Integrated Combined Heat and Power with Biomass Gasification and …Fryda-view... · CHP with Biomass Gasification and MGT Kurzfassung Integrierte Wärme- und Strom-erzeugung mit

66 VGB PowerTech 4/2008

CHP with Biomass Gasification and MGT

Kurzfassung

Integrierte Wärme- und Strom-erzeugung mit Biomassevergasung

und SOFC-Mikrogasturbine

Autotherme Biomassevergasung erfordert kei-

ne externe Wärmezufuhr, da diese mit Luft als

Vergasungsmittel betrieben wird, und produ-

ziert einen gasförmigen Brennstoff (Synthese-

gas), der gereinigt und konditioniert wird und

anschließend in eine oxidkeramischen Brenn-

stoffzelle (SOFC) genutzt werden kann. Vom

Konzept her kann das integrierte System unter

atmosphärischem oder unter etwas erhöhtem

Druck betrieben werden, und ermöglicht so

die Kombination mit einer Mikrogasturbine

(MGT). In dem vorliegenden Beitrag werden

drei kleine Kraft-Wärme-Kopplungssystme

(KWK) miteinander verglichen, die diese Tech-

nologien integrieren: a) Vergasung unter 4 bar

plus MGT, b) Vergasung unter 1,4 bar plus

SOFC und c) Vergasung unter 4 bar plus

SOFC-MGT. Jedes Haupt- und Zusatzbe-

standteil des KWK wurde mit Hilfe der As-

penplus™-Prozeßsimulationssoftware simu-

liert. Das interessanteste Ergebnis war, dass

sich das MGT-System gegenüber dem atmos-

phärischen System mit SOFC als effizienteres

herausgestellt hat. Beide Systeme wurden je-

doch durch die Leistung des kombinierten

SOFC-MGT-Systems übertroffen, das eine

exergetische elektrische Leistungsfähigkeit

von 35,6 % aufweist, mit einer aktiven SOFC-

Fläche von 100 m2 und einen nominalen Bio-

massedurchsatz von 200 kg/h. Die angewand-

te Exergieanalyse erlaubte die Optimierung

des SOFC-Brennstoff-Anwendungsfaktors (Uf)

und die Ermittlung der Auswirkung der Pro-

duktgaseanfeuchtung vor der Brennstoffzelle,

auf das Niveau der Leistungsfähigkeit und Ka-

pazität des Systems.

Introduction

There is an increasing trend to develop moreefficient biomass-fuelled energy systems.Steam cycles driven by biomass combustionin the range of 5 to 20 MWe currently pro-duce most of the bioelectricity worldwidewith electrical efficiencies around 20 %. Sol-id biofuel gasification was recently estab-lished as a significant option for power plantsbased on combined cycles that can achieveefficiencies of above 35 % in the ranges of20 to 40 MWe [1, 2]. In parallel, R&Dmoved towards efficient, small-scale CHPsystems incorporating gasification and use ofproduct gas in internal combustion engines,micro gas turbines or fuel cells. More specifi-cally there is a growing interest and promis-ing works on the combination of biomassgasification and Solid Oxide Fuel Cells(SOFCs) [3 to 6], mainly because these canutilise the product gas CO content and haveinherent higher tolerances towards severalproduct gas contaminants compared to otherfuel cell types.

The paper at hand presents an investigationon the combination of an air-blown fluidisedbed biomass gasifier with a high-temperatureSOFC and/or MGT in a CHP system of lessthan 1 MWe, which could respectively oper-ate at two pressure levels, near atmosphericand ~ 4 bar. The analysis is based on realisticperformance estimations by taking into ac-count the functionality of the proposed con-figurations without overestimations or grossassumptions. For this purpose, accurate mod-els were incorporated into the AspenplusTM

process simulation software for all the inte-grated unit operations, followed by an exer-getic analysis; second law efficiencies for themajor process steps and for the overall CHPsystem were evaluated and discussed, en-couraging comparisons with existing exer-getic analyses on biomass gasification [7]and SOFCs [8 to 14].

System Description and Modelling Aspects

Three possible configurations for a Com-

bined Heat and Power (CHP) system areshown in F i g u r e 1 , incorporating a flu-idised bed air-blown gasifier, a warm gascleaning train, an SOFC stack and its powerconditioning, and/or an MGT. These are:

a) near atmospheric gasification plus SOFC,

b)pressurised gasification (at 4 bar) plusSOFC-MGT,

c) pressurised gasification (at 4 bar) and MGT.

The peripheral unit operations include twoair compressors/blowers allowing flexibilityof operation of the gasifier and SOFC, gas-to-gas heat exchangers (HX 1, 2 and 3), aheat recovery steam generator (HRSG –HX4) producing saturated steam at around418 K, and (HX5) to make use of the flue gasthermal content. Olive kernel was consideredas the biomass fuel for the gasifier. Particu-lates are removed from the hot raw productby a ceramic filter while downstream tar re-moval takes place in a fixed catalytic bed. Tofacilitate halogen and sulphide removal insorbent beds, the gas is cooled down toaround 430 K [15]. Heat losses during gascleaning were accounted as 30 % losses fromthe heat transferred across HX1, in which theclean gas is reheated up to 900 K with thehelp of the hot product gas. For trouble-freeSOFC operation without carbon deposition,steam is added to increase the steam to car-bon ratio (STCR) of the product gas. SOFCair is preheated up to some 880 K by fluegas heat exchanger HX2. In the MGT option,HX2 serves as the recuperator air preheater.Depleted fuel and air from the SOFC react inthe post-cell combustion chamber providingheat for further cathode air preheating up to900 K. If the SOFC operates at increasedpressure, it is followed by an MGT expander,which produces additional power and coversthe compressor work demand. In the atmos-pheric operation an additional flue gas heatexchanger HX3 is employed to preheat airfor the gasifier; in the pressurised options airis preheated only through its compression.Flue gas HX4 provides saturated steamwhich is added to the product gas for its hu-midification before the SOFC stack. HX5produces useful heat through a heat transfermedium at the available off gas temperatureminus the minimum allowed �T across theheat exchanger. The SOFC system is basedon the existing tubular concept by Siemens-Westinghouse [16], with a 100 m2 active sur-face and nominal power output in the rangeof 200 to 250 kWe.

Dr. Mech.-Eng. Lydia Emilie Fryda*

Energy Research Centre of the Netherlands (ECN), Petten/The Netherlands

Dr.-Ing. Kyriakos D. Panopoulos

Laboratory of Steam Boilers and ThermalPlants, School of Mechanical Engineering,Thermal Engineering Section, National Technical University of Athens,Athens/Greece

Professor Dr.-Ing. Emmanuel Kakaras

Laboratory of Steam Boilers and ThermalPlants, School of Mechanical Engineering,Thermal Engineering Section, National Technical University of Athens, Athens/Greece

AutoorsAuthors

Integrated Combined Heat and Power with BiomassGasification and SOFC-micro Gas Turbine

* Dr. Fryda was awarded the Heinrich-Mandel-Prize 2007 for her work on the utilisation of biomass in decentral generation stations <1 MW with fluidised bed combustion systemsand the application of oxide-ceramic fuel cells.

066-074_PT4_08.qxd 15.04.2008 12:00 Uhr Seite 66

VGB PowerTech 4/2008 67


Steady state flow sheet models of the threeconfigurations were developed in Aspen-plusTM process simulation software togetherwith special FORTRAN blocks for the mod-elling of gasifier and SOFC sections. TheRedlich-Kwong-Soave cubic equation of statewas employed for the gas properties estima-tion [17]. The pressure drop across each unitoperation was assumed 1%. Heat exchangerswere allowed a significant minimum temper-ature difference of �T > 50 K in countercur-rent mode. This is a rather conservative as-sumption if higher efficiencies are demandedbut further reducing �Ts would result incostly heat exchangers. Heat losses in thegasifier, the HRSG and the heat exchangerswere assumed ≥ 2 % of the heat transferred.

Methodology for the Exergetic Analysis

Steady state flow sheet results were used toperform an exergy analysis on each sub-process, i.e. work potential of material andheat streams at any point in a series of energyconversion devices were evaluated in exergyterms. Exergy is the maximum possibleamount of work that can be obtained from amaterial or heat stream that eventually equili-

brates with the reference environment, whichconsists of reference components and is char-acterised by absence of pressure and temper-ature gradients.

The exergy of a material stream is given asthe sum of molar physical and chemical exer-gy:

E = N (�ph + �ch) (W) (1)

The molar physical exergy of a materialstream is evaluated using the data on physicalproperties, temperature (T), pressure (p),enthalpy (h) and entropy (s), calculated byAspenplusTM and its properties in referenceenvironmental conditions (To = 298.15 K,po = 1.013 bar) using the following expres-sion:

�ph = (h – ho) – To (s – so) (J mol-1) (2)

The molar chemical exergy is obtained whenthe components of the energy carrier are con-verted to reference compounds and diffuseinto the environment:

�ch = ��i�oi + RTo��i ln �i (J mol-1) (3)i i

where �i is the mole fraction and �oi is thestandard molar chemical exergy (J mol-1) ofeach component i, assuming a reference at-mospheric composition given by Kotas [18].

The chemical exergy of the solid fuel wascalculated 17689 kJ kg-1 with the help of the statistical correlation �, proposed bySzargut [19]:

�ch, fuel = (LHVfuel, dry + whfg) � + (�o

s – Huso zs (kJ kg-1) (4)

T a b l e 1 presents the proximate and ulti-mate analysis of the biomass fuel (dried olivekernel). The oxygen to carbon mass fractionof the solid fuel is calculated between 0.667≤ (zO2 / zc) ≤ 2.67, and the formula for woodis applied [19], which gives � =1.1182.

The exergy of a heat stream Q is given with the help of the Carnot factor: EQ

T =Q (1 – To / T), where T is the temperature at which Q is available. Exergy of power output, EW, equals power itself.

AspenPlusTM flow sheet calculations providevalues for mole flows N and mole fractions�i of all streams, as well as their physicalproperties (h, s, T, �). The evaluation of themolar reference enthalpy and entropy (ho, so)of every material stream was obtained bya duplicate of each stream expanded andcooled to reference conditions.

For the modeling of a realistic system, heat losses were introduced in several unit operations. These losses do not increase the

HCl, H2Sremoval

FB biomassgasifier 1080 K

Particulate filter

Char/Ash

Clean Product gas

900 K

CV 1

HX1

Biomass

Catalytic tardestruction

Flue gas outlet 363 K

SOFC 1173K

A C Inverter

Power

Flue gas to expander

1173 K

Post

combustor

Hot air 900K

Flue gas ~ 850 K

HX2 HX5

Usefulheat

Hot air 880 K

CV 2

Combustor

HX3

MGT option

Hot air

480 K

Atmosph. SOFC option

Air 293 K

Hot air 880 K

G

SOFCoptions

Pressurisedoptions

SOFCoptions

Atmosph.SOFCoption

SOFCoption

HX4

Water

Saturated steam

418 K

Air

Biomass/syngas

H2O

Flue gas

TIT temperature

Control air

CV 3

CV 4

Heatlosses

Figure 1. Flow sheet diagram of the biomass gasification CHP with SOFC and/or MGT.

066-074_PT4_08.qxd 15.04.2008 12:00 Uhr Seite 67



streams' entropy but contribute to exergy lossout of the system. Additionally, exergy lossesdue to mechanical and electrical inefficien-cies were considered; these deteriorate poweroutputs from equipment such as turboma-chinery and the inverter. Non-ideal condi-tions and real gas equations were used for thesimulation of each unit operation and exergylosses due to mixing were accounted; thesecannot be avoided and contribute to entropyincrease and therefore exergy destructionwithin the boundary of the system due to dis-sipation (irreversible exergy destruction).Both exergy losses and exergy destructionhave been summed up under the term 'irre-versibilities' IR, and an exergy balance for acontrol volume is expressed:

�Ej + �EQT = �Ek + �EQ

T, useful +IN IN OUT

+ �EW + IR (W) (5)OUT

Modelling of the Combined Heat and Power System

Model l ing Biomass Air Gasi f ica t ion

The Gibbs free energy minimisation methodfor the C-H-O atom blend of the biomass fueland oxidant mixture was applied for predict-ing the thermodynamic equilibrium composi-tion of product gas major components: H2,CO, CH4, CO2, H2O, N2, as well as char,which was considered as solid graphite (Cs).This thermodynamic calculation underesti-mates methane and unreacted char amountsin biomass gasification [1]. Therefore, non-equilibrium corrections were taken into ac-count to bring these product componentscloser to experimentally derived values. Theunreacted char was assumed to consist onlyof carbon and to be 5 % of the total fuel car-bon content [20], and this amount did notparticipate in the thermodynamic equilibriumcalculations. In a similar fashion, CH4 con-tent (mainly deriving from the decomposi-tion/pyrolysis of biomass) was assumed toreach 4 % v/v in the final nitrogen-free prod-uct gas [21].

The oxidising agent throughput determinesthe gasifier operating temperature. In order toachieve autothermal gasifier operation at adesired temperature, the air input was adjust-ed with a FORTRAN calculator allowing2 % heat loss. The air ratio with respect to

the gasifier fuel input is expressed consider-ing the amount required for stoichiometriccombustion of the fuel, as Equivalence Ratio

Air input (kg s-1)ER = –––––––––––––––––––––

Stoichiometric air (kg s-1) (6)

The gasification cold gas efficiency neglectsthe sensible heat of the gas and char pro-duced and is defined as [1]:

LHV in cold product gascg = –––––––––––––––––––––

LHV in feedstock (7)

In order to evaluate the degree of biomasswork potential conservation in the gaseousproduct fuel, exergy analysis of air gasifica-tion is applied for the control volume 1 (CV1) shown in Figure 1. Based on the generaldefinition of the degree of perfection for aprocess by Szargut [19], the exergetic effi-ciency of air gasification is:

Egas + Echarex, gas = –––––––––––––––––––––

Ebiomass + EOT + Eair (8)

Here Egas includes the product gas sensibleheat. The gasifier operates autothermally,therefore EO

T is zero. Since in this applicationthe physical and chemical exergy of char isof no use it was not included in the nomina-tor of equation (8).

Mode l l i ng t he SOFC

The SOFC configuration and its control vol-ume are depicted in Figure 1 (CV 2). A typi-cal tubular cathode supported SOFC similarto the Siemens Westinghouse system [16]was modelled. Both anode and cathode de-pleted fuel and air are assumed to exit theSOFC stack compartment at 1173 K and areintroduced to the post-cell combustor whichserves as a final air pre-heater. The flue gastemperature could increase significantly, butfor their use in an MGT expander furtheramount of compressed preheated air is intro-duced to suppress the turbine inlet tempera-ture (TIT) to its maximum allowed value, as-sumed 1173 K.

Two operation pressure levels were studiedcorresponding to the two gasifier operationpressure levels taking into account pressurelosses, i.e. pSOFC = 1.2 bar and 3.56 bar. Dueto very low methane and hydrocarbons con-centrations in the clean product gas, the com-mon internal pre-reformer was not employedhere [22]. Also contrary to natural gas fuelledSOFC configurations [23], no recirculation

of depleted anode gas is considered, due tohigh nitrogen content of the fuel gas thatwould significantly dilute the anode gas.Supplementary steam is added to the productgas before it reaches the catalytic SOFC an-ode, to ensure carbon deposition-free opera-tion. In the base case calculations, the addi-tional quantity of steam is specified toachieve a Steam to Carbon Ratio (STCR)equal to 2:

nH2OSTCR = –––––––––––––––– (9)

nCH4 + nCO + nCO2

Nevertheless the requirement for largeamounts of steam raises two problems: findingavailable water quantities and coping with ef-ficiency penalty deriving from exergy destruc-tion associated with steam production. Thelevel of the latter negative effect was investi-gated by setting STCR as a parameter rangingfrom 0.5 (which is the least thermodynamicrequirement to avoid carbon deposition) up tothe value of 2. Similarly high STCR valueshave been suggested when partially pre-re-formed methane is fed to SOFCs to assure nocarbon deposition will occur [23, 24].

The SOFC model was built in AspenplusTM

using available blocks and a calculator with aFORTRAN routine for the electrochemicalproperties estimation. The electrochemically-reacted oxygen is separated from the cathodeand fed to the anode, which is modelled byan RGIBBS reactor model that brings the an-ode mixture into chemical equilibrium. Thelow methane content justifies this equilibri-um assumption rather than using somemethane reforming rate reactions. The fuelutilisation factor of the stack is:

nH2,REACTUf = –––––––––––––––– (10)

nINN2 + nIN

CO + 4 · n INCH4

where n iIN refers to the anode's fuel species

input and nH2,REACT is the H2 (mol s-1) react-ing in the hydrogen electrochemical reaction,which was solely considered:

H2 + (1/2) O2 ↔ H2O (11)

The output voltage of the cell is:

V = VOC – VOHM – VACT – VPO (V) (12)

The Nernst open circuit cell voltage VOC wasevaluated at a corrected average operatingtemperature TSOFC, i.e. the average betweenthe mixed anode and cathode inlet flow (~ 900 K) and the outlet of the SOFC at 1173 K:

ΔGO R · TSOFC pH2out · (pO2

out)12

VOC = –––– + –––––––– · ln ––––––– (13)2 · F 2 · F pH2O

out

where F = 6.023�1023 �1.602�10-19 Cmol-1

is the Faraday constant, 2 is the number of e- produced per H2 mole that reactsthrough reaction (11) of which the molarGibbs free energy change is expressed as

Proximate analysis Ultimate analysis (%w/w dry basis)

Volatiles (%w/w dry) 72.64 C 51.19

Fixed carbon (%w/w dry) 24.78 H 6.06

Moisture (%w/w) 10.0 O 39.32

Heating values N 0.76

HHV (kJ/kg dry) 18 900 S 0.09

LHV 15 567 Ash 2.58

Table 1. Biomass fuel data.

066-074_PT4_08.qxd 15.04.2008 12:00 Uhr Seite 68


ΔGo

= ΔHo

– TSOFCΔSo

, calculated at TSOFC

and standard pressure. Finally, pouti are the

SOFC-exit partial pressures of the participat-ing components in reaction (11). Using theabove partial pressures and temperature data,a FORTRAN calculator was used for the esti-mation of the overpotentials due to Ohmic(VOHM), activation (VACT), and polarisation(VPO) losses. This calculator, the details of which were presented elsewhere [14], isbased on works from Campanari et al.[24, 25], Chan et al. [26, 27], Costamagna etal. [23], and Selimovic [28].

The SOFC stack's power output is:

PSOFC = V · I (W) (14)

where the current is evaluated as I =2FnH2,REACT. The corresponding current den-sity is J = I /ASOFC (Am-2) where ASOFC is theactive cell surface area (m-2).

By specifying the utilisation factor, the anodeflow throughput and composition, iterativecalculations of the overall energy balance areperformed over the SOFC stack control vol-ume to result in the air throughput adjustmentto reach an almost adiabatic operation (al-lowing ~ 5 kWth thermal losses) at the de-sired SOFC temperature. The electrical effi-ciency of the stack is:

PSOFCSOFC = –––––––––––––––––––––––––––––––– (15)

n INCH4

·LHVCH4 + nINH2 · LHVH2 + n IN

CO· LHVCO

where n iN

is the molar input of each gas com-ponent and LHVi their respective lower heat-ing value, while the exergetic electrical effi-ciency [19] is:

PSOFCex,SOFC = –––––––––– (16)

Egas + Eair

Mode l l i ng o f Hea t Exchange r s andMic ro Gas Tu rb ine

Heat exchangers were modelled with As-penplusTM HEATER modules. The microgas turbine was modelled as an expander us-ing common values for the maximum tur-bine inlet temperature and an isentropic effi-ciency value of 84 %. Mechanical efficien-cies were also taken into account. T a b l e 2shows the input data for all the peripheralequipments, including the power condition-ing inverter.

Results and Discussion

Gas i f i e r Ana ly s i s Resu l t s

F i g u r e 2 a shows the calculated Equiva-lence Ratio (ER) and the cold gas efficiencyof air gasification vs. gasification tempera-ture. The two gasification pressure levelsconsidered were 1.4 bar and 4 bar. It must benoted that for both operating pressure optionsthe air inlet temperatures are equal, approxi-

0

10

20

30

40

50

60

70

80

90

100

850 900 950 1000 1050 1100 1150 1200 12500.0

0.2

0.4

0.6

0.8

pressurisedatmospheric

ηcg

ER

ER

ηcg

in %

Tgas in K

0

5

10

15

20

25

30

vol i

n %

(dry

& N

2 fre

e)

850 900 950 1000 1050 1100 1150 1200 12500

20

40

60

80

100

Gasifier operating point

CO2

CH4

H2

CO


ηex, gas

ηex

, gas

in %

Tgas in K

Figure 2. (a) cold gas efficiency (%) and equivalence ratio (ER) of product gas, and (b) hot moist gas composition (% vol dry & N2 free) and pro-duct gas exergetic efficiency, for atmospheric and pressurised operation vs. gasifier temperature.

Equipment Input data

Temperatures (K) Heat losses

Heat exchangers Hot steam Cold steam % of heat

In Out In Out transfered

HX1 – – 430 900 30*

~ 940

HX2 or – – 880 2

> 1200**

HX2

HX3 hot steam – – 480 2

outlet

HX2/HX3 hot Water Steam

HX4 steam outlet – T = 293 K T = 418 2

P = 4 bar P = 4 bar

HX4 Heat transfer

HX5 hot stream 363 – medium 2

outlet at HX4 hot

stream outlet –

ΔTmin, pinch

(i.e – 50 K)

Isentropic Inlet Outlet Mechanical

Turbomachinery Efficiency temperature pressure Efficiency

(%) (K) (bar) (%)

MGT 84 1173 (max) 1.1 99.5

Air compressor/ 75 293 Downstream 98

blowers requirements

Inverter Efficiency

(%) 95

* accounting for losses through the gas cleaning system

** atmospheric cases (i.e. no MGT), temperature varies on Uf.

Table 2. Input data for peripheral equipment.


066-074_PT4_08.qxd 15.04.2008 12:00 Uhr Seite 69



mately 480 K. After the conversion of avail-able char, the gasification model results in al-most similar product gas composition forboth pressures ( F i g u r e 2 b ). There is al-ways some methane left even at higher tem-peratures and ERs because of methane cor-rection. From a thermodynamic point ofview, biomass air gasification processesshould be accomplished with the minimumair (ER) necessary above the requirement tomaximise carbon conversion. Increasing thegasifier temperature and therefore ER has anoverall negative effect on the exergetic effi-ciency because major chemical exergy carriercomponents, i.e. combustibles in the productgas, are minimised (Figure 2b). Nevertheless,kinetic reasons such as advancement of tarreforming reactions, fluidisation limitationsor heat losses might impose higher ER valuesin practice. The gasifier temperature waschosen 1080 K, and the corresponding ERvalue in both atmospheric and pressurisedmode of operation is 0.37. The model pre-dicts a very slight exergetic effectiveness in-crease in the case of pressurised gasification.

A higher moisture fuel would result in apenalty on the gasification efficiency becauseof dilution of the product gas with watervapour and requirements of higher ERs tosustain autothermal operation.

SOFC Ana ly s i s Resu l t s

The clean product gas main composition atboth pressure levels, before and after humidi-fication up to STCR = 2, are shown inT a b l e 3 . For these anode gas composi-tions, several characteristic curves are drawnfor TSOFC = 1173 K and pSOFC = 1.2 bar/3.56bar, respectively. Pressurised operation is im-proved over atmospheric, for a given fuelthroughput and utilisation factor, because (a)slightly increased current densities areachieved, (b) the SOFC voltage (V) is in-creased due to increased open circuit voltage(Voc), (c) activation overpotential is lesswhile (d) ohmic and cathode overpotentialsremain almost constant. F i g u r e 3 showsthe cell power output vs. current density forthree fuel utilisation ratios and the two pres-sure levels. SOFC power is increased forsmall Uf values because the partial pressuresof the reactants remain high until exitingthe stack. The 100 m2 SOFC stack, fed withproduct gas from the gasifier with biomassthroughput of 200 kg/h, operates with currentdensities J around 4000 A/m2; this region ofoperation was chosen for the rest of the cal-culations.

For finalising the configuration of the CHPsystem the preferred fuel utilisation has to bedetermined. The stack's air requirement (andtherefore off-gas volume) increases with Uf

because of greater cooling load required fromthe stack. Furthermore, the post-SOFC com-

bustor temperature de-creases with increas-ing Uf because theair/fuel mixture isleaner. F igu re 4shows the combustortemperature withoutadditional air, for thethree Uf values in thepressurised operationat the above-men-tioned system through-

put. Even with very high Uf = 0.85 some ad-ditional air is required to suppress the combus-tor temperature down to the maximum al-lowed turbine inlet temperature (TIT)(1173 K). Figure 4 also shows the required ad-ditional preheated air as percentage of the pri-mary SOFC air. Higher Uf (above 0.85) werenot considered because of risk of SOFC anodeoxidation. No additional air option was takeninto account in the atmospheric option whereno MGT-TIT limitation is posed.

F i g u r e 5 shows the comparison of SOFCexergetic efficiencies vs. SOFC power outputat different fuel utilisation ratios for the at-mospheric and pressurised operation. A firstremark is that SOFC stack efficiencies areconsiderably lower at atmospheric operation.Furthermore it is obvious that maximisingthe Uf does not necessarily result in higherstack efficiencies. Nevertheless this wrongassumption is very commonly taken forgranted in SOFC power cycles presentationswhich are based on gross assumptions aboutthe SOFC behaviour. Higher efficiencies aregained by lowering Uf at higher electric de-mands from an SOFC stack of a given active

surface area, and the Uf choice has to bebased on the examination of the overall CHPefficiency.

CHP Sys t em Ana ly s i s Resu l t s

After correcting the power outputs of theSOFC, PSOFC, and MGT, PGT, with mechani-cal and inverter efficiencies respectively, andsubtracting the power for compressors opera-tion, PCOMP, the energetic electrical efficien-cy of the system is defined as:

PSOFC + PGT – PCOMPel = ––––––––––––––––– (17)

(Input biomass) LHV

while the power and thermal energetic effi-ciency of the system is defined as:

PSOFC + PGT – PCOMP + QusefullCHP = ––––––––––––––––––––––––– (18)

(Input biomass) LHV

The system exergetic efficiency for electrici-ty production is:

PSOFC + PGT – PCOMPex,el = –––––––––––––––––– (19)

Ebiomass + Eair

and the combined electrical and thermal(CHP) exergetic efficiency is:

Product gas composition

(% vol)

Component Before After

steam addition steam addition

(1.4 bar/4 bar) (1.4 bar/4 bar

CH4 2.10 / 2.17 1.39 / 1.41

CO 16.80 / 18.1 11.03 / 11.76

CO2 11.16 / 10.65 7.40 / 6.95

H2 13.90 / 14.96 9.15 / 9.73

H2O 8.70 / 8.3 39.67 / 40.25

N2 47.20 / 45.7 31.33 / 29.87

Table 3. Product gas composition before and after steam addition.

Atmospheric Pressurised Pressurised

gasifier/SOFC gasifier – gasifier –

SOFC/MGT MGT

Biomass throughput

(kg/h) 200

Energetic/exergetic biomass

throughput 864 / 982

(kW)

Anode STCR 2 –

SOFC stack activa area

(m2) 100 –

Optimised Uf ~ 0.75 0.85 –

J

(A/m2) 3774 4280 –

Pnet

(kWel) 170.3 349.9 225.7

(PGT – PCOMP) / Pnet

(%) – 41.9 100

V / VOC

(Volt) 0.51 / 0.80 0.52 / 0.80 –

el / CHP

(%) 20.0 / 62.3 40.6 / 58.1 26.1 / 70.7

ex,el / ex,CHP

(%) 17.6 / 39.7 35.6 / 40.6 23.0 / 38.5

Table 4. Base case results for the three configurations studied.

066-074_PT4_08.qxd 15.04.2008 12:00 Uhr Seite 70


PSOFC + PGT – PCOMP + EQuseful

ex,CHP = ––––––––––––––––––––––––– (20)Ebiomass + Eair

The electrical exergetic efficiencies of thethree examined configurations for a range offuel utilisation ratios versus biomass through-put are shown in F i g u r e 6 . For a greatrange of fuel throughputs, and in specificaround the base case of 1000 kWth fuel exer-gy input (i.e. 200 kg biomass/h) the MGT op-tion has higher electrical efficiencies com-pared to the atmospheric SOFC system, de-spite its optimised Uf for that region of oper-ation (at Uf = 0.75). The atmospheric SOFCsuffers from low effectiveness to producepower together with great exergy losses asso-ciated with nil power extraction from thehigh temperature SOFC flue gases. The larg-er air compression power consumption of the

SOFC-MGT system is more than offset byincreased SOFC performance and additionalpower produced from the MGT resulting inex,el ≥ 35.6 % at biomass exergy input1000 kW. For the three systems studied andthe base case of 200 kg biomass/h, the effi-ciency ratios (i.e. equations 17 to 20) and im-portant results are presented in T a b l e 4 .The pressurised MGT configuration offersincreased thermal output but at low tempera-tures ( F i g u r e 7 c ), thus not reflected in theex,CHP.

Part of the flue gas energy of the SOFC con-figurations is consumed to produce steamfor product gas moistening, therefore, dete-riorates the quality of available heat ( F i g -u r e 7 a , b ). The influence of the requiredSTCR on the electrical and CHP system

exergetic efficiencies was studied paramet-rically. In order to present the exergy de-struction associated with steam production,two exergy loss ratios, steam and compr weredefined:

IRsteamsteam = ––––– (21)

EIN

where IRsteam is the cumulative irreversibilityassociated to steam production and its mixingwith the syngas prior to entering the anode,and was calculated from a combined exergybalance applied to control volumes CV3 andCV4 in Figure 1:

Ewater + Eflue gs IN + Esyngas IN = CV3 CV3 CV4

Eanode input + Eflue gas OUT + IRsteamCV4 CV3 (22)

0

50

100

150

200

250

300

0 1000 2000 3000 4000 5000 6000 7000


Uƒ = 0.650.75

0.85

0.650.75

0.85

P SOFC

in k

We

J in A/m2

Figure 3. SOFC power output (kWel) vs. current density (A/m2) foratmospheric and pressurised operation and three fuel utilisation ratios.

0

200

400

600

800

1000

1200

1400

1600

0.850.750.65

Uƒ

0

50

100

150

200

250

300

350

400

450

combustor T

additional air

TIT

Air a

dditi

on in

% o

f SOF

C re

quire

d

Post

-cel

l com

bust

or T

in K

(with

out a

dditi

onal

air)

Figure 4. Post-SOFC combustor temperature (K) and additional airexpressed as (%) of SOFC air for the pressurised operationvs. fuel utilisation ratio.

0

5

10

15

20

25

30

35

0 50 100 150 200 250 300


0.65

0.75

Uƒ = 0.85

0.85

0.750.65

ηex

, SOF

C in

%

PSOFC in kWe

Figure 5. SOFC exergetic efficiency vs. SOFC power output (kWe)for atmospheric and pressurised operation and three fuelutilisation ratios.

EIN in kW

0

5

10

15

20

25

30

35

40

45

200 400 600 800 1000 1200 1400

0.75

0.65

0.65


Uƒ = 0.85

Uƒ = 0.85

0.75

ηex

, el

in %

ηex, el (MGT)

Figure 6. Electric exergetic efficiencies of the three system configu-rations vs. biomass exergy throughput (kW).


066-074_PT4_08.qxd 15.04.2008 12:00 Uhr Seite 71



The second exergy loss ratio is associatedwith fuel cell compressor power consump-tion, because it is indirectly affected from thelevel of steam addition since the laterchanges the cooling demands of the stack:

Pcomprcompr = ––––– (23)

Ein

F i g u r e 8 a (pres-surised system)shows that the elec-trical efficiencymarginally deterio-rates with STCR be-cause despite steam

grows there is acounterbalancing re-duction of the compr,since the additionalsteam cools thestack replacing con-siderable amountsof air for this pur-pose. In all cases,the thermal through-put from the MGT

and therefore its contribution to power outputis constant (line PGT/Ein in Figure 8a) and theminor decrease in the electrical efficiency withSTCR is attributed to lower partial pressuresof reactants within the stack. Similarly, the lat-ter effect is evident in Figure 8b for the atmos-pheric operation. In both atmospheric andpressurised configurations the effect of the

level of additional steam requirement is re-flected on the CHP efficiencies (lines ex,CHP

and steam almost follow parallel trends) be-cause this affects the temperature at whichuseful heat is available from HX5.

Conclusions

The combination of biomass air gasificationwith SOFC and/or MGT for small-scale CHPwas assessed by modelling in AspenplusTM

process simulation software. Two system op-eration pressures were studied, atmosphericand ~ 4 bar. This small pressure shift doesnot have significant effect on the product gascomposition or on the exergetic efficiency ofthe gasification process. On the contrary, thepressurised SOFC operation is greatly im-proved, and with the additional power froman MGT expander, achieves the highest effi-ciencies ex,el ≥ 35 % at J values around 4000A/m2. The Uf was optimised at 0.85 for thepressurised and at 0.75 at atmospheric SOFCoperation for biomass throughputs of around200 kg/h. The atmospheric SOFC configura-

0

200

400

600

800

1000

1200

HX2

COM

PR.

AirQ

Flue gasoutlet

HX5

MGT

T in

K

Figure 7. Heat exchangers temperature profile for (a) pressurisedSOFC-MGT, (b) atmospheric SOFC and (c) pressurisedMGT configuration.

50

40

30

20

10

0

– 10

– 20

– 30

0.0 0.5 1.0 1.5 2.0 2.5

η &

ζ ra

tios

%

STCR

- ζ steam

- (ζ steam + ζ compr)

ηex, el

ηex, CHP

50

40

30

20

10

0

– 10

– 20

– 300.0 0.5 1.0 1.5 2.0 2.5

η &

ζ ra

tios

%

STCR

PGT / E in

- ζ steam

- (ζ steam + ζ compr)

ηex, CHP

ηex, CHP

Figure 8. Exergetic efficiencies and loss ratios vs. STCR for (a) pressurised, and (b) atmospheric operation with Uf = 0.85 (biomass throughput 200 kg/h).

0

200

400

600

800

1000

1200

HX2Post-cell

com-buster

Air

Q

Flue gasoutlet

HX5

Evap. Preheater

H2OT in

K

HX4 (HSRG)

0

200

400

600

800

1000

1200

MGT

HX2Post-cell

com-buster

HX4 (HSRG)

COM

PR.

Air Q

Flue gasoutlet

HX5

Evap. Preheater

H2OT in

K

a a

b b

c

066-074_PT4_08.qxd 15.04.2008 12:00 Uhr Seite 72


tion results in considerably lower efficienciesthan the simpler pressurised gasificationMGT configuration which gives ex,el =23 %, and could only surpass this efficiencyif very low power densities were employed.Such a system would probably not be eco-nomic, i.e. to have the high SOFC-related in-vestment costs without significant revenuesfrom the power production. Through a de-tailed exergetic parametric analyses it wasshown that the increase of additional steamproduction to achieve a desired STCR doesnot greatly affect the electrical efficiencies ofthe SOFC configurations but is negatively re-flected on the combined thermal exergetic ef-ficiencies of the CHP systems because it dete-riorates the temperature at which off gases arefinally available for useful heat production.

References

[1] Higman, C., and Van Der Burg, M.: Gasifi-cation, Gulf Publishing, 2003.

[2] Maniatis, K., and Millich, E.: Energy frombiomass and waste: the contribution of utili-ty scale biomass gasification plants, Bio-mass and Bioenergy 1998; 15(3), 195-200.

[3] Baron, S., Brandon, N., Atkinson, A., Steele,B., and Rudkin, R.: The impact of wood-deri-ved gasification gases on Ni-CGO anodes inintermediate temperature solid oxide fuelcells, J. Power Sources 2004; 126(1-3), 58-66.

[4] Omosun, O., Bauen, A., Brandon, N. P., Adjiman, C. S., and Har,t D.: Modellingsystem efficiencies and costs of two bio-mass-fuelled SOFC systems, J. Power Sour-ces 2004; 131(1-2), 96-106.

[5] Vasileiadis, S., and Vasileiadoum Z. Z.:Biomass reforming process for integratedsolid oxide-fuel cell power generation, Che-mical Engineering Science 2004; 59(22-23),4853-4859.

[6] Singh, D., Hernández-Pacheco, E., Hutton,P. N., Patel N., and Mann, M. D.: Carbondeposition in an SOFC fuelled by tar-ladenbiomass gas: a thermodynamic analysisJournal of Power Sources 2005; 142(1-2),194-199.

[7] Prins, M.J., and Ptasinski, K.J.: Energy and exergy analyses of the oxidation and ga-sification of carbon, Energy 2005; 30(7),9821002.

[8] Bedringås, K. W., Ertesvåg, I. S., ByggstøylS., and Magnussen Bj. F.: Exergy analysisof solid-oxide fuel-cell (SOFC) systems,Energy 1997; 22(4), 403-412.

[9] Monanteras, N. C., and Frangopoulos, C.A.: Towards synthesis optimization of a fu-el-cell based plant, Energy Conversion andManagement 1999; 40(15-16), 1733-1742.

[10] Van den Oosterkamp, P. F., Goorse, A.A.,and Blomen, L.J.M.J.: Review of an energyand exergy analysis of a fuel cell system, J.Power Sources 1993; 41(3), 239-252.

[11] De Groot, A.: Advanced exergy analysis ofhigh temperature fuel cell systems, PhDThesis, Petten: Energy Research Centre ofthe Netherlands; 2004.

[12] Chan, S. H., Low, C. F., and Ding, O. L.:Energy and exergy analysis of simple solid-

oxide fuel-cell power systems, J. PowerSources 2002; 103(2), 188-200.

[13] Hotz, N., Senn, S. M., and Poulikakos, D.:Exergy analysis of a solid oxide fuel cellmicropowerplant, Journal of Power Sources,Volume 158, Issue 1, 14 July 2006, pp 333-347.

[14] Panopoulos, K.D., Fryda, L., Karl, J.,Poulou, S., and Kakaras, E.: High tempera-ture solid oxide fuel cell integrated with no-vel allothermal biomass gasification: Part II:Exergy analysis, Journal of Power Sources,Volume 159, Issue 1, 13 September 2006,pp. 586-594.

[15] Kapfenberger, J., Sohnemann, J., Schleitzer,D., and Loewen, A.: Acid gas removal bycustomised sorbents for integrated gasifica-tion fuel cell systems, 5th InternationalSymposium on Gas Cleaning at High Tem-peratures. U.S. Department of Energy Na-tional Energy Technology Laboratory, Mor-gantown, WV, September 17-20, 2002(available from: http://www.netl.doe.gov/publications/).

[16] Veyo, St. E.: Siemens Westinghouse PowerCorporation. Tubular SOFC Hybrids: Pre-sent and Prospect. In: Second DOE/UN In-ternational Conference and Workshop onHybrid Power Systems April 16-17, 2002.(available from: http://www.netl.doe.gov/).

[17] AspenPlus® Physical Property Methods andModels Reference Manual, Aspentech®, 1999.

[18] Kotas, T.J.: The Exergy Method of ThermalPlant Analysis, Krieger Publishing Com-pany, Malabar, Florida, 1995.

[19] Szargut, J., Morris, D. R., and Steward, F.R.: Exergy Analysis of Thermal, Chemical,and Metallurgical Processes, Taylor & Fran-cis Inc, 1988.

®

SUPERBOLT®

SUPERBOLT® – stark und clever

Für Schraubverbindungen,

die länger halten:

– sicher

– dauerfest

– lösbar

Mechanische Spannelemente

· für grosse Gewinde· torsionsfrei· ohne Spezialwerkzeug· höchste Spannkräfte· einfach und schnell, auch beim Lösen

· ohne Setzungen· höchste Sicherheit· niedrige Kosten· Standardprogramm: M20–M160

· Spezialausführungenbis M1000 und grösser

P&S Vorspannsysteme AGCH-8735 St. Gallenkappel · Schweiz

T: +41 55 284 64 64 · F: +41 55 284 64 [email protected] · www.p-s.ch

Dampfturbine – Verspannung einer

Teilfuge

Dampfturbine – Verspannung der

Wellenkupplung

066-074_PT4_08.qxd 15.04.2008 12:01 Uhr Seite 73

Heinrich-Mandel-Preis für Kraftwerkstechnik der VGB-FORSCHUNGSSTIFTUNGDie VGB-FORSCHUNGSSTIFTUNG wurde 1970 zur Förderung vonGemeinschaftsforschung auf dem Gebiet der Strom- und Wärmeerzeu-gung gegründet. Sie ist eine gemeinnützige Stiftung nach deutschem Pri-vatrecht. Der Vorsitzende der VGB-FORSCHUNGS-STIFTUNG ist Dr. Gerd Jäger. Das Stiftungskuratorium besteht aus insgesamt zehn Mit-gliedern, die Betreiber und Hersteller von Anlagen zur Erzeugung vonStrom- und Wärme sowie auf diesem Gebiet tätige Hochschulinstitute re-präsentieren.

Mit dem Heinrich-Mandel-Preis für Kraftwerkstechnikzeichnet die VGB-FORSCHUNGS-STIFTUNG seitüber 25 Jahren herausragende Leistungen junger Inge-nieurinnen und Ingenieure aus. Der Preis ist mit 10.000Euro dotiert und wird jährlich vergeben. Ausgezeich-net werden Arbeiten aus allen Gebieten der Kraftwerkstechnik. Dabeiwerden Kraftwerksbetrieb (einschließlich Planung und Bau) und zuzuord-nende Forschung als gleichwertig beurteilt. Das Höchstalter der Kandida-tinnen und Kandidaten liegt bei 35 Jahren.

Das Kuratorium der VGB-FORSCHUNGSSTIFTUNG vergab den Heinrich-Mandel-Preis für Kraftwerkstechnik 2007 zu gleichen Teilen an

– Dr. Lydia-Emilia Fryda für ihre Arbeiten zur Nutzung von Biomasse indezentralen Stro-merzeugungsanlagen < 1 MWel unter Einsatz einerWirbelschichtvergasung und von oxid-keramischen Brennstoffzellenund

– Dr. Jens Hampel für die Entwicklung eines Turbogenerators mit mechatronischer Netz-kopplung zur Wirkungsgraderhöhung kleinerDampfturbinen.

Die beiden Preisträger stellen die ausgezeichneten Arbeiten in dieserAusgabe der VGB PowerTech vor.

Heinrich-Mandel Prize for Power Plant Technology of the VGB Research FoundationThe VGB Research Foundation (VGB-FORSCHUNGSSTIFTUNG) wasfounded in 1970 to foster joint research in the field of power and heat generation. As a foundation of German private law it pursuessolely non-profit purposes. Chairman of the foundation is Dr. Gerd Jäger.The curatorship consists of ten members representing power plant opera-tors, the power plant manufacturing industry as well as university insti-tutes working in this field.

The VGB Research Foundation has rewarded outstand-ing performances of young engineers in the area ofpower plant engineering with the Heinrich-MandelPrize for more than 25 years. The prize is endowedwith 10,000 and is awarded annually. Eligible areworks from all fields of power plant engineering.

Equal significance is attributed to power plant operation (including plan-ning and construction) and related research. Limiting age of the candi-dates is at 35 years.In 2007, the Board of Trustees of the VGB-FORSCHUNGSSTIFTUNGequally awarded the Heinrich-Mandel Prize for Power Plant Technologyto– Dr. Lydia-Emilia Fryda for her works on the utilisation of biomass in

decentral electricity generation plants < 1 MWe with fluidised bedgasification and oxide ceramic fuel cells and to

– Dr. Jens Hampel for the development of a turbo generator with mecha-tronic link to the electrical grid for increasing the efficiency of smallersteam turbines.

Both prize winners present the awarded works here.


[20] H. Morita, F., Yoshiba, N., Woudstra, K.,Hemmes, and Spliethoff, H.: Feasibility stu-dy of wood biomass gasification/molten car-bonate fuel cell power system -comparativecharacterization of fuel cell and gas turbinesystems, J. of Power Sources 2004; 138(1-2); 31-40.

[21] Kakaras, E., Vourliotis, P., Panopoulos, K.D., and Fryda, L.: Cotton residue gasificati-on tests in lab scale fluidised bed. Clean Air2003. Seventh International Conference onEnergy for a Clean Environment, 7 - 10 July2003, Calouste, Gulbenkian Foundation,Lisbon, Portugal.

[22] Peters, R., Riensche, E., and Cremer, P.:Pre-reforming of natural gas in solid oxidefuel-cell systems, J. Power Sources 2000;86(1-2), 432-441.

[23] Costamagna, P., Magistri, L., and Massar-do, A. F.: Design and part-load performanceof a hybrid system based on a solid oxidefuel cell reactor and a micro gas turbine, J.Power Sources 2001; 96(2), 352-368.

[24] Campanari ,S.: Thermodynamic model andparametric analysis of a tubular SOFC module, J. Power Sources 2001; 92(1-2), 26-34.

[25] Campanari, S., and Iora, P.: Definition and sensitivity analysis of a finite volumeSOFC model for a tubular cell geometry,Journal of Power Sources 2004; 132(1-2),113-126.

[26] Chan, S. H., Khor, K. A., and Xia, Z. T.: Acomplete polarization model of a solid oxidefuel cell and its sensitivity to the change ofcell component thickness, J. Power Sources2001, 93(1-2), 130-140.

[27] Chan, S. H., Low, C. F., and Ding, O. L.:Energy and exergy analysis of simple solid-oxide fuel-cell power systems, J. PowerSources 2002; 103 (2), 188-200.

[28] Selimovic, A.: Modelling of Solid OxideFuel Cells Applied to the Analysis of Integra-ted Systems with Gas Turbines, DoctoralThesis, Department of Heat and Power Engineering, Lund University, Sweden, 2002.

NomenclatureASOFC SOFC active surface (m-2)E Total exergy of a material stream (W)EQ

T Thermal exergy of a heat streamavailable at temperature T (W)

EW Work or power output (W)ER Equivalence ratioF Faraday constant

6.023�1023 �1.602�10-19 Cmol-1

h Enthalpy of a stream (J mol-1)ho Standard enthalpy at environmental

conditions (J mol-1)hfg Latent heat of water vaporisation

(2.442 kJ/kg)Ho

us Sulphur lower heating value (9.259kJ/kg)

I, J SOFC current and current density (A, A m-2)

IR Irreversibility of a process (W)LHV Fuel Low Heating Value (kJ kg-1 for

solids / MJ mn-1 for gases)

LHVfuel,dry Lower heating value, dry (17,567kJ/kg)

N Mole flow rate (mole s-1)ni Mole flow rate (mol s-1) of compo-

nent ip Pressure (bar) po Standard pressure = 1 (atm) PSOFC Direct current electric power

produced from the SOFC (W)PCOMP Air compressor power (W)PGT Gas turbine power output (W)Q Heat stream (W) R Universal gas constant

(8.314 kJ kmol-1 K-1)s Entropy of a stream (Jmole-1K)so Standard entropy at environmental

conditions (J mol-1)STCR Steam to carbon ratio

(refers to product gas) T Temperature (K)

To Standard temperature (K)Uf SOFC fuel utilisation factor VOC Open circuit (Nernst)

SOFC voltage (V)VOHM Ohmic SOFC voltage

over potential (V)VACT Activation SOFC voltage

over potential (V)VPO Polarisation SOFC voltage

over potential (V)�i Mole fraction of component i w Moisture mass fraction in fuel

(0.1 w/w)zs Sulphur mass fraction in fuel,

dry (0.09 w/w)

Subscripts / Superscripts cg cold gas CHP Combined heat and power overall

system el electrical ex exergetic gas Gasification IN Input OUT Output TIT Turbine Inlet TemperatureGreek symbols� Statistical correlation for solid fuel

exergy calculation�oi Standard chemical exergy of a com-

ponent i in a mixture (J mol-1)�ph Specific physical exergy of a material

stream (J mole-1)�ch Specific chemical exergy of a materi-

al stream (J mole-1)�o

s Chemical exergy of sulphur (18,676 kJ/kg)

�ch,fuel Fuel chemical exergy (kJ/kg) Exergy loss ratio associated with a

sub process IR over total Ein Efficiency �

066-074_PT4_08.qxd 15.04.2008 12:01 Uhr Seite 74

VGB-PowerTech-DVDDVDMehr als 10 000 Seiten Daten, Fakten und Kompetenz digital

(einschließlich Recherchefunktion über alle Dokumente)

Bitte ausfüllen und per Post oder Fax zusenden

Ich möchte die „VGB-PowerTech-DVD1990 bis 2007“ (Einzelplatzversion) bestellen.

950,– Euro (Abonnent VGB PowerTech Journal 1)

1950,– Euro (Nicht-Abonnent VGB PowerTech Journal 2)

zuzüglich Versand, Inland 7,50 Euro, und MwSt.

Netzwerklizenz (Corporate License), VGB-Mitglieder-version (InfoExpert) sowie Lizenz Forschung und Lehreauf Anfrage (Tel. +49 201 8128-200)

Jährliches Update: 1) 150,– Euro; 2) 350,– EuroDas Update muss jährlich bestellt werden.

Per Fax oder im Fensterkuvert an die

VGB PowerTech Service GmbHFax-Nr. +49 201 8128-329

InfoExpert

Name, Vorname

Straße

PLZ Wohnort

Telefon/Fax

Datum 1. Unterschrift

Widerrufsrecht: Diese Vereinbarung kann ich innerhalb von 14 Tagen bei der

VGB PowerTech Service GmbH, Essen, widerrufen. Zur Wahrung der Frist

genügt die rechtzeitige Absendung.

Ich bestätige dies durch meine 2. Unterschrift.

Datum 2. Unterschrift

�

VGB PowerTech – www.vgb.orgIn allen Teilen der Welt ist die Erzeugung von Strom und die Bereitstellung von Wärme

ein zentrales Thema von Technik, Wirtschaft und Politik sowie des alltäglichen Lebens.

Experten sind dabei zuständig für den Bau und Betrieb von Kraftwerken, deren Ent-

wicklung und Überwachung sowie die vielfältigen Aufgaben von Service und Manage-

ment.

Hier ist die VGB PowerTech die kompetente internationale Fachzeitschrift für den Kraft-

werksbereich. Jährlich 11 Ausgaben der deutsch-englischen Gemeinschaftsausgabe in-

formieren mit Fachbeiträgen und Nachrichten zu allen wichtigen Fragen

der Strom- und Wärmeerzeugung.

Die VGB PowerTech erscheint bei VGB PowerTech Service GmbH,

Verlag technisch-wissenschaftlicher Schriften.

Herausgeber ist der europäische Fachverband VGB PowerTech e.V.

VGB-PowerTech-DVD 1990 bis 2007:Fachbeiträge der VGB Kraftwerkstechnik und VGB PowerTech digital

Auf mehr als 10000 Seiten finden Sie das kompetente Fachwissen aus 18 Jahren

VGB Kraftwerkstechnik (dt. Ausgabe bis 2000) und der Internationalen Fachzeitschrift

VGB PowerTech (ab 2001) mit:

– über 2300 Fachartikeln,

– allen Dokumenten im PDF-Format (bis zum Jahr 2000 aus technischen Gründen als

nachbearbeiteter, recherchierbarer S/W-Scan),

– einer komfortablen Suchfunktion über alle Beiträge, als Volltextsuche bzw. gezielter

Suche nach Autoren und Dokumententiteln.

Navigieren Sie schnell zu den gesuchten Beiträgen mit wenigen Mausklicks.

Die VGB-PowerTech-DVD ist erhältlich als Einzelplatzversion oder Netzwerklizenz für

Unternehmen, Forschung und Behörden.

Die Einzelplatzversion können Sie per Formular auf dem Post-/Faxweg bestellen, oder

Sie nutzen unseren Online-Shop unter www.vgb.org.

Ein Angebot über eine Netzwerklizenz unterbreiten wir Ihnen gerne auf Anfrage.

Mit dem VGB-PowerTech-Update können Sie Ihre DVD jährlich aktualisieren.

Ihr Ansprechpartner bei der VGB PowerTech Service GmbH:

Jürgen Zimander, Tel.: +49 201 8128-200, E-Mail: [email protected]

VGB PowerTech Service GmbHPostfach 10 39 32

45039 Essen

DEUTSCHLAND

InfoExpert

Documents

Integrated Combined Heat and Power with Biomass Gasification and …Fryda-view... · CHP with Biomass Gasification and MGT Kurzfassung Integrierte Wärme- und Strom-erzeugung mit