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Performance of a spark ignition engine fuelled with reformate gas produced on-board vehicle
International Journal of Hydrogen Energy 32 (2007) 2532 – 2538 www.elsevier.com/locate/ijhydene
Performance of a spark ignition engine fuelled with reformate gas produced on-board vehicle Enzo Galloni, Mariagiovanna Minutillo ∗ Department of Industrial Engineering, University of Cassino Via G. Di Biasio, 43, 03043 Cassino, FR, Italy Received 3 October 2006; accepted 3 October 2006 Available online 28 November 2006
Abstract In recent years, the interest in the use of hydrogen, as an alternative fuel for spark-ignition engines, has grown according to energy crises and pollution problems. By comparing the properties of hydrogen and gasoline, it is possible to underline the possibilities, for hydrogen–gasoline fuelled engines, of operating with very lean mixtures, thus obtaining interesting fuel economy and emission reductions. In this paper, the performance of a spark-ignition engine, fuelled by hydrogen enriched gasoline, has been evaluated by using a numerical model. A multidimensional code (KIVA-3V) has been modified in order to model the engine combustion process using a hybrid combustion model adapted for dual fuelling. Based on computed results, the performance of the engine has been evaluated in different operating conditions. Furthermore, for the hydrogen enriched gasoline engine fuelling, the hydrogen production on-board the vehicle has been considered. A thermochemical model of a reforming system has been developed by means of the Aspen Plus code. The conversion of gasoline to hydrogen has been investigated and thermodynamic analysis of the reforming system has been conducted. The thermal efficiency of the fuel processor and the efficiency of the integrated reformer/SI engine system have been calculated. 䉷 2006 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. Keywords: Hydrogen enriched gasoline engine; Hydrogen production; Plasma reforming; Numerical modelling
1. Introduction In recent years, the interest in the use of hydrogen, as an alternative fuel for spark-ignition engines, has grown according to energy crises and environmental pollution. For the introduction of hydrogen in the automotive market, an interesting procedure is the employment of hydrogen as an additional fuel to gasoline in spark ignition engines. The premixed charges, based on gasoline enriched with small hydrogen amounts, are characterized by wide flammability limits and high flame velocities leading to a high thermal efficiency, good engine performance and reduced pollutant emissions. Furthermore, the high molecular diffusivity allows a better mixing in the cylinder or in the ducts, thus guaranteeing homogeneous mixtures.
∗ Corresponding author. Tel.: +39 0776 2993 655.
E-mail address: [email protected] (M. Minutillo).
In previous works [1,2] the authors analysed different operating conditions of an engine fuelled with very lean hydrogen–gasoline mixtures at the unthrottled operation. The idea is that hydrogen can allow the leaning of the gasoline–air mixture in order to vary the engine load varying the fuel flow rate as in a diesel engine. The results of this study encouraged the authors to continue this research activity on an internal combustion engine fuelled with alternative fuels (dual fuelling). Problems associated with the use of hydrogen are fuel production and storage; using small hydrogen mass flows these problems can be overcome by producing the hydrogen on-board the vehicles by means of hydrocarbon reforming systems. In a previous work [3] a compact and simple reforming system has been studied and modelled by means of the Aspen Plus code [4]. The choice of studying this reformer depended on its characteristics which are: rapid response time, no sensitivity and degradation of the catalyst and its small size. These properties can meet the requirements for its application on-board the vehicle.
0360-3199/$ - see front matter 䉷 2006 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2006.10.008
E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538
2533
Nomenclature A, B Cf
IRES k Li LHVj mfuel
constants in the EBU model forward pre-exponential factor in the kinetic model concentration of species j indicated mean effective pressure (IMEP is obtained by dividing the work per cycle by the cylinder volume displaced) integrated reformer/engine system turbulent kinetic energy indicated engine work per cycle low heating value of the species j engine gasoline mass per engine cycle
mR fuel
reformer fuel mass per engine cycle
mADD fuel msyngas
additional gasoline mass per engine cycle syngas (or reformate gas) mass per engine cycle molecular weight of species j number of moles of the species j air ratio
cj IMEP
mwj nj O/C
The hydrogen generator (called GlidArc plasma) is a noncatalytic reformer based on high-voltage discharges that assist the exothermic partial oxidation (POX) process using air as oxidant [5]. The performance of the system has been investigated and the chemical composition of the reformate gas has been calculated. The purpose of the present work is to predict the performance of an integrated reformer/engine system (IRES), in which the reformer produces the syngas burned together with straight gasoline in a spark ignition engine. In this way, the performance of IRES will be compared to the performance of the same spark ignition engine fuelled with pure gasoline. CFD calculations have been carried out in order to predict the engine performance. In particular, a modified multidimensional KIVA-3V code release [6,7] has been used to model the turbulent combustion process of a multicomponent fuel, while one-dimensional calculations have been done to estimate the gas exchange processes. The investigation has been carried out considering different torque values at a fixed speed, 1500 rpm. The transition from high to low torque is obtained by varying the gasoline–air equivalence ratio ((G)), while the hydrogen–air equivalence ratio ((H2 )) is set constant. The hydrogen amount feeding the engine has been assumed according to the lowest hydrogen–air equivalence ratio used in experimental tests carried out on a pure hydrogen engine.
2. Reformer/engine system modelling The performance of the IRES has been evaluated by theoretical analysis. In Fig. 1 the IRES flowsheet is shown.
Ta
activation temperature
Greek symbols , , ε (G) (CO) (H2 ) R Engine System j j
r
m
k
constants in the kinetic combustion model turbulent kinetic energy dissipation engine air–gasoline equivalence ratio engine air–carbon monoxide equivalence ratio engine air–hydrogen equivalence ratio reformer thermal efficiency engine thermal efficiency overall thermal efficiency density of species j constant in the hybrid combustion model stoichiometric coefficient of species j in the combustion reaction combustion reaction rate EBU model reaction rate kinetic model reaction rate
The overall thermal efficiency of the IRES can be calculated as the following: System =
(mADD fuel
Li , + mR fuel ) · LHVfuel
(1)
where the reformer auxiliaries’ work is neglected. Therefore, to evaluate the efficiency of this system, numerical investigations have been carried out using both a thermochemical model for the plasma reformer and a CFD model for the engine. 2.1. Reformer modelling In a previous paper [3] a numerical model was developed to reproduce the operating conditions of a plasma reactor that assists the exothermic POX of hydrocarbons (gaseous or liquid). The reforming system was modelled by means of Aspen Plus code [4]. The conversion of gasoline to hydrogen was investigated and a thermodynamic analysis of the reformer was conducted. The thermochemical model was tested by comparing numerical results with experimental data. The model of the fuel processor consists of: a plasma reformer, a water separation unit and some auxiliary units; the flowsheet is shown in Fig. 1. It has been achieved by considering the integration of the reformer with an internal combustion engine. Gasoline (mR fuel ), after vaporizing, is fed into the reactor, combined with air and converted into syngas. Then syngas is cooled and the water is separated. Syngas and additional gasoline (mADD fuel ) are fed into the engine. A brief description of the main blocks used for the modelling is listed in Table 1.
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E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538
AIRADD
mfuel
C8H15
msyngas WATER
Li
SYNGAS
SEPARAT
MIXTURE
HEATER
ENGINE
EXHAUST
4 R
mfuel 1
GASOLINE
OUT-GAS 2
PUMP
REACTOR
EVAP
AIR
COMPR 3 Fig. 1. Integrated reformer/engine system model.
Table 1 Main blocks description
Table 2 Reformate gas from gasoline reforming [3]
Phenomena
Block
Name
Gas composition
Numerical results (vol%)
Plasma reforming Cooling Water separation
RGIBBS HEAT-EX SEPARATOR
REACTOR HEATER SEPARAT
CH4 N2 CO2 H2 CO LHV (kJ/kg)
1% 55% 3% 21% 20% 5023
In the following, some of the simulation hypotheses are listed. • The lack of detailed kinetic data (reaction rates, residence time, known intermediate species, etc.) has deterred a kinetic modelling of the reforming system. Therefore, the simulation has been run with a Gibbs reactor that simulates a chemical reactor by solving the heat and mass balances on minimizing the free energy of the components that can be produced during a reforming process. The method of minimizing the Gibbs free energy is normally preferred in the fuel reforming analysis [4,8] because the results provide a good prediction on the chemical composition of the reformate gas. On the other hand, the Gibbs reactor successfully emulated the experimental results [3]. • The possible species that might be found in the reformate gas are: N2 , CH4 , CO, CO2 , H2 O, H2 .
To analyse the operating condition of the reforming system, the air ratio (O/C) parameter has been defined: nO2 nO2 (O/C) = . (2) nfuel nfuel complete_combustion The O/C ratio is 0.37 and it has been assigned according to the experimental test on gasoline reforming. The temperature of the Gibbs reactor, that affects significantly the equilibrium compositions, is assumed equal to 968 K. In correspondence with this temperature, a good agreement between the syngas composition calculated by the model and the experimental data was obtained. The simulation results in terms of reformate gas composition are shown in Table 2. The reformer thermal efficiency has been calculated by considering both hydrogen and carbon monoxide in the syngas
E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538 Table 3 Characteristic constants for the combustion model
(Eq. (3)). The calculated efficiency is about 80%. R =
msyngas · LHVsyngas mR fuel · LHVfuel
.
(3)
2.2. Engine modelling The numerical approach has been described in detail in [1,9]. Briefly, both 1D and 3D calculations have been carried out: 3D simulations, starting at intake valve closing and ending at exhaust valve opening, have been performed by means of the Kiva3V code. Twelve chemical species have been considered: gasoline (C8 H15 ), H2 , O2 , N2 , CO2 , H2 O, H, O, N, OH, CO and NO. One-dimensional calculations have been done in order to estimate the initial conditions for CFD runs: at intake valve closing all the in-cylinder scalar variables have been considered uniform, while an initial tumble motion has been considered. Considering the reformate gas calculated previously, the Kiva3V code has been modified in order to simulate the combustion of hydrogen, gasoline and carbon monoxide (the methane combustion has been neglected). A serial mechanism based on the following overall reactions has been considered: 2H2 + O2 ⇒ 2H2 O,
(a)
4C8 H15 + 47O2 ⇒ 32CO2 + 30H2 O, (b) 2CO + O2 ⇒ 2CO2 .
(c)
Each oxidation rate ( r ) has been estimated by the following hybrid model [9] (Eq. (4)), in order to consider both turbulent ( m , Eq. (5)) and chemical kinetic aspects ( k , Eq. (6)): ⎧ for m k , ⎨ r = m (4) ⎩ 1 = + 1 − for m < k ,
r
m
k O2 ε f , ,
m = A · min k f · mw f O2 · mw O2 B · p , (5) f · mw f + O2 · mw O2
k = Cf · cf · cO2 · cH2 O · e−Ta /T .
2535
Reaction
A, B,
Cf
Ta (K)
a b c
18, 0.5, 0.44
1.8 × 1013 46 × 1010 1.9 × 106
1.0 0.25 1
1.0 1.5 0.3
0.0 0.0 0.5
1.76 × 104 3 × 104 8.056 × 103
gas enriched gasoline. The engine is the FIAT FIRE 1200 16 V, and features four cylinders (total piston displacement 1242 cm3 ) and a geometrical compression ratio equal to 10.6. The engine performance has been calculated at 1500 rpm and at different loads. The wide flammability limits and the high flame speed characterising the hydrogen–air mixtures allow the engines fuelled with pure hydrogen to run with very lean mixtures [15,16]. In this paper, it has been assumed that the hydrogen amount contained in the syngas is able to allow stable combustion phase if the relative air–hydrogen equivalence ratio is equal to a minimum value. This minimum value has been assumed according to the lowest hydrogen–air equivalence ratio mentioned in the technical literature [17–19]. Thus, the engine load can be controlled by varying the fuel flow-rate at wide open throttle, limiting the pumping losses at part load. In particular, by keeping constant the minimum air–hydrogen equivalence ratio, it is possible to increase the engine load, by dispatching a variable gasoline amount directly to the engine. In this way, both the reformer size is curbed and the engine power is not damaged too much because of the poor syngas heating value. Two kinds of fuelling have been considered: • gasoline–reformate gas, in which the syngas flow rate is able to allow a hydrogen–air equivalence ratio (H2 ) equal to 0.2 (Case A); • gasoline–reformate gas, in which the syngas flow rate is able to allow a hydrogen–air equivalence ratio equal to 0.14 (Case B).
(6)
The model has been tuned in order to reproduce the behaviour of the FIAT FIRE 1200 16 V engine fuelled with pure gasoline; assuming that , A and B depend on the turbulent structure characterizing the engine alone, they have been kept constant for each oxidation mechanism. The chemical property of the fuels has been considered in Eq. (6). All parameters of the combustion models are summarized in Table 3 [10–12]. To estimate the pollutant production, the CO2 dissociation and the NO production have been calculated, in the burned gas zone alone, according to the Meintjes–Morgan model [13] and the extended Zeldovich mechanism [14], respectively. 2.3. Engine performance Numerical investigations have been carried out to predict the performance of a spark ignition engine fuelled with reformate
Considering the reformate gas composition calculated by the reformer modelling, for Case A the carbon monoxide–air equivalence ratio (CO) is equal to 0.19, while for Case B it is equal to 0.13. As mentioned above, different loads are obtained by varying the gasoline amount (G). The highest air–gasoline ratio considered is equal to 0.6. The main numerical results are shown in Table 4. In each case studied the spark-timing has been optimized by considering the maximum brake thermal efficiency. To consider knock risk and mechanical stresses, spark-advances producing a pressure rise less than 2 bar/◦ have been assumed. The engine thermal efficiency has been evaluated as following: Engine =
mADD fuel
Li . · LHVfuel + msyngas LHVsyngas
(7)
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E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538
It can be observed that the highest efficiency (Engine = 0.43) has been obtained for Case B, at (G) = 0.2; moreover for (G) ranging between 0.1 and 0.5 the efficiency is higher than 0.4. The lowest efficiency (Engine = 0.28) has been calculated for Case B, at (G) = 0.0. This low efficiency is due to the very slow combustion caused by the ultra-lean mixture considered. The calculated efficiencies for the dual fuel (reformate gas and gasoline) engine have been compared to those of the same engine fuelled with straight gasoline (conventional engine). The comparison is shown in Fig. 2. Obviously, for the conventional gasoline engine different loads are obtained by keeping a stoichiometric air–fuel ratio (except at full load, IMEP≈ 10 bar) and reducing the volumetric efficiency by throttling. The lower energy density of the air–reformate gas/gasoline mixture, compared to the air–gasoline mixture, significantly reduces the maximum engine power output (about 10% for (H2 ) = 0.2 and 14% for (H2 ) = 0.14). Nevertheless, at the same load, the dual fuel engine shows higher efficiencies than the gasoline engine. At middle load the efficiency increases about 25%. Fewer pumping losses, due to the adopted load control strategy, as well as thermal advantages due to the lean burning, justify this trend. It is worth nothing that in Case A, the dual fuel engine shows a minimum load (IMEP = 3.85 bar) higher than that
Table 4 Engine performance (speed 1500 rpm, throttle 100%) Case A
Case B
(H2 )=0.20; (CO) = 0.19 (H2 ) = 0.14; (CO) = 0.13 (kg/kg) 0 0.1 0.2 0.35 0.5 0.6
IMEP (bar) 3.85 5.35 6.32 7.72 8.74 9.06
Engine
Engine
IMEP (bar) 2.17 4.08 5.23 6.66 7.87 8.60
(dimensionless) 0.36 0.42 0.42 0.42 0.40 0.38
(dimensionless) 0.28 0.41 0.43 0.42 0.41 0.40
3. System performance analysis The CFD numerical results have shown very interesting efficiencies for the stand alone hydrogen enriched gasoline engine, higher than that of the gasoline engine. However, these results are not complete. To evaluate the performance of the dual fuel engine, it is also necessary to consider the performance of the hydrogen generation system. The thermochemical model of the reformer, by evaluating the gasoline mass flow used to generate reformate gas for engine fuelling, allows to reach this target. As mentioned above, the overall thermal efficiency of the IRES has been calculated as Eq. (1). Eq. (8) shows the correlation between the three considered efficiencies: System Engine
=
ADD 1 + (mR fuel /mfuel ) · R ADD 1 + (mR fuel /mfuel )
.
• at low loads the efficiency of the gasoline engine is higher than the overall efficiency of the IRES (Cases A and B); • at middle loads the efficiency of the gasoline engine is very close to the IRES efficiency;
0.4 0.3 0.2 Conventional Engine (Gasoline) Dual Fuel Engine (Case A) Dual Fuel Engine (Case B)
0.1 0 0
1
2
3
(8)
The efficiency of the IRES depends on both the reformer effiADD ciency and the mR fuel /mfuel ratio. The calculated results for the IRES and the conventional engine are shown in Fig. 3. Of course, the overall thermal efficiency of the IRES is lower than that of the dual fuel engine especially for Case A, where the highest hydrogen concentration means a higher reformer ADD fuel consumption, i.e. a higher mR fuel /mfuel ratio. Comparing the IRES with the gasoline engine efficiencies, it is worth noting that
Engine Thermal Efficiency
0.5
ηengine [-]
(G)
of the gasoline engine (IMEP = 2.3 bar). Of course, lower loads can be reached by throttling. In this way, above IMEP∼ 4 bar, the engine load is regulated by varying the gasoline amount at wide open throttle, while below IMEP∼ 4 bar the throttle opening controls the engine load (at (H2 )=0.2, (CO)=0.19, (G) = 0.0).
4
5 6 IMEP [bar]
7
Fig. 2. Engine thermal efficiency.
8
9
10
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E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538
2537
Overall Thermal Efficiency
0.5
ηSystem [-]
0.4 0.3 0.2 Conventional Engine (Gasoline) Reformer/Engine (Case A) Reformer/Engine (Case B)
0.1 0 0
1
2
3
4
5 6 IMEP [bar]
7
8
9
10
11
Fig. 3. Overall thermal efficiency of the integrated reformer/engine system.
9.E-03
• at middle and high loads the calculated CO production of the IRES is much lower than that of the engine fuelled with straight gasoline, while at low loads the CO increases because of the very slow combustion, due to ultra-lean mixtures burned in the dual fuel engine; • for the IRES, the calculated NO production is very close to zero at low and middle loads, while at high loads the NO production increases because of oxygen overabundance and high temperature inside the cylinder, characterising the dual fuel engine.
Engine (Gasoline) Reformer/Engine (Case A) Reformer/Engine (Case B)
8.E-03 7.E-03 CO [g]
6.E-03 5.E-03 4.E-03 3.E-03 2.E-03 1.E-03 0.E+00 0
1
2
3
4
5 6 7 IMEP [bar]
8
9
10 11
4. Conclusions
Fig. 4. Overall CO emissions per engine cycle.
5.E-04
Engine (Gasoline) Reformer/Engine (Case A) Reformer/Engine (Case B)
4.E-04 4.E-04 NO [g]
3.E-04 3.E-04 2.E-04 2.E-04 1.E-04 5.E-05 0.E+00 0
1
2
3
4
5 6 7 IMEP [bar]
8
9
10 11
Fig. 5. Overall NO emissions per engine cycle.
• at high loads the gasoline engine efficiency is lower than that of the IRES. Therefore, it is worth noting that the mean overall efficiency of the systems seems to be very similar. With regard to emissions, it is possible to underline that comparable efficiencies mean comparable CO2 emissions. However, considering pollutant emissions, as shown in Figs. 4 and 5
In this paper numerical investigations have been conducted to predict the performance of an IRES, in which the reformer produces the syngas burned together with straight gasoline in a spark ignition engine. The thermochemical modelling of the reformer has allowed to calculate the syngas composition and the efficiency of the fuel processor. The calculated efficiency of the gasoline reforming is equal to 80%. The considered engine, fuelled with syngas and gasoline, has shown a much higher efficiency than the engine fuelled with pure gasoline, according to CFD analysis. The performance of this integrated system has been compared to the performance of the same spark ignition engine fuelled with pure gasoline. The results of the computational analysis have pointed out that the mean overall efficiency of IRES is close to that of the conventional spark ignition engine. Regarding pollutant emissions, the numerical analysis has shown advantages at middle load operation. At the end, it is appropriate to highlight that these results have been obtained considering an existing gasoline engine; therefore, further improvement in the IRES performance can be achieved by designing a dual fuel engine, able to effect the full potential of hydrogen in internal combustion engines.
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References [1] Fontana G, Galloni E, Jannelli E, Minutillo M. Performance and fuel consumption estimation of a hydrogen enriched gasoline engine at partload operation, SAE Paper N. 2002-01-2196, Paris; 2002, reprinted In: Holt DJ editor. Hydrogen and its future as a transportation fuel, SAE Progress in Technology Series; 2002. p. 105–10. [2] Fontana G, Galloni E, Jannelli E, Minutillo M. Different technologies for hydrogen engine fuelling. International symposium hypothesis V, Porto Conte; 2003. [3] Minutillo M. On-board fuel processor modelling for hydrogen enriched gasoline fuelled engine. Int J Hydrogen Energy 2005;30:1483–90. [4] Aspen Plus 12.1. User guide. Aspen Technology Inc.; 2000. [5] Czernichowski A, Czernichowski M, Czernichowski P. GlidArc-assisted reforming of gasoline and diesel oils into synthesis gas. European hydrogen energy conference (EHEC), Grenoble; 2003. [6] Amsden AA. KIVA-3V: a block-structured KIVA program for engines with vertical or canted valves. Los Alamos National Laboratory Report LA-13313-MS; 1997. [7] Amsden AA, O’Rourke PJ, Butler TD. KIVA-II: a computer program for chemically reactive flows with sprays. Los Alamos National Laboratory Report LA-11560-MS; 1989. [8] Seo Y-S, Shirley A, Kolaczkowski ST. Evaluation of thermodynamically favourable operating conditions for production of hydrogen in three different reforming technologies. J Power Sources 2002;108:213–25. [9] Fontana G, Galloni E, Jannelli E. Multidimensional modeling of premixed turbulent combustion: models and experimental test of a small spark-ignition engine, ASME Spring Conference, Philadelphia; 2001.
[10] Marinov N, Westbrook CK, Pitz WJ. Detailed and global chemical kinetics model for hydrogen. In: Chan SH, editor. Transport Phenomena in Combustion, vol. 1. Washington, DC: Taylor & Francis; 1996. [11] Westbrook CK, Dryer FL. Chemical kinetic modeling of hydrocarbon combustion. Prog Energy Combust Sci 1984;10:1–57. [12] Bradley RA, Marc AC, Dave HW. Modeling non-equilibrium CO oxidation in combustion systems. In: International mechanical engineering congress and exposition, Orlando, FL, 2000. [13] Ramos JI. Internal combustion engine modeling. Hemisphere Publishing Corporations; 1989. [14] Heywood JB. Internal combustion engine fundamentals. New York: Mc Graw-Hill; 1988. [15] Das LM. Hydrogen–Oxygen reaction mechanism and its implication to hydrogen engine combustion. Int J Hydrogen Energy 1996;21(8): 703–15. [16] Milton BE, Keck JC. Laminar burning velocities in stoichiometric Hydrogen and Hydrogen–Hydrocarbon gas mixtures. Combust Flame 1984;58:13–22. [17] Varde KS, Frame GM. A study of combustion and engine performance using electronic hydrogen fuel injection. Int J Hydrogen Energy 1984;9(4):327–32. [18] Gentili R. Lean air–fuel mixtures supplemented with Hydrogen for S.I. engines: a possible way to reduce specific fuel consumption? Int J Hydrogen Energy 1985;10(7/8):491–5. [19] Tang X, Kabat DM, Natkin RJ, Stockhausen WF, Heffel J. Ford P2000 Hydrogen engine dynamometer development. SAE Paper, 2002-01-242; 2002.
Performance of a spark ignition engine fuelled with reformate gas produced on-board vehicle Enzo Galloni, Mariagiovanna Minutillo ∗ Department of Industrial Engineering, University of Cassino Via G. Di Biasio, 43, 03043 Cassino, FR, Italy Received 3 October 2006; accepted 3 October 2006 Available online 28 November 2006
Abstract In recent years, the interest in the use of hydrogen, as an alternative fuel for spark-ignition engines, has grown according to energy crises and pollution problems. By comparing the properties of hydrogen and gasoline, it is possible to underline the possibilities, for hydrogen–gasoline fuelled engines, of operating with very lean mixtures, thus obtaining interesting fuel economy and emission reductions. In this paper, the performance of a spark-ignition engine, fuelled by hydrogen enriched gasoline, has been evaluated by using a numerical model. A multidimensional code (KIVA-3V) has been modified in order to model the engine combustion process using a hybrid combustion model adapted for dual fuelling. Based on computed results, the performance of the engine has been evaluated in different operating conditions. Furthermore, for the hydrogen enriched gasoline engine fuelling, the hydrogen production on-board the vehicle has been considered. A thermochemical model of a reforming system has been developed by means of the Aspen Plus code. The conversion of gasoline to hydrogen has been investigated and thermodynamic analysis of the reforming system has been conducted. The thermal efficiency of the fuel processor and the efficiency of the integrated reformer/SI engine system have been calculated. 䉷 2006 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. Keywords: Hydrogen enriched gasoline engine; Hydrogen production; Plasma reforming; Numerical modelling
1. Introduction In recent years, the interest in the use of hydrogen, as an alternative fuel for spark-ignition engines, has grown according to energy crises and environmental pollution. For the introduction of hydrogen in the automotive market, an interesting procedure is the employment of hydrogen as an additional fuel to gasoline in spark ignition engines. The premixed charges, based on gasoline enriched with small hydrogen amounts, are characterized by wide flammability limits and high flame velocities leading to a high thermal efficiency, good engine performance and reduced pollutant emissions. Furthermore, the high molecular diffusivity allows a better mixing in the cylinder or in the ducts, thus guaranteeing homogeneous mixtures.
∗ Corresponding author. Tel.: +39 0776 2993 655.
E-mail address: [email protected] (M. Minutillo).
In previous works [1,2] the authors analysed different operating conditions of an engine fuelled with very lean hydrogen–gasoline mixtures at the unthrottled operation. The idea is that hydrogen can allow the leaning of the gasoline–air mixture in order to vary the engine load varying the fuel flow rate as in a diesel engine. The results of this study encouraged the authors to continue this research activity on an internal combustion engine fuelled with alternative fuels (dual fuelling). Problems associated with the use of hydrogen are fuel production and storage; using small hydrogen mass flows these problems can be overcome by producing the hydrogen on-board the vehicles by means of hydrocarbon reforming systems. In a previous work [3] a compact and simple reforming system has been studied and modelled by means of the Aspen Plus code [4]. The choice of studying this reformer depended on its characteristics which are: rapid response time, no sensitivity and degradation of the catalyst and its small size. These properties can meet the requirements for its application on-board the vehicle.
0360-3199/$ - see front matter 䉷 2006 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved. doi:10.1016/j.ijhydene.2006.10.008
E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538
2533
Nomenclature A, B Cf
IRES k Li LHVj mfuel
constants in the EBU model forward pre-exponential factor in the kinetic model concentration of species j indicated mean effective pressure (IMEP is obtained by dividing the work per cycle by the cylinder volume displaced) integrated reformer/engine system turbulent kinetic energy indicated engine work per cycle low heating value of the species j engine gasoline mass per engine cycle
mR fuel
reformer fuel mass per engine cycle
mADD fuel msyngas
additional gasoline mass per engine cycle syngas (or reformate gas) mass per engine cycle molecular weight of species j number of moles of the species j air ratio
cj IMEP
mwj nj O/C
The hydrogen generator (called GlidArc plasma) is a noncatalytic reformer based on high-voltage discharges that assist the exothermic partial oxidation (POX) process using air as oxidant [5]. The performance of the system has been investigated and the chemical composition of the reformate gas has been calculated. The purpose of the present work is to predict the performance of an integrated reformer/engine system (IRES), in which the reformer produces the syngas burned together with straight gasoline in a spark ignition engine. In this way, the performance of IRES will be compared to the performance of the same spark ignition engine fuelled with pure gasoline. CFD calculations have been carried out in order to predict the engine performance. In particular, a modified multidimensional KIVA-3V code release [6,7] has been used to model the turbulent combustion process of a multicomponent fuel, while one-dimensional calculations have been done to estimate the gas exchange processes. The investigation has been carried out considering different torque values at a fixed speed, 1500 rpm. The transition from high to low torque is obtained by varying the gasoline–air equivalence ratio ((G)), while the hydrogen–air equivalence ratio ((H2 )) is set constant. The hydrogen amount feeding the engine has been assumed according to the lowest hydrogen–air equivalence ratio used in experimental tests carried out on a pure hydrogen engine.
2. Reformer/engine system modelling The performance of the IRES has been evaluated by theoretical analysis. In Fig. 1 the IRES flowsheet is shown.
Ta
activation temperature
Greek symbols , , ε (G) (CO) (H2 ) R Engine System j j
r
m
k
constants in the kinetic combustion model turbulent kinetic energy dissipation engine air–gasoline equivalence ratio engine air–carbon monoxide equivalence ratio engine air–hydrogen equivalence ratio reformer thermal efficiency engine thermal efficiency overall thermal efficiency density of species j constant in the hybrid combustion model stoichiometric coefficient of species j in the combustion reaction combustion reaction rate EBU model reaction rate kinetic model reaction rate
The overall thermal efficiency of the IRES can be calculated as the following: System =
(mADD fuel
Li , + mR fuel ) · LHVfuel
(1)
where the reformer auxiliaries’ work is neglected. Therefore, to evaluate the efficiency of this system, numerical investigations have been carried out using both a thermochemical model for the plasma reformer and a CFD model for the engine. 2.1. Reformer modelling In a previous paper [3] a numerical model was developed to reproduce the operating conditions of a plasma reactor that assists the exothermic POX of hydrocarbons (gaseous or liquid). The reforming system was modelled by means of Aspen Plus code [4]. The conversion of gasoline to hydrogen was investigated and a thermodynamic analysis of the reformer was conducted. The thermochemical model was tested by comparing numerical results with experimental data. The model of the fuel processor consists of: a plasma reformer, a water separation unit and some auxiliary units; the flowsheet is shown in Fig. 1. It has been achieved by considering the integration of the reformer with an internal combustion engine. Gasoline (mR fuel ), after vaporizing, is fed into the reactor, combined with air and converted into syngas. Then syngas is cooled and the water is separated. Syngas and additional gasoline (mADD fuel ) are fed into the engine. A brief description of the main blocks used for the modelling is listed in Table 1.
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AIRADD
mfuel
C8H15
msyngas WATER
Li
SYNGAS
SEPARAT
MIXTURE
HEATER
ENGINE
EXHAUST
4 R
mfuel 1
GASOLINE
OUT-GAS 2
PUMP
REACTOR
EVAP
AIR
COMPR 3 Fig. 1. Integrated reformer/engine system model.
Table 1 Main blocks description
Table 2 Reformate gas from gasoline reforming [3]
Phenomena
Block
Name
Gas composition
Numerical results (vol%)
Plasma reforming Cooling Water separation
RGIBBS HEAT-EX SEPARATOR
REACTOR HEATER SEPARAT
CH4 N2 CO2 H2 CO LHV (kJ/kg)
1% 55% 3% 21% 20% 5023
In the following, some of the simulation hypotheses are listed. • The lack of detailed kinetic data (reaction rates, residence time, known intermediate species, etc.) has deterred a kinetic modelling of the reforming system. Therefore, the simulation has been run with a Gibbs reactor that simulates a chemical reactor by solving the heat and mass balances on minimizing the free energy of the components that can be produced during a reforming process. The method of minimizing the Gibbs free energy is normally preferred in the fuel reforming analysis [4,8] because the results provide a good prediction on the chemical composition of the reformate gas. On the other hand, the Gibbs reactor successfully emulated the experimental results [3]. • The possible species that might be found in the reformate gas are: N2 , CH4 , CO, CO2 , H2 O, H2 .
To analyse the operating condition of the reforming system, the air ratio (O/C) parameter has been defined: nO2 nO2 (O/C) = . (2) nfuel nfuel complete_combustion The O/C ratio is 0.37 and it has been assigned according to the experimental test on gasoline reforming. The temperature of the Gibbs reactor, that affects significantly the equilibrium compositions, is assumed equal to 968 K. In correspondence with this temperature, a good agreement between the syngas composition calculated by the model and the experimental data was obtained. The simulation results in terms of reformate gas composition are shown in Table 2. The reformer thermal efficiency has been calculated by considering both hydrogen and carbon monoxide in the syngas
E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538 Table 3 Characteristic constants for the combustion model
(Eq. (3)). The calculated efficiency is about 80%. R =
msyngas · LHVsyngas mR fuel · LHVfuel
.
(3)
2.2. Engine modelling The numerical approach has been described in detail in [1,9]. Briefly, both 1D and 3D calculations have been carried out: 3D simulations, starting at intake valve closing and ending at exhaust valve opening, have been performed by means of the Kiva3V code. Twelve chemical species have been considered: gasoline (C8 H15 ), H2 , O2 , N2 , CO2 , H2 O, H, O, N, OH, CO and NO. One-dimensional calculations have been done in order to estimate the initial conditions for CFD runs: at intake valve closing all the in-cylinder scalar variables have been considered uniform, while an initial tumble motion has been considered. Considering the reformate gas calculated previously, the Kiva3V code has been modified in order to simulate the combustion of hydrogen, gasoline and carbon monoxide (the methane combustion has been neglected). A serial mechanism based on the following overall reactions has been considered: 2H2 + O2 ⇒ 2H2 O,
(a)
4C8 H15 + 47O2 ⇒ 32CO2 + 30H2 O, (b) 2CO + O2 ⇒ 2CO2 .
(c)
Each oxidation rate ( r ) has been estimated by the following hybrid model [9] (Eq. (4)), in order to consider both turbulent ( m , Eq. (5)) and chemical kinetic aspects ( k , Eq. (6)): ⎧ for m k , ⎨ r = m (4) ⎩ 1 = + 1 − for m < k ,
r
m
k O2 ε f , ,
m = A · min k f · mw f O2 · mw O2 B · p , (5) f · mw f + O2 · mw O2
k = Cf · cf · cO2 · cH2 O · e−Ta /T .
2535
Reaction
A, B,
Cf
Ta (K)
a b c
18, 0.5, 0.44
1.8 × 1013 46 × 1010 1.9 × 106
1.0 0.25 1
1.0 1.5 0.3
0.0 0.0 0.5
1.76 × 104 3 × 104 8.056 × 103
gas enriched gasoline. The engine is the FIAT FIRE 1200 16 V, and features four cylinders (total piston displacement 1242 cm3 ) and a geometrical compression ratio equal to 10.6. The engine performance has been calculated at 1500 rpm and at different loads. The wide flammability limits and the high flame speed characterising the hydrogen–air mixtures allow the engines fuelled with pure hydrogen to run with very lean mixtures [15,16]. In this paper, it has been assumed that the hydrogen amount contained in the syngas is able to allow stable combustion phase if the relative air–hydrogen equivalence ratio is equal to a minimum value. This minimum value has been assumed according to the lowest hydrogen–air equivalence ratio mentioned in the technical literature [17–19]. Thus, the engine load can be controlled by varying the fuel flow-rate at wide open throttle, limiting the pumping losses at part load. In particular, by keeping constant the minimum air–hydrogen equivalence ratio, it is possible to increase the engine load, by dispatching a variable gasoline amount directly to the engine. In this way, both the reformer size is curbed and the engine power is not damaged too much because of the poor syngas heating value. Two kinds of fuelling have been considered: • gasoline–reformate gas, in which the syngas flow rate is able to allow a hydrogen–air equivalence ratio (H2 ) equal to 0.2 (Case A); • gasoline–reformate gas, in which the syngas flow rate is able to allow a hydrogen–air equivalence ratio equal to 0.14 (Case B).
(6)
The model has been tuned in order to reproduce the behaviour of the FIAT FIRE 1200 16 V engine fuelled with pure gasoline; assuming that , A and B depend on the turbulent structure characterizing the engine alone, they have been kept constant for each oxidation mechanism. The chemical property of the fuels has been considered in Eq. (6). All parameters of the combustion models are summarized in Table 3 [10–12]. To estimate the pollutant production, the CO2 dissociation and the NO production have been calculated, in the burned gas zone alone, according to the Meintjes–Morgan model [13] and the extended Zeldovich mechanism [14], respectively. 2.3. Engine performance Numerical investigations have been carried out to predict the performance of a spark ignition engine fuelled with reformate
Considering the reformate gas composition calculated by the reformer modelling, for Case A the carbon monoxide–air equivalence ratio (CO) is equal to 0.19, while for Case B it is equal to 0.13. As mentioned above, different loads are obtained by varying the gasoline amount (G). The highest air–gasoline ratio considered is equal to 0.6. The main numerical results are shown in Table 4. In each case studied the spark-timing has been optimized by considering the maximum brake thermal efficiency. To consider knock risk and mechanical stresses, spark-advances producing a pressure rise less than 2 bar/◦ have been assumed. The engine thermal efficiency has been evaluated as following: Engine =
mADD fuel
Li . · LHVfuel + msyngas LHVsyngas
(7)
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It can be observed that the highest efficiency (Engine = 0.43) has been obtained for Case B, at (G) = 0.2; moreover for (G) ranging between 0.1 and 0.5 the efficiency is higher than 0.4. The lowest efficiency (Engine = 0.28) has been calculated for Case B, at (G) = 0.0. This low efficiency is due to the very slow combustion caused by the ultra-lean mixture considered. The calculated efficiencies for the dual fuel (reformate gas and gasoline) engine have been compared to those of the same engine fuelled with straight gasoline (conventional engine). The comparison is shown in Fig. 2. Obviously, for the conventional gasoline engine different loads are obtained by keeping a stoichiometric air–fuel ratio (except at full load, IMEP≈ 10 bar) and reducing the volumetric efficiency by throttling. The lower energy density of the air–reformate gas/gasoline mixture, compared to the air–gasoline mixture, significantly reduces the maximum engine power output (about 10% for (H2 ) = 0.2 and 14% for (H2 ) = 0.14). Nevertheless, at the same load, the dual fuel engine shows higher efficiencies than the gasoline engine. At middle load the efficiency increases about 25%. Fewer pumping losses, due to the adopted load control strategy, as well as thermal advantages due to the lean burning, justify this trend. It is worth nothing that in Case A, the dual fuel engine shows a minimum load (IMEP = 3.85 bar) higher than that
Table 4 Engine performance (speed 1500 rpm, throttle 100%) Case A
Case B
(H2 )=0.20; (CO) = 0.19 (H2 ) = 0.14; (CO) = 0.13 (kg/kg) 0 0.1 0.2 0.35 0.5 0.6
IMEP (bar) 3.85 5.35 6.32 7.72 8.74 9.06
Engine
Engine
IMEP (bar) 2.17 4.08 5.23 6.66 7.87 8.60
(dimensionless) 0.36 0.42 0.42 0.42 0.40 0.38
(dimensionless) 0.28 0.41 0.43 0.42 0.41 0.40
3. System performance analysis The CFD numerical results have shown very interesting efficiencies for the stand alone hydrogen enriched gasoline engine, higher than that of the gasoline engine. However, these results are not complete. To evaluate the performance of the dual fuel engine, it is also necessary to consider the performance of the hydrogen generation system. The thermochemical model of the reformer, by evaluating the gasoline mass flow used to generate reformate gas for engine fuelling, allows to reach this target. As mentioned above, the overall thermal efficiency of the IRES has been calculated as Eq. (1). Eq. (8) shows the correlation between the three considered efficiencies: System Engine
=
ADD 1 + (mR fuel /mfuel ) · R ADD 1 + (mR fuel /mfuel )
.
• at low loads the efficiency of the gasoline engine is higher than the overall efficiency of the IRES (Cases A and B); • at middle loads the efficiency of the gasoline engine is very close to the IRES efficiency;
0.4 0.3 0.2 Conventional Engine (Gasoline) Dual Fuel Engine (Case A) Dual Fuel Engine (Case B)
0.1 0 0
1
2
3
(8)
The efficiency of the IRES depends on both the reformer effiADD ciency and the mR fuel /mfuel ratio. The calculated results for the IRES and the conventional engine are shown in Fig. 3. Of course, the overall thermal efficiency of the IRES is lower than that of the dual fuel engine especially for Case A, where the highest hydrogen concentration means a higher reformer ADD fuel consumption, i.e. a higher mR fuel /mfuel ratio. Comparing the IRES with the gasoline engine efficiencies, it is worth noting that
Engine Thermal Efficiency
0.5
ηengine [-]
(G)
of the gasoline engine (IMEP = 2.3 bar). Of course, lower loads can be reached by throttling. In this way, above IMEP∼ 4 bar, the engine load is regulated by varying the gasoline amount at wide open throttle, while below IMEP∼ 4 bar the throttle opening controls the engine load (at (H2 )=0.2, (CO)=0.19, (G) = 0.0).
4
5 6 IMEP [bar]
7
Fig. 2. Engine thermal efficiency.
8
9
10
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E. Galloni, M. Minutillo / International Journal of Hydrogen Energy 32 (2007) 2532 – 2538
2537
Overall Thermal Efficiency
0.5
ηSystem [-]
0.4 0.3 0.2 Conventional Engine (Gasoline) Reformer/Engine (Case A) Reformer/Engine (Case B)
0.1 0 0
1
2
3
4
5 6 IMEP [bar]
7
8
9
10
11
Fig. 3. Overall thermal efficiency of the integrated reformer/engine system.
9.E-03
• at middle and high loads the calculated CO production of the IRES is much lower than that of the engine fuelled with straight gasoline, while at low loads the CO increases because of the very slow combustion, due to ultra-lean mixtures burned in the dual fuel engine; • for the IRES, the calculated NO production is very close to zero at low and middle loads, while at high loads the NO production increases because of oxygen overabundance and high temperature inside the cylinder, characterising the dual fuel engine.
Engine (Gasoline) Reformer/Engine (Case A) Reformer/Engine (Case B)
8.E-03 7.E-03 CO [g]
6.E-03 5.E-03 4.E-03 3.E-03 2.E-03 1.E-03 0.E+00 0
1
2
3
4
5 6 7 IMEP [bar]
8
9
10 11
4. Conclusions
Fig. 4. Overall CO emissions per engine cycle.
5.E-04
Engine (Gasoline) Reformer/Engine (Case A) Reformer/Engine (Case B)
4.E-04 4.E-04 NO [g]
3.E-04 3.E-04 2.E-04 2.E-04 1.E-04 5.E-05 0.E+00 0
1
2
3
4
5 6 7 IMEP [bar]
8
9
10 11
Fig. 5. Overall NO emissions per engine cycle.
• at high loads the gasoline engine efficiency is lower than that of the IRES. Therefore, it is worth noting that the mean overall efficiency of the systems seems to be very similar. With regard to emissions, it is possible to underline that comparable efficiencies mean comparable CO2 emissions. However, considering pollutant emissions, as shown in Figs. 4 and 5
In this paper numerical investigations have been conducted to predict the performance of an IRES, in which the reformer produces the syngas burned together with straight gasoline in a spark ignition engine. The thermochemical modelling of the reformer has allowed to calculate the syngas composition and the efficiency of the fuel processor. The calculated efficiency of the gasoline reforming is equal to 80%. The considered engine, fuelled with syngas and gasoline, has shown a much higher efficiency than the engine fuelled with pure gasoline, according to CFD analysis. The performance of this integrated system has been compared to the performance of the same spark ignition engine fuelled with pure gasoline. The results of the computational analysis have pointed out that the mean overall efficiency of IRES is close to that of the conventional spark ignition engine. Regarding pollutant emissions, the numerical analysis has shown advantages at middle load operation. At the end, it is appropriate to highlight that these results have been obtained considering an existing gasoline engine; therefore, further improvement in the IRES performance can be achieved by designing a dual fuel engine, able to effect the full potential of hydrogen in internal combustion engines.
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