiency.

MEE and METVC desalination units with atop brine temperature (TBT) lower than 70EChave attracted attention in recent years [20].Compared with the most widely used MSFdesalination, MEE has the advantages of lowercorrosion and scaling rates, lower capital cost,longer operation life and less pumping powerconsumption [21]. When moderate-pressure(around 3 bar or higher [16]) steam is available,it can be used effectively for entraining and com-pressing the vapor produced in the last effect of


Fig. 1. Schematic diagram of a four-effect TVC system. ECON, end condenser; EVA1–EVA4, evaporators; FLA1–FLA4,flashing boxes; H1–H3, preheaters; SJE, steam jet ejector; WST, water storage tank.

the MEE plant by using a steam jet ejector.Compared with a stand-alone MEE plant, such aMETVC arrangement needs less cooling water,and thereby lower pumping power and pretreat-ment costs [22].

Fig. 1 schematically illustrates a four-effect(EVA1–EVA4 with associated FLA1–FLA4)TVC unit with seawater preheaters (H1, H2, H3),which is apparently the combination of a steamjet ejector and a conventional MEE unit. Run bythe motive steam (1), the steam jet ejector (SJE)entrains and compresses part of the water vapor(2) produced in the last effect (EVA4). The steam(3), which is called heating steam, leaves the SJEand condenses in EVA1 providing energy forseawater evaporation. Part of the condensate (5)returns to the boiler or HRSG, and the remainingpart (6) is introduced into the associated flashingbox (FLA1) where a small amount of vapor (7)flashes off because of a pressure drop. The vapor(8) evaporated from the seawater in EVA1 passes

through the preheater (H1) to preheat the feedseawater (9), and is then routed into the secondeffect (EVA2) together with the flashing vapor(7) from FLA1, serving as the heat source inEVA2. The balanced brine (12) from EVA1flows into the second effect (EVA2) and producesvapor by flashing. This process is repeated for alleffects until the last one. Part of the vapor (2)formed in the last effect is entrained by the steamjet ejector, and the remainder (13) is introducedinto the end condenser (ECON) where it releasesits latent heat by heating seawater (15). Part ofthe heated seawater (16) is used as the feed of theMETVC unit, and the balance (17) is rejectedback to the sea.

2.2. Performance criteria

To understand the water production and para-meter selection of the METVC unit used in thispaper, the performance of the METVC was


studied first and compared with the performanceof a MEE unit using the following criteria.

1. Performance ratio (PR) — the ratio betweenthe mass of the produced fresh water mw to that ofthe consumed motive steam mm:

2. Specific heat transfer area (a) — the heattransfer area needed to produce 1 kg fresh water:

where A is the total heat transfer area of thedesalination unit, composed of the area of theeffects Aef and the condensation area of the endcondenser Aecon.

3. Specific exergy consumption (ec) — theexergy consumed for producing 1 kg fresh water:

where em is the specific exergy of the motivesteam, and ehc that of the condensate (stream 5 inFig. 1) of the motive steam flowing out of thedesalination unit; T0 is the temperature of thesurroundings.

4. Exergy efficiency of the desalination unit(gD):

where WP is the pumping work consumed bydesalination, and wmin the minimum work neededin a reversible separation process for producing1 kg of fresh water. The calculation of the mini-mum work is given in Appendix A.

5. Exergy efficiency (gE) and exergy loss rate

(χE) of the steam jet ejector — The steam jetejector is the basic component in the METVCunit that distinguishes METVC from MEE. Twoindices are used to evaluate the performance ofthe SJE, its exergy efficiency gE and exergy lossrate χE:

The numerator and the denominator in Eq. (5)represent the exergy the entrained steam gainsand the motive steam loses in the steam jet ejectorprocess, respectively, and those in Eq. (6) repre-sent the exergy destruction in the steam jet ejectorand the exergy provided for the METVC unit.

2.3. Modeling and simulation

The modeling and simulation were made usingthe Engineering Equation Solver (EES) software[23]. The properties of seawater and brine aretaken from Husain [11]. The heat transfer coeffi-cients of evaporation and condensation, theboiling point elevation of brine, as well as thenon-equilibrium allowance of flashing evapora-tion in the flashing box were taken from El-Dessouky and Hisham [24]. The performance ofthe steam jet ejector was taken from Power [25].In the modeling and simulation, the distillateproduced in each effect was considered to be saltfree, and, in accordance with industrial practice,each evaporator had the same heat transfer area[24]. Table 1 shows the calculation conditions.More detailed information on MEE and METVCcan be found in many publications (cf. [24,26]).

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Table 1Calculation conditions for MEE and METVC

Temp. of seawater Tsw, ECTemp. of rejected cooling seawater Tcool, ECTBT, ECSalinity of seawater Xsw, ppmSalinity of rejected brine Xbr, ppmState of motive steamCondensation temp. of heating steam Thc, ECPressure of heating steam ph, MPaMin. condensation temp. Tcon in end condenser, EC

305E lower than Tcon6936,00070,000Saturated72

0.03397a

40

aSaturated temperature is 72E.

Using the model developed by the authors,PR, specific area of effects aef , which is the ratioof Aef and mw, and ec of four cases are calculatedunder the same conditions as those previouslygiven [26–28], and the results are shown inTable 2. aef accounts for the main part of a, in asix-effect METVC unit studied below; for in-stance, the typical value of aef /a is around 94%.It was observed that the model predictions com-pare well with the data in the literature (Table 2),and the relative differences are within 3%, exceptaef. The larger difference of aef is due to thedifferent correlation of heat-transfer coefficientsand the calculation model used in this paper andthe literature [26]. The heat transfer coefficientcorrelations used are shown in Appendix B; theywere reported by their authors to have been vali-dated through comparison against other corre-lations and available experiments [24].

A parametric analysis was carried out toinvestigate three important factors: compressionratio (CR), pressure of the motive steam (pm), andseawater preheating on the performance of theMETVC unit.

2.4. Influence of the compression ratio (CR)The compression ratio CR, which is the pres-

sure ratio of the heating steam entering the first

effect and the entrained steam from the last effect

and the expansion ratio ER, which is the pressureratio of the motive and entrained steam,

are two parameters which determine the perfor-mance of the steam jet ejector and then of theMETVC unit, under the calculation conditionsshown in Table 1. When the pressure of theheating steam ph is specified (0.034 MPa in thispaper), a certain CR corresponds to a certainrange of operating temperatures of the MEE inthe METVC unit, since a change of CR meansthat there is a change of the pressure and tem-perature of the entrained steam, therefore, theoperation pressure and temperature of the lasteffect. If pm is also given, ER and then the per-formance of the METVC unit can be calculated.

Fig. 2 shows the influence of CR on PR and awhen pm is fixed. Fig. 3 shows the variations ofec, gE and χE with CR, and Fig. 4 shows thevariation of gD of METVC. For a specifiednumber of effects n, increasing CR causes thedecrease of the parameters a, PR and gD, and anincrease of ec. An increase in CR implies a lower

Fig. 2. Dependence of PR and a of METVC on CR and n.

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Table 2Comparison of model predictions against available data for METVC units

Calculation condition Case 1 [26] Case 2 [26] Case 3 [27] Case 4 [28]

nThc, ECTbr, ECTBT, ECPm, MPaTcool, ECTsw, EC∆Tph, ECmm /men

Xsw, ppmXbr, ppm

66542.861.82.540303.81.3647,800a

71,500a

66542.861.82.54030No preheaters1.3647,800a

71,500a

46246.858.82.540—No preheaters0.8647,800a

71,500a

462.748.4—2.44433No preheaters0.847,80071,500

PR ReferenceModel

10.05 [26]10.11

8.87 [26]9.10

7.65 [27]7.73

8.61 [28]8.49

aef, m2/(kg/s) ReferenceModel

326.2 [26]353.4

341 [26]354.8

—347

—363.7

ec, kJ/kg ReferenceModel

87.91 [26]89.44

99.26 [26]99.3

—117.09

—106.05

aAssumed values based on the context of the references.

Fig. 3. Dependence of ec, gE and χE on CR and n.

pen (Eq. 7) and thus a lower Ten, leading to anincreased temperature difference between the firstand the last effect and thus an increased tem-perature difference across each effect. Thisobviously results in a decrease of the needed heattransfer area. At the same time, the increased

Fig. 4. Dependence of gD of METVC on CR and n.

temperature difference increases the heat transferirreversibility in each effect, causing a higher ecand lower PR and gD. For the same reason, for aspecified CR, a higher number of effects n resultsin higher PR, a and gD, but a lower ec.


It is noteworthy that the exergy efficiency ofthe METVC unit is very low, <3.5%, as shown inFig. 4, indicating great potential for performanceimprovement of the desalination process.

As well know, the steam jet ejector is not anexergy-efficient component. From Fig. 3, with amotive pressure of 0.5 MPa and CR between 1.8and 4.5, gE is only 16–25%, and about 45% of theexergy provided for desalination is lost in theSJE.

2.5. Influence of the motive steam pressure (pm)

As analyzed in Section 2.4, specification ofthe CR and ph determines pen, and thus the per-formance of the steam jet ejector and of the entireMETVC unit can be determined given pm. Fig. 5illustrates the effect of pm on the performance ofthe METVC unit. It is clear that PR increases ata diminishing rate with the increase of pm. Whenthe steam jet ejector is run by higher-pressuremotive steam, it draws more vapor from the lasteffect, thus increasing the amount of the heatingsteam, consequently the amount of product water,resulting in an increased PR. The increase rate ofPR diminishes because (1) the increase rate of theexergy of the motive steam provided for desali-nation diminishes as pm increases, as determinedby the properties of the saturated steam; and(2) the increased exergy destruction in the steamjet ejector and the evaporators at higher pm [26]cause the specific exergy consumption ec toincrease (Fig. 5).

Calculations also show that pm has only aslight influence on a (Fig. 5). The reason is thatthe higher enthalpy of the motive steam at thehigher pm results in an increased heating steamtemperature Th, which increases the temperaturedifference for the sensible heat transfer in the firsteffect, and consequently slightly reduces aef. Atthe same time, higher-pressure motive steamincreases the amount of vapor entrained from thelast effect, and thus reduces the amount of vapor

Fig. 5. Impact of pm on PR, a, and ec of METVC.

Fig. 6. Performance of the MEE unit.

that needs to be condensed in the end condenser,leading to a slightly reduced aecon.

Fig. 6 shows the performance of a conven-tional MEE unit operated by saturated steam atdifferent conditions: 72EC (0.034 MPa), 151.8EC(0.5 MPa) and 242.6EC (3.5 MPa). Only one areacurve is shown due to the small influence of pm(Tm) on a. Compared with the data in Figs. 2 and5, one can see that, although the exergy efficiencyof the steam jet ejector gE is very low (Fig. 3), theMETVC unit has a significant advantage overMEE running with the same heat source when pmis high enough to run the steam jet ejector. For


Fig. 7. Exergy efficiency of METVC and MEE.

instance, a seven-effect MEE unit run by satu-rated steam at 3.5 MPa has a PR of 5.5 for aspecific area of 310 m2/kg; however, a six-effectMETVC unit with a CR of 3.5 run by the sameheat source has a PR of 8.6 for a specific area of301 m2/kg. Different from the MEE unit in whichno work is produced, mechanical work is per-formed in the METVC unit by the motive steamin the steam jet ejector process by compressingthe entrained steam. This is more efficientexergy-wise because the steam is available at apressure significantly higher than needed for sup-plying heat to the MEE; instead of wastefullythrottling it to the heat supply pressure, theejector is used for vapor compression in theMETVC process.

Fig. 7 shows the exergy efficiency, gD, of thesix-effect METVC and seven-effect MEE unitsmentioned above for different pm. It is revealedthat a higher pm leads to a lower gD for bothMETVC and MEE units due to the increasedexergy consumption of water production ec(Figs. 5 and 6); run by the same heat source,METVC has a higher gD than MEE due to thework process in the steam jet ejector in METVC.

Table 1 shows the calculation conditions used

for METVC and MEE, including the TBT of69EC. Assuming that the minimal heat transfertemperature difference between the heating steamand seawater is 3EC, the temperature of theheating steam is 72EC, for which the saturationpressure is 0.034 MPa. The irreversibility in thesteam jet ejector process and the heat transferprocess between heating steam and seawater isreduced as the temperature of the steam providedfor desalination reaches closer to 72EC and itspressure to 0.034 MPa, thus resulting in highergD. Thus, under the same calculation conditions,the MEE unit run at 72EC (0.034 MPa) saturatedsteam has a higher gD than the METVC or MEEunits run by higher-temperature/pressure steam,although the latter has higher PR; this higher PRis at the expense of higher ec. This can be seenclearly from Figs. 5–7. For instance, the seven-effect MEE run at 72EC (0.034 MPa) saturatedsteam has a gD of 3.8%, PR of 5.5 and ec of51 kJ/(kg distillate), while the six-effect METVCrun at motive steam of 240.6EC (3.5 MPa) has agD of 1.9%, PR of 8.6 and ec of 109 kJ/(kgdistillate). Consequently, when the pressure/temperature of the steam provided for desalina-tion is higher than 0.034 MPa/72EC, a moreexergy-efficient way of utilizing it is to lower itstemperature/pressure to 72EC/0.034MPa by pro-ducing work, and only then use as the heat sourceof MEE.

2.6. Influence of seawater preheating

The influence of feed seawater preheating inthe preheaters (H1, H2, H3 in Fig. 2) is shown inFig. 8. High preheating ∆Tph, which is thetemperature rise of the feed seawater in eachpreheater, results in a high PR with slightlyincreased specific area a compared with aMETVC unit without preheaters. In MEE orMETVC units, the feed seawater is first heated toboiling temperature (the seawater heating pro-cess), and is then boiled to produce vapor (the


Fig. 8. Impact of preheating on PR and a.

seawater evaporation process), by absorbing thecondensation heat of the vapor from the formereffect (heating steam is used for the first effect).Seawater preheating helps reduce the temperaturedifference between the heated seawater and thecondensing vapor in each effect except the lasteffect, leading to a more exergy-efficient heattransfer process, which increases the water pro-duction but also increases the specific areaneeded in the seawater heating process. Since theseawater evaporation process is the main heattransfer process in the effects, and the evapora-tion area is the dominant factor of the total area,the increased area in the seawater heating processhas a minor influence on a.

2.7. Parameter selection and the performance ofthe METVC unit

In a STIG-based combined system, thepressure of the saturated steam for desalination isthe same as that of the injected steam, 0.5 MPahigher than the operation pressure of the com-bustor in this paper. In an EvGT-based system (tobe discussed in Part 2), the steam pressure ishigher than 0.3 MPa. Consequently, an METVCunit is preferred to MEE, both in STIG- andEvGT-based systems, based on the analysis inSection 2.5.

Fig. 9. PR of a six-effect TVC unit.

In an MEYVC unit, pm, which is determinedby the heat source (boiler, power plant or indus-trial process) has a great influence on PR, espe-cially when pm <1.5 MPa (Fig. 5), but littleinfluence on the specific area a, while the com-pression ratio CR and the number of effects nhave significant influence on both PR and a. Thespecific area a, important in influencing equip-ment cost, is determined therefore mainly by CRand n under the calculation conditions shown inTable 1. In lieu of detailed economic andtechnological optimization studies needed forparameter selection, we have chosen here, afterreferring to the information from commercialMETVC units previously described [20], is a six-effect TVC unit with a CR of 3.5, which has aspecific area a of about 301 m2/(kg/s), changingonly slightly with pm (Fig. 5). The PR of the unitwith and without preheaters is shown in Fig. 9.

3. Calculation conditions and performancecriteria for the analysis of combined powerand desalination systems

3.1. Calculation conditions and assumptions

The main calculation conditions and assump-tions are summarized in Table 3. The commercial


Table 3Main conditions and assumptions for the modeling of HGT-based systems

Inlet air conditionsDead state of exergy analysisFuelIsentropic efficiency of compressorIsentropic efficiency of turbinePressure drop of compressor inletPressure drop in combustorPressure drop in stackPressure drop in HRSGPressure drop in gas-gas heat exchangerPressure drop in water-gas heat exchangerPressure drop in direct-contact gas-water heat exchangerCombustion chamber efficiencyMinimum pinch point temperature difference of HRSG, ECMinimum temperature difference at hot side of HRSG, ECMinimum exhaust temperature leaving the HRSG, ECMinimum driving force of enthalpy difference in humidifier, kJ/kgTemperature effectiveness of gas-gas heat exchangerPressure of steam injected in STIG cycle

25EC, 1 atm air with 100% relative humidity25EC, 1a tm air with 100% relative humidityCH4

0.880.91% of inlet pressure3% of inlet pressure1% of inlet pressure3% of inlet pressure3% of inlet pressure2% of inlet pressure2% of inlet pressure0.991550140250.820.5 MPa higher than the combustor pressure

Aspen Plus code [29] was used to carry out thesimulation of the HGT cycle.

This study focused on the power and waterproduction of the combined systems. A detailedperformance analysis and comparison of thepower-only cycles is available elsewhere [30–32].

The computerized models of the HGT-basedcombined systems were validated by (1) examin-ing the physical sensibility of the calculationresults for each component and the entire cycle,(2) allowing a relative error tolerance of only 10!4

in the Aspen Plus program, (3) comparison withavailable results (Table 2), (4) examining therelative errors in the mass and energy balance ofthe computerized model of the METVC unitwhere the former was found to be <10!5 and thelatter <10!13. It would have been good to validatethe results by using experimental data, but theonly case we found was a conceptual design of aSTIG–METVC system [19] with insufficientinformation to run our model.

3.2. Performance criteria

The performance criteria used to analyze theenergy and exergy utilization and the water andpower production of the combined systems,which are very helpful in helping understand theprocess and ways by which it may be improved,are described below.

1. Thermal efficiency ηt and exergy efficiencyge of the cycles:

where W is the net power output from the gasturbine plant, mf is the mass of fuel input to thecombustor, and qf and ef are the low heat valueand the specific exergy of fuel, respectively.

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2. Heat loss ratio χt and exergy loss ratio χe ofeach component. The heat loss ratio χt is the heatloss Ql in each component as a fraction of theinput fuel energy Qf, and exergy loss rate χe theexergy loss El to the fuel exergy Ef :

3. Heat recovery rate ξt and exergy recoveryrate ξe. In the two combined systems studied, partof the exhaust gas energy is recovered by theinjection steam in the HRSG in the STIG-basedsystem or by the humid air in the regenerator inthe EvGT-based system, and is then returned tothe combustor; part of the energy is recovered bythe motive steam used as the heat source fordesalination. Obviously, improved heat recoveryimproves the process efficiency, and the criteriaused are ξt, the heat recovered Qr as a fraction ofthe input fuel energy Qf, and ξe the exergyrecovered Er as a fraction of the input fuel exergyEf :

4. Power-to-water ratio Rpw. Power-to-waterratio Rpw is the ratio of the generated power w tothe mass of the produced water mw, which is akey factor in determining the performance of thecombined power and water system.

Table 4Reference system performance for normalization

Cycle pattern

Desalination unit

Pressure ratio, βFiring temperature (TIT), ECPressure of motive steam for desalination, pm, MPaGross power output, w0,gross kJ/(kg fuel)Net power outputa, w0,net kJ/(kg fuel)Water production, mw0, kg/(kg fuel)

Simple gas turbine cycleSix-effect TVC, CR = 3.5, ∆Tph = 4EC1013001.5

16,390

15,692

97

aBy taking the pumping work for water production as7.2 kJ/(kg distillate) [16].

MW/MIGD is often used as the unit of Rpw(1 MIGD = 52.662 kg/s).

5. Normalized power and water production.To exhibit more clearly the sensitivity of powerand water production in the combined systems,mw and w are normalized by the simple-cyclewater production mw0 and power output w0 shownin Table 4. The flowsheet of the combined systembased on the simple cycle is the same as that ofthe STIG-based system shown in Fig. 10, exceptthat no steam is injected into the combustor in thesimple cycle.

4. STIG-based power and water system

4.1. System configuration of STIG-based com-bined system

Fig. 10 schematically shows a STIG-basedintegrated power and water production system inwhich part of the saturated steam produced in theHRSG is used to operate a thermal desalinationunit (TDC), and the balance is superheated andthen injected into the combustor (CC) to enhancepower output. A system configuration that uses

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(14)

(15)


Fig. 10. STIG-based power and water combined system.C, compressor; CC, combustor; FC, fuel compressor;G, generator; HRSG, heat recovery steam generator;P, pump; T, turbine; TDU, thermal desalination unit.

the exhaust gas to heat seawater was not used toavoid potential corrosion and scaling problemscaused by the higher temperature seawater thatwould be generated.

In a conventional STIG cycle, all the heat ofthe gas turbine exhaust is used to produce steamfor injection, and the optimum performance canbe obtained when both the temperature and themass of the steam reach the maximum valueallowed by the technical condition of the HRSG.If part of the energy is used to produce water, theefficiencies ηt and ge of the cycle, and thereby thepower output, decrease. The power output is evenlower when the pumping work consumed by thedesalination unit is also considered. Water pro-duction is at the expense of the reduction of pro-duced power, which is a common characteristicof integrated power and water systems.

The pressure ratio β of the compressor, theturbine inlet temperature (the “firing” tempera-

ture, TIT), and the steam injection rate xj are themost important parameters influencing the per-formance of the STIG cycle and the combinedsystem.

4.2. Influence of steam injection rate xj

In a STIG-based system at a specified β andTIT, the mass ratio of the injected steam to thecompressed air, xj,

is the dominant factor influencing Rpw because itdetermines the energy and exergy distribution ofthe exhaust gas in the HRSG between the injec-tion steam and motive steam, and thus the fuelenergy and exergy distribution between the powerand water production.

Fig. 11 shows the influence of xj and Fig. 12is an example of energy and exergy utilizationunder different xj. Increasing xj increases theamount of the thermal energy and exergy of thegas turbine exhaust gas that is used to producesteam for injection, thus reducing the amountavailable for desalination (Fig. 12). This results inhigher power output and lower water production(Fig. 11). Increasing xj obviously also increasesthe water consumption, thus reducing the netwater production even further (Fig. 11). At acertain xj, the net water production mw,net is zero(Fig. 11), which is called “self-water production”.

Self-water production is a way to solve theproblem of injection water supply for a power-only STIG plant. A saline water desalinationsystem including a seawater pretreatment unit,desalination unit and brine feed and drain pipe-lines, etc, is required, and part of the steam, about10% in the case shown in Fig. 11, produced in theHRSG is needed to serve as the heat source forthis desalination. The premise of this course isthat saline water is available. Adding a condenserafter the HRSG is another way to provide waterfor injection, as previously discussed [3–6]. In the

(16)


Fig. 11. Normalized power and water production of theSTIG-based system as a function of xj.

Fig. 12. Exergy and energy utilization of the STIG-basedcombined system for different xj. β = 10; TIT = 1300EC.

case shown in Fig. 11, the flue gas should becooled to 53EC to recover the injected water. Acondenser, a coolant circulation and cooling sys-tem, and a condensate treatment unit are neededin this situation. The method of choice is deter-mined by an economic analysis for the concretecondition of the power plant.

(a)

(b)

Fig. 13. Influence of β and TIT on ge, ηt, ξ e,D and ξ t,Din the STIG-based combined system. (a) Variations of geand ξ e,D with β and TIT; (b) Variations of ηt and ξ t,D withβ and TIT.

4.3. Influence of pressure ratio β and tubine inlettemperature (TIT) of the cycle

Fig. 13 shows the exergy efficiency ge andthermal efficiency ηt of the STIG cycle, as well asthe exergy recovery rate ξ e,D and energy recoveryrate ξ t,D for desalination, for different β (from 10to 30), TIT (1100EC and 1300EC) and xj (0.05


Fig. 14. Normalized power and water production of theSTIG-based combined system.

Fig. 15. Exergy and energy utilization of the STIG-basedcombined system for different β. TIT = 1300 EC; xj =0.05.

and 0.1). We can see the percentage of fuelenergy/exergy converted into power and con-sumed by desalination. Although ξe,D and ξt,D onlyrepresent the thermal energy/exergy provided fordesalination, they determine water production, asshown below, since that pumping work is only a

Fig. 16. Heat exchanger T–Q diagram of the HRSG fordifferent TIT.

small fraction of the desalination energy con-sumption. Fig. 14 shows the normalized powerand water production, and Fig. 15 the exergy andenergy utilization for different β.

These figures reveal that ge and ηt have asimilar trend, as the power output w, consistentwith the definition of ge and ηt by Eqs. (9) and(10). Within the range of β studied, for a fixed xj,a higher β increases ge, ηt and w, as known [30].Increasing TIT, as well known, also improvespower production efficiency.

Figs.13 and 14 also show that ξe,D and ξt,D andthe water production mw have a similar trend,consistent with the definition of ξe,D and ξt,D byEqs. (11) and (12). Opposite to the influence of βon power, mw decreases with the increase of βbecause the temperature of the exhaust gas fromthe gas turbine is lower for higher β; therefore,less energy and exergy can be recovered fordesalination (Fig. 15).

A higher TIT is not only beneficial tothe power generation in the STIG-based systembut also to the water production rate (Fig. 14)because the increased turbine outlet temperaturemakes more energy and exergy available fordesalination.


It is noteworthy that ξe,D, ξt,D and mw dropmore sharply with β at TIT = 1100EC than at TIT= 1300EC, which is determined by the heattransfer process in the HRSG. As shown in Fig.16, when TIT = 1300EC and β = 30, the exhaustgas can be cooled to the minimal temperature weallowed, 140EC, but when TIT = 1100EC, thepinch point temperature ∆Tp of the HRSGrestricts the heat exchange process. To meet theminimum ∆Tp we allowed, the exhausttemperature leaving the HRSG must in this casebe increased to 167.7EC, causing a sharperdecrease of ξe and ξt, and thereby of mw.

4.4. Energy recovery from the stack gasFig. 12 shows that the energy rejected to the

environment rises significantly with the increaseof xj due to the increased steam fraction in theflue gas and the corresponding large latent heat ofwater. For example, the heat and exergy losseswith the stack gas are 19.3% and 5.2%, respec-tively, when β = 10, TIT = 1300EC and xj = 0.05,and 34.9% and 7.1%, respectively, when xj =0.15.

It is difficult to use the heat of the rejectedstack gas because of its low temperature. The gascontains water vapor, and thus, when cooled,loses first its sensible heat until the temperaturedrops to the water vapor condensation level, atwhich point the latent heat of condensation isreleased too. From Fig. 17, which shows the heatreleased in the process of cooling the stack gasfrom 140EC to 40EC, we can see that whenxj = 0.05, more than 60% of the total heat isreleased in the temperature range of 57.5EC to40EC during the condensation process of thewater vapor in the flue, and when xj = 0.15, about80% of the heat is released in the temperaturerange of 68.5EC to 40EC. It is hence impossibleto use this heat to produce steam for waterdesalination.

There are three ways to use the stack gas heat.The first one is for heating service water, and at

Fig. 17. Heat release process of the stack gas leaving theHRSG.

the same time the water in the stack gas isrecovered by condensation. The second is topreheat the seawater (part of stream 16 in Fig. 1)feed to the desalination unit to recover the heat,and at the same time recover the vapor in thestack gas. A water–gas indirect contact heatexchanger is needed in both cases, and if therecovered water is to be used, it should be treatedto separate it from undesirable flue gas com-ponents. The third way is by using a direct-contact gas–seawater heat exchanger to heatseawater for the desalination unit, but only if theflue gas does not contain species that would beharmful if present in the product water or thatmay impair plant performance. The second wayis the most energy-profitable for the STIG-basedcombined system because heating the seawatercan help increase the production of fresh water,and the reuse of the vapor in the stack gas resultsin higher net water production. The heat exchan-ger must be designed to withstand the presence ofboth corrosive fluids. The third way, in which thegas–seawater direct-contact heat exchanger isvery similar to the long commercialized wet-gasdesulfurization system, is the simplest and thusseems to be technically feasible [19]. This third


Fig. 18. Net water production of STIG-based combinedsystem with and without the direct-contact heat exchan-ger for stack gas heat recovery.

way is chosen in this paper to recover the stackheat.

Part of the seawater from the end condenser isfed to the direct-contact heat exchanger, heated to63EC, and then, after mixing with the seawaterfrom the end condenser according to the massflow and temperature required by each effect,used as the feed of a six-effect METVC unit inwhich no preheaters are used because the feedseawater has been preheated by the stack gas. Thenet water production with and without stack heatrecovery are shown in Fig. 18. The effect of heatrecovery is significant, and the results indicatethat more than a 16% water production gain canbe obtained.

5. Conclusions

The analysis performed improves the under-standing of the combined STIG power and waterdesalination process and of ways to improve andoptimize it. Some specific conclusions are that:(1) a METVC desalination system is preferred toa MEE one when the pressure of the motivesteam is high enough, > ~3 bar, to run a steam jet

ejector, as in the two combined systems studied;(2) the steam injection rate in the STIG cycle hasa strong effect on water and power production,offering good flexibility for design and operation;(3) higher pressure ratios and higher steam injec-tion rates of the STIG cycle increase powergeneration, but decrease water production rates,and higher turbine inlet temperatures increaseboth power and water production; (4) a distinctwater production gain can be obtained by reco-vering the stack gas energy.

6. Symbols

a — Specific heat transfer area, m2/(kg/s)A — Heat transfer area, m2

c — Specific heat, kJ/( kg K)CR — Compression ratioe — Specific exergy, kJ/kgE — Exergy, kJec — Specific exergy consumption, kJ/(kg

distillate)ER — Expansion ratiog — Specific free enthalpy, kJ/kgh — Specific enthalpy, kJ/kgm — Mass, kgM — Molar mass, kg/kmoln — Number of effectsp — Pressure, MPaPR — Performance ratioq — Low heat value, kJ/kgQ — Heat energy, kJR — Gas constant, kJ/(kg K)Rpw — Power-to-water ratio, MW/MIGDs — Specific entropy, kJ/(kgK)S0 — Absolute entropy at 25EC and 1 atm,

kJ/(kmol K)T — Temperature, EC or KTc — Condensation temperature, ECTe — Evaporation temperature, ECTIT — Turbine inlet temperature (firing

temperature), ECTBT — Top brine temperature, EC


Uc — Heat-transfer coefficient of conden-sation, kW/(m2 EC)

Ue — Heat-transfer coefficient of evapora-tion, kW/ (m2 EC)

w — Specific work, kJ/kgW — Power, kJxj — Steam injection rate in STIG cycle xh — Humidification rate in EvGT cycleX — Salinity of saline water, ppmy — Mass fractionz — Molar fraction

Greek

β — Pressure ratio of gas turbine cyclege — Exergy efficiency of thermal cycle,

%gE — Exergy efficiency of steam jet ejec-

tor, %gD — Exergy efficiency of thermal desali-

nation unit, %ηt — Thermal efficiency of thermal cycle,

%χt — Heat loss rate, %χe — Exergy loss rate, %χE — Exergy loss rate of steam jet ejector,

%ξt — Heat recovery rate, %ξe — Exergy recovery rate, %

— Formation of enthalpy at 25EC and1 atm, kJ/ kmol

∆Tp — Pinch point temperature difference ofHRSG, EC

∆Tph — Temperature rise of seawater in pre-heater, EC

Subscripts

a — Airbr — Rejected brinecon — Condensate at end condensercool — Rejected cooling seawaterD — DesalinationE — Steam jet ejector

econ — End condenseref — Effects in METVC or MEE uniten — Entrained steamf — Fuelh — Heating steam hc — Condensate of heating steaml — Lossm — Motive steam min — Minimum net — Net productiono — Surroundingsr — Recoverysaline — Saline watersw — Seawaterw — Fresh water0 — Reference parameter for water and

power production

Superscripts

0 — Standard reference state in chemicalthermodynamics

Acknowledgement

The authors gratefully acknowledge the sup-port of the State Scholarship Fund of China to thefirst author.

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Appendix A: Calculation of the minimal workneeded in seawater desalination process

In a reversible process, the separation of saltand fresh water from seawater occurs at constanttemperature T and pressure p equal to that of thesurroundings, 30EC and 1 atm in this paper. Theminimal work required for the separation of waterfrom the saline solution is the difference in theGibbs energy [33]:

(A1)

where subscripts br, w and sw represent rejectedbrine, produced fresh water and feed seawaterrespectively, and g is specific Gibbs energy,

(A2)

In the desalination process, saline water,including rejected brine and feed seawater, can beconsidered as an ideal solution (at least as a firstapproximation). To simplify calculation further,the salt in the saline water is considered to be onlyNaCl. If the mass fractions of NaCl and water insaline water are yNaCl and yw, and mole fractionszNaCl and zw, are then based on the properties ofideal solution [34],

(A3)

(A4)

where RNaCl and Rw are the gas constants for NaCland water, respectively.

To ensure that different components in theseparation process have the same referencestandard, T0 = 298.15 K and p0 = 1 atm are takenas the reference state, and [35]:

(A5)

(A6)

(A7)

(A8)

where and S0 are enthalpy of formation andabsolute entropy at T0 and p0, and M is the molarmass. Then

(A9)

(A10)

(A11)


(A12)

Since NaCl and water are incompressible,

(A13)

(A14)

where c is the specific heat.

Appendix B: Heat-transfer coefficient corre-lations used in the METVC and MEE models[24]

(B1)

(B2)

where Ue and Uc are heat-transfer coefficients ofevaporation and condensation [in kW/(m2EC)],respectively, and Te and Tc (in EC) the tem-peratures of evaporation and condensation.

performance analysis of combined humidified gas...

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