Article pubs.acs.org/EF
Fuel-Substitution Method for Investigating the Kinetics of LowVolatility Fuels under Enginelike Operating Conditions Daniel Janecek,* David Rothamer, and Jaal Ghandhi Engine Research Center, University of Wisconsin-Madison, 1500 Engineering Drive, Madison, Wisconsin 53706, United States ABSTRACT: The combustion properties of two low-volatility diesel fuels were characterized under engine-relevant conditions using a new fuel-substitution strategy that minimizes biases associated with unknown boundary conditions. The engine is operated in a homogeneous charge compression ignition mode with port fuel injection of the primary reference fuels, n-heptane and isooctane, and an upstream prevaporizer for the diesel-like heavy test fuels. The engine is first operated on the primary reference fuels (PRFs). The intake conditions are then fixed, and the heavy fuel is introduced in increasing amounts while the mass of port-injected fuel is reduced to maintain engine load and the n-heptane-to-isooctane ratio is adjusted to maintain constant combustion phasing. It is shown that this provides nearly constant in-cylinder thermodynamic and boundary conditions. The baseline condition, which uses well-characterized fuels, can be used to adjust the trapped-gas initial conditions and the heat transfer rates of a computational model, but these parameters do not change with the subsequent addition of heavy test fuel. The test fuel combustion characteristics are described in terms of the PRF number of the port-injected mixture. A simple linear blending method was found to adequately represent the heavy test fuel in terms of an effective PRF number. Testing under different operating conditions showed no significant change in the effective PRF number of the test fuels. The fuels investigated in this study were F-76, a high-sulfur-content diesel fuel, and HRD, a hydroprocessed renewable diesel fuel composed primarily of alkanes. Effective PRF numbers of 35 and −25 were found for F-76 and HRD, respectively.
1. INTRODUCTION By eliminating the complications of a fuel spray, homogeneous charge compression ignition (HCCI) can be a useful combustion strategy to investigate the chemical kinetic characteristics of fuels. Injecting the fuel upstream of the engine intake port allows a homogeneous air/fuel mixture to enter the combustion chamber. However, homogeneous fueling becomes difficult for low-volatility fuels that do not readily vaporize at standard engine intake temperatures. Three different strategies have been employed to overcome this limitation. First, there have been numerous attempts to use an early direct-injection method for mixture preparation.1,2 With this strategy the hope is that the elevated temperature due to partial compression will aid the vaporization. However, direct injection gives rise to fuel stratification, i.e., the mixture is less homogeneous, and depending on the in-cylinder conditions at injection time, significant wall wetting can occur. Puschmann et al.3 employed what they called a cool flame vaporizer upstream of the intake port. The low-temperature heat release of the fuel was used to promote fuel vaporization without requiring an external heat source. However, this method involves partial reaction of the fuel prior to its induction into the combustion chamber, which affects the kinetic characteristics of the fuel. A third strategy is to use a diesel prevaporizer to heat the fuel upstream of the intake port using air temperatures in the range of 130 to 200 °C to completely vaporize the fuel. The intake air temperature and residence time must be correctly selected to ensure that a homogeneous mixture of fuel vapor and air enters the combustion chamber. Numerous examples of this prevaporizer system can be found in the literature.4−6 Even with the assumption that the fuel/air mixture is completely homogeneous, there are still a number of unknown parameters that make HCCI experiments difficult to model © 2016 American Chemical Society
computationally. Since there is no spark or direct fuel injection, the start of the combustion process is completely dependent on chemical kinetics, which makes the process highly dependent on boundary and initial conditions such as the wall temperatures and the bulk-gas temperature at intake valve closure. Experiments to quantify and remove the effects of in-cylinder conditions, such as wall and piston temperatures and exhaust residuals, on the charge temperature have been undertaken by Sjöberg and Dec,7,8 who used direct fuel injection and a skipfiring strategy wherein the engine was operated for a large number of cycles under the baseline condition followed by just a few cycles at the condition of interest. This alternate-firing method (they used 19−1 and 18−2 sequences) was effective at maintaining consistent wall temperatures and exhaust residuals. Sjöberg and Dec’s method, however, relies on direct injection of the fuel, which was sufficient for the relatively high volatility fuels in which they were interested, but the alternate-firing method is not possible when heavy fuels are used because of incylinder mixing and vaporization limitations. Boundary and initial conditions are critical to any model that will utilize engine data to evaluate chemical kinetic mechanisms. The wall heat transfer rate, in particular, is difficult to model accurately and is equally difficult to measure experimentally. As a result, the wall heat transfer rate and initial conditions are often used as “tuning” parameters to match simulations to experimental data. Ad hoc tuning, however, makes it impossible to draw conclusions about the fuel chemistry. Received: October 28, 2015 Revised: January 8, 2016 Published: January 11, 2016 1400
DOI: 10.1021/acs.energyfuels.5b02542 Energy Fuels 2016, 30, 1400−1406
Article
Energy & Fuels The objective of this study was to develop a methodology for performing HCCI engine combustion experiments so that they provide useful data for chemical kinetics evaluations under engine-relevant conditions with minimal bias from boundary and initial condition variations when using low-volatility fuels. A baseline condition that uses well-characterized gasoline primary reference fuels (PRFs) is provided to allow global tuning of the computational model, but after this no adjustment of the initial or boundary conditions is needed and should not be performed. The fuel of interest is then substituted for the PRFs in increments, which elucidates the chemical behavior of the fuel in reference to that of the PRFs. The fuels investigated in this study were F-76, a high-sulfurcontent diesel fuel, and HRD, a hydroprocessed renewable diesel fuel composed primarily of alkanes. A summary of the fuels’ measured physical and chemical properties is shown in Table 1, and the fuels’ distillation curves are displayed in Figure
Table 2. Engine Geometry compression ratio displacement (L) stroke (mm) bore (mm) con. rod length (mm) intake valve closing exhaust valve opening swirl ratio piston bowl type
suite. NI hardware was also used for all high- and low-speed controls such as intake air pressures and temperatures, coolant temperatures, and the exhaust gas recirculation (EGR) valve. The in-cylinder pressure was measured using a Kistler 6125A pressure transducer in series with a Kistler 5010 charge amplifier. A 0.25° resolution encoder fixed to the crankshaft was used as the timing clock for the high-speed data acquisition, giving a resolution of about 22 μs at the tested engine speed. Pressure data were acquired for 300 cycles for each test condition. A five-gas Horiba emissions measurement system was used to perform relevant gaseous emissions measurements. The mean piston temperature and heat flux were characterized using a fast-response surface-mounted J-type thermocouple made by Medtherm mounted in the squish region, as indicated by the star in Figure 2. Crank-angle-resolved temperature data were transmitted to a
Table 1. Fuel Property Information method D1319
D5291 CH D2622_07 D4052M D4809 gross D4809 net D613
description
units
F-76
HRD
aromatic olefins saturate carbon hydrogen sulfur specific gravity heating value heating value cetane number
% % % wt % wt % ppm
27 2.3 70.7 86.4 13.32 3292.4 0.8222 45.576 42.749 52
0.6 0.9 98.5 85.28 15.17 74.7
MJ/kg MJ/kg
9.5 0.477 90.4 82 145.54 −132° aTDC 112° aTDC 1.5 stock (re-entrant)
1. Because both fuels had relatively low volatilities, an upstream prevaporizer system was implemented, which will be described below.
Figure 2. Piston thermocouple layout. receiver and recorded via a wireless telemetry system made by IRTelemetrics. Figure 3 shows an overview of the thermocouple power and transmission system. More information about the piston temperature measurement technique can be found in ref 9. 2.2. Homogeneous Prevaporized Fuel System. To ensure complete vaporization of low-volatility fuels, an air-assisted fuel injector was used to inject a highly atomized fuel spray into the main intake air stream directly downstream of the in-line intake air heater. Tape heaters were utilized to maintain the downstream wall temperatures at the gas temperature up to the engine intake port to prevent condensation. At the engine operating speed of 1900 rpm, which was used in all of the tests, the fuel had a residence time of approximately 3.5 s. Viewing ports were installed in the system to allow visual verification that the fuel was completely vaporized. Initial testing of the system with the F-76 fuel was performed to define the temperature limits required to fully vaporize the fuels for the range of desired engine loads. A laser was passed through the viewing ports to visually detect whether fuel droplets were present. For each tested fuel/air ratio, the intake temperature was increased until full fuel vaporization was achieved. Figure 4 shows the required temperature to achieve full fuel vaporization as a function of the fuel
Figure 1. Distillation curves for HRD, F-76, and PRF21.
2. EXPERIMENTAL SETUP 2.1. Lab Overview. The engine experiments were carried out in a single-cylinder version of the General Motors/Fiat JTD 1.9 L fourcylinder diesel engine. Relevant engine parameters can be seen in Table 2. The stock four-cylinder engine head was mounted on a Ricardo Hydra single-cylinder block. The air flow rate was controlled using choked flow orifices, and the intake air temperature was controlled using heaters both upstream of and surrounding the outer walls of the intake surge tank. Control of the fuel pressure and injection timings was achieved using the National Instruments (NI)−Drivven hardware/software 1401
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3. RESULTS AND DISCUSSION 3.1. Overview. Two types of engine testing were performed. Intake temperature sweeps were performed as the heavy fuel was substituted for the PRFs (with constant fuel energy) to investigate the temperature-dependent kinetic behavior of the heavy fuels and to compare this to the kinetic behavior of the well-characterized PRFs. However, these tests uncovered issues associated with the engine boundary conditions that confound the results and motivated the development of a new testing method. In the new testing method, the heavy fuel was substituted for the PRFs while holding a constant combustion phasing (maintained by varying the ratio of the PRFs delivered to the engine). This fixes the engine boundary conditions throughout the sweep, providing results that are not confounded with boundary condition variation. 3.2. HCCI Intake Temperature Sweeps. An initial study was undertaken to determine the effect of temperature on the F-76 and HRD kinetics; intake temperature sweeps were performed at constant engine speed, intake pressure, and fuel energy. These results provide the basis for the new methodology that will be described subsequently. A stable operating point was found using port injection of a combination of nheptane and isooctane, and then the intake temperature was swept in 3 °C increments. It is important to note that the high intake temperatures were required in order to ensure vaporization of the heavy fuel, as discussed above. The intake temperature sweep was then repeated with progressively larger fractions of the heavy fuels, while the PRFs were decreased proportionally in order to maintain a constant injected total fuel energy. Table 3 details the conditions investigated. Tests were run at intake pressures of 1.2 and 2 bar to investigate whether any pressure-dependent kinetics trends existed. Examples of the pressure and apparent heat release rate (AHRR) data obtained for an intake temperature sweep with HRD substituted for 10% of the original PRF mixture are shown in Figure 5. As expected, as the intake temperature increases, the fuel reaction rate increases, causing advancement of both the peak pressure and the location of 50% heat release (CA50) and an increase in the peak pressure. It is interesting to note that the temperature increase can be seen to advance both the low- and high-temperature heat release. While the magnitude of the peak high-temperature heat release rate strongly increases as combustion phasing advances, the magnitude of the peak low-temperature heat release rate is seen to remain approximately constant, indicating that for HRD the temperature affects the phasing but not the quantity or duration of the low-temperature heat release. The effect of substituting F-76 in place of isooctane at an intake temperature of 165 °C and an intake pressure of 1.2 bar can be seen in Figure 6. As the fraction of F-76 increases, the
Figure 3. Thermocouple telemetry system.
Figure 4. Heavy fuel vaporization limits.
partial pressure (concentration). The observed vaporization limits appear to be well-predicted by the Antoine equation for molecules between C16 (hexadecane) and C18 (octadecane).10 On the basis of these data, the intake operating temperatures were chosen to ensure full fuel vaporization. The actual range of operating conditions tested is also shown in Figure 4. Because of the relatively high intake temperatures combined with the long residence time, tests were performed to verify that no fuel decomposition reactions occurred before the fuel entered the combustion chamber. n-Heptane, which has ignition characteristics similar to those of the diesel fuels tested, was injected both through the upstream prevaporizer system and directly at the intake port. Pressure and heat release data were analyzed, and no measurable effect was found for n-heptane, verifying that the residence time of 3.5 s at temperatures up to 200 °C likely had little to no effect on the combustion characteristics. In addition, a review of the literature regarding autoignition of diesel-like fuels showed negligible reactivity at temperatures below 300 °C.11−13
Table 3. Operating Conditions for Intake Temperature Sweeps (Φ = Equivalence Ratio) fuel composition (mass %) intake pressure (bar)
gross IMEP (bar)
Φ
HRD
F76
n-heptane
isooctane
intake temperature (°C)
1.2
3.5
0.3
7.3
0.35
0 0 10→45 0 0 10→20
55 45→0 55 20 10→0 10→0
45 45 35→0 80 80 80
165−189
2
0 10→55 0 0 10→20 0 1402
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Figure 5. Pressures and AHRRs for an intake temperature sweep for a 10 mass % HRD mixture.
Figure 8. CA50 vs intake temperature for all of the fuel mixtures at an intake pressure of 2 bar.
the intake temperature increases, as expected. In both Figure 7 and Figure 8 the blue line represents the baseline PRF case, the solid lines are the HRD-substitution cases, and the dashed lines are the F-76-substitution cases (it is important to note that for the high-intake-pressure cases both HRD and F-76 were substituted in place of n-heptane, whereas for the low-intakepressure cases HRD was substituted in place of n-heptane but F-76 was substituted in place of isooctane in order to maintain stable combustion phasing; see Table 3). From the high-intakepressure data, it is clear that HRD is more reactive than nheptane and causes an advance in CA50 as an increasing fraction is added, while F-76 is less reactive than n-heptane and causes a retardation in CA50 as the F-76 fraction increases. At the low intake pressure it can be seen that HRD and F-76 are more reactive than n-heptane and isooctane, respectively. The intake temperature sweep data show the kinetic behavior of the test fuels in relation to the well-studied fuels n-heptane and isooctane, but the test results also show some other effects. The intake temperature sweep data show an approximately 0.15 crank angle degree advance in CA50 per 1 °C increase in intake temperature, except for HRD at the low intake pressure, which shows a 3-fold higher sensitivity of approximately 0.45 crank angle degree advance in CA50 per 1 °C increase in intake temperature. These results show a substantially smaller effect of intake temperature on CA50 than has been seen in other studies.14 This suggests that the magnitude of the change in the intake temperature in the current study is not necessarily representative of the magnitude of the change in the bulk gas temperature at intake valve closing, most likely as a result of heat transfer losses in the intake runner and during the intake stroke. The mean piston temperatures for all of the intake temperature sweeps performed at the lower intake pressure are shown in Figure 9. It can be seen that there is a strong correlation between increased mean piston temperature and advanced CA50. This indicates that the wall and piston boundary conditions change as the combustion phasing varies. Thus, the trends observed in Figure 5 cannot be solely attributed to changes in the intake temperature, and likewise, the trends in Figure 6 are not solely attributable to the changing F76 fraction. 3.3. HCCI Constant Phasing Fuel Blend Studies. The results of the intake temperature sweep experiments showed that, at a minimum, the combustion phasing and load must be kept constant in order to keep the boundary conditions, such as the wall and piston temperatures, constant. However, achieving
Figure 6. Pressures and AHRRs for the F-76 substitution sweep at 165 °C and an intake pressure of 1.2 bar.
combustion advances substantially; this trend was expected since F-76 was substituted in place of the less reactive isooctane. It is interesting to note that similar to the intake temperature sweeps, the magnitude of the high-temperature peak heat release rate increases as combustion phasing advances; however, unlike the intake temperature sweeps, the magnitude of the low-temperature release rate also increases as combustion phasing advances, i.e., as the F-76 fraction increases. A summary of combustion phasing for all of the fuelsubstitution intake temperature sweeps can be seen in Figure 7 for the 1.2 bar intake pressure data and in Figure 8 for the 2 bar intake pressure data. For all of the test cases, CA50 advances as
Figure 7. CA50 vs intake temperature for all of the fuel mixtures at an intake pressure of 1.2 bar. 1403
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that nearly identical in-cylinder thermal conditions and combustion performance were achieved for the entirety of the sweep. Thus, to the best of our ability, we have been able to independently change the mass fraction of the low-volatility test fuel without affecting the initial or boundary conditions. Figure 11 shows the PRF number of the port-injected PRF mixture (the PRF number is calculated on the basis of the mass
Figure 9. Mean piston temperature vs CA50 for all of the fuel mixtures at an intake pressure of 1.2 bar.
this by varying the intake temperature leaves another significant unknown: the intake valve closing (IVC) temperature. As can be seen in Figures 7 and 8, the combustion phasing change with intake temperature shows significantly less sensitivity than expected, suggesting that the trapped-gas temperature is not directly related to the intake port temperature through a simple relationship. A new fuel-substitution method was devised to overcome this problem. An operating point is first established using only the portinjected PRFs. The test fuel is then introduced through the upstream prevaporizer system at the desired mass fraction; the port fuel mass is reduced to maintain the engine load, and the isooctane-to-n-heptane ratio is adjusted to maintain the same combustion phasing. Data points are taken at different test fuel mass fractions, and the isooctane-to-n-heptane ratio is adjusted for each test fuel mass fraction to give the same combustion phasing. Utilizing separate fuel injectors for the isooctane and n-heptane fuel streams allows this ratio to easily be adjusted in real time. The main benefit of using the PRF mixture rather than the intake temperature to control the CA50 is that load and phasing can be matched without changing any of the inlet conditions, so that the boundary conditions are minimally affected too. Figure 10 shows the cylinder pressure and heat release for one of the F-76 substitution sweeps, where yHF is the heavy-fuel mass fraction. For the data in Figure 10, the indicated mean effective pressure (IMEP) varied by less than 1%, the CA50 varied from the mean by less than 0.5°, and the IVC temperature determined using the trapped mass and the ideal gas law varied about the mean by less than 3 °C. It can be seen
Figure 11. Measured PRF number vs F-76 mass fraction.
fraction of isooctane) required to keep the phasing and load constant as a function of the heavy fuel fraction (yHF) for F-76. Since the IVC temperature, load, and combustion phasing are kept constant, the total fuel mixture reactivity, described by the overall effective PRF number, is constant throughout the sweep and is defined as being equal to the PRF number of the baseline PRF mixture when no heavy fuel is injected upstream, e.g., 52 for the case shown in Figure 11. When heavy fuel is injected, the total fuel mixture reactivity is a function of the individual reactivities of the heavy fuel and the PRF mixture. A review of the literature yielded no universally accepted blending rule for the reactivity of fuel blends, although the PRF blends are defined on a liquid volume basis when used for engine knock testing.15 Because fuel mass flow rates are directly measured during testing using Coriolis meters and there is an uncertainty in the liquid volumes as the ambient temperature fluctuates, the fuel blends herein are defined on a mass basis. For PRFs, calculating fuel blends on a mass basis rather than a liquid volume basis results in a maximum difference of 0.22% in the calculated fuel blends because of the similarity in the densities of isooctane and n-heptane. Another alternative, calculating fuel blends on an energy basis, results in a maximum difference of 0.31% in the calculated fuel blends because of the similarity in the lower heating values of isooctane and nheptane. If the resulting mixture reactivity obeys a mass-based linear blending rule, the required PRF versus heavy fuel mass fraction data can be used to calculate an effective PRF number for the heavy fuel, PRFHF. The effective overall PRF number of the three-component fuel mixture (n-heptane/isooctane/heavy fuel), PRFmix, which is constant as discussed above, can be written as PRFmix = (1 − yHF )PRFmeas + yHF PRFHF
(1)
where PRFmeas is the PRF number of the port-injected mixture. This equation has two unknowns: PRFmix and PRFHF, which can be determined using a least-squares approach. Figure 12 shows an example of a curve fit for a sweep using F-76. The graph shows the individual data points, the curve fit to the data,
Figure 10. Pressures for an F-76 substitution sweep. 1404
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In general, for fuels containing components other than alkanes, such as olefins and aromatics, one would expect the effective PRF number of the fuel to change with the engine operating conditions because of the differing sensitivities of these components to temperature and pressure relative to the PRFs used as a reference. This is the source of the sensitivity (S = RON − MON) seen for gasoline fuels for the RON and MON octane number tests. In the current study, no statistically significant difference is seen in the effective PRF number for the different operating conditions tested. This does not mean that there is definitively no change in the effective PRF number for the fuels among the different operating conditions but simply indicates that for the data that were taken it is not possible to claim a change that is statistically significant. The relatively small sensitivity to the operating conditions suggests that the fuels behave reasonably closely to a mixture with a fixed PRF number for the range of thermodynamic conditions tested in this work. If a wider range of conditions were tested or more data points were taken to reduce the uncertainty, it is likely that statistically significant changes in the effective PRF number with the operating conditions would be found. Current research in our lab that has yet to be published demonstrates this result. The effective PRF value of 35 calculated for F-76 indicates that under similar engine operating conditions F-76 exhibits kinetic behavior similar to a mixture of 35% isooctane and 65% n-heptane by mass. This suggests that the well-known chemistry of isooctane and n-heptane could be used to model F-76 rather than trying to accurately model the kinetics of all of the large number of different molecules that constitute the fuel for the range of engine conditions tested. It should be noted that while these data represent engine-relevant thermodynamic conditions (see Table 4 for the in-cylinder conditions at CA10) that are difficult to achieve in other systems, the lean equivalence ratios may differ from combustion systems of interest, e.g., diesel ignition. Kinetic models validated with these data should be used with caution for test conditions that differ significantly. The effective PRF number of −25 calculated for HRD demonstrates that the developed fuel substitution strategy can be used to characterize fuels having reactivity outside of the limits set by isooctane and n-heptane at the low- and highreactivity thresholds, respectively. The negative effective PRF number for HRD is a quantitative indication of how much more reactive HRD is than n-heptane. The fuel-substitution data provide a direct comparison of a test fuel’s kinetic behavior in terms of the (hopefully) well-characterized primary reference fuels, which means that the data can be used for validation outside the typical 0−100 range of PRF numbers. The use of the PRF terminology and the fact that comparisons are made to isooctane and n-heptane should not cause the PRF number to be confused with the octane number. The octane tests (RON and MON) are performed on-engine, include other physics (flame propagation), are performed under
Figure 12. Predicted and measured PRF numbers vs F-76 mass fraction.
and the least-squares fit value for PRFHF along with its uncertainty. It can be seen that the fit to the data is excellent and correctly captures the curvature of the data points, suggesting that the simple linear blending rule applied in eq 1 adequately represents the results. To investigate whether PRFHF is dependent on the engine operating conditions, four different operating conditions were chosen; two overall mixture PRF numbers (21 and 48) causing substantially different combustion phasing were run at intake pressures of 1.2 and 2 bar. A summary of the operating conditions can be seen in Table 3. The heavy fuel PRF results, including confidence intervals, for all of the F-76 and HRD tests are shown in Figure 13 for
Figure 13. Calculated heavy fuel PRF numbers for all of the sweep conditions.
each operating condition (1−4). The calculated effective PRF numbers vary slightly among the different operating conditions, but the values are quite similar, and most of the observed variation is within the error range of the results. It can be seen that F-76 has an effective PRF number of ∼35, while HRD has an effective PRF number of approximately −25. Table 4. Operating Conditions for Fuel Substitution Sweeps operating condition
intake pressure (bar)
intake temperature (°C)
gross IMEP (bar)
CA50 (deg aTDC)
Φ
TCA10 (K)
ρCA10 (kg/m3)
1 2 3 4
1.2 1.2 2 2
165 165 165 165
4.2 4.6 5.4 5.6
−8 3 −15 −11
0.36 0.36 0.28 0.28
850 925 775 800
7.8 10.2 9.9 10.5
1405
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(4) Singh, A.; Agarwal, A. An Experimental Investigation of Combustion, Emissions and Performance of a Diesel Fuelled HCCI Engine; SAE Technical Paper 2012-28-0005; SAE International: Warrendale, PA, 2012; DOI: 10.4271/2012-28-0005. (5) Hosseini, V.; Neill, W.; Chippior, W. Influence of Engine Speed on HCCI Combustion Characteristics Using Dual-Stage Autoignition Fuels; SAE Technical Paper 2009-01-1107; SAE International: Warrendale, PA, 2009; DOI 10.4271/2009-01-1107. (6) Gray, A.; Ryan, T. Homogeneous Charge Compression Ignition (HCCI) of Diesel Fuel; SAE Technical Paper 971676; SAE International: Warrendale, PA, 1997; DOI: 10.4271/971676. (7) Sjöberg, M.; Dec, J. An Investigation of the Relationship between Measured Intake Temperature, BDC Temperature, and Combustion Phasing for Premixed and DI HCCI Engines; SAE Technical Paper 2004-01-1900; SAE International: Warrendale, PA, 2004; DOI: 10.4271/2004-01-1900. (8) Dec, J.; Sjöberg, M. Isolating the Effects of Fuel Chemistry on Combustion Phasing in an HCCI Engine and the Potential of Fuel Stratification for Ignition Control; SAE Technical Paper 2004-01-0557; SAE International: Warrendale, PA, 2004; DOI: 10.4271/2004-010557. (9) Gingrich, E.; Ghandhi, J.; Reitz, R. Experimental Investigation of Piston Heat Transfer in a Light Duty Engine under Conventional Diesel, Homogeneous Charge Compression Ignition, and Reactivity Controlled Compression Ignition Combustion Regimes; SAE Technical Paper 201401-1182; SAE International: Warrendale, PA, 2014; DOI: 10.4271/ 2014-01-1182. (10) Yuan, W.; Hansen, A.; Zhang, Q. Vapor Pressure and Normal Boiling Point Predictions for Pure Methyl Esters and Biodiesel Fuels. Fuel 2005, 84, 943−950. (11) Aligrot, C.; Champoussin, J.; Guerrassi, N.; Claus, G. A Correlative Model To Predict Autoignition Delay of Diesel Fuels; SAE Technical Paper 970638; SAE International: Warrendale, PA, 1997; DOI: 10.4271/970638. (12) Stein, Y.; Yetter, R.; Dryer, F.; Aradi, A. The Autoignition Behavior of Surrogate Diesel Fuel Mixtures and the Chemical Effects of 2Ethylhexyl Nitrate (2-EHN) Cetane Improver; SAE Technical Paper 1999-01-1504; SAE International: Warrendale, PA, 1999; DOI: 10.4271/1999-01-1504. (13) Yamada, H.; Yoshii, M.; Tezaki, A. A Chemical Mechanistic Analysis on Compression Ignition Process of Straight Chain Alkanes; SAE Technical Paper 2004-01-1912; SAE International: Warrendale, PA, 2004; DOI: 10.4271/2004-01-1912. (14) Iverson, R.; Herold, R.; Augusta, R.; Foster, D.; Ghandhi, J.; Eng, J.; Najt, P. The Effects of Intake Charge Preheating in a GasolineFueled HCCI Engine; SAE Technical Paper 2005-01-3742; SAE International: Warrendale, PA, 2005; DOI: 10.4271/2005-01-3742. (15) ASTM D2699: Standard Test Method for Research Octane Number of Spark-Ignition Engine Fuel; ASTM International: West Conshohocken, PA.
near-stoichiometric conditions, and involve significantly different temperatures and pressures. The utility of these data will be in evaluating kinetic models for real fuels under engine-relevant thermodynamic conditions. The baseline condition uses only PRFs, which are reasonably well-characterized kinetically. Under the baseline operating condition, the unknown aspects of a computational model, for example the wall heat transfer rate, residual fraction, IVC temperature, etc., can be adjusted so that the model matches the experimental data. All of these parameters can then be fixed and left unchanged for model calculations with nonzero test fuel fractions.
4. CONCLUSIONS It was seen that sweeping the intake temperature while attempting to hold all of the other boundary conditions constant did not result in the desired simple single-parameter variation. For example, the piston temperature, which affects the fuel ignition characteristics, was found to change with combustion phasing. Decoupling the intake temperature effect from the piston/wall temperature effect is not possible without fully characterizing the entire system. To address these issues, a fuel-substitution method was introduced that is well-suited for comparison to computational models. The method involves initially running an HCCI engine with a port-injected mixture of PRFs. The intake conditions are then fixed for the remainder of the testing, while the test fuel of interest is introduced in increasing fractions; the total mass of the port-injected PRFs is adjusted to maintain the engine load, and the ratio of the two PRFs is adjusted to maintain constant combustion phasing. It was demonstrated that constant in-cylinder conditions can be achieved over a wide range of test fuel mass fractions. A simple mass-based linear blending method was found to describe the data well and allowed the effective PRF numbers of the two low-volatility fuels tested to be determined under four different engine operating conditions. The effective PRF numbers of the test fuels were not found to be strongly affected by the operating conditions for the range of conditions tested in this study.
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AUTHOR INFORMATION
Corresponding Author
*E-mail:
[email protected]. Funding
Support for this work was provided by the Office of Naval Research (ONR) through Contract N00014-12-1-0650 (Sharon Beermann-Curtin, grant monitor). Notes
The authors declare no competing financial interest.
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REFERENCES
(1) Takeda, Y.; Keiichi, N.; Keiichi, N. Emission Characteristics of Premixed Lean Diesel Combustion with Extremely Early Staged Fuel Injection; SAE Technical Paper 961163; SAE International: Warrendale, PA, 1996; DOI: 10.4271/961163. (2) Iwabuchi, Y.; Kawai, K.; Shoji, T.; Takeda, Y. Trial of New Concept Diesel Combustion SystemPremixed Compression-Ignited Combustion; SAE Technical Paper 1999-01-0185; SAE International: Warrendale, PA, 1999; DOI: 10.4271/1999-01-0185. (3) Puschmann, H.; Buchwald, R.; Pannwitz, M.; Sommer, A. Homogeneous Diesel Combustion with External Mixture Formation by a Cool Flame Vaporizer; SAE Technical Paper 2006-01-3323; SAE International: Warrendale, PA, 2006; DOI: 10.4271/2006-01-3323. 1406
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