The hydrogen-fueled internal combustion engine: a technical review
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International Journal of Hydrogen Energy 31 (2006) 1292 – 1305
www.elsevier.com/locate/ijhydene
The hydrogen-fueled internal combustion engine: a technical review
C.M. White∗ , R.R. Steeper, A.E. Lutz
Sandia National Laboratories, Combustion Research Facility, P.O. Box 969, MS 9053, Livermore, CA 94551-0969, USA
Received 12 May 2005; received in revised form 21 November 2005
Available online 30 January 2006
Abstract
A review is given of contemporary research on the hydrogen-fueled internal combustion engine. The emphasis is on light-
to medium-duty engine research. We first describe hydrogen-engine fundamentals by examining the engine-specific properties
of hydrogen and surveying the existing literature. Here it will be shown that, due to low volumetric efficiencies and frequent
preignition combustion events, the power densities of premixed or port-fuel-injected hydrogen engines are diminished relative to
gasoline-fueled engines. Significant progress has been made in the development of advanced hydrogen engines with improved
power densities. We discuss several examples and their salient features. Finally, we consider the overall progress made and
provide suggestions for future work.
䉷 2006 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved.
Keywords: Internal combustion engine; Review
1. Introduction out emissions,1 formed by the thermal dissociation and
oxidation of N2 in atmospheric air during combustion,
The potential for hydrogen-fueled internal combus- and (ii) the low lean-flammability limit of hydrogen
tion engines (H2 ICEs) to operate as clean and efficient allows stable combustion at highly dilute conditions.
power plants for automobiles is now well-established. The coupled effect is that during ultra-lean operation,
In particular, H2 ICEs with near-zero emissions and combustion temperatures are low enough such that NOx
efficiencies in excess of conventional gasoline-fueled formation rates are too slow and engine-out emissions
ICEs have been demonstrated (e.g. [1–3]). The abil- are near zero (for a summary, see [4]). The ability to
ity for H2 ICEs to operate with near-zero engine-out operate efficiently is in part also owed to (ii) because
emissions is primarily owed to the coupled effect of unthrottled operation is possible at low-loads.
two characteristics unique to hydrogen: (i) in principle, The unique combustion characteristics of hydrogen
nitrogen oxides (NOx ) are the only undesirable engine that allow clean and efficient operation at low engine
loads present difficulties at high engine loads. Here, the
low ignition energies of hydrogen–air mixtures cause
∗ Corresponding author. Tel.: +1 925 294 4904; 1 In practice, the burning of lubricating oil in the combustion
fax: +1 925 294 1004. chamber produces carbon oxides (COx ) and hydrocarbons (HC) at
E-mail address: cmwhit@sandia.gov (C.M. White). near-zero levels [4].
0360-3199/$30.00 䉷 2006 International Association for Hydrogen Energy. Published by Elsevier Ltd. All rights reserved.
doi:10.1016/j.ijhydene.2005.12.001C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305 1293
frequent unscheduled combustion events, and high com- but the focus here is to expand the understanding in the
bustion temperatures of mixtures closer to the stoichio- context of a survey of recent literature. Advanced hy-
metric composition lead to increased NOx production. drogen engines and their salient features are reviewed in
Both effects, in practical application, limit the power Section 3. The intent of these engines is to overcome the
densities of H2 ICEs. The recent research thrust and limiting effects at high engine loads for naturally aspi-
progress on this front is the development of advanced rated premixed or port-fuel-injection (PFI) H2 ICEs. We
hydrogen engines with improved power densities and conclude with some remarks regarding the overall ben-
reduced NOx emissions at high engine loads. efit of H2 ICEs and possible future work in Section 4.
There are several reasons for reviewing recent
developments in H2 ICE engine-specific research: recent 2. H2 ICE Fundamentals
progress in control strategies for NOx reduction, devel-
opment of advanced hydrogen engines with improved The ability for H2 ICEs to burn cleanly and operate
power densities, and recent progress in engine model- efficiently is owed to the unique combustion charac-
ing and cycle analysis. Perhaps the most compelling teristics of hydrogen that allow ultra-lean combustion
reason is the ever-increasing likelihood that the H2 ICE with dramatically reduced NOx production and effi-
will serve as a transitional hydrogen powertrain during cient low-engine load operation. In contrast, the same
the initial development of a hydrogen economy. This combustion characteristics impose technical challenges
view is based on the fact that the implementation of a at high engine-loads due to an increased propensity to
production ready hydrogen proton exchange membrane preignite the hydrogen–air mixture and increased NOx
fuel cell (PEMFC) is at least 10 years away [5]. In production. In this section we review the benefit and
contrast, the H2 ICE offers the potential to utilize technical challenge of H2 ICE operation at low and high
manufacturing infrastructure already developed for engine-load, through an examination of hydrogen prop-
petroleum-fueled ICEs, and can serve as an economi- erties relevant to engine operation and control. Hydro-
cal near- to mid-term option for a transportation power gen properties of interest are listed in Table 1 [11–13].
plant in a hydrogen economy while fuel cells undergo The same properties for typical gasoline [11,14] and
continued development. This view has been the accel- compressed natural gas (CNG) [11,15] are also pro-
erant behind a renewed interest and recent progress in vided. For direct comparison, H2 ICE engine properties
the research and development of the H2 ICE. will be compared to those of a PFI gasoline engine.
1.1. Scope 2.1. Preignition and knock
The literature on the H2 ICE is voluminous, and dates The high autoignition temperature of hydrogen
back over one-hundred years [6]. H2 ICE research prior (858 K) means that hydrogen is most suitable as a
to the early 1980s has been reviewed by Escher [7–9],
and an excellent historical perspective of the hydrogen Table 1
Fuel properties at 25 ◦ C and 1 atm
engine and technical review of H2 ICE research prior to
1990 can be found in Das [10]. For the most part, we Property Hydrogen CNG Gasoline
consider here more recent H2 ICE research, with an em-
Density (kg/m3 ) 0.0824 0.72 730a
phasis on the last 10 years. The review is intended for Flammability limits (volume % in air) 4–75 4.3–15 1.4–7.6
those who have a familiarity with internal combustion Flammability limits ( ) 0.1–7.1 0.4–1.6 ≈ 0.7–4
engines but are not familiar with the hydrogen internal Autoignition temperature in air (K) 858 723 550
combustion engine. While it is not possible to be inclu- Minimum ignition energy (mJ)b 0.02 0.28 0.24
sive of all subject matter in the space provided here, the Flame velocity (m s−1 )b 1.85 0.38 0.37–0.43
Adiabatic flame temperature (K)b 2480 2214 2580
present study provides a representative overview of the
Quenching distance (mm)b 0.64 2.1c ≈2
“state-of-the-art” engine-specific H2 ICE research. The Stoichiometric fuel/air mass ratio 0.029 0.069 0.068
scope of the review is as follows: H2 ICE fundamentals Stoichiometric volume fraction % 29.53 9.48 ≈ 2d
are described in Section 2 by examination of the engine- Lower heating value (MJ/kg) 119.7 45.8 44.79
specific properties of hydrogen. These properties will Heat of combustion (MJ/kgair )b 3.37 2.9 2.83
be used to show, with reference to ICE operation and
a Liquid at 0 ◦ C.
control, both the advantages and disadvantages of hy- bAt stoichiometry.
drogen compared to conventional gasoline-fueled ICEs. c Methane.
Past and recent studies cover some of the same ground, d Vapor.1294 C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305
10 1 preignition-limit is defined as the upper bound of this
envelope.
Complementary information to Fig. 1, and relevant to
minimum ignition energy, mJ
the above discussion, is that studies at pressures vary-
10 0 ing from 0.2–1 atm and mixture temperatures varying
from 273–373 K have found that the minimum ignition
energies of hydrogen–air mixtures vary inversely with
the square of the pressure and inversely with tempera-
10-1
ture [12]. If these trends hold at high pressure and tem-
perature, extrapolations indicate that for typical engine
temperatures and pressures, the minimum ignition ener-
gies of hydrogen–air mixtures will be much lower than
that shown in Fig. 1. Similar dependencies on pressure
10-2
0 0.5 1 1.5 2 2.5 3 3.5 4 and temperature for the minimum ignition energies of
φ hydrocarbon–air mixtures have also been reported [13].
Knock, or spark knock [14], is defined as autoignition
Fig. 1. Minimum ignition energies of (•) hydrogen–air, ()
methane–air and () heptane-air mixtures in relation to at atmo- of the hydrogen–air end-gas ahead of the flame front
spheric pressure [13]. that has originated from the spark. The high autoignition
temperature, finite ignition delay and the high flame ve-
locity of hydrogen (i.e., the latter two effects translate to
fuel for spark ignition (SI) engines, though compres- less residence time for the end-gas to ignite) means that
sion ignition (CI) has been studied in some detail knock, as defined, is less likely for hydrogen relative to
(e.g. [16–19]). Despite the high autoignition temper- gasoline, and hence the higher research octane number
ature, the ignition energies of hydrogen–air mixtures (RON) for hydrogen (RON > 120 [2,20]) in compari-
are approximately an order-of-magnitude lower than son to gasoline (RON = 91 − 99 [14]). The global effect
that of hydrocarbon–air mixtures. This is observed of both knock and preignition (i.e., an audible pinging
in Fig. 1 where the minimum ignition energies for or “knocking”) is nearly indistinguishable, and the pri-
hydrogen–air, propane–air and heptane–air mixtures mary reason for the lack of a clear distinction between
at atmospheric pressure are plotted as a function of the two in the hydrogen literature [2]. However, a dis-
equivalence ratio, , where is defined as the ratio tinction is necessary since the controlling phenomena
of the actual fuel/air mass ratio to the stoichiomet- are very different: preignition can be avoided through
ric fuel/air mass ratio. The low ignition energies of proper engine design, but knock is an inherent limit on
hydrogen–air mixtures mean that H2 ICEs are predis- the maximum compression ratio that can be used with a
posed towards the limiting effects of preignition. Here, fuel [14]. The differentiating factor between knock and
preignition is defined as combustion prior to spark preignition is that knock can be controlled by retarding
discharge, and in general, results from surface ignition the spark timing, while, in general, preignition cannot
at engine hot spots, such as spark electrodes, valves be controlled by adjusting spark timing.
or engine deposits. The limiting effect of preignition Observed in Fig. 1 is that as the stoichiometric con-
is that a preignition event will advance the start of dition ( = 1) is approached from the lean side ( < 1),
combustion and produce an increased chemical heat- the minimum ignition energy for hydrogen is a strongly
release rate. In turn, the increased heat-release rate decreasing function of the equivalence ratio with the
results in a rapid pressure rise, higher peak cylinder minimum at ≈ 1. This trend is responsible for the
pressure, acoustic oscillations and higher heat rejection experimentally observed fact that it is extremely diffi-
that leads to higher in-cylinder surface temperatures. cult to operate an H2 ICE at or near the stoichiometric
The latter effect can advance the start of combustion condition in the absence of frequent preignition events.
further, which in turn can lead to a runaway effect, Therefore, for practical application, the maximum
and if left unchecked will lead to engine failure [14]. and, consequently, peak power output can be limited
It is therefore a necessity for practical application that by the preignition-limit. Stockhausen et al. [21] report
preignition is avoided. The preignition-limited oper- a preignition-limit of ≈ 0.6 for a 4-cylinder 2.0-l
ating envelope defined by Tang et al. [2] consists of engine at an engine speed of 5000 rpm. Consequently,
operating conditions where 1% or fewer of the com- engine peak power output was reduced by 50% com-
bustion cycles experience a preignition event. The pared to engine operation with gasoline.C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305 1295
Although the preignition-limit is engine specific, 300
consistent trends with variations in engine proper- [O2]/[O2 + N2]=0.21
ties and operational conditions have been found: the 250
−1
preignition-limited decreased monotonically with 0.175
laminar flame velocity, cms
increased compression ratio (CR) [2,22,23] and in- 200
creased mixture temperature [22]. Engine speed has 0.15
also been shown to have an effect [2,23] but the trend is 150
more complicated due to the coupled effect of residual 0.125
mass fraction (i.e., mixture-temperature effect). 100
2.1.1. Preignition control 50
The diminished peak power output, set by the
preignition-limit, will decrease the performance of a
0
H2 ICE-powered vehicle in comparison to its gasoline 0 0.5 1 1.5 2 2.5 3 3.5 4
φ
equivalent. Therefore, determining the mechanism of
preignition, practical operational limits, and control Fig. 2. Laminar flame velocity for (—) hydrogen, oxygen and
strategies has been a primary focus of many research nitrogen mixtures [13] and (◦, - -) gasoline and air mixtures [14]
studies (e.g. [1,22–26]). Unfortunately, despite much at room temperature and atmospheric pressure. The dashed line is
effort, there exist no guaranteed preventive steps. How- a least squares fit polynomial. [O2 ]/[O2 + N2 ] is the oxygen mole
fraction in the ambient.
ever, identification of preignition sources, such as in-
cylinder hot-spots [1,24], oil contaminants [21], com-
bustion in crevice volumes [27] and residual energy
in the ignition system [26], has provided the neces- 2.2. Flammability range, flame velocity and adiabatic
sary minimizing steps. These include use of cold-rated flame temperature
spark plugs, low coolant temperature and optimized
fuel-injection timing. The flammability range, in fuel volume fraction in
Advanced control strategies include intake charge air at 298 K and 1 atm, for hydrogen is 4–75% and
cooling [28], variable valve timing for effective scav- for gasoline is 1–7.6%. For ICEs, it is more mean-
enging of exhaust residuals [29], advanced ignition sys- ingful to give the flammability range in terms of
tems [26], and hydrogen direct injection (DI) [24,30]. equivalence ratio. Then, the flammability range of hy-
These advanced strategies can be quite effective: drogen is 0.1 7.1, and gasoline is approximately
Berckmüller et al. [29] using a single-cylinder 0.5-l 0.7 < < 4. In recasting the flammability range in
PFI-H2 ICE at CR of 12:1 used variable cam phasing terms of equivalence ratio, it is evident that the H2 ICE
and computational fluid dynamics (CFD) simulations is amenable to stable operation under highly dilute
to optimize coolant flow, injection location, and in- conditions, which allows more control over engine op-
jection timing to operate at = 1 over a speed range eration for both emissions reduction and fuel metering.
2000–4000 rpm. Homan et al. [30] used DI late in the In practical application, the diluent can either be excess
compression stroke and near-simultaneous spark to air (lean operation) or recycled exhaust gas. An impor-
eliminate preignition events. However, to prevent mis- tant distinction between the two is that the latter allows
fires, the hydrogen jet had to be directed towards the stoichiometric operation at low to medium load and the
spark. Kondo et al. [26] used an ignition system specif- use of a three-way catalyst (TWC) for NOx reduction.
ically designed to prevent residual energy and a water However, problems due to water condensation using
cooled spark plug. In the absence of any advanced con- recycled exhaust gas dilution in ICEs is exacerbated in
trol, maximum ≈ 0.35; by the elimination of residual H2 ICEs due to the large fraction of H2 O in the exhaust
energy in the ignition system, maximum ≈ 0.6; with stream.
the addition of the water-cooled spark plug, maximum Flame velocity and adiabatic flame temperature are
≈ 0.8. With liquid-hydrogen fueling, Furuhama important properties for engine operation and control, in
et al. [31] and Knorr et al. [32] report that preignition particular thermal efficiency, combustion stability and
can be eliminated without too much effort simply due emissions. Laminar flame velocity and flame temper-
to the cooling effect of the cold hydrogen. A further ature, plotted as a function of equivalence ratio, are
discussion of liquid-hydrogen-fueled engines follows shown in Figs. 2 and 3, respectively. These figures will
in Section 3.2. be further referenced in subsections to follow.1296 C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305
2600 104
Adiabatic flame temperature, K
2400
2200
103
NOx, ppm
2000
102 Tier II fleet
average
1800
SULEV
1600
101
1400
NOx limited φ
1200 100
0 0.5 1 1.5 2 2.5 3 3.5 4 0 0.25 0.5 0.75 1 1.25 1.5
φ φ
Fig. 3. Adiabatic flame temperature for hydrogen-air mixtures [12]. Fig. 4. NOx emissions as a function of equivalence ratio for en-
gine-out (closed symbols) and tail-pipe with exhaust gas after treat-
ment (open symbols) from various studies: • [2], for CR=12.5, 14.5
and 15.3; [29]; [4] for CR = 11 : 1; [36]; [31], for intake
2.3. Nitrogen oxides (NOx ) temperature of 298 K; [36]; [34], for fuel input of 1.39 kg/h.
The dashed and solid lines represent the US Federal Tier II man-
H2 ICE emissions and control techniques have been ufacturer fleet average NOx standard and CARB LEV II SULEV
thoroughly reviewed by Das [4]. Here we provide a brief NOx standard, respectively, for a fuel efficiency of 35 mpg.
survey of more recent NOx research.
Ultra-lean combustion (i.e., 0.5), which is ade-
quately synonymous with low temperature combustion, NOx -limited equivalence ratio for H2 ICE operation is
is an effective means for minimizing NOx emissions in essentially < 0.5.
ICEs. As described in Section 2.2, H2 ICEs can operate We have also compiled in Fig. 4 data from various
ultra-lean, and are therefore amenable to low temper- sources for tailpipe emissions with exhaust after treat-
ature NOx reduction strategies. Engine-out NOx con- ment (open symbols). These data show that NOx emis-
centration verses equivalence ratio is plotted in Fig. 4 sions at > 0.95 are near zero with the use of a three-
(closed symbols) from various studies. The claim of way catalyst (TWC). Berckmüller et al. [29] and Heffel
near-zero emissions can be made more robust by ana- [34,35] investigated charge dilution with exhaust gas
lyzing Fig. 4 in reference to the NOx limit of 0.07 and recirculation (EGR) to provide a stoichiometric air/fuel
0.02 g/mile to satisfy, respectively, US Federal Tier II ratio at low- to medium-load operation. With EGR dilu-
manufacturer fleet average standard and California Air tion, a TWC, and =1, tail-pipe emissions were signif-
Resources Board (CARB) Low Emissions Vehicle II icantly reduced for engine loads that would otherwise
(LEV II) standard for Super Ultra Low-Emissions Vehi- produce high NOx concentrations for dilution with air
cles (SULEV).2 Conversion of NOx concentrations in (i.e., 0.5 < < 1.02). For instance, with excess air at
g/mile to ppm (or vice-versa) depends on vehicle fuel ef- =0.78, Heffel [34] measured 7000 ppm of NOx ; hold-
ficiency, drive cycle and equivalence ratio, among other ing the fueling constant and using EGR as the diluent,
factors [33]. Here we consider a vehicle with a fuel ef- such that = 1, NOx emissions were reduced to less
ficiency of 35 miles per gallon (mpg) operated at steady than 1 ppm at the tailpipe. Lee et al. [36] using a TWC
state. Then, as a function of equivalence ratio, the Tier measured 0.04 g/mile NOx emissions for a H2 ICE ve-
II fleet average and SULEV NOx limits are shown in hicle on a chassis dynamometer on the U.S. Federal
Fig. 4 as the dashed and solid lines, respectively. For Test Procedure (FTP75) drive cycle. Berckmüller et al.
different fuel efficiency these lines would effectively be [29] found that the maximum level of EGR while main-
translated vertically (i.e., lower for lower mpg or higher taining smooth engine operation was 50%. Maximum
for higher mpg). However, due to the rapid increase in EGR levels for gasoline operation are 15–30% [14]. The
NOx production with increasing near ≈ 0.5, the higher EGR tolerance of the H2 ICE is due to the high
flame velocities of hydrogen–air mixtures as shown in
2 The SULEV NO standard is equal to the now-abandoned
x the comparison of hydrogen and gasoline data in Fig. 2.
Equivalent Zero Emissions Vehicle (EZEV) standard. The hydrogen curve labeled [O2 ]/[O2 + N2 ] = 0.125 isC.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305 1297
equivalent to 40% mole fraction dilution of air with ni- large passageways with low pressure drop, possible with
trogen, providing an approximation of flame speeds for hydrogen fueling since high intake velocities required
the case of 40% EGR dilution. This approximation ig- for fuel atomization at low engine speeds are not nec-
nores the effect of water contained in hydrogen-engine essary [37]. With the use of a large diameter manifold,
EGR, but at an equivalence ratio of 1, Fig. 2 indicates a Swain et al. reported a 2.6% increase in peak power
significantly higher flame speed for the diluted hydro- output compared to that for a small diameter manifold.
gen mixture compared with the undiluted gasoline case. However, the improvement was lower than the estimated
In summary, without aftertreatment there is a trade- 10% that was expected. One possible explanation for
off between H2 ICE power output and NOx emissions. the less-than-expected performance improvement was
Considering that the primary benefit of an H2 ICE is that the intake flow dynamics with hydrogen fueling
near-zero emissions, a practical limit of engine opera- are more complex than for gasoline-fueled engines. In
tion is < 0.5 (assuming no EGR). This restriction will this context, Sierens and Verhelst [39] found that the
translate to a large loss in the effective power density start and duration of injection influences volumetric
of an H2 ICE. The potential to expand the power band efficiency due to the interaction between the injected
while maintaining near-zero NOx emissions is possible hydrogen and the intake pressure waves.
by charge dilution with EGR and use of a TWC or by While PFI-H2 ICEs suffer from low volumetric effi-
improving lean power density with pressure boosting, ciency at high loads, at low- to medium-loads H2 ICEs
as will be discussed in Section 3.1. offer the benefit of being able to operate unthrottled.
The advantage is that the pumping loss due to the pres-
2.4. Power output and volumetric efficiency sure drop across the throttle plate is eliminated, and fuel
efficiency is improved. The ability for the H2 ICE to op-
H2 ICE peak power output is primarily determined by erate unthrottled is owed to the low lean-flammability
volumetric efficiency, fuel energy density, and preigni- limit and high flame-velocity of hydrogen. However,
tion. For most practical applications, the latter effect due to increasing amounts of unburned hydrogen at
has been shown to be the limiting factor that determines ultra-dilute conditions, some throttling is required at
peak power output. idle conditions [1].
Premixed or PFI-H2 ICEs inherently suffer from a loss
in volumetric efficiency due to the displacement of in- 2.5. Thermal efficiency
take air by the large volume of hydrogen in the intake
mixture. For example, a stoichiometric mixture of hy- The high RON and low lean-flammability limit of hy-
drogen and air consists of approximately 30% hydrogen drogen provides the necessary elements to attain high
by volume, whereas a stoichiometric mixture of fully- thermal efficiencies in an ICE. Brake thermal efficiency
vaporized gasoline and air consists of approximately (BTE) versus brake mean effective pressure (BMEP) for
2% gasoline by volume. The corresponding power den- various sources is plotted in Fig. 5. For direct compar-
sity loss is partially offset by the higher energy content ison between the various studies, we have normalized
of hydrogen. The stoichiometric heat of combustion per BMEP by maximum BMEP for equivalent gasoline op-
standard kg of air is 3.37 MJ and 2.83 MJ, for hydrogen eration. Dashed lines through the data are best-fit poly-
and gasoline, respectively. It follows that the maximum nomials and the solid line is the BTE of a 4-cylinder
power density of a pre-mixed or PFI-H2 ICE, relative 1.6-l gasoline engine with a CR of 9.0:1 [38]. The data
to the power density of the identical engine operated of Tang et al. [2] at CR of 14.5:1 are illustrative of the
on gasoline, is approximately 83% [31]. For applica- increase in BTE with higher CR that is possible with
tions where peak power output is limited by preigni- hydrogen. Tang et al. [2] and Nagalingam et al. [40]
tion, H2 ICE power densities, relative to gasoline oper- found CR of approximately 14.5:1 to be optimal due to
ation, can be significantly below 83%. Furuhama et al. heat transfer losses at higher CR.
[31] and Tang et al. [2] report preignition-limited power Aside from the increase in BTE by increasing CR,
densities of 72% and 50%, respectively, relative to op- H2 ICEs have higher efficiencies than gasoline ICEs at
eration with gasoline. similar CR. This is observed by comparing the gasoline
Presently, without exception, H2 ICEs are modified and hydrogen data sets of [38]. Compared to gasoline
conventional gasoline (or Diesel) engines, with vary- operation (i.e., solid curve), the BTE with hydrogen op-
ing degrees of modification. Swain et al. [37] designed eration (triangles) is higher across the entire operating
an intake manifold to take advantage of the character- range, with the relative increase maximum at medium
istics of hydrogen. The important feature is the use of loads. The drop-off in the relative difference in BTE1298 C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305
0.4 density, NOx emissions, and thermal efficiency. For
the most part, we consider only advanced reciprocating
0.35
engine concepts.
0.3
3.1. Pressure-boosted H2 ICE
0.25
BTE
0.2 3.1.1. Overview of boosted H2 ICE research
Intake-air pressure-boosting is an effective and
0.15
proven strategy for increasing peak engine power in
0.1 conventional petroleum-fueled ICEs. For premixed
or PFI-H2 ICEs, pressure boosting is likely necessary
0.05
to achieve power densities comparable to gasoline
0 engines and, fortunately, application is straightforward.
0 0.2 0.4 0.6 0.8 1 1.2
Although research examining the strategy has been
BMEP*
active for decades, optimization of boosted H2 ICEs is
Fig. 5. Brake thermal efficiency (BTE) as a func- far from complete.
tion of normalized brake mean effective pressure, where Early work testing boosted H2 ICEs has been carried
BMEP∗ = BMEP/BMEPmax .gasoline . ◦, [2] for CR = 14.5 and out by Nagalingam et al. [40], Furuhama and Fukuma
12.5, respectively; , — [38] for hydrogen and gasoline, respec-
tively, at CR = 9 : 1. The dashed lines through the hydrogen data
[42] and Lynch [43]. Nagalingam et al. worked with
are least squares fit polynomials. a single-cylinder research engine and simulated tur-
bocharged operation by pressurizing inlet air to 2.6 bar
and throttling the exhaust to mimic a turbine’s backpres-
between gasoline and hydrogen at low loads is due to sure. Researchers at the Musashi Institute of Technol-
the need for some throttling, as discussed in Section 2.4. ogy turbocharged a liquid-hydrogen, two-stroke diesel
The drop-off at high loads is likely due to increasing engine and tested its performance on the bench and in a
heat transfer losses. Shudo et al. [41] showed that for an vehicle [42]. In early tests of turbocharged hydrogen en-
H2 ICE the relative fraction of the heat release lost by gines in commercial vehicles, Lynch converted gasoline
heat transfer to the cylinder walls increases monotoni- and Diesel engines to spark-ignited hydrogen operation
cally with increasing equivalence ratio. The trend is ex- at maximum inlet pressures of 1.5 bar absolute [43].
plained as a consequence of increasing flame velocity, More recently, substantial development has been
increasing flame temperature and decreasing quenching brought about by research efforts from BMW [29] and
distance with increasing equivalence ratio that leads to Ford [44,45]. Berckmüller et al. [29] have reported
narrow thermal boundary layers. Shudo et al. [41] re- results from a single-cylinder engine supercharged to
ported that at = 0.4 the energy lost by heat transfer to 1.8 bar that achieves a 30% increase in specific power
the wall accounted for 25% of the total heat release, and output compared to a naturally aspirated gasoline en-
at =1 this percentage increased to 45%. An important gine. Natkin et al. [44] report results for a supercharged
conclusion is that improvements in H2 ICE efficiencies 4-cylinder 2.0-l Ford Zetec engine and a 4-cylinder
will require strategies to minimize heat transfer losses 2.3-l Ford Duratec engine that is used for conventional
to the cylinder walls (e.g., charge stratification). and hybrid vehicles [45]. Two Nissan engines tested
for hydrogen hybrid vehicle use at Musashi Institute of
Technology showed a similar 35% increase in power due
3. Advanced hydrogen engines to boosting while holding NOx emissions at 10 ppm [7].
Because boosting pressure increases charge pressure
In the previous section, we have shown that the power and temperature the problems of preignition, knock and
density of a naturally aspirated premixed or PFI-H2 ICE NOx control are heightened during boosted operation.
is inferior to the power density of the identical engine In addition Nagalingam et al. [40] reported that the
fueled with PFI gasoline. In practical application, the preignition-limited equivalence ratio decreased from 1
loss in power density can be as high as 50% [21]. It is down to 0.5 when they increased intake pressure from
not surprising then that much effort has been put forth 1 bar to 2.6 bar. In their work, water injection was used
in the development of advanced hydrogen engines with to mitigate the effects of higher charge temperatures
improved power densities. In this section, we review associated with boosted operation. Berckmüller et al.
several examples. Particular attention is paid to power [29] used other methods to mitigate increased chargeC.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305 1299
temperatures including optimizing coolant flow to the 200
knock limit
150
exhaust valve seats and spark plugs, and varying cam
peak charge pressure
100
phasing to improve residual scavenging. They reported
unburnt
a decrease in the preignition-limited equivalence ratio gas burnt
gas
from 1 to 0.6 when inlet pressure was increased from
1 bar to 1.85 bar. Intercooling is a commonly used strat-
Pressure [bar]
egy for addressing this problem as well: Ford’s experi- peak
charge
mental hydrogen engines incorporate dual intercoolers temperature
to maximize cooling of the supercharged air [44]. With 10 mean
gas
proper ignition-timing control, they were able to oper-
ate free of preignition at equivalence ratios up to 0.8. Speed = 1900 rpm
Phi = 0.325
As with preignition, NOx production is sensitive to Brake power = 105kW
BMEP = 11.3 bar
charge temperature, and pressure boosting aggravates 2.8 bar boost Brake torque = 525 Nm
the problem. Berckmüller et al. [29] required equiva- (abs) Brake eff= 39.0%
EGR = 0
lence ratios lower than 0.45 to operate below the knee 1
200 300 500 1000 2000 3000
of the NOx curve (occurring at NOx ≈ 100 ppm, by Temperature [K]
their definition). Similarly, Nagalingam et al. [40] pub-
Fig. 6. A simulated cycle representing optimized results from the
lished results showing NOx levels below 100 ppm for
WAVE model are displayed on log(P)–log(T) axes. Path of cycle is
equivalence ratios less than 0.4 when operating at su- clockwise on plot.
percharged intake pressures of 2.6 bar absolute. Much
the same, Natkin et al. [44] reported 90 ppm NOx emis-
sions at = 0.5. To achieve emissions of 3–4 ppm
(levels likely required to attain SULEV standards, see engineering simplicity: robust H2 PFI injectors have
Fig. 4), Ford’s supercharged engine was run at a leaner been identified [46] and turbochargers and intercoolers
of 0.23 [44]. The Musashi tests mentioned above are proven technologies. Clearly, the simplest strategy
achieved NOx levels below 10 ppm for equivalence ra- for boosted H2 ICEs is to operate lean enough to meet
tios leaner than 0.4 [42]. The scatter in the NOx -limited a given NOx limit, and available modeling tools are
equivalence ratio between the various studies is a con- useful for predicting expected performance. To explore
sequence of the dependency of NOx formation rates the limits of the lean-operation strategy, we have sim-
on CR, intercooling and in-cylinder mixing. In general ulated a medium-duty truck-sized turbocharged hydro-
the scatter is equivalent to that observed in Fig. 4 for gen engine using the one-dimensional modeling tool,
naturally-aspirated H2 ICEs. WAVE [47]. Operating limits are defined by setting
Multi-mode operating strategies are often adopted constraints on knock, NOx production, and maximum
for boosted H2 ICEs. Berckmüller et al. [29] recom- cylinder pressure and then performance is optimized at
mended the following strategy for increasing loads: (1) 1900 rpm for maximum BMEP. Details of the model are
unthrottled lean operation at low enough loads to avoid provided in Appendix A.
NOx production, then (2) unthrottled stoichiometric op- A simulated cycle representing optimized results
eration with EGR dilution to the naturally aspirated from the WAVE model is displayed on log(P)–log(T)
full-load condition, and finally (3) supercharged stoi- axes in Fig. 6. This graphical format has been proposed
chiometric operation with EGR dilution. The lean oper- as an effective aid for visualizing engine operation
ating mode is clean enough to require no aftertreatment, boundaries [48]. Several important constraints that
while the = 1.0 modes enable the use of low-cost limit operating conditions are displayed on the graph:
TWC to deal with the associated higher NOx produc- (1) a peak charge-pressure line, (2) a knock-limit
tion. Natkin et al. [44] used the expedient, but less ef- line, and (3) a peak charge-temperature line for NOx
ficient approach of throttling their supercharged H2 ICE control. The peak pressure line (150 bar) is simply a
at low loads while holding fixed at 0.5. For high loads, reasonable mechanical limit for a medium-duty truck
they too proposed an EGR strategy enabling = 1.0 engine.
operation. The knock limit is computed using the chemical
kinetics code Senkin [49] and the hydrogen reac-
3.1.2. Predictions of turbocharged H2 ICE performance tion mechanism of Miller and Bowman [50]. Given
The turbocharged H2 ICE is an attractive advanced charge conditions at intake valve closing, the code
H2 ICE option in the near-term because of its relative establishes were on the adiabatic compression line1300 C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305 the hydrogen–air mixture autoignites. To avoid knock, 3.2. Liquid-hydrogen-fueled internal combustion fuel-air packets must not be compressed to temper- engine (L-H2 ICE) ature pressure conditions falling to the right of this line. The use of liquid hydrogen as an automotive fuel has The final limit, the peak charge-temperature line, is been recently reviewed by Peshka [54]. The L-H2 ICE selected to control NOx emissions. The use of a temper- label means that hydrogen is stored as a liquid, but not ature limit for this purpose is based on the demonstrated necessarily injected as a liquid. While the primary ben- strong correlation between peak cylinder gas temper- efit of the L-H2 ICE is the higher stored-energy density atures and NOx production [51]. To achieve SULEV- of hydrogen available with liquefaction, it is not the sole level emissions, a limit of 1800 K has been selected as benefit. The charge-cooling effect of the cold hydrogen estimated from published correlations [3,52,53]. provides for several advantages compared to conven- In Fig. 6, the solid curve represents 1900 rpm mean tional gaseous port fuel injection. cycle data for operation at the highest load consis- Intake-charge cooling improves volumetric effi- tent with the above limits. At these operating con- ciency, minimizes preignition and lowers NOx emis- ditions, inlet pressure is boosted to 2.8 bar absolute sions. The increase in volumetric efficiency and, sub- and peak cylinder pressure reaches 125 bar, satisfy- sequently, power density is a simple consequence of ing the pressure limit of 150 bar. The mean cylinder intake mixture density varying inversely with tem- data do not reveal peak temperatures since in-cylinder perature. Furuhama et al. [31] calculated that with a temperatures are far from uniform; to visualize peak hydrogen temperature of 120 K, the peak power output temperatures, unburnt and burnt-gas curves (dashed) of an L-H2 ICE can equal that of the identical engine are also plotted in Fig. 6. At these operating condi- fueled with gasoline. Similarly, Wallner et al. [55] tions, the unburnt-gas curve remains well below the estimated that with intake charge cooling to 210 K, knock-limit line, indicating that ignition of the last- the power density of a L-H2 ICE will be 15% higher to-burn packet should occur by flame propagation and compared to fueling with PFI gasoline. Furthermore, not by autoignition. The burnt gas curve provides an the lower charge temperature will also mitigate preig- estimate of the hottest in-cylinder temperatures, ap- nition events (similar to water injection) and thereby propriate for estimates of NOx production. In the case increase the preignition-limited maximum equivalence of Fig. 6, the burnt-gas curve just touches the peak- ratio. For example, stoichiometric operation in the charge-temperature line, establishing the maximum absence of preignition events has been reported by load consistent with SULEV NOx emissions as esti- Furuhama et al. [31] and Knorr et al. [32]. The net mated here. Predicted performance specifications of effect is that not only are high power densities possi- the turbocharged H2 ICE engine are listed in Fig. 6. At ble with L-H2 ICE they are also attainable in practice. the key operating speed of 1900 rpm and NOx -limited An improvement in BTE and NOx emissions are also equivalence ratio of 0.33, the engine achieves a brake realized with charge cooling: Furuhama et al. [31] re- power of 105 kW, BMEP of 11.3 bar, brake torque of port that, relative to gaseous hydrogen fueling, BTE 525 Nm, and brake efficiency of 39%. Given that the increases and specific NOx emissions decrease with model overestimates performance by not considering decreasing injection temperatures. The latter effect incomplete combustion (see Appendix A), the estimated was described as being primarily due to the fact that BTE of 39% agrees well with Ford’s H2 ICE measured with liquid hydrogen fueling a leaner mixture could efficiency of 37% for unthrottled operation [44]. Al- be used to produce the same power that would other- though the brake power levels are substantially lower wise require a richer mixture with gaseous hydrogen than delivered by current turbocharged Diesel technol- fueling. ogy, they are sufficient for medium-duty applications such as urban buses and delivery trucks. For example, 3.2.1. L-H2 ICE hardware the engine geometry used in the simulation matches For practical application a L-H2 ICE fueling system Cummins Westport’s BG195 naturally-aspirated natu- typically requires a vacuum-jacketed fuel line, heat ex- ral gas engine, and the latter operates with just 10% changer and cryogenic pumps, and injectors. A detailed more power and torque at 1900 rpm. The simulation description of a L-H2 ICE fueling system is given by results suggest that a boosted lean-burn H2 ICE engine Peschka [56]. The practical difficulties of liquid stor- can deliver acceptable performance while achieving age include the energy penalty of liquefaction, evapora- engine-out emission levels low enough to eliminate tion during long-term storage, and the cost of on-board aftertreatment. cryogenic dewars.
C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305 1301
Several experimental and prototype L-H2 ICE pow- 10000
ered vehicles have been demonstrated over approxi- 9000
mately the last twenty-five years: Students at Musashi 8000
University converted a gasoline-powered automobile
7000
to operate on liquid hydrogen [31].3 The automo-
NOx, ppm
bile was entered in, and completed, a student road 6000
rally in California totaling roughly 1800 miles. The 5000
reported fuel economy was the gasoline equivalent of 4000
36.3 miles/gal. MAN over a two year test period
3000
(1996–1998) operated a dual-fueled liquid hydrogen
and gasoline bus in regular public transport service [32]. 2000
The bus was powered by a 6-cylinder 12-l engine with 1000
a CR of 8:1. The liquid hydrogen storage capacity was 0
120 100 80 60 40 20 0
570 l, of which 6% of this total was reported as lost to
SOI, oBTDC
evaporation per day. BMW has developed a small fleet
of dual fueled liquid hydrogen and gasoline-powered Fig. 7. NOx concentration verses start of injection (SOI): ◦ [59] at
vehicles under the model name 750hL. The engine = 0.5; (, ♦, , , , +) from [3] at = 0.35, 0.43, 0.51, 0.60,
has 12 cylinders and a displacement of 5.4 l. Liquid 0.69 and 0.79, respectively; (×, ∗) from [30] at = 0.35 and 0.52,
respectively. Dashed lines through the data are for visual aid.
hydrogen storage capacity is 140 l providing a range
of 400 km. Acceleration is respectable at 0–60 mph in
9.6 s. 1000–5000 rpm, respectively. In practice, to avoid
preignition, start of injection (SOI) is retarded with
3.3. Direct-injection hydrogen-fueled internal respect to IVC, and mixing times are further reduced.
combustion engine (DI-H2 ICE) To evaluate the plausibility of complete mixing in a
DI-H2 ICE, Homan [58], using experimental correla-
The direct injection H2 ICE has long been viewed tions for air-entrainment rates in free turbulent jets
as one of the most attractive advanced H2 ICE options and order of magnitude expressions for turbulent mix-
(for a summary, see [58]). The view is based on the high ing times, estimated that a free-hydrogen jet with
volumetric efficiency, (since hydrogen is injected after sonic velocity at the orifice issuing into air will en-
intake valve closing) and the potential to avoid preigni- train a stoichiometric amount of air in approximately
tion. The latter effect is controlled by timing injection 1 ms. However, contrary to this optimistic estimate,
to both minimize the residence time that a combustible the overwhelming experimental evidence [30,58–61]
mixture is exposed to in-cylinder hot-spots (i.e., late in- demonstrates that complete mixing in an engine takes
jection) and allow for improved mixing of the intake air approximately 10 ms. Homan et al. [30] conjectured
with the residual gases. The improved volumetric effi- that the order-of-magnitude difference in mixing times
ciency (equal to PFI gasoline or higher) and the higher between the estimate for a free jet and that measured in
heat of combustion of hydrogen compared to gasoline an engine is a result of fluid flow interaction between
provides the potential for DI-H2 ICE power density to in-cylinder flow and the hydrogen jet. A strong inter-
be approximately 115% that of the identical engine op- action will cause the free-jet analysis to break down.
erated on gasoline. This estimate is consistent with that As evidence, Homan et al. [30] demonstrated, in an
measured by Eichlseder et al. [3]. In particular, they engine with side injection and side spark (separated by
measured a 15% increase in IMEP for engine operation 180◦ ), that the frequency of misfire for late injection
with DI hydrogen compared to engine operation with decreased significantly when the injected hydrogen was
PFI gasoline. in coflow with the swirl instead of in crossflow.
The challenge with DI-H2 ICE operation is that Given the high probability of incomplete mixing with
in-cylinder injection requires hydrogen–air mixing in late injection, much effort has been devoted to under-
a very short time. For early injection (i.e., coincident standing the effect of injection timing on DI-H2 ICE
with IVC) maximum available mixing times range properties. The effect of SOI on NOx emissions has
from approximately 20–4 ms across the speed range been investigated by Homan et al. [30], Glasson and
Green [59] and Eichlseder et al. [3]. The data of NOx
3 See [57] for a summary of Musashi University’s twenty years versus SOI for these studies are shown in Fig. 7, where
of L-H2 ICE research. it is observed that the effect of SOI on NOx emissions1302 C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305
is not simple: NOx emissions increase with retard of charge stratification such that the local mixture fraction
SOI in several data sets, and decrease in others. These in the near-wall region is expected to be lean compared
conflicting trends can be explained by separating the to the bulk mixture, cooling losses can be decreased
data sets of [59] and [3] into two groups: (i) data sets significantly and thermal efficiency increased.
with a global equivalence ratio at or below the NOx - In order to take full advantage of the potential of the
limited equivalence ratio ( 0.5, open symbols), and DI-H2 ICE, a high-pressure, high flow-rate hydrogen in-
(ii) data sets with a global equivalence ratio above the jector is required for operation at high engine speeds and
limit (, , +). For all data sets, retarding SOI is as- to overcome the in-cylinder pressure for injection late
sumed to increase mixture inhomogeneity. For the data in the compression stroke. Here we define high pressure
that fall below the NOx -limited equivalence ratio, the as greater than 80 bar to ensure sonic injection veloci-
increase in mixture inhomogeneity leads to locally rich ties and high enough mass flow rates for SOI throughout
regions with high combustion temperatures and subse- the compression stroke. The need for rapid mixing ne-
quent higher NOx production. For the data sets above cessitates the use of critical flow injectors and the short
the limit, with increasing mixture inhomogeneity, rich time duration with late injection requires high mass flow
regions with local > 1 are produced with a local reduc- rates. The development of high pressure injectors has
tion in NOx formation. Subsequently, the overall heat been reported by Green and Glasson [64] and Jorach
release rate is decreased (in the locally rich regions most et al. [60]. The development of hydrogen injectors for
of the excess H2 will eventually burn but in the slower injection pressures lower than 80 bar have been reported
mixing-controlled mode). Then, the overall reduction in by Homan et al. [30] and Varde and Frame [65]. In ad-
NOx emissions is thought due to the coupled effect of a dition, a DI-L-H2 ICE injector has been developed by
slower heat release rate and in-cylinder heat losses that Furuhama et al. [66]. To prevent valve leakage at the
leads to lower peak in-cylinder temperatures and slower valve seat a seal made of an elastomer material has been
NOx formation rates. used with success [30,64]. Typical flow rates required
The data of [30] are unique in that a modified diesel are 1–10 g s−1 .
head was used to locate the spark plug close to the in- Multi-mode operating strategies have also been
jector. This arrangement allowed DI-H2 ICE operation proposed for DI-H2 ICEs [62,67], similar to pressure
with late injection and near-simultaneous spark. Homan boosted H2 ICEs. Here, a dual-injector strategy would
et al. termed this operating scheme LIRIAM (late injec- be used to take advantage of the high thermal effi-
tion, rapid ignition and mixing). The observed increase, ciencies at low- to medium-loads with PFI fueling and
then decrease in NOx production with SOI retard is also the high peak power with DI fueling. Rottengruber
explained as an effect of charge stratification. However, et al. [67] propose the following strategy: (1) unthrot-
due to the engine geometry and late-injection strategy, tled lean operation with external mixture at loads below
it is not surprising that the data follows a different trend the NOx formation limit, then (2) throttled stoichio-
with . metric operation with external mixture, and finally (3)
Similar to the NOx emissions described above, the stoichiometric internal mixture formation. Lee et al.
effect of SOI on thermal efficiency is not straight- [62] report, for a dual injector H2 ICE, improvements
forward. Eichlseder et al. [3] found that at low loads in thermal efficiency of approximately 15–30% at low-
(or similarly, low equivalence ratio), indicated ther- to medium-loads with PFI fueling compared to DI
mal efficiency (ITE) increases with retard of SOI. The fueling, and an increase in peak power output of ap-
increase was shown to be due to a decrease in the com- proximately 60–70% with DI fueling compared to PFI
pression work caused by differences in mixture gas fueling. The disadvantage of the dual-injector strategy
properties and charge mass with retarded SOI. At high is the added cost.
loads, the same authors found that ITE first increases
and then decreases with retard of SOI. The reversing 3.4. H2 ICE-electric hybrid
trend is assumed to be a consequence of an unfavor-
able mixture formation. However, Lee et al. [62] report A hybrid-electric version of an H2 ICE offers the
results contradictory to [3], where they find that, for potential for improved efficiencies and reduced emis-
both low and high load, thermal efficiency decreases sions without the need for aftertreatment. In a hybrid
monotonically with retard of SOI. These contradic- electric vehicle (HEV), the ICE operates either in
tory findings may be a result of differences in mixture series or parallel with an electric motor. There are
formation. The view is based on the recent results of advantages and disadvantages for both configurations.
Shudo et al. [63], in which it was shown that with Recent hydrocarbon-fueled HEVs have incorporated aC.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305 1303
configuration that combines features of both series and economically viable in the near-term. Consequently,
parallel.4 The series hybrid configuration for H2 ICEs there is little doubt that the hydrogen-fueled internal
has been investigated by Van Blarigan and Keller [68] combustion engine can serve as a near-term option for a
and Fiene et al. [28]. In this configuration, the H2 ICE transportation power plant in a hydrogen economy. This
is used to drive an alternator that generates electricity. is best illustrated in the fact that at least one H2 ICE-
The electricity is used either to charge the batteries powered vehicle, the BMW 745 h, is expected to reach
or power the electric motor that powers the vehicle the market in the next few years.
drivetrain. The advantage is the ability to operate, However, the long-term future of the H2 ICE is less
and optimize, the H2 ICE for single-speed operation at certain and hard to predict, as is the future of the hy-
maximum power. For this purpose, Van Blarigan and drogen economy itself. The uncertainty is in part due
Keller [68] have extensively studied the optimization to the multiple H2 ICE options available, as described
of H2 ICE thermal efficiencies and NOx emissions for in Section 3. There are good prospects for increased ef-
single-speed operation. Peak ITEs of 44–47% were re- ficiencies, high power density, and reduced emissions
ported [68]. To assess the potential for H2 ICE vehicles with hybridization, multi-mode operating strategies, and
to meet EZEV (i.e., SULEV, see footnote 2) emission advancements in ICE design and materials. The com-
standards, Aceves and Smith [69] modeled a non-hybrid mercial viability of these advanced H2 ICE options re-
H2 ICE vehicle, and parallel and series H2 ICE HEVs. quires: continued advancements in fundamental H2 ICE
They reported that for the input parameters selected, all research, reduction of NOx formation and emissions,
three vehicles can satisfy EZEV emissions standards. research and development of advanced engine compo-
An attractive feature of the HEV is that the peak power nents, and highly advanced control and optimization
output of the ICE can be significantly lower than that strategies. Provided that these efforts produce a H2 ICE
required for a non-hybrid ICE without any sacrifice option that is highly efficient, with near-zero emissions,
in vehicle performance. Then, with respect to H2 ICE and a drivability that surpasses present day gasoline-
HEVs, the obvious operating strategy would be to op- fueled ICEs, then competition will dictate the transition
erate the H2 ICE lean enough such that engine-out NOx to the PEMFC.
emissions are near-zero (see Fig. 4). It follows that
since lean power density can be improved with intake
pressure boosting, a boosted H2 ICE HEV would be an Acknowledgments
attractive option, as demonstrated by the Ford H2 RV
[45]. The H2 RV is an HEV that uses a supercharged The technical guidance of Pat Flynn of Cummins
4-cylinder 2.3-l H2 ICE. The boosted H2 ICE has a peak Engine Co. (retired) regarding H2 ICE cycle design is
power of 110 HP at 4500 rpm and the electric motor gratefully acknowledged. Support for this research
provides an additional 33 HP, and is used primarily was provided by the U.S. Department of Energy, Of-
for power assist. Acceleration is0–60 mph in 11 s. Fuel fice of FreedomCAR and Vehicle Technologies. The
economy is 45 miles per kg of H2 (gasoline equivalent research was performed at the Combustion Research
of 45 mpg) and driving range is 125 miles. Facility, Sandia National Laboratories, Livermore,
The potential for improved efficiencies and reduced California. Sandia is a multiprogram laboratory
emissions with H2 ICE hybridization appears quite operated by Sandia Corporation, a Lockheed Martin
promising. In particular, in comparing a H2 ICE HEV Company, for the United States Department of
to a PEMFC vehicle, Keller and Lutz [70] found that Energy’s National Nuclear Security Administration
the H2 ICE HEV compares favorably. under contract DE-AC04-94AL85000.
4. Concluding remarks Appendix A. Wave H2 ICE engine model details
The last decade has produced significant advance- The 1-dimensional WAVE model was used to sim-
ments in the development of the hydrogen-fueled in- ulate a medium-duty engine with turbocharger and in-
ternal combustion engine and H2 ICE-powered vehicles. tercooler. The 6-cylinder engine displaces 5.8 l, and its
Undoubtedly aided by the technological advancements geometric compression ratio is set at 13 as a compro-
of the ICE, simple H2 ICE options are convenient and mise between the opposing trends of efficiency and pre-
ignition. The 4-valve geometry and lift profiles are the
4 For example, see Toyota Hybrid System (THS). same as a Cummins 6B engine, with exhaust valves1304 C.M. White et al. / International Journal of Hydrogen Energy 31 (2006) 1292 – 1305
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