- Email: [email protected]
An energetic–exergetic analysis of a residential CHP system based on PEM fuel cell
Applied Energy 88 (2011) 4334–4342
Contents lists available at ScienceDirect
Applied Energy journal homepage: www.elsevier.com/locate/apenergy
An energetic–exergetic analysis of a residential CHP system based on PEM fuel cell L. Barelli ⇑, G. Bidini, F. Gallorini, A. Ottaviano Department of Industrial Engineering, University of Perugia, Via G. Duranti 1/A4, Perugia 06125, Italy
a r t i c l e
i n f o
Article history: Received 13 December 2010 Received in revised form 23 March 2011 Accepted 11 April 2011 Available online 13 July 2011 Keywords: PEM fuel cell Exergy analysis Residential CHP
a b s t r a c t The use of fuel cell systems for distributed residential power generation represents an interesting alternative to traditional thermoelectric plants due to their high efficiency and the potential recovering of the heat generated by the internal electrochemical reactions. In this paper the study of a micro cogenerative (CHP) energy system based on a Proton Exchange Membrane fuel cell (PEMFC) is reported. With the aim to evaluate the performance and then the feasibility of this non-conventional energy system, in consideration of thermal and electrical basic demand of a multifamily apartment blocks, a zerodimensional PEMFC model in Aspen Plus environment has been developed. A simulations sequence has been carried out at different operating conditions of the fuel cell (varying temperature, pressure and relative humidity). Subsequently, on the basis of the obtained results, an energy/exergy analysis has been conducted to define the optimal operating conditions of the PEMFC that ensures the most efficient use of the energy and exergy inputs. Ó 2011 Elsevier Ltd. All rights reserved.
1. Introduction In the actual energetic background in which a better use of primary resources is the principal problem, the cogeneration (CHP) can be a practicable solution for users who are sensitive to energy savings. In this context, to maximize the use of primary energy and avoid any waste of money, economic feasibility studies which take account of the energy market [1] and the type of application required by the market are of utmost importance [2]. Specifically, the CHP plants based on fuel cells (FCs) represent an interesting alternative to traditional technologies for the combined production of electrical energy and heat; this is due to: the high electrical efficiency of the FCs in part load, the modularity, the ability to ensure substantial autonomy of the user from the electrical distribution network and the possible reduction of the environmental impact. As regards fuel cells, the chemical and thermodynamic aspects of the characteristic processes [3,4] and the basic research on materials to enhance performance and durability [5,6] are under study in many research activities. At the same time, even the design and analysis of performance of CHP systems based on FCs (mostly prototypes are used to demonstrate the potential of such a technology) are subjects of many publications [7–10]. In view of a micro-cogeneration distributed production destined to residential users who require thermal loads characterized by low temperatures, the subject of this research has been identified in a high-efficiency CHP energy system based on a polymer ⇑ Corresponding author. Tel.: +39 075 5853740; fax: +39 075 5853736. E-mail address: [email protected] (L. Barelli). 0306-2619/$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.apenergy.2011.04.059
electrolyte fuel cell (PEMFC). This type of cell is currently the one with the most advanced technological development and some cogenerative units are already commercialized [11,12]. The purpose of the present paper is to evaluate the performance and, hence, the feasibility of a non-conventional energy system that can help to meet the thermal and electrical demand of residential consumers. To this aim a zero-dimensional model of such a system has been developed in the Aspen Plus environment, with the intent to analyse its performance under different operating conditions (temperature, pressure and relative humidity). The model simulations have been conducted by setting as target the fulfilment of a fixed electrical load (3 kW) that can help to cover the basic demand of a multifamily apartment block, such as the electrical auxiliary facilities, and by evaluating, as a result, the recoverable heat for the user. The first part of the study, described in Section 2 of the paper (system layout), has been focused on the identification of the optimal plant lay-out for the CHP unit under study and on the modelling of such a system in the Aspen Plus environment [13]. Subsequently, the performance analysis methods, based on the first and second thermodynamic law efficiency, have been defined in Sections 3 and 4. At this step, particular attention has been devoted to exergetic analysis by referring to the literature guidelines [14,15]. In Section 5 the simulations made in Aspen Plus have been presented and then the results, achieved by means of the analysis methodology cited above, have been discussed. In future studies, it will be interesting to analyse the behaviour of the PEMFC system at variable electrical and thermal load using typical load curves of residential users.
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
4335
Nomenclature Cp e eCH eCH n ePH EX ExQ h HHVH2 K P PEMFC Psat Pvap P0 Q Qair QCOOL1
specific heat capacity at constant pressure (kJ/(kmol K)) molar stream exergy (kJ/(kmol K)) molar stream chemical exergy (kJ/(kmol K)) molar chemical exergy of a single chemical species (kJ/ (kmol K)) molar stream physical exergy (kJ/(kmol K)) stream total exergy (kW) exergy associated to the recovered thermal power (kW) molar stream enthalpy (kJ/kmol) hydrogen higher heating value (kJ/kmol) ratio between specific heat capacity at constant pressure and specific heat capacity at constant volume pressure (atm) polymer electrolyte membrane fuel cell water saturation pressure at a certain temperature (atm) water partial pressure (atm) restricted dead state pressure (atm) total thermal power recovered in cogenerative arrangement (kW) thermal power delivered to feeding air flow (kW) thermal power recovered in the heat exchanger COOL1 (kW)
2. System layout The system layout considered for the simulations is depicted in Fig. 1. It is characterized, as well as the PEM electrochemical system described in [3], by heat exchangers that allow the FC cooling and the exhaust heat recovery for both the internal regeneration, the preheating of the cell feeding flows and the cogenerative purposes. Specifically the model developed by the authors in [3] and adopted in this paper is valid for the following fuel cell operating conditions:
Cell temperature: 70–120 °C. Pressure: 1–3 atm. Relative humidity: 35–100%. CO content in the feeding gas: 0–200 ppm.
Qfuel R RH s T Tm T0 Ua Uf xw xn Wfuel We n_ g(I) g(I)cog
g(II) g(II)cog gel
thermal power delivered to feeding fuel flow (kW) ideal gas constant (kJ/(kmol K)) relative humidity (%) molar stream entropy (kJ/(kmol K)) temperature (K) average temperature inside an heat exchanger (K) restricted dead state temperature (K) air utilization factor fuel utilization factor water molar fraction in the feeding streams mole fraction of a single chemical species chemical power supplied through the fuel (hydrogen) (kW) produced electric power (kW) molar flow rate (kmol/s) first Law of Thermodynamic efficiency first Law of Thermodynamic efficiency in cogenerative arrangement second Law of Thermodynamic efficiency second Law of Thermodynamic efficiency in cogenerative arrangement electric conversion efficiency
The countercurrent water heat exchanger, called COOL1, allows the maintenance of the FC operating temperature by varying the hot water mass flow sent to the user and, therefore, reproducing the fuel cell cooling system. This system has been modelled as a heat exchange between the exhaust gases of the fuel cell and the water cooling, although it is just a simplified cooling system that compensates the modelled adiabatic behaviour of the CATHODEblock [3]. The gases leaving the heat exchanger COOL1, that are therefore at the operating conditions of the fuel cell, move through further three countercurrent heat exchangers in series. The first two, shown in Fig. 1 and called respectively EXAIR and EXFUEL, guarantee the preheating of the cell feeding flows; moreover, the third heat exchanger, COOL2, provides a preliminary warming up to the input water flow intended for users.
Fig. 1. PEMFC cogenerative layout.
4336
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
The condensation enthalpy for water vapour gases has been taken into account for all four heat exchangers (COOL1, COOL2, EXAIR and EXFUEL). The electrochemical model [3], implemented in Aspen Plus through Fortran procedures, uses semi-empirical correlations to evaluate the cell voltage and current in function of the operating conditions to satisfy a particular electrical power demand. After determining the operating point, the model calculates the feeding hydrogen necessary to the cell (flow 1F), taking into account the recirculation (flow 4R) of the fuel not used at each single passage (determined by the utilization coefficient, Uf). Furthermore, based on the air utilization coefficient Ua, defined as the ratio of the stoichiometric air and the total air input, the model calculates the amount of oxidant introduced (flow 1A) and the water flow needed for the humidification demand. In particular the percentage of relative humidity (RH) at which the cell works has been set through a Fortran iterative procedure with, as the target, the RH value set by the user for the feeding streams. The procedure, shown in [3], at any iteration varies the water molar fraction in the stream 1A according to:
Pv ap ¼ xw P RH ¼
Pv ap 100 Psat
ð1Þ ð2Þ
where: Pvap [atm] indicates the water partial pressure in the feeding streams. xw represents the water molar fraction in the feeding streams. P [atm] is the operating pressure of the fuel cell. Psat [atm] represents the saturation pressure of water at a certain temperature defined as ‘‘P sat ¼ 0:00987 eð16:6534030:13=ðT273:15þ235ÞÞ ’’, with T [K] the operating temperature of the fuel cell. RH [%] is the relative humidity percentage in the feeding streams. Finally, the heat exchangers AIRHOT and FUELHOT have been introduced to evaluate the external thermal power [kW], for example provided by the exhaust of a condominium boiler, needed to bring the feed gas to the FC temperature operating conditions; while HEATER1 has been used to simulate the thermal power [kW] generated by the cathode exothermic reactions. The electrical power (We), produced by the cell, has been modelled such an output (LOAD) of the CATHODE block. Therefore the analysis has been focused on the behaviour of the system in the CHP arrangement, evaluating the First and the Second Thermodynamic Law efficiency of the plant as a function of the operating conditions. The main hypotheses assumed in the simulation are: – – – –
Ideal gas. Ideal heat exchangers (heat losses have not been considered). Pressure drop neglected. The free Gibbs energy is minimized in all the chemical reactions.
The electric efficiency defined, according to the First Law of Thermodynamic, as the ratio between the electric power and the total power provided in input to the system, has been calculated using the following equation:
We W fuel þ Q air þ Q fuel
We is the produced electric power. Wfuel represents the power delivered to the system through the fuel (hydrogen), defined as W fuel ¼ HHV H2 n_ H2 , with n_ H2 the molar hydrogen flow rate. Qair and Qfuel are the thermal powers necessary to guarantee, respectively, the air temperature and humidification and the fuel temperature needed at the FC input. To determine Qair and Qfuel it has been taken into account the contribution provided by the heat recovery operated on the exhaust gases and particularly on the stream 8 of Fig. 1 and, subsequently, on the stream E1. Anyway, in the general expression of the energetic efficiency (Eq. (3)), all the contributions have been considered, because the heat recovering through the exhaust gases could be not sufficient to guarantee the achievement of the required operating feeding conditions. In this occurrence, therefore, in the following calculations (detailed in Section 5.2) the Qair and Qfuel contributions have nonzero values, otherwise, they assume values equal to zero. In the CHP arrangement, the calculation of the First Thermodynamic Law efficiency has been performed with the following expression (4):
gðIÞcog ¼
We þ Q W fuel þ Q air þ Q fuel
ð3Þ
ð4Þ
where Q represents the total thermal power recovered by the exchangers, COOL1 and COOL2, by the water flow sent to the user. 4. Second law of thermodynamic efficiency In order to completely evaluate the thermodynamic performance of the PEMFC system, it is necessary to characterize the system efficiency, according to the Second Law of Thermodynamic, through the exergetic analysis of all molar flows [kmol/s] and thermal and electrical power flow rates [kW]. The purpose of the exergetic analysis is to assess whether the energy resources, consumed by the system under consideration, are used efficiently. Therefore a model able to quantify the exergetic contributions, both physically and chemically, has been developed. So the exergetic efficiency has been quantified in relation to various operating conditions. In particular the Second Law of Thermodynamic efficiency has been calculated as follows:
gðIIÞ ¼
We Ex IN
ð5Þ
where We indicates the electric power [kW] produced by the cell and EX_IN the total exergy input to the system [kW]. In particular, the EX term relates to the feeding flows (H2 and humid air) and the respective preheating heat flow rate (QAIR and QFUEL in Fig. 1). The calculation of the total flow exergy [14,15] has been determined with the following expression:
Ex ¼ n_ e
ð6Þ
where n_ and e represent respectively the molar flow and the sum of the chemical (eCH) and physical exergy (ePH) of the stream considered. In fact:
e ¼ eCH þ ePH
3. First law efficiency
gðIÞ ¼
where:
ð7Þ
In the expression (7) the contribution of the kinetic and the potential exergies, considering the FC system for these aspects at rest to the environment, has been neglected. Specifically, the physical exergy relates to temperature and pressure of the PEMFC reactants and products. Therefore, for its calculation, the enthalpy and entropy differences, for every mass flow of the system, have been determined between its thermodynamic state and the restricted dead state conditions
4337
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
(T0 = 298 K and P0 = 1 atm). Consequently the expression of the term ePH is:
ePH ¼ ðh h0 Þ T 0 ðs s0 Þ
ð8Þ
Assuming the hypothesis of ideal gas with constant specific heat Cp, then it is possible to determine the physical exergy as:
" ePH ¼ C p T 0
# K1 T T P K þ ln 1 ln T0 T0 P0
ð9Þ
Furthermore, in order to estimate how much the chemical composition of a generic system deviates from the reference environmental conditions, the chemical exergy of the cell input and output flows has been determined through the following expression:
eCH ¼
X
xn eCH n þ RT 0
X
xn ln xn
ð10Þ
where xn and eCH n (Table 1) are respectively the mole fraction of a single chemical species and its specific chemical exergy (kJ/kmol). Specifically, the chemical exergy is the maximum theoretical useful work obtained as the system passes from the restricted dead state, where mechanical and thermal equilibrium are satisfied, to the dead state where it is in complete equilibrium with the environment. When evaluating chemical exergy, substance comprising the system must be referred to the properties of a suitable selected set of environmental substance in mutual equilibrium (Table 1). Then in the subsequent calculation of the chemical exergy, concerning the feeding humid air, made on the basis of the values in Table 1, its composition and thus its levels of relative humidity have been taken in account. Fig. 2 shows the plan adopted for the exergetic analysis of the fuel cell system. As input to the system, the fuel (hydrogen) and wet air flows, both characterized by ambient temperature conditions, have been considered. Moreover the analysis includes, as inputs, the exergy [kW] associated with the thermal powers [kW] (QAIR and QFUEL) necessary to the achievement of the FC operating temperatures and, as regards the air flow, the required humidification degree.
Table 1 Standard molar chemical exergy of various substance at T0 = 298.15 k and P0 = 1.0 atm. [14].
So considering the molar composition and the molar flow rate of the streams 1F and 1A applying the Eqs. (9) and (10), the term Ex_IN can be expressed as follows:
Ex
IN
¼ n_ 1F e1F þ n_ 1A e1A þ ExQ
QAIR
þ ExQ
QFUEL
ð11Þ
As output, the electric power produced by FC and, in CHP arrangement, the contribution of the exergy flow rate (ExQ) relative to the recovered heat in the heat exchanger COOL1 and COOL2 have been considered. Consequently, the expression (4) becomes:
gðIIÞcog ¼
W e þ ExQ Ex IN
ð12Þ
with:
ExQ ¼ ExQ
COOL1
þ ExQ
COOL2
ð13Þ
For example, the evaluation of the exergy related to the COOL1 heat exchanger (ExQ_COOL1) has been made as follows (same considerations are valid for the ExQ_COOL2 term):
ExQ
COOL1
T0 ¼ Q COOL1 1 Tm
ð14Þ
where: QCOOL1 is the thermal heat recovered by COOL1. T0 is the restricted dead state conditions temperature (298.15 K). Tm represents the average temperature of the water flow inside the heat exchanger. 5. Tests 5.1. Conditions investigated The analysis of the fuel cell performance in CHP arrangement has been carried out through simulations by varying: the pressure conditions of the fuel and air feeding flows (1–2 atm), the FC operating temperature (353.15 K, 378.15 K and 393.15 K) and the relative humidity of the inlet air flow (35–58–100%), maintaining a constant electric power demand (3 kW). Moreover, considering as subject of the study a CHP system for residential consumers, in agreement with the operating temperatures and the produced power chosen, the water delivery and return temperatures (for the user feeding) have been set respectively to 338.15 K and 318.15 K [11]. Such temperatures have been kept constant for the entire campaign of simulations.
Chemical exergy (kJ/kmol) H2 O2 N2 CH4 CO CO2 C2H6 C3H8 C4H10 H2O liquid H2O vapour
236,100 3970 720 831,650 275,100 19,860 14,95,840 21,54,000 28,05,800 900 9500
Fig. 2. Fuel cell system plan.
5.2. Analysis results 5.2.1. System efficiency in terms of the relative humidity After setting the temperature and the operating pressure of the cell, several tests have been carried out varying the relative humidity in order to calculate the efficiency parameters described in the Sections 3 and 4. Table 2 shows the results obtained in the simulations at 1 atm and 353.15 K as a function of RH (Eqs. (1) and (2)). Fig. 3 shows the relative trends, varying the RH degree, of the efficiency values relative both to the First and the Second Laws in CHP and no-CHP arrangement, together with the trend of the maximum flow of the produced hot water. The data of electric conversion efficiency gel, reported in the 6th column of Table 2, express the relationship between the produced electric power and the one corresponding to the FC fuel input, without considering the thermal inputs needed to achieve the operating conditions of the feeding flows. Therefore, in the following, the considerations are made in reference to the g(I) parameter, which includes such contributions.
4338
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
Table 2 Simulation results at T = 353.15 K and P = 1 atm. RH (%)
g(I) (%)
T = 353.15 K; P = 1 atm 35 27 40 32 50 34 58 35 65 35 70 35 80 34 90 32 100 29
g(II) (%)
g(I)cog (%)
g(II)cog (%)
gel (%)
Hot water mass flow rate produced (kg/h)
Molar liquid fraction in exhaust gas (%)
36 43 49 52 54 55 57 58 57
87 87 87 87 87 88 88 89 89
43 50 56 59 62 63 66 68 69
30 35 40 43 45 46 47 48 49
287 231 202 196 197 200 212 232 282
15 18 22 26 29 32 36 41 46
Fig. 3. Efficiency and hot water mass flow trends at T = 353.15 K and P = 1 atm.
The trend of the First Thermodynamic Law efficiency, g(I), is maximum at RH equal to 58%. This reflects the positive effect of humidification on the cell electrochemical performance which balances, as more as the RH increases up to 58%, the disadvantage due to the higher thermal power necessary to achieve the operating air humidification. In the CHP arrangement, the efficiency g(I)cog is high and almost constant due to the combined effect of the electric production and the heat recovery. The last column of Table 2 shows the molar liquid fraction percentage present in the exhaust flow, in order to highlight the thermal recovery made also through condensation. With the increasing of RH percentage, the Second Thermodynamic Law efficiency g(II) has a generally improving trend due to the reduction of the fuel input exergy. At the same time the thermal power rate needed for the humidification increases, but its negative influence is negligible. This last effect is evident in g(I) trend for RH values greater than 58%. This because in the exergy calculation, the thermal power needed for the humidification is multiplied by the Carnot factor. Moreover, the g(II) trend is maintained also on values greater than g(I). The Second Law efficiency in CHP arrangement g(II)cog has, at the increasing of the humidity content in the input flow, a similar trend to g(II), with slightly greater values. Compared instead to g(I)cog, g(II)cog presents lower values due to the low Carnot factor corresponding to the heat recovery at low temperature, typical of the user thermal demand. Furthermore, it is interesting to underline the trend of the maximum flow of hot water producible from the cell, that results in an initial reduction for RH values up to 58%, due to:
the growing of the thermal power demand for the humidification of the feeding air; consequently it follows a reduction in the thermal power recovered for cogenerative purposes, the low steam partial pressure in the exhaust gases, which does not allow a substantial condensation of water vapour in the exchangers COOL1 and COOL2. To confirm this, for low RH values, the liquid fraction of the discharge flow is small (Table 2) and it is produced exclusively in the heat exchanger at low temperature (COOL2). With the increasing of RH percentage in the feeding air, for values above 58%, there is a concrete increase in the produced flow of hot water. This effect is due to the achievement of a partial pressure value of the water vapour that permits the heat recovery through condensation also in the COOL1 exchanger. This is evidenced by the greater condensate fraction in the discharge (Table 2). In Table 3 the heat exergy concerning the thermal power QAIR and QFUEL and the ones recovered in the exchangers COOL1 and COOL2 are, for example, shown in Table 3 for the case of 353.15 K and 1 atm varying RH value. In a second phase the PEMFC performance have been investigated, at the RH variation, in reference to different operating conditions. In particular the operating temperature has been varied, considering the values of 378.15 K and 393.15 K and, subsequently, the operating pressure has been increased up to 2 atm. In Table 4 are indicated the simulation results, in terms of the PEMFC efficiency parameters and the produced hot water mass flow, for different RH values and, separately, for the operating temperatures of 378.15 K and 393.15 K.
4339
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342 Table 3 Exergy values of thermal power and inlet fuel at T = 353.15 K and P = 1 atm. Relative humidity (%)
Exergy associated to thermal power streams and recovered heat (kW) COOL2
Exergy associated to the input fuel (kW) 1F
Second law cogenerative efficiency (%) g(II)cog (%)
RH
QAIR
QFUEL
COOL1
T = 353.15 K, P = 1 atm 35 58 100
0.137 0.260 0.747
0.006 0.003 0.002
0.532 0.296 0.363
0.082 0.136 0.256
8.202 5.511 4.545
43 59 68
Table 4 Simulation results at T = 378.15 K and 393.15 K both at P = 1 atm. RH (%)
g(I) (%)
T = 378.15 K; P = 1 atm 35 24 58 16 T = 393.15 K; P = 1 atm 35 14
g(II) (%)
g(I)cog (%)
g(II)cog (%)
gel (%)
Hot water mass flow rate produced (kg/h)
Molar liquid fraction in exhaust gas (%)
42 42
89 91
53 60
36 45
360 604
41 69
34
91
52
38
739
68
Table 5 Simulation results at T = 353.15 K, 378.15 K e 393.15 K and P = 2 atm.
g(II) (%)
g(I)cog (%)
g(II)cog (%)
gel (%)
Hot water mass flow rate produced (kg/h)
Molar liquid fraction in exhaust gas (%)
T = 353.15 K; P = 2 atm 35 33 58 42 100 40
38 50 54
92 92 93
44 55 61
34 45 50
238 157 175
13 17 27
T = 378.15 K; P = 2 atm 35 33 58 34 100 23
42 49 43
93 92 93
49 58 55
39 47 52
237 226 395
24 37 61
T = 393.15 K; P = 2 atm 35 29 58 24 75 16
41 40 30
92 93 93
50 52 45
41 48 51
285 388 660
37 59 75
RH (%)
g(I) (%)
It can be noted as, in these temperature and pressure conditions, it is not possible to reach high values of RH, as the steam production for hot water mass flow humidification is not feasible. Comparing the results of Table 4 with the ones previously discussed and relative to a low operating temperature, it can be evidenced as, at the temperature increment, the First Thermodynamic Law efficiency g(I) decreases (due to the higher thermal power request for the feeding flows pre-treatment), while the efficiency parameters g(I)cog, g(II)cog in the CHP arrangement increase. Table 5 shows the results of the tests carried out at 2 atm. Under these pressure conditions it has been possible to reach all the values of relative humidity except for the case of high operating temperature (393.15 K), in which a maximum of 75% RH has been reached. The results are graphically shown in Figs. 4–6. The trend of the produced hot water mass flow changes mainly at the PEMFC operating temperature increase. In particular, for medium–low temperatures, it reflects the trend described before in relation to the tests at 1 atm and 353.15 K. In the case of high operating temperature, the trend always grows with the RH, because the condensation occurs (only in the heat exchanger COOL2) with a greater intensity at the RH increase (last column of Table 5). At low values of RH, instead, the thermal power needed for the air humidification has an influence on the thermal recovery greater than the condensation phenomenon, resulting in a global negative effect.
With increasing the cell operating temperature, the First Thermodynamic Law efficiency, g(I), assumes a decreasing trend when RH increases; this can be explained considering the greatest thermal power necessary for the air feeding flow to achieve the operating temperature conditions. In the CHP case, the efficiency g(I)cog is characterized by values and trend similar to the ones obtained at 1 atm. Regarding the Second Thermodynamic Law efficiency, at low temperature the g(II) and the g(II)cog trends (Fig. 4) are similar to the ones discussed previously for the simulations at 1 atm. For medium–high temperatures (Figs. 5 and 6) a decreasing trend results at the relative humidity increment; this behaviour is justified by the greater weight of the Carnot factor, with regard to the input heat flow rates, to which corresponds an increasing of the input exergy. 5.2.2. System efficiency in terms of the operating temperature A further analysis of the data obtained by simulations has been carried out, focusing on the efficiencies variation in function of the PEMFC operating temperature (from 353.15 K up to 393.15 K), at fixed pressure and RH conditions. In particular, it is interesting to underline how, in conditions characterized by RH = 35% and P = 2 atm, the heated water flow rate is quite constant until 378.15 K (Fig. 7). In fact, the quantity of produced hot water depends on the cell operating temperature, on the input heat flow rate for the achievement of the required feeding conditions and
4340
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
Fig. 4. Efficiency and hot water flow trends at T = 353.15 K and P = 2 atm.
Fig. 5. Efficiency and hot water mass flow trends at T = 378.15 K and P = 2 atm.
Fig. 6. Efficiency and hot water mass flow trends at T = 393.15 K and P = 2 atm.
on the thermal power recoverable from the exhaust gases (inclusive of the vapour condensation contribution). In this range of temperature these contributions balance out; instead, for further temperature increasing, the thermal power recoverable through condensation and addressed to the users becomes predominant and, therefore, the heated water flow rate increases. This
phenomenon is enlarged when the relative humidity grows: in such conditions, in fact, the producible hot water quantity increases in the analysed range of temperature (Figs. 8 and 9). The g(I) trend decreases when the temperature rises and this phenomenon is enlarged by the increment of the relative humidity. This can be explained considering that the thermal power, needed
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
4341
Fig. 7. Efficiency and hot water mass flow trends at RH = 35% and P = 2 atm.
Fig. 8. Efficiency and hot water mass flow trends at RH = 58% and P = 2 atm.
Fig. 9. Efficiency and hot water mass flow trends at RH = 100% and P = 2 atm.
for the pre-processing of the feeding flows to achieve the operating conditions, increases with the temperature and the relative humidity. The efficiencies g(II), g(II)cog show trends similar to the ones already discussed and depicted in Figs. 3–6.
6. Conclusions In this research activity a zero-dimensional model implemented in Aspen Plus environment has been developed to analyse the
4342
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
performance of PEMFC cogenerative systems, for residential application, varying the operating conditions (temperature, pressure and relative humidity). Through the analysis of the efficiency according to both the First and the Second Laws of Thermodynamic, the optimum fuel cell operating conditions have been identified, fixed the produced electric power (3 kW) and the delivery and return temperatures of the user feeding water, respectively set to 338.15 K and 318.15 K. In the initial analysis, the optimal conditions have been detected in 2 atm, 378.15 K and 58%, respectively in terms of the system pressure, temperature and RH of the feeding air. So the corresponding efficiency values have been calculated, specifically: g(I) = 34%, g(II) = 49%, g(I)cog = 92% and g(II)cog = 58%. This conclusion has been achieved on the basis of the electric First Thermodynamic Law efficiency, which presents the maximum values in the case of pressurization at 2 atm. Furthermore, taking into consideration the medium–high operating temperatures (378.15–393.15 K), with the aim to obtain water delivery temperatures to the user even greater than 338.15 K, the optimal operating conditions of temperature and RH, cited above, have been individuated on the basis of the Second Law efficiency in the CHP arrangement (g(II)cog = 58%). Anyway, it is important to underline that at operating conditions of 2 atm and 353.15 K, not considered in the case described above, it corresponds the lowest value of the hot water mass flow sent to the user. In fact, for the same produced electric power (3 kW), the recoverable thermal power is equal to 3.95 kW (low temperature case) compared with 5.10 kW producible in the case of the best operating conditions. Nevertheless analysing the cogenerative system aiming to maximize the Second Law efficiency g(II)cog (69%) and, in the same time, to guarantee a delivery hot water temperature peculiar of the commercial CHP fuel cell systems detailed in [11], a different set of optimal operating conditions has been detected: 1 atm, 353.15 K and 100% RH. This configuration, even though it is not characterized by high First Law efficiency, is also advantageous for the constructive simplicity and for the consequent reduced production cost of low temperature systems at atmospheric pressure. This has an important account specially in relation with the small size of the system object of study, addressed to residential applications. Moreover, due
to its low temperature, this solution is also advantageous considering the drying problems of the PEMFC membrane at high operating temperature. In future studies, it will be interesting to analyse the behaviour of the PEMFC system at variable electrical and thermal load, specially using typical load curves of residential users. By applying variable loads, possible transients can be analysed in order to find appropriate control logics for the micro-CHP unit based on PEMFC. References [1] Streckiene G, Martinaitis V, Andersen AN, Katz J. Feasibility of CHP-plants with thermal stores in the German spot market. Appl Energy 2009;86:2308–16. [2] Giaccone L, Canova A. Economical comparison of CHP systems for industrial user with large steam demand. Appl Energy 2009;86:904–14. [3] Barelli L, Bidini G, Gallorini F, Ottaviano A. Analysis of the operating conditions influence on PEM fuel cell performance by means of a novel semi-empirical model. Int J Hydrogen Energy 2010. doi:10.1016j.ijhydene.2010.06.032. [4] Kunusch C, Puleston PF, Mayosky MA, More JJ. Characterization and experimental results in PEM fuel cell electrical behavior. Int J Hydrogen Energy 2010;35:5876–81. [5] Zhang J, Xie Z, Zhang J, Tang Y, Song C, Navessin T, et al. High temperature PEM fuel cells. J Power Sources 2006;160:872–91. [6] Li MQ, Shao Z-G, Scott K. A high conductivity Cs2.5H0.5PMo12O40/ polybenzimidazole (PBI)/H3PO4 composite membrane for proton-exchange membrane fuel cells operating at high temperature. Journal of Power Sources 2008;183:69–75. [7] Gigliucci G, Petruzzi L, Cerelli E, Garzisi A, La Mendola A. Demonstration of a residential CHP system based on PEM fuel cells. J Power Sources 2004;131:62–8. [8] Zabalza I, Aranda A, de Gracia MD. Feasibility analysis of fuel cells for combined heat and power systems in the tertiary sector. Int J Hydrogen Energy 2007;32:1396–403. [9] Bagnoli M, De Pascale A. Performance evaluation of a small size cogenerative system based on a PEM fuel cell stack. In: Proceedings of GT2005 ASME Turbo Expo 2005: power for land, sea and air. Reno-Tahoe, Nevada, USA; June 6–9, 2005. [10] Santangelo PE, Tartarini P. Fuel cell systems and traditional technologies. Part I: Experimental CHP approach. Appl Therm Eng 2007;27:1278–84. [11] Nuvera Fuel Cell Avanti Data Sheet. (http://www.nuvera.com/products/ avanti.php). [12] Wanga Y, Chen KS, Mishler J, Cho SC, Adroher XC. A review of polymer electrolyte membrane fuel cells: technology, applications, and needs on fundamental research. Appl Energy 2011;88:981–1007. [13] Aspen Plus Version 2006 (20.0.3595) Users Guide 2006. Aspen Tech Ltd, Cambridge MA, USA. [14] Bejan A, Tsatsaronis G, Moran M. Thermal design & optimization. New York: John Wiley and sons Inc; 1996. [15] Kazim A. Exergy analysis of a PEM fuel cell at variable operating conditions. Energy Convers Manage 2004;45:1949–61.
Contents lists available at ScienceDirect
Applied Energy journal homepage: www.elsevier.com/locate/apenergy
An energetic–exergetic analysis of a residential CHP system based on PEM fuel cell L. Barelli ⇑, G. Bidini, F. Gallorini, A. Ottaviano Department of Industrial Engineering, University of Perugia, Via G. Duranti 1/A4, Perugia 06125, Italy
a r t i c l e
i n f o
Article history: Received 13 December 2010 Received in revised form 23 March 2011 Accepted 11 April 2011 Available online 13 July 2011 Keywords: PEM fuel cell Exergy analysis Residential CHP
a b s t r a c t The use of fuel cell systems for distributed residential power generation represents an interesting alternative to traditional thermoelectric plants due to their high efficiency and the potential recovering of the heat generated by the internal electrochemical reactions. In this paper the study of a micro cogenerative (CHP) energy system based on a Proton Exchange Membrane fuel cell (PEMFC) is reported. With the aim to evaluate the performance and then the feasibility of this non-conventional energy system, in consideration of thermal and electrical basic demand of a multifamily apartment blocks, a zerodimensional PEMFC model in Aspen Plus environment has been developed. A simulations sequence has been carried out at different operating conditions of the fuel cell (varying temperature, pressure and relative humidity). Subsequently, on the basis of the obtained results, an energy/exergy analysis has been conducted to define the optimal operating conditions of the PEMFC that ensures the most efficient use of the energy and exergy inputs. Ó 2011 Elsevier Ltd. All rights reserved.
1. Introduction In the actual energetic background in which a better use of primary resources is the principal problem, the cogeneration (CHP) can be a practicable solution for users who are sensitive to energy savings. In this context, to maximize the use of primary energy and avoid any waste of money, economic feasibility studies which take account of the energy market [1] and the type of application required by the market are of utmost importance [2]. Specifically, the CHP plants based on fuel cells (FCs) represent an interesting alternative to traditional technologies for the combined production of electrical energy and heat; this is due to: the high electrical efficiency of the FCs in part load, the modularity, the ability to ensure substantial autonomy of the user from the electrical distribution network and the possible reduction of the environmental impact. As regards fuel cells, the chemical and thermodynamic aspects of the characteristic processes [3,4] and the basic research on materials to enhance performance and durability [5,6] are under study in many research activities. At the same time, even the design and analysis of performance of CHP systems based on FCs (mostly prototypes are used to demonstrate the potential of such a technology) are subjects of many publications [7–10]. In view of a micro-cogeneration distributed production destined to residential users who require thermal loads characterized by low temperatures, the subject of this research has been identified in a high-efficiency CHP energy system based on a polymer ⇑ Corresponding author. Tel.: +39 075 5853740; fax: +39 075 5853736. E-mail address: [email protected] (L. Barelli). 0306-2619/$ - see front matter Ó 2011 Elsevier Ltd. All rights reserved. doi:10.1016/j.apenergy.2011.04.059
electrolyte fuel cell (PEMFC). This type of cell is currently the one with the most advanced technological development and some cogenerative units are already commercialized [11,12]. The purpose of the present paper is to evaluate the performance and, hence, the feasibility of a non-conventional energy system that can help to meet the thermal and electrical demand of residential consumers. To this aim a zero-dimensional model of such a system has been developed in the Aspen Plus environment, with the intent to analyse its performance under different operating conditions (temperature, pressure and relative humidity). The model simulations have been conducted by setting as target the fulfilment of a fixed electrical load (3 kW) that can help to cover the basic demand of a multifamily apartment block, such as the electrical auxiliary facilities, and by evaluating, as a result, the recoverable heat for the user. The first part of the study, described in Section 2 of the paper (system layout), has been focused on the identification of the optimal plant lay-out for the CHP unit under study and on the modelling of such a system in the Aspen Plus environment [13]. Subsequently, the performance analysis methods, based on the first and second thermodynamic law efficiency, have been defined in Sections 3 and 4. At this step, particular attention has been devoted to exergetic analysis by referring to the literature guidelines [14,15]. In Section 5 the simulations made in Aspen Plus have been presented and then the results, achieved by means of the analysis methodology cited above, have been discussed. In future studies, it will be interesting to analyse the behaviour of the PEMFC system at variable electrical and thermal load using typical load curves of residential users.
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
4335
Nomenclature Cp e eCH eCH n ePH EX ExQ h HHVH2 K P PEMFC Psat Pvap P0 Q Qair QCOOL1
specific heat capacity at constant pressure (kJ/(kmol K)) molar stream exergy (kJ/(kmol K)) molar stream chemical exergy (kJ/(kmol K)) molar chemical exergy of a single chemical species (kJ/ (kmol K)) molar stream physical exergy (kJ/(kmol K)) stream total exergy (kW) exergy associated to the recovered thermal power (kW) molar stream enthalpy (kJ/kmol) hydrogen higher heating value (kJ/kmol) ratio between specific heat capacity at constant pressure and specific heat capacity at constant volume pressure (atm) polymer electrolyte membrane fuel cell water saturation pressure at a certain temperature (atm) water partial pressure (atm) restricted dead state pressure (atm) total thermal power recovered in cogenerative arrangement (kW) thermal power delivered to feeding air flow (kW) thermal power recovered in the heat exchanger COOL1 (kW)
2. System layout The system layout considered for the simulations is depicted in Fig. 1. It is characterized, as well as the PEM electrochemical system described in [3], by heat exchangers that allow the FC cooling and the exhaust heat recovery for both the internal regeneration, the preheating of the cell feeding flows and the cogenerative purposes. Specifically the model developed by the authors in [3] and adopted in this paper is valid for the following fuel cell operating conditions:
Cell temperature: 70–120 °C. Pressure: 1–3 atm. Relative humidity: 35–100%. CO content in the feeding gas: 0–200 ppm.
Qfuel R RH s T Tm T0 Ua Uf xw xn Wfuel We n_ g(I) g(I)cog
g(II) g(II)cog gel
thermal power delivered to feeding fuel flow (kW) ideal gas constant (kJ/(kmol K)) relative humidity (%) molar stream entropy (kJ/(kmol K)) temperature (K) average temperature inside an heat exchanger (K) restricted dead state temperature (K) air utilization factor fuel utilization factor water molar fraction in the feeding streams mole fraction of a single chemical species chemical power supplied through the fuel (hydrogen) (kW) produced electric power (kW) molar flow rate (kmol/s) first Law of Thermodynamic efficiency first Law of Thermodynamic efficiency in cogenerative arrangement second Law of Thermodynamic efficiency second Law of Thermodynamic efficiency in cogenerative arrangement electric conversion efficiency
The countercurrent water heat exchanger, called COOL1, allows the maintenance of the FC operating temperature by varying the hot water mass flow sent to the user and, therefore, reproducing the fuel cell cooling system. This system has been modelled as a heat exchange between the exhaust gases of the fuel cell and the water cooling, although it is just a simplified cooling system that compensates the modelled adiabatic behaviour of the CATHODEblock [3]. The gases leaving the heat exchanger COOL1, that are therefore at the operating conditions of the fuel cell, move through further three countercurrent heat exchangers in series. The first two, shown in Fig. 1 and called respectively EXAIR and EXFUEL, guarantee the preheating of the cell feeding flows; moreover, the third heat exchanger, COOL2, provides a preliminary warming up to the input water flow intended for users.
Fig. 1. PEMFC cogenerative layout.
4336
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
The condensation enthalpy for water vapour gases has been taken into account for all four heat exchangers (COOL1, COOL2, EXAIR and EXFUEL). The electrochemical model [3], implemented in Aspen Plus through Fortran procedures, uses semi-empirical correlations to evaluate the cell voltage and current in function of the operating conditions to satisfy a particular electrical power demand. After determining the operating point, the model calculates the feeding hydrogen necessary to the cell (flow 1F), taking into account the recirculation (flow 4R) of the fuel not used at each single passage (determined by the utilization coefficient, Uf). Furthermore, based on the air utilization coefficient Ua, defined as the ratio of the stoichiometric air and the total air input, the model calculates the amount of oxidant introduced (flow 1A) and the water flow needed for the humidification demand. In particular the percentage of relative humidity (RH) at which the cell works has been set through a Fortran iterative procedure with, as the target, the RH value set by the user for the feeding streams. The procedure, shown in [3], at any iteration varies the water molar fraction in the stream 1A according to:
Pv ap ¼ xw P RH ¼
Pv ap 100 Psat
ð1Þ ð2Þ
where: Pvap [atm] indicates the water partial pressure in the feeding streams. xw represents the water molar fraction in the feeding streams. P [atm] is the operating pressure of the fuel cell. Psat [atm] represents the saturation pressure of water at a certain temperature defined as ‘‘P sat ¼ 0:00987 eð16:6534030:13=ðT273:15þ235ÞÞ ’’, with T [K] the operating temperature of the fuel cell. RH [%] is the relative humidity percentage in the feeding streams. Finally, the heat exchangers AIRHOT and FUELHOT have been introduced to evaluate the external thermal power [kW], for example provided by the exhaust of a condominium boiler, needed to bring the feed gas to the FC temperature operating conditions; while HEATER1 has been used to simulate the thermal power [kW] generated by the cathode exothermic reactions. The electrical power (We), produced by the cell, has been modelled such an output (LOAD) of the CATHODE block. Therefore the analysis has been focused on the behaviour of the system in the CHP arrangement, evaluating the First and the Second Thermodynamic Law efficiency of the plant as a function of the operating conditions. The main hypotheses assumed in the simulation are: – – – –
Ideal gas. Ideal heat exchangers (heat losses have not been considered). Pressure drop neglected. The free Gibbs energy is minimized in all the chemical reactions.
The electric efficiency defined, according to the First Law of Thermodynamic, as the ratio between the electric power and the total power provided in input to the system, has been calculated using the following equation:
We W fuel þ Q air þ Q fuel
We is the produced electric power. Wfuel represents the power delivered to the system through the fuel (hydrogen), defined as W fuel ¼ HHV H2 n_ H2 , with n_ H2 the molar hydrogen flow rate. Qair and Qfuel are the thermal powers necessary to guarantee, respectively, the air temperature and humidification and the fuel temperature needed at the FC input. To determine Qair and Qfuel it has been taken into account the contribution provided by the heat recovery operated on the exhaust gases and particularly on the stream 8 of Fig. 1 and, subsequently, on the stream E1. Anyway, in the general expression of the energetic efficiency (Eq. (3)), all the contributions have been considered, because the heat recovering through the exhaust gases could be not sufficient to guarantee the achievement of the required operating feeding conditions. In this occurrence, therefore, in the following calculations (detailed in Section 5.2) the Qair and Qfuel contributions have nonzero values, otherwise, they assume values equal to zero. In the CHP arrangement, the calculation of the First Thermodynamic Law efficiency has been performed with the following expression (4):
gðIÞcog ¼
We þ Q W fuel þ Q air þ Q fuel
ð3Þ
ð4Þ
where Q represents the total thermal power recovered by the exchangers, COOL1 and COOL2, by the water flow sent to the user. 4. Second law of thermodynamic efficiency In order to completely evaluate the thermodynamic performance of the PEMFC system, it is necessary to characterize the system efficiency, according to the Second Law of Thermodynamic, through the exergetic analysis of all molar flows [kmol/s] and thermal and electrical power flow rates [kW]. The purpose of the exergetic analysis is to assess whether the energy resources, consumed by the system under consideration, are used efficiently. Therefore a model able to quantify the exergetic contributions, both physically and chemically, has been developed. So the exergetic efficiency has been quantified in relation to various operating conditions. In particular the Second Law of Thermodynamic efficiency has been calculated as follows:
gðIIÞ ¼
We Ex IN
ð5Þ
where We indicates the electric power [kW] produced by the cell and EX_IN the total exergy input to the system [kW]. In particular, the EX term relates to the feeding flows (H2 and humid air) and the respective preheating heat flow rate (QAIR and QFUEL in Fig. 1). The calculation of the total flow exergy [14,15] has been determined with the following expression:
Ex ¼ n_ e
ð6Þ
where n_ and e represent respectively the molar flow and the sum of the chemical (eCH) and physical exergy (ePH) of the stream considered. In fact:
e ¼ eCH þ ePH
3. First law efficiency
gðIÞ ¼
where:
ð7Þ
In the expression (7) the contribution of the kinetic and the potential exergies, considering the FC system for these aspects at rest to the environment, has been neglected. Specifically, the physical exergy relates to temperature and pressure of the PEMFC reactants and products. Therefore, for its calculation, the enthalpy and entropy differences, for every mass flow of the system, have been determined between its thermodynamic state and the restricted dead state conditions
4337
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
(T0 = 298 K and P0 = 1 atm). Consequently the expression of the term ePH is:
ePH ¼ ðh h0 Þ T 0 ðs s0 Þ
ð8Þ
Assuming the hypothesis of ideal gas with constant specific heat Cp, then it is possible to determine the physical exergy as:
" ePH ¼ C p T 0
# K1 T T P K þ ln 1 ln T0 T0 P0
ð9Þ
Furthermore, in order to estimate how much the chemical composition of a generic system deviates from the reference environmental conditions, the chemical exergy of the cell input and output flows has been determined through the following expression:
eCH ¼
X
xn eCH n þ RT 0
X
xn ln xn
ð10Þ
where xn and eCH n (Table 1) are respectively the mole fraction of a single chemical species and its specific chemical exergy (kJ/kmol). Specifically, the chemical exergy is the maximum theoretical useful work obtained as the system passes from the restricted dead state, where mechanical and thermal equilibrium are satisfied, to the dead state where it is in complete equilibrium with the environment. When evaluating chemical exergy, substance comprising the system must be referred to the properties of a suitable selected set of environmental substance in mutual equilibrium (Table 1). Then in the subsequent calculation of the chemical exergy, concerning the feeding humid air, made on the basis of the values in Table 1, its composition and thus its levels of relative humidity have been taken in account. Fig. 2 shows the plan adopted for the exergetic analysis of the fuel cell system. As input to the system, the fuel (hydrogen) and wet air flows, both characterized by ambient temperature conditions, have been considered. Moreover the analysis includes, as inputs, the exergy [kW] associated with the thermal powers [kW] (QAIR and QFUEL) necessary to the achievement of the FC operating temperatures and, as regards the air flow, the required humidification degree.
Table 1 Standard molar chemical exergy of various substance at T0 = 298.15 k and P0 = 1.0 atm. [14].
So considering the molar composition and the molar flow rate of the streams 1F and 1A applying the Eqs. (9) and (10), the term Ex_IN can be expressed as follows:
Ex
IN
¼ n_ 1F e1F þ n_ 1A e1A þ ExQ
QAIR
þ ExQ
QFUEL
ð11Þ
As output, the electric power produced by FC and, in CHP arrangement, the contribution of the exergy flow rate (ExQ) relative to the recovered heat in the heat exchanger COOL1 and COOL2 have been considered. Consequently, the expression (4) becomes:
gðIIÞcog ¼
W e þ ExQ Ex IN
ð12Þ
with:
ExQ ¼ ExQ
COOL1
þ ExQ
COOL2
ð13Þ
For example, the evaluation of the exergy related to the COOL1 heat exchanger (ExQ_COOL1) has been made as follows (same considerations are valid for the ExQ_COOL2 term):
ExQ
COOL1
T0 ¼ Q COOL1 1 Tm
ð14Þ
where: QCOOL1 is the thermal heat recovered by COOL1. T0 is the restricted dead state conditions temperature (298.15 K). Tm represents the average temperature of the water flow inside the heat exchanger. 5. Tests 5.1. Conditions investigated The analysis of the fuel cell performance in CHP arrangement has been carried out through simulations by varying: the pressure conditions of the fuel and air feeding flows (1–2 atm), the FC operating temperature (353.15 K, 378.15 K and 393.15 K) and the relative humidity of the inlet air flow (35–58–100%), maintaining a constant electric power demand (3 kW). Moreover, considering as subject of the study a CHP system for residential consumers, in agreement with the operating temperatures and the produced power chosen, the water delivery and return temperatures (for the user feeding) have been set respectively to 338.15 K and 318.15 K [11]. Such temperatures have been kept constant for the entire campaign of simulations.
Chemical exergy (kJ/kmol) H2 O2 N2 CH4 CO CO2 C2H6 C3H8 C4H10 H2O liquid H2O vapour
236,100 3970 720 831,650 275,100 19,860 14,95,840 21,54,000 28,05,800 900 9500
Fig. 2. Fuel cell system plan.
5.2. Analysis results 5.2.1. System efficiency in terms of the relative humidity After setting the temperature and the operating pressure of the cell, several tests have been carried out varying the relative humidity in order to calculate the efficiency parameters described in the Sections 3 and 4. Table 2 shows the results obtained in the simulations at 1 atm and 353.15 K as a function of RH (Eqs. (1) and (2)). Fig. 3 shows the relative trends, varying the RH degree, of the efficiency values relative both to the First and the Second Laws in CHP and no-CHP arrangement, together with the trend of the maximum flow of the produced hot water. The data of electric conversion efficiency gel, reported in the 6th column of Table 2, express the relationship between the produced electric power and the one corresponding to the FC fuel input, without considering the thermal inputs needed to achieve the operating conditions of the feeding flows. Therefore, in the following, the considerations are made in reference to the g(I) parameter, which includes such contributions.
4338
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
Table 2 Simulation results at T = 353.15 K and P = 1 atm. RH (%)
g(I) (%)
T = 353.15 K; P = 1 atm 35 27 40 32 50 34 58 35 65 35 70 35 80 34 90 32 100 29
g(II) (%)
g(I)cog (%)
g(II)cog (%)
gel (%)
Hot water mass flow rate produced (kg/h)
Molar liquid fraction in exhaust gas (%)
36 43 49 52 54 55 57 58 57
87 87 87 87 87 88 88 89 89
43 50 56 59 62 63 66 68 69
30 35 40 43 45 46 47 48 49
287 231 202 196 197 200 212 232 282
15 18 22 26 29 32 36 41 46
Fig. 3. Efficiency and hot water mass flow trends at T = 353.15 K and P = 1 atm.
The trend of the First Thermodynamic Law efficiency, g(I), is maximum at RH equal to 58%. This reflects the positive effect of humidification on the cell electrochemical performance which balances, as more as the RH increases up to 58%, the disadvantage due to the higher thermal power necessary to achieve the operating air humidification. In the CHP arrangement, the efficiency g(I)cog is high and almost constant due to the combined effect of the electric production and the heat recovery. The last column of Table 2 shows the molar liquid fraction percentage present in the exhaust flow, in order to highlight the thermal recovery made also through condensation. With the increasing of RH percentage, the Second Thermodynamic Law efficiency g(II) has a generally improving trend due to the reduction of the fuel input exergy. At the same time the thermal power rate needed for the humidification increases, but its negative influence is negligible. This last effect is evident in g(I) trend for RH values greater than 58%. This because in the exergy calculation, the thermal power needed for the humidification is multiplied by the Carnot factor. Moreover, the g(II) trend is maintained also on values greater than g(I). The Second Law efficiency in CHP arrangement g(II)cog has, at the increasing of the humidity content in the input flow, a similar trend to g(II), with slightly greater values. Compared instead to g(I)cog, g(II)cog presents lower values due to the low Carnot factor corresponding to the heat recovery at low temperature, typical of the user thermal demand. Furthermore, it is interesting to underline the trend of the maximum flow of hot water producible from the cell, that results in an initial reduction for RH values up to 58%, due to:
the growing of the thermal power demand for the humidification of the feeding air; consequently it follows a reduction in the thermal power recovered for cogenerative purposes, the low steam partial pressure in the exhaust gases, which does not allow a substantial condensation of water vapour in the exchangers COOL1 and COOL2. To confirm this, for low RH values, the liquid fraction of the discharge flow is small (Table 2) and it is produced exclusively in the heat exchanger at low temperature (COOL2). With the increasing of RH percentage in the feeding air, for values above 58%, there is a concrete increase in the produced flow of hot water. This effect is due to the achievement of a partial pressure value of the water vapour that permits the heat recovery through condensation also in the COOL1 exchanger. This is evidenced by the greater condensate fraction in the discharge (Table 2). In Table 3 the heat exergy concerning the thermal power QAIR and QFUEL and the ones recovered in the exchangers COOL1 and COOL2 are, for example, shown in Table 3 for the case of 353.15 K and 1 atm varying RH value. In a second phase the PEMFC performance have been investigated, at the RH variation, in reference to different operating conditions. In particular the operating temperature has been varied, considering the values of 378.15 K and 393.15 K and, subsequently, the operating pressure has been increased up to 2 atm. In Table 4 are indicated the simulation results, in terms of the PEMFC efficiency parameters and the produced hot water mass flow, for different RH values and, separately, for the operating temperatures of 378.15 K and 393.15 K.
4339
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342 Table 3 Exergy values of thermal power and inlet fuel at T = 353.15 K and P = 1 atm. Relative humidity (%)
Exergy associated to thermal power streams and recovered heat (kW) COOL2
Exergy associated to the input fuel (kW) 1F
Second law cogenerative efficiency (%) g(II)cog (%)
RH
QAIR
QFUEL
COOL1
T = 353.15 K, P = 1 atm 35 58 100
0.137 0.260 0.747
0.006 0.003 0.002
0.532 0.296 0.363
0.082 0.136 0.256
8.202 5.511 4.545
43 59 68
Table 4 Simulation results at T = 378.15 K and 393.15 K both at P = 1 atm. RH (%)
g(I) (%)
T = 378.15 K; P = 1 atm 35 24 58 16 T = 393.15 K; P = 1 atm 35 14
g(II) (%)
g(I)cog (%)
g(II)cog (%)
gel (%)
Hot water mass flow rate produced (kg/h)
Molar liquid fraction in exhaust gas (%)
42 42
89 91
53 60
36 45
360 604
41 69
34
91
52
38
739
68
Table 5 Simulation results at T = 353.15 K, 378.15 K e 393.15 K and P = 2 atm.
g(II) (%)
g(I)cog (%)
g(II)cog (%)
gel (%)
Hot water mass flow rate produced (kg/h)
Molar liquid fraction in exhaust gas (%)
T = 353.15 K; P = 2 atm 35 33 58 42 100 40
38 50 54
92 92 93
44 55 61
34 45 50
238 157 175
13 17 27
T = 378.15 K; P = 2 atm 35 33 58 34 100 23
42 49 43
93 92 93
49 58 55
39 47 52
237 226 395
24 37 61
T = 393.15 K; P = 2 atm 35 29 58 24 75 16
41 40 30
92 93 93
50 52 45
41 48 51
285 388 660
37 59 75
RH (%)
g(I) (%)
It can be noted as, in these temperature and pressure conditions, it is not possible to reach high values of RH, as the steam production for hot water mass flow humidification is not feasible. Comparing the results of Table 4 with the ones previously discussed and relative to a low operating temperature, it can be evidenced as, at the temperature increment, the First Thermodynamic Law efficiency g(I) decreases (due to the higher thermal power request for the feeding flows pre-treatment), while the efficiency parameters g(I)cog, g(II)cog in the CHP arrangement increase. Table 5 shows the results of the tests carried out at 2 atm. Under these pressure conditions it has been possible to reach all the values of relative humidity except for the case of high operating temperature (393.15 K), in which a maximum of 75% RH has been reached. The results are graphically shown in Figs. 4–6. The trend of the produced hot water mass flow changes mainly at the PEMFC operating temperature increase. In particular, for medium–low temperatures, it reflects the trend described before in relation to the tests at 1 atm and 353.15 K. In the case of high operating temperature, the trend always grows with the RH, because the condensation occurs (only in the heat exchanger COOL2) with a greater intensity at the RH increase (last column of Table 5). At low values of RH, instead, the thermal power needed for the air humidification has an influence on the thermal recovery greater than the condensation phenomenon, resulting in a global negative effect.
With increasing the cell operating temperature, the First Thermodynamic Law efficiency, g(I), assumes a decreasing trend when RH increases; this can be explained considering the greatest thermal power necessary for the air feeding flow to achieve the operating temperature conditions. In the CHP case, the efficiency g(I)cog is characterized by values and trend similar to the ones obtained at 1 atm. Regarding the Second Thermodynamic Law efficiency, at low temperature the g(II) and the g(II)cog trends (Fig. 4) are similar to the ones discussed previously for the simulations at 1 atm. For medium–high temperatures (Figs. 5 and 6) a decreasing trend results at the relative humidity increment; this behaviour is justified by the greater weight of the Carnot factor, with regard to the input heat flow rates, to which corresponds an increasing of the input exergy. 5.2.2. System efficiency in terms of the operating temperature A further analysis of the data obtained by simulations has been carried out, focusing on the efficiencies variation in function of the PEMFC operating temperature (from 353.15 K up to 393.15 K), at fixed pressure and RH conditions. In particular, it is interesting to underline how, in conditions characterized by RH = 35% and P = 2 atm, the heated water flow rate is quite constant until 378.15 K (Fig. 7). In fact, the quantity of produced hot water depends on the cell operating temperature, on the input heat flow rate for the achievement of the required feeding conditions and
4340
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
Fig. 4. Efficiency and hot water flow trends at T = 353.15 K and P = 2 atm.
Fig. 5. Efficiency and hot water mass flow trends at T = 378.15 K and P = 2 atm.
Fig. 6. Efficiency and hot water mass flow trends at T = 393.15 K and P = 2 atm.
on the thermal power recoverable from the exhaust gases (inclusive of the vapour condensation contribution). In this range of temperature these contributions balance out; instead, for further temperature increasing, the thermal power recoverable through condensation and addressed to the users becomes predominant and, therefore, the heated water flow rate increases. This
phenomenon is enlarged when the relative humidity grows: in such conditions, in fact, the producible hot water quantity increases in the analysed range of temperature (Figs. 8 and 9). The g(I) trend decreases when the temperature rises and this phenomenon is enlarged by the increment of the relative humidity. This can be explained considering that the thermal power, needed
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
4341
Fig. 7. Efficiency and hot water mass flow trends at RH = 35% and P = 2 atm.
Fig. 8. Efficiency and hot water mass flow trends at RH = 58% and P = 2 atm.
Fig. 9. Efficiency and hot water mass flow trends at RH = 100% and P = 2 atm.
for the pre-processing of the feeding flows to achieve the operating conditions, increases with the temperature and the relative humidity. The efficiencies g(II), g(II)cog show trends similar to the ones already discussed and depicted in Figs. 3–6.
6. Conclusions In this research activity a zero-dimensional model implemented in Aspen Plus environment has been developed to analyse the
4342
L. Barelli et al. / Applied Energy 88 (2011) 4334–4342
performance of PEMFC cogenerative systems, for residential application, varying the operating conditions (temperature, pressure and relative humidity). Through the analysis of the efficiency according to both the First and the Second Laws of Thermodynamic, the optimum fuel cell operating conditions have been identified, fixed the produced electric power (3 kW) and the delivery and return temperatures of the user feeding water, respectively set to 338.15 K and 318.15 K. In the initial analysis, the optimal conditions have been detected in 2 atm, 378.15 K and 58%, respectively in terms of the system pressure, temperature and RH of the feeding air. So the corresponding efficiency values have been calculated, specifically: g(I) = 34%, g(II) = 49%, g(I)cog = 92% and g(II)cog = 58%. This conclusion has been achieved on the basis of the electric First Thermodynamic Law efficiency, which presents the maximum values in the case of pressurization at 2 atm. Furthermore, taking into consideration the medium–high operating temperatures (378.15–393.15 K), with the aim to obtain water delivery temperatures to the user even greater than 338.15 K, the optimal operating conditions of temperature and RH, cited above, have been individuated on the basis of the Second Law efficiency in the CHP arrangement (g(II)cog = 58%). Anyway, it is important to underline that at operating conditions of 2 atm and 353.15 K, not considered in the case described above, it corresponds the lowest value of the hot water mass flow sent to the user. In fact, for the same produced electric power (3 kW), the recoverable thermal power is equal to 3.95 kW (low temperature case) compared with 5.10 kW producible in the case of the best operating conditions. Nevertheless analysing the cogenerative system aiming to maximize the Second Law efficiency g(II)cog (69%) and, in the same time, to guarantee a delivery hot water temperature peculiar of the commercial CHP fuel cell systems detailed in [11], a different set of optimal operating conditions has been detected: 1 atm, 353.15 K and 100% RH. This configuration, even though it is not characterized by high First Law efficiency, is also advantageous for the constructive simplicity and for the consequent reduced production cost of low temperature systems at atmospheric pressure. This has an important account specially in relation with the small size of the system object of study, addressed to residential applications. Moreover, due
to its low temperature, this solution is also advantageous considering the drying problems of the PEMFC membrane at high operating temperature. In future studies, it will be interesting to analyse the behaviour of the PEMFC system at variable electrical and thermal load, specially using typical load curves of residential users. By applying variable loads, possible transients can be analysed in order to find appropriate control logics for the micro-CHP unit based on PEMFC. References [1] Streckiene G, Martinaitis V, Andersen AN, Katz J. Feasibility of CHP-plants with thermal stores in the German spot market. Appl Energy 2009;86:2308–16. [2] Giaccone L, Canova A. Economical comparison of CHP systems for industrial user with large steam demand. Appl Energy 2009;86:904–14. [3] Barelli L, Bidini G, Gallorini F, Ottaviano A. Analysis of the operating conditions influence on PEM fuel cell performance by means of a novel semi-empirical model. Int J Hydrogen Energy 2010. doi:10.1016j.ijhydene.2010.06.032. [4] Kunusch C, Puleston PF, Mayosky MA, More JJ. Characterization and experimental results in PEM fuel cell electrical behavior. Int J Hydrogen Energy 2010;35:5876–81. [5] Zhang J, Xie Z, Zhang J, Tang Y, Song C, Navessin T, et al. High temperature PEM fuel cells. J Power Sources 2006;160:872–91. [6] Li MQ, Shao Z-G, Scott K. A high conductivity Cs2.5H0.5PMo12O40/ polybenzimidazole (PBI)/H3PO4 composite membrane for proton-exchange membrane fuel cells operating at high temperature. Journal of Power Sources 2008;183:69–75. [7] Gigliucci G, Petruzzi L, Cerelli E, Garzisi A, La Mendola A. Demonstration of a residential CHP system based on PEM fuel cells. J Power Sources 2004;131:62–8. [8] Zabalza I, Aranda A, de Gracia MD. Feasibility analysis of fuel cells for combined heat and power systems in the tertiary sector. Int J Hydrogen Energy 2007;32:1396–403. [9] Bagnoli M, De Pascale A. Performance evaluation of a small size cogenerative system based on a PEM fuel cell stack. In: Proceedings of GT2005 ASME Turbo Expo 2005: power for land, sea and air. Reno-Tahoe, Nevada, USA; June 6–9, 2005. [10] Santangelo PE, Tartarini P. Fuel cell systems and traditional technologies. Part I: Experimental CHP approach. Appl Therm Eng 2007;27:1278–84. [11] Nuvera Fuel Cell Avanti Data Sheet. (http://www.nuvera.com/products/ avanti.php). [12] Wanga Y, Chen KS, Mishler J, Cho SC, Adroher XC. A review of polymer electrolyte membrane fuel cells: technology, applications, and needs on fundamental research. Appl Energy 2011;88:981–1007. [13] Aspen Plus Version 2006 (20.0.3595) Users Guide 2006. Aspen Tech Ltd, Cambridge MA, USA. [14] Bejan A, Tsatsaronis G, Moran M. Thermal design & optimization. New York: John Wiley and sons Inc; 1996. [15] Kazim A. Exergy analysis of a PEM fuel cell at variable operating conditions. Energy Convers Manage 2004;45:1949–61.