Investigation of Fuel Effects on Dilute, Mixing-Controlled Combustion

Energy & Fuels 2007, 21, 1989-2002

1989

Investigation of Fuel Effects on Dilute, Mixing-Controlled Combustion in an Optical Direct-Injection Diesel Engine A. S. (Ed) Cheng,*,† Ansis Upatnieks,‡ and Charles J. Mueller‡ School of Engineering, San Francisco State UniVersity, San Francisco, California 94132, and Combustion Research Facility, Sandia National Laboratories, LiVermore, California 94550 ReceiVed December 18, 2006. ReVised Manuscript ReceiVed March 23, 2007

Effects of fuel type on dilute diesel combustion and emissions were investigated using an optically accessible diesel engine. A number 2 diesel fuel, a primary reference fuel blend, and a soy-derived biodiesel were evaluated at an engine speed of 1200 rpm, and for intake-oxygen mole fractions ranging from 21% down to 8%. Diagnostics included conventional heat-release analysis; the measurement of in-cylinder natural luminosity; in-cylinder imaging of the start of fuel injection; and the measurement of engine-out emissions of nitrogen oxides (NOx), smoke, HC, and CO. Data previously obtained using the fuel diethylene glycol diethyl ether are presented to provide further insight. Results reveal that reduced oxygen mole fractions produce a slower heat release and lower levels of soot formation. However, engine-out particulate matter levels are also dramatically influenced by subsequent soot oxidation, which is greatest at higher oxygen mole fractions. NOx emissions are observed to drop dramatically as oxygen mole fractions are reduced, due to the associated reductions in combustion temperatures. Overall, changes in oxygen mole fractions produce similar trends in emissions results for each of the fuels investigated; however, differences do exist that may provide opportunities for implementing effective fuel-specific emissions-control strategies. A parameter called the overlimit function is introduced to help evaluate and optimize operating strategies for engine systems that must comply with multiple simultaneous constraints (e.g., emissions levels, efficiency, peak cylinder pressure, and ringing intensity) as operating conditions are varied.

1. Introduction In-cylinder strategies to reduce emissions from compressionignition engines have historically been governed by the wellknown nitrogen oxides/particulate matter (NOx/PM) tradeoff. Under the limitations of this tradeoff, reductions in NOx can be achieved only at the expense of increases in PM, and vice versa. However, within the past decade, researchers have shown that strategies employing high rates of exhaust gas recirculation (EGR) can be used to reduce emissions of NOx and PM simultaneously.1-9 In the modulated kinetics, or “MK”, combustion strategy described by Kimura et al.,1,2,7 EGR rates of ∼40% (intake-O2 * Corresponding author. Phone: 415-405-3486. Fax: 415-338-0525. E-mail: [email protected]. † San Francisco State University. ‡ Sandia National Laboratories. (1) Kimura, S.; Aoki, O.; Ogawa, H.; Muranaka, S.; Enomoto, Y. New Combustion Concept for Ultra-Clean and High-Efficiency Small DI Diesel Engines. SAE Tech. Pap. Ser. 1999, 1999-01-3681. (2) Kimura, S.; Aoki, O.; Kitahara, Y.; Aiyoshizawa, E. Ultra-Clean Combustion Technology Combining a Low-Temperature and Premixed Combustion Concept for Meeting Future Emission Standards. SAE Tech. Pap. Ser. 2001, 2001-01-0200. (3) Akihama, K.; Takatori, Y.; Inagaki, K.; Sasaki, S.; Dean, A. M. Mechanism of the Smokeless Rich Diesel Combustion by Reducing Temperature. SAE Tech. Pap. Ser. 2001, 2001-01-0655. (4) Kitamura, T.; Ito, T.; Senda, J.; Fujimoto. H. Mechanism of Smokeless Diesel Combustion with Oxygenated Fuels Based on the Dependence of the Equivalence Ratio and Temperature on Soot Particle Formation. Int. J. Engine Res. 2002, 3 (4), 223-248. (5) Wagner, R. M.; Green, J. B., Jr.; Dam, T. Q.; Edwards, K. D.; Storey, J. M. Simultaneous Low Engine-Out NOx and Particulate Matter with Highly Diluted Diesel Combustion. SAE Tech. Pap. Ser. 2003, 2003-01-0262. (6) Sluder, C. S.; Wagner, R. M.; Lewis, S. A.; Storey, J. M. E. Exhaust Chemistry of Low-NOx, Low-PM Diesel Combustion. SAE Tech. Pap. Ser. 2004, 2004-01-0114.

mole fractions of ∼15%) are combined with fuel-air premixing to achieve low NOx and PM emissions. Numerous other researchers have investigated highly dilute combustion under mixing-controlled conditions. Akihama et al. have described a “smokeless rich” diesel combustion strategy3 in which lowtemperature (and thus low-NOx) operation is achieved by increasing EGR rates to as high as 60%. Other researchers have also investigated similar high-EGR operating strategies that produce low levels of both NOx and PM.5,6,8-11 These innovative combustion strategies are similar in that they rely on high rates of EGR (high dilution levels) to reduce combustion temperatures and even to alter the diesel combustion process on a more fundamental level. Recent work at Sandia National Laboratories has provided important insight into the nature of highly dilute diesel combustion and includes experimental investigations carried out in an optically accessible constant-volume combustion vessel12,13 and in an optical diesel (7) Kawamoto, K.; Araki, T.; Shinzawa, M.; Kimura, S.; Kiode, S.; Shibuya, M. Combination of Combustion Concept and Fuel Property for Ultra-Clean DI Diesel. SAE Tech. Pap. Ser. 2004, 2004-01-1868. (8) Sluder, C. S.; Wagner, R. M.; Storey, J. M. E.; Lewis, S. A. Implications of Particulate and Precursor Compounds Formed During HighEfficiency Clean Combustion in a Diesel Engine. SAE Tech. Pap. Ser. 2005, 2005-01-3844. (9) Ogawa, H.; Miyamoto, N.; Shimizu, H.; Kido, S. Characteristics of Diesel Combustion in Low Oxygen Mixtures with Ultra-High EGR. SAE Tech. Pap. Ser. 2006, 2006-01-1147. (10) Ogawa, H.; Li, T.; Miyamoto, N.; Kido, S.; Shimizu, H. Dependence of Ultra-High EGR and Low Temperature Diesel Combustion on Fuel Injection Conditions and Compression Ratio. SAE Tech. Pap. Ser. 2006, 2006-01-3386. (11) Li, T.; Okabe, Y.; Izumi, H.; Shudo, T.; Ogawa, H. Dependence of Ultra-High EGR Low Temperature Diesel Combustion on Fuel Properties. SAE Tech. Pap. Ser. 2006, 2006-01-3387.

10.1021/ef0606456 CCC: $37.00 © 2007 American Chemical Society Published on Web 05/25/2007

1990 Energy & Fuels, Vol. 21, No. 4, 2007

engine based on the Cummins N-series production engine.14 Flame liftoff lengths and the extent of charge-gas entrainment during the quasi-steady mixing-controlled phase of diesel combustion have been shown to increase with decreasing intakeO2 mole fractions. Decreases in soot formation rates and increases in the residence times available for soot formation have also been correlated with decreasing intake-O2 mole fractions.13 For the dilute and early-injection operating conditions investigated in ref 14, the absence of a well-defined diffusion flame and the formation of soot farther downstream from the fuel injector were observed as marked differences in comparison to conventional diesel combustion. Recent studies of dilute diesel combustion have also been conducted with the oxygenated fuel diethylene glycol diethyl ether (DGE), using an optical diesel engine based on a Caterpillar on-highway truck engine.15,16 Results showed that levels of NOx and PM emissions below the 2010 regulated limits were achieved during certain steady-state operating modes employing intake-O2 mole fractions of ∼14% or lower. The extremely low emission levels were possible primarily due to the high oxygen content and low soot formation potential of DGE. The goal of the current study is to further investigate the effects of fuel type on dilute diesel combustion and emissions. Results from the study have the potential to produce two important benefits: (1) identification of a fuel that has better combustion characteristics under dilute operating conditions than conventional diesel fuel and (2) guidance for developing practical combustion strategies for fuels that are more attractive than conventional diesel from energy-security and renewableenergy standpoints (e.g., biodiesel). Additionally, data obtained from this study provide fundamental insights into NOx and PM formation mechanisms in diesel engines. For the study, an optical engine based on a Caterpillar onhighway truck engine is used to investigate three different fuels under dilute diesel combustion conditions. The fuels investigated are a number 2 diesel, a primary reference fuel blend, and a soy-derived biodiesel. Diagnostics include conventional heatrelease analysis; the measurement of in-cylinder natural luminosity; in-cylinder imaging of the start of fuel injection; and the measurement of engine-out emissions of NOx, smoke, HC, and CO. Previously reported results for DGE also are included in the discussion to provide additional insight into the effects of fuel type on dilute diesel combustion. 2. Materials and Methods Sandia Compression-Ignition Optical Research Engine. The Sandia compression-ignition optical research engine (SCORE) is a single-cylinder version of a Caterpillar 3176/C-10 engine that has been modified to provide extensive optical access to the combustion chamber. The production engine has ratings from 310 to 425 hp and is commonly used in Class 7-8 heavy-duty trucks. (12) Pickett, L. M.; Siebers, D. L. Non-Sooting, Low Flame Temperature Mixing-Controlled DI Diesel Combustion. SAE Tech. Pap. Ser. 2004, 200401-1399. (13) Idicheria, C. A.; Pickett, L. M. Soot Formation in Diesel Combustion under High-EGR Conditions. SAE Tech. Pap. Ser. 2005, 2005-01-3834. (14) Musculus, M. P. B. Multiple Simultaneous Optical Diagnostic Imaging of Early-Injection Low-Temperature Combustion in a Heavy-Duty Diesel Engine. SAE Tech. Pap. Ser. 2006, 2006-01-0079. (15) Upatnieks, A.; Mueller, C. J. Clean, Controlled DI Diesel Combustion Using Dilute, Cool Charge Gas and a Short-Ignition-Delay, Oxygenated Fuel. SAE Tech. Pap. Ser. 2005, 2005-01-0363. (16) Upatnieks, A.; Mueller, C. J.; Martin, G. C. The Influence of ChargeGas Dilution and Temperature on DI Diesel Combustion Processes Using a Short-Ignition-Delay, Oxygenated Fuel. SAE Tech. Pap. Ser. 2005, 200501-2088.

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Figure 1. Schematic of Sandia compression-ignition optical research engine (SCORE). Table 1. Sandia Compression-Ignition Optical Research Engine (SCORE) Specifications engine type cycle valves per cylinder bore stroke intake valve opena intake valve closea exhaust valve opena exhaust valve closea connecting rod length piston bowl diameter piston bowl depth swirl ratio displacement compression ratiob simulated compression ratioc

1-cylinder version of Caterpillar 3176/C-10 4-stroke CIDI 4 125 mm 140 mm 32° BTDC exhaust 153° BTDC compression 116° ATDC compression 11° ATDC exhaust 225 mm 90 mm 16.4 mm 0.59 1.72 L 11.75:1 16.00:1

a All valve timings are for lift = 0.03 mm. b Compression ratio during DGE experiments was 11.3:1. c Motored TDC temperature, pressure, and density in the production engine are matched in the optical engine by preheating and boosting the pressure of the intake air.

A schematic of the SCORE is shown in Figure 1, and the specifications of the engine are provided in Table 1. The fuel injector used in the engine is a Caterpillar hydraulically actuated, electronically controlled unit injector (HEUI) capable of injection pressures up to 142 MPa. Specifications of the fuel-injection system are provided in Table 2. The SCORE differs from the production engine in that the bottom of the piston bowl (top of the piston-crown window) is flat rather than contoured. In addition, the piston rings are positioned lower on the optical engine to allow for the placement of piston bowlrim windows and to prevent the rings from riding over windows in the cylinder wall. These modifications result in an increased clearance volume and decreased compression ratio in the SCORE relative to the production engine. To address this issue, the temperature and pressure of the intake air are increased to match

Fuel Effects on Mixing-Controlled Combustion

Energy & Fuels, Vol. 21, No. 4, 2007 1991

Table 2. Fuel Injection System Specifications injector type injector model nozzle style number of orifices orifice diameter (nom.) hydro-erosion orifice L/D included spray angle oil rail pressure max. fuel injection pressure pressure intensification ratio valve opening pressure

Caterpillar HEUI A HIA-450 single-guided VCO 6 0.163 mm 13% 8.0 140° 20.8 MPa (3000 psig) 142 MPa (20 600 psig) 6.85:1 31 MPa (4500 psig)

motored top-dead-center (TDC) conditions in the production engine. Further details of the SCORE can be found in refs 17 and 18. While the engine configuration used for this study was consistent for all three fuels tested, the engine configuration for the previous DGE experiments used a slightly different piston without bowlrim windows. Because of this, the clearance volume of the combustion chamber during the DGE experiments was approximately 2% larger. The effects of this difference in clearance volume on the DGE data will be addressed in the Results and Discussion section. Test Fuels. During the study, two pure hydrocarbon fuels and a neat, soy-derived biodiesel were investigated. The two hydrocarbon fuels were a Phillips Petroleum diesel fuel and a primary reference fuel blend. The Phillips diesel fuel, denoted D2, is a number 2 diesel fuel with 11 ppm of sulfur and 34.8% aromatics. The primary reference fuel blend consists of 76.5% by volume n-hexadecane and a balance of 2,2,4,4,6,8,8-heptamethylnonane, to produce a blend that has a defined cetane number of 80. This fuel is denoted CN80 and was previously identified to provide matched combustion phasing with neat biodiesel19 (although under different engine operating conditions than those used in the current study). The biodiesel test fuel was obtained from Peter Cremer North America, the distributor for biodiesel manufactured by Procter & Gamble Chemicals. The fuel has the trade name NEXSOL BD-0100 and is denoted B100 in this study. The fuel meets ASTM D 6751 specifications for biodiesel fuel blend stock20 and has a measured cetane number of 49.9 and an oxygen content of 11% by mass. Properties for D2, CN80, B100, and DGE are provided in Table 3. Engine Operating Conditions. The SCORE is operated in a skip-fired mode to reduce the required frequency of window cleaning and the risk of window failure due to thermal and mechanical stresses. For this study, engine operating conditions were selected to facilitate comparison with data obtained previously with DGE.16 The engine was operated at 1200 rpm, and the fuel injector was fired on every 12th engine cycle only (i.e., 11 motored cycles occur between two consecutive fired cycles). Intake-O2 mole fractions were varied from 21% to 8%, which simulates conditions ranging from no EGR to ∼75% EGR at the current load condition. The use of conventional EGR was not possible in this study due to the skip-fired operating mode that was used. Instead, EGR was simulated by diluting the intake air with pure nitrogen to achieve the same intake-O2 mole fraction present in the intake gases of the production engine when real EGR is used. The desired mixture of air and nitrogen was metered using sonic-flow orifices. Flow rates were determined from the orifice diameters and measured upstream (17) Mueller, C. J.; Musculus, M. P. Glow Plug Assisted Ignition and Combustion of Methanol in an Optical DI Diesel Engine. SAE Tech. Pap. Ser. 2001, 2001-01-2004. (18) Mueller, C. J.; Martin,.G. C. Effects of Oxygenated Compounds on Combustion and Soot Evolution in a DI Diesel Engine: Direct Luminosity Imaging. SAE Tech. Pap. Ser. 2002, 2002-01-1631. (19) Cheng, A. S.; Upatnieks, A.; Mueller, C. J. Investigation of the Impact of Biodiesel Fuelling on NOx Emissions Using an Optical Direct Injection Diesel Engine. Int. J. Engine Res. 2006, 7 (4), 297-318. (20) ASTM D 6751. Standard Specification for Biodiesel Fuel (B100) Blend Stock for Distillate Fuels; ASTM International: West Conshohocken, PA.

Figure 2. Relationship between intake-O2 mole fraction and simulated EGR. Values differ slightly from fuel to fuel due to differences in combustion products (EGR mixture composition), and thus a range of data (shaded area) is shown.

pressures using established procedures.21 Figure 2 shows the relationship between the intake-O2 mole fraction and simulated EGR rate for the intake-mixture compositions investigated in this study. For each test fuel, the indicated duration of fuel injection (DOIi) was selected such that a load of 6.7-bar gross indicated mean effective pressure (IMEP) was obtained at the 21% O2 condition. This DOIi value was maintained at all other intake-O2 mole fractions, regardless of the effect that changing the intake-O2 mole fraction had on combustion efficiency and load. Start of combustion (SOC) was controlled by adjusting the injection timing to achieve SOC within (0.25 crank angle degrees (CAD) of TDC. While varying SOC for different intake-O2 mole fractions may have provided lower emissions and higher efficiency, the goal of this study was to isolate the effects of dilution, and thus engine combustion parameters were allowed to vary with the intake-O2 mole fraction. Table 4 provides a detailed summary of the experimental engine operating conditions. Each engine run consisted of 60 fired cycles, and no fewer than three runs were conducted at each operating condition with each test fuel. Diagnostics. In-Cylinder Pressure. In-cylinder pressure was measured with 0.5-CAD resolution using a water-cooled quartz piezoelectric pressure transducer (AVL Model QC32C). Cylinderpressure data were used to determine engine parameters such as IMEP, apparent heat-release rate (AHRR), SOC, premixed-burn fraction, and combustion duration. SOC corresponds to the CAD at which the AHRR first becomes positive after the start of fuel injection. The premixed-burn fraction represents the fraction of the total heat release that occurs between SOC and the end of the premixed burn (defined by the CAD corresponding to the first local minimum in the AHRR after SOC). Combustion duration is determined on the basis of the end of combustion being defined as the CAD at which the heat release rate first becomes negative after the peak AHRR. Additional details of the cylinder-pressure data acquisition system for the SCORE can be found in ref 17. Note that, throughout this paper, CAD is referenced in a manner such that it ranges from -360 to +360, with 0 CAD corresponding to TDC of the compression stroke. Actual Start of Injection and Ignition Delay. The trigger signal to fire the SCORE fuel injector is sent to the electronic control module of the engine at an operator-selected crank angle. This is denoted the indicated start of injection, SOIi. However, the mechanical response of the injector to the trigger signal is such that fuel does not actually begin to enter the combustion chamber until a later crank angle. This is the actual start of injection, denoted (21) ASME/ANSI Standard MFC-7M - 1987. Measurement of Gas Flow by Means of Critical Flow Venturi Nozzles; ASME International: New York.


Cheng et al. Table 3. Test Fuel Properties

fuel

description

B100 CN80

soy-derived biodiesel primary reference fuel blend number 2 diesel (secondary reference fuel) diethylene glycol diethyl ether

D2 DGE a

cetane number

oxygen content [mass %]

49.9 80.0b

11.0 0.0

45.9

0.0

113-151c

29.6

sulfur [ppm]

11

aromatics [%]

1 ring: 23.1 2+ rings: 11.7

densitya [kg/m3]

lower heating value [MJ/kg]

volumetric lower heating valuea [MJ/L]

877 778

37.4 43.9

32.8 34.2

834

42.8

35.9

907

28.9

26.2

At 25 °C. b Calculated value based upon volume concentrations of primary reference fuel components. c From ref 24. Table 4. Engine Operating Conditions

engine speed engine load indicated duration of injection (DOIi) start of combustion (SOC) exhaust gas recirculation (EGR)

engine coolant temperature intake air temperature intake air pressure simulated intake air temperaturea simulated intake air pressurea motored TDC temperaturea motored TDC pressurea motored TDC densitya exhaust pressure

1200 rpm 6.7 bar IMEP at 21% O2 B100: 1565 µs CN80: 1510 µs D2: 1495 µs DGE: 2000 µs TDC ((0.25 CAD) simulated using nitrogen dilution; intake-O2 mole fractions between 21 and 8% 95 °C 95 °C 2.8 bar 53.2 °C 1.82 bar 962 K 86.0 bar 31.1 kg/m3 1.06 bar

a Assuming polytropic compression with a polytropic exponent of 1.39 obtained from motored pressure data.

SOIa. The difference in time between these two events (SOIa SOIi) is termed the injector lag. To measure ignition delay, a precise determination of SOIa is required. SOIa was determined by imaging the injected fuel spray during engine runs. Images were obtained by flood-illuminating the combustion chamber using a 532-nm, 10-Hz Nd:YAG laser beam passed through a diffuser and entering the combustion chamber through a window in the cylinder wall. The laser illumination, which lasted ∼7 ns, was pulsed at a series of delays with respect to the SOIi trigger signal. Elastically scattered light from the fuel droplets was imaged through a 532-nm bandpass filter and a Nikkor 50-mm, f/1.4 lens connected to a Roper Scientific CCD camera. Three images were acquired in 0.5 CAD increments starting from just prior to the entry of fuel into the combustion chamber. Linear extrapolation (backward in time) of fuel spray penetration lengths was carried out to determine SOIa to the nearest 0.01 CAD. SOIa was measured for each test fuel, but for the 21% O2 condition only; SOIa for a given fuel at other intake-O2 mole fractions was assumed to be unchanged. Engine-Out Emissions. Exhaust-gas samples were drawn to emissions analyzers through a heated Teflon sampling line installed downstream of an exhaust surge tank. The surge tank has a volume of 125 L and serves to dampen pressure oscillations in the exhaust system as well as to mix the gases leaving the engine from both fired and motored cycles. NOx was measured using a California Analytical Instruments (CAI) Model 600 heated chemiluminescence detector (HCLD); unburned hydrocarbons (HC) were measured using a CAI Model 600 heated flame ionization detector (HFID), and CO was measured using a CAI Model 602-P nondispersive infrared (NDIR) analyzer. Smoke emissions were measured using an AVL Model 415S smokemeter. The procedure for acquiring the emissions data was as follows. First, the engine was motored at 1200 rpm for 45 s to allow the

conditions in the intake manifold to stabilize, at which time a datalogging computer began measuring emissions levels from the HCLD, HFID, and NDIR at a rate of approximately 20 Hz. One minute after the start of motoring, the smokemeter began acquiring its sample and the injector began firing. (The first smoke from firing is captured by the smokemeter due to the finite transit time of the sample through the line.) Each engine run consisted of 60 skipfired cycles, which lasted 72 s at 1200 rpm. Engine motoring continued well after the end of firing to allow the exhaust surge tank to purge and to allow the measured emissions levels to return to their baseline values. The smokemeter and the gaseous-emissions analyzers ended their sample collection 120 and 140 s after the start of engine firing, respectively, at which time the engine was stopped. Because the SCORE is skip-fired for relatively short durations, pollutant concentrations in its exhaust do not match those that would be measured if the engine were continuously fired until emissions reached steady-state levels. The differences are due to (1) the large volume of the exhaust surge tank introducing a transient in the measured emissions such that steady-state concentrations are not reached during the 72 s duration of engine firing and (2) dilution of the exhaust gas from the fired cycles by that from the interspersed motored cycles. To address differences in gaseous emissions levels due to the first issue noted above (i.e., to determine concentrations that appropriately represent those that would be measured during steadystate skip-fired operation), the entire transient history of each gaseous species is integrated (to provide a ppm‚s result), and the result is divided by the duration of engine firing. This yields the measured average mole fraction of the species in the skip-fired exhaust. To address the difference in gaseous emissions due to skip-firing (i.e., the second issue noted above), the measured average mole fraction of the species in the skip-fired exhaust can be multiplied by the total number of engine cycles per fired cycle, which is 12 for the current conditions. Another way to remove the dilution effect of skip-firing is to use the measured average skip-fired mole fraction to compute an indicated specific emission level (e.g., in g/hp‚h) by dividing by the indicated power output during the fired cycles only. The latter approach is used in this paper, with indicated specific NOx emissions calculated using the molar mass of NO2 (46 g/mol) as per conventional practice.22 Finally, filter smoke numbers (FSNs) were corrected for dilution due to skip-firing by assuming that the sample volume is that which would pass through the smokemeter over the duration of the fired cycles only. In other words, corrected sample volume ) duration of fired cycles × actual sample volume (1) sample duration For the sample duration of 120 s, and 60 fired cycles at 1200 rpm, the correction factor becomes 0.05. (22) Code of Federal Regulations, Title 40, Part 86: Control of Emissions from New and In-Use Highway Vehicles and Engines; Environmental Protection Agency: Washington, DC, 2003.



Figure 3. Effect of intake-O2 mole fraction on apparent heat release rates (AHRR) for each of the four fuels: #2 diesel fuel (D2), 80-cetane primary reference fuel blend (CN80), neat soy biodiesel (B100), and neat diethylene glycol diethyl ether (DGE).

Quantity of Fuel Injected. Fuel consumption was quantified by measuring the mass of fuel injected as a function of the DOIi. This calibration procedure was carried out for each test fuel by using a stainless steel capture vessel placed under the cylinder head and sealed around the tip of the injector. With the engine static but heated to normal operating temperature, the fuel injector was fired a known number of times into the vessel at approximately atmospheric pressure. The rate of fuel injector firing was identical to that used during actual skip-fired engine operation. The number of injections was selected in order to inject a total volume of approximately 16 mL into the fuel capture vessel (corresponding to a mass of 12.37-14.03 g, depending on the density of the fuel). An Acculab AL-1502 balance with a resolution of 0.01 g was used to record the difference in the weight of the vessel before and after the injector firings. This result was then used to quantify the average mass of fuel introduced into the combustion chamber per injection event. Several measurements were made for each fuel at each duration of injection. The calculated average mass of fuel per injection was very repeatable, with a variation of less than 1% between measurements.

3. Results and Discussion

Figure 4. Combustion duration versus intake-O2 mole fraction for each of the four test fuels.

Apparent Heat Release Rate and Fuel Conversion Efficiency. Figure 3 provides a graphical summary of the effects of intake-O2 mole fraction on the AHRR for each of the three test fuels, along with DGE. A number of trends can be observed from the AHRR data. First, for all of the fuels, combustion duration increases with decreasing intake-O2 mole fraction, a

trend shown quantitatively in Figure 4. This is presumably a result of limiting the oxygen available for combustion. That is, assuming a fixed rate of charge gas entrainment/mixing, the rate at which O2 is supplied for fuel oxidation becomes a direct function of the intake-O2 mole fraction. Comparing across fuels,


Cheng et al.

Figure 5. Premixed burn duration versus intake-O2 mole fraction for each of the four fuels (data not shown for DGE as AHRR profiles for DGE did not exhibit a well-defined premixed burn spike).

the rate of combustion duration increase (with intake-O2 mole fraction decrease) is the greatest for the pure hydrocarbon fuels D2 and CN80, followed by B100 (11.0% fuel oxygen by mass) and then DGE (29.6% fuel oxygen by mass). Thus, reducing the charge-gas O2 mole fraction appears to have a less significant impact on combustion duration for more highly oxygenated fuels. Note that, at any given intake-O2 setpoint, fuel-to-fuel differences in combustion duration are in part due to the different fuel injection durations that are necessary to compensate for variations in fuel energy density (see Tables 3 and 4). For the test fuels D2, CN80, and B100, the duration of the premixed burn follows a similar trend as the overall combustion duration, as shown in Figure 5, in that both durations increase with increasing dilution. It should be noted that, for D2 and CN80 at the lowest intake-O2 mole fractions, a large, characteristic premixed burn spike no longer exists, and the premixed burn duration simply represents the duration of the first small heat release bump. The AHRR curves for DGE do not show premixed burn behavior at any intake-O2 mole fraction, which is likely a consequence of the fuel’s rapid autoignition behavior (cetane number ≈ 130 23). Nevertheless, DGE does produce a premixed burn spike under other engine operating conditions, as reported in refs 15 and 16. Figure 6 shows the response of peak AHRR during the premixed and mixing-controlled phases of combustion to changes in the intake-O2 mole fraction. It is evident from the figure that peak AHRRs during both phases of combustion decrease with dilution, and that fuel type has a much stronger effect on the premixed phase than on the mixing-controlled phase. Differences in peak AHRR during the premixed burn can be linked to differences in ignition delays for the fuels, as will be discussed in a subsequent section. Trends in the magnitude of the premixed burn spike can also be assessed in terms of the total amount of heat released during the premixed burn (integral of the premixed burn portion of the AHRR curve); these data are shown in Figure 7. Values are roughly constant down to intake-O2 mole fractions of 14%, then they appear to slightly increase (especially for D2 and B100) before falling off at the lowest intake-O2 mole fractions (∼8-10%). Maximum cylinder-pressure data are presented in Figure 8. A steady decrease in maximum cylinder pressure is seen with decreasing intake-O2 mole fraction for all fuels, and the (23) Murphy, M. J.; Taylor, J. D.; McCormick, R. L. Compendium of Experimental Cetane Data; National Renewable Energy Laboratory Report No. NREL/SR-540-36805; National Renewable Energy Laboratory: Golden, CO, September 2004.

Figure 6. Peak heat release during the premixed and mixing-controlled phases of combustion versus intake-O2 mole fraction. Data are not shown for DGE since AHRR profiles for DGE did not exhibit a welldefined premixed burn spike.

Figure 7. Integrated premixed burn heat release versus intake-O2 mole fraction (data not shown for DGE since AHRR profiles for DGE did not exhibit a well-defined premixed burn spike).

maximum cylinder pressure is significantly lower for DGE than for the other test fuels. One important factor limiting the maximum cylinder pressure for DGE is its lower heat release during the early stages of combustion (near TDC). Another contributor to the deficit in maximum cylinder pressure for DGE is the marginally larger clearance volume that was present during the DGE experiments. It is interesting to note that, while the peak AHRR and integrated heat release for the premixed burn are higher for D2 than for B100 (Figures 6 and 7), the maximum cylinder pressures are comparable for the two fuels. However, peak



Figure 8. Maximum measured in-cylinder pressure versus intake-O2 mole fraction for each of the four test fuels.

Figure 10. Fuel-conversion efficiency (ηf) versus intake-O2 mole fraction for each of the four test fuels.

Figure 9. Cylinder pressure traces for D2 and B100 at the 21% O2 condition. While the initial pressure rise is larger for D2, pressure increases more rapidly for B100 between ∼3-10 CAD. The result is a comparable maximum cylinder pressure at ∼10-11 CAD.

Figure 11. Gross indicated mean effective pressure (IMEP) versus intake-O2 mole fraction for each of the four test fuels.

cylinder pressures are achieved after the onset of the mixingcontrolled phase of combustion, and the pressure data indicate that B100 produces higher heat release rates relative to D2 during the initial stages of mixing-controlled combustion. This is illustrated for the 21% O2 condition in Figure 9. In terms of the effect of the intake-O2 mole fraction on heatrelease parameters, trends are largely similar for all of the four fuels. However, the intake-O2 mole fraction appears to have a smaller effect on the combustion duration for DGE than for the other fuels (see Figure 4). Also, since peak heat-release rates are highest for D2, the O2 mole fraction effect on that parameter appears more dramatic for D2 than it does for CN80 and B100 (see Figure 6). Fuel-conversion efficiencies are presented in Figure 10 and are defined as

ηf )

Wig mf qLHV

(2)

where Wig is the average gross indicated work per cycle (derived from the measured pressure data), mf is the mass of fuel injected per cycle, and qLHV is the lower heating value of the fuel. Fuelconversion efficiencies remained relatively constant down to intake-O2 mole fractions of ∼12%, then they dropped significantly at the lowest intake-O2 mole fractions. The lower ηf values at higher dilution levels are primarily a result of lower combustion efficiencies but can also be attributed in part to heat release being delayed until later in the expansion stroke (i.e., lower thermal efficiency). It is believed that fuel-conversion

efficiencies remained high at low intake-O2 mole fractions for DGE because of its exceptional ignition quality and its high oxygen content. The penalty to ηf from dilution was also slightly less significant for B100, which supports the hypothesis that increased fuel oxygen content reduces the amount of fuel/chargegas mixing that is required for efficient combustion. Fueling rates were kept constant, and thus IMEP data, shown in Figure 11, show similar trends to those observed for fuel-conversion efficiency. Actual Start of Injection and Ignition Delay. As described in section 2, in-cylinder imaging of the injected fuel spray provided a precise determination of injector lag, that is, the elapsed time between indicated and actual starts of injection. Figure 12 presents the injector lag results for each of the four fuels, which differed by no more than 0.125 ms (0.9 CAD). Ignition delays differed more significantly for the fuels, as shown in Figure 13. A timeline plot illustrating the relative differences between the injector lags and the ignition delays for the 21% O2 condition is provided in Figure 14. For lower intake-O2 mole fractions, the actual start of injection (injector lag) is assumed to be unchanged, but ignition delay increases, pushing the start of combustion to as late as 1.4 CAD, and a more significant variation in start of combustion exists between the fuels. Returning to Figure 13, for all of the fuels, ignition delays generally increase in a similar manner as the intake-O2 mole fraction decreases. For D2 and B100, the slight leveling off or drop in ignition delay at the lower (8-12%) intake-O2 mole fractions is a result of a slow, cool-flame heat release that begins to appear as an initial and distinct phase of combustion. This can be seen in the zoomed-in AHRR curves shown in


Cheng et al.

Figure 12. Injector lag for each of the four fuels (measured at the 21% O2 condition).

Figure 14. Timeline of preignition (precombustion) events for each fuel for the 21% intake-O2 mole fraction condition. Labels are provided on the B100 bar to indicate points corresponding to the indicated start of injection (SOIi), actual start of injection (SOIa), and start of combustion (SOC). For lower oxygen concentrations, SOIa is assumed to be unchanged, but ignition delay increases push the SOC to as late as 1.4 CAD.

Figure 13. Ignition delay versus intake-O2 mole fraction for each of the four test fuels.

Figure 15. Although the AHRR curve begins to cross zero at earlier crank angles, the steep, rapidly rising portion of the AHRR curve does continue to shift to later and later crank angles. The differences in ignition delay between the fuels shown in Figure 13 produce the fuel-to-fuel differences in premixed-burn heat-release parameters shown in Figures 6-7. Longer ignition delays (e.g., for D2 relative to B100) result in more fuel being available for combustion when ignition occurs, and thus moreenergetic premixed burns. However, for a given fuel, the longer ignition delays that occur as the intake-O2 mole fraction is reduced (Figure 13) are not associated with more-energetic premixed burns (Figures 6-7). Thus, for a given fuel, and over the range of engine operating conditions used in this study, the reduced intake-O2 mole fraction affects premixed burn heat release more significantly than increased ignition delay. Spatially Integrated Natural Luminosity. Spatially integrated natural luminosity (SINL) data for each of the four fuels are provided in Figure 16, while Figure 17 presents the peak values from each of the SINL curves. (Note that a logarithmic scale is used in Figure 16 while a linear scale is used in Figure 17.) For D2, CN80, and B100, peak SINL is observed to steadily decrease as the intake-O2 mole fraction decreases, in almost exactly the same manner for all three fuels. In addition, SINL curves become stretched out (in a temporal sense) in correlation with combustion durations, and peak SINL values occur at later and later crank angles. The DGE data show much lower SINL levels (at least an order of magnitude lower than those of the three test fuels), a result of the high oxygen content and thus low soot-formation potential of DGE. The rate at which the

Figure 15. Zoomed-in plots of the initial stage of heat release for the four fuels.

SINL signal decreases during the expansion stroke also appears to be greater for DGE, suggesting a faster oxidation rate of the soot produced by this fuel.24 For D2, CN80, and B100, the SINL data correspond to visual observations made during the test runs, where in-cylinder flames were a very bright orange-white at 21% O2, then grew gradually dimmer as intake-O2 mole fractions were decreased. In addition, for CN80 and B100, the flames transitioned to a very dim blue luminosity at the 8% O2 setpoint, where essentially no SINL signal was measured. (24) Vander Wal, R. L.; Mueller, C. J. Initial Investigation of Effects of Fuel Oxygenation on Nanostructure of Soot from a Direct-Injection Diesel Engine. Energy Fuels 2006, 20 (6), 2364-2369.



Figure 18. Filter smoke number versus intake-O2 mole fraction for each of the four test fuels.

Figure 16. Effect of oxygen concentration on spatially integrated natural luminosity (SINL) for each of the four fuels.

Figure 19. NOx emissions versus intake-O2 mole fraction for each of the four test fuels.

Figure 17. Peak SINL versus intake-O2 mole fraction for each of the four test fuels.

The SINL data are consistent with the following interpretations of the combustion process as intake-O2 mole fractions are reduced: first, combustion temperatures are reduced, which lowers soot temperatures and/or volume fractions and thus soot luminosity; second, reduced rates of oxygen entrainment prolong combustion duration and extend the time over which soot formation can occur. Engine-Out Emissions. Engine-out filter smoke numbers and emissions of NOx, HC, and CO are presented in Figures 1822. As shown in Figure 18, as intake-O2 mole fractions decrease, filter smoke numbers for D2, CN80, and B100 exhibit a significant increase followed by a sharp decrease. The data replicate the well-documented trends (e.g., refs 3, 8, and 9) where low PM emissions are achieved at very low intake-O2 mole fractions. For DGE, filter smoke numbers were extremely low across all intake-O2 mole fractions due to the low sootformation potential of this highly oxygenated fuel. For D2, CN80, and B100, comparison of the engine-out FSN data with the SINL data presented in Figure 17 suggests that soot emissions are low for the higher O2 levels because most of the soot formed in the reaction zone is oxidized before entering the exhaust system. A likely location for a large amount

Figure 20. NOx emissions for B100, CN80, and DGE relative to D2 at 18% and 21% O2.

of soot oxidation is at the outer diffusion flame of the reaction zone, at least for those higher O2 conditions where a significant outer diffusion flame exists. However, as dilution is increased, a well-defined diffusion flame gives way to a more volumetric combustion phenomenon,14 and even though soot concentrations may be lower in the reaction zone, a smaller fraction of this soot may be subsequently oxidized, resulting in a sharp rise in engine-out soot levels (following curves from 21% down to 10% O2 in Figure 18). From 10% to 8% O2 levels, the data suggest that, in the absence of effective soot oxidation by a diffusion flame, the FSN curve is governed more by the trend in reactionzone soot concentrations (i.e., decreasing with increasing


Figure 21. HC emissions versus intake-O2 mole fraction for each of the test fuels (data unavailable for DGE).

Figure 22. CO emissions (on a logarithmic scale) versus intake-O2 mole fraction for each of the four test fuels (data unavailable for DGE).

Figure 23. Conceptual representation of possible reaction-zone and engine-out soot concentrations as a function of intake-O2 mole fraction, based upon measured SINL and FSN data.

dilution). The preceding discussion is illustrated conceptually in Figure 23, where the top curve represents the soot formed in the reaction zones (corresponds to peak SINL curves in Figure 17) and the bottom curve or dark gray area represents engineout soot (corresponds to FSN curves in Figure 18). The light gray region in between the two curves would then represent the amount of soot formed in the reaction zone that is subsequently oxidized. Although this interpretation of soot formation and oxidation is consistent with the SINL and FSN data, more research is required to more conclusively determine its validity.

Cheng et al.

In terms of NOx emissions, the characteristic effect of a decreasing intake-O2 mole fraction (increasing simulated EGR) on NOx was observed for all fuels, as shown in Figure 19. With increasing dilution and reduced combustion temperatures, NOx emissions drop dramatically for 18% and 16% O2 levels, and they fall to below 5-6% of their undiluted levels for intake-O2 mole fractions of 14% and below. Comparing fuels, notable differences in NOx emissions are evident at the 21% O2 level. Similar fuel-to-fuel differences are also evident at 18% O2, although less significant in terms of absolute magnitude. Figure 20 summarizes the NOx emissions for B100, CN80, and DGE, relative to D2, at 18% and 21% O2. The observed NOx differences may be due, at least in part, to the significant differences that exist in AHRR between the fuels (see Figures 3 and 5-7). Measurements to quantify local in-cylinder temperatures were not made, but the AHRR data were used to calculate bulk (average) in-cylinder gas temperatures, assuming ideal gas behavior and a constant number of moles (the actual number of moles in the combustion chamber changes by no more than 2% as a result of fuel injection and combustion). The results for 18% and 21% O2 are shown in Figure 24 and illustrate the magnitude of the bulk-temperature difference that could exist between the fuels at these intake-O2 levels. Peak bulk temperatures and the rate of bulk-temperature rise are comparable for D2, CN80, and DGE. However, the bulk temperature rise is delayed for CN80 compared to D2 and is even further delayed for DGE (due to AHRR differences as well as the larger clearance volume of the engine during the DGE tests). For B100, the bulk temperature rises more rapidly than for D2 and reaches a higher peak value. The trends in bulk temperature are therefore consistent with the NOx results and the known dependence of NOx formation on temperature. However, caution must be exercised when interpreting the data in Figure 24, as bulk-temperature differences do not necessarily represent the magnitude of local temperature differences in the diffusion-flame zones where significant NOx formation occurs. Additional factors can be identified that could also contribute to the observed NOx differences. For B100, higher NOx emissions might be a result of a leaner mixture stoichiometry at the liftoff length and/or decreased radiative heat transfer away from the flame zone, as proposed in ref 19. For CN80, which is a paraffinic fuel, the stoichiometric adiabatic flame temperature may be lower than that for D2. More specific conclusions regarding the cause(s) of the NOx differences between the fuels cannot be drawn solely upon the basis of the experimental data. Modeling efforts are currently underway to further investigate NOx formation as fuel and engine conditions are varied and should provide more insight into the factors that influence NOx emissions. HC and CO emissions data are presented in Figures 21 and 22. While HC and CO emissions are low for dilution levels down to 11-12% O2, sharp increases are observed at the lowest (8-10%) intake-O2 mole fractions (HC and CO emissions were not measured for DGE). The trends in HC and CO are a consequence of incomplete combustion at the highest dilution conditions, and the sharp increases coincide with the sharp drops in fuel-conversion efficiency and IMEP shown in Figures 10-11. Figures 25-28 present summaries of the effect of the intakeO2 mole fraction on FSN, NOx, and HC on a fuel-by-fuel basis. CO data are not included because the HC data can be considered representative of trends in CO as well. Penalties (increases) in HC emissions can also be viewed as representing penalties (decreases) in fuel-conversion efficiency. Upon the basis of the



Figure 24. Bulk cylinder temperatures calculated from AHRR data assuming ideal gas behavior. (a) 18% O2 and (b) 21% O2.

Figure 25. NOx, HC, and filter smoke number versus intake-O2 mole fraction for D2.

Figure 27. NOx, HC, and filter smoke number versus intake-O2 mole fraction for B100.

Figure 26. NOx, HC, and filter smoke number versus intake-O2 mole fraction for CN80.

Figure 28. NOx and filter smoke number versus intake-O2 mole fraction for DGE. (HC emissions were not measured during the engine tests with DGE.)

figures, the impact of the intake-O2 mole fraction appears similar for D2, CN80, and B100. Low-NOx and low-smoke operation is achievable at the lowest (8%) O2 setpoint, but at the expense of dramatically increased HC emissions. DGE maintains low FSN and high efficiency values across all intake-O2 mole fractions. Thus, although HC emissions were not measured for DGE, the fuel should be able produce low NOx and low smoke with oxygen dilution before HC penalties begin to occur.

Figure 29 presents NOx/FSN tradeoff curves (as intake-O2 mole fractions are varied) for the four fuels. As previously discussed, DGE produced low smoke emissions across all dilution levels, largely due to its high oxygen content. For the other three fuels, a notable difference can be identified between the oxygenated B100 and the pure hydrocarbon fuels D2 and CN80 in that the FSN peak (sideways peak in the figure) is


Cheng et al.

value below its specified limit value does not further decrease the value of F. This behavior of the overlimit function accurately represents constraints originating from emissions legislation, since “extra credit” is not given for achieving emissions levels lower than legislated limits. The merit function has been used to serve a similar purpose to the overlimit function in previous studies (e.g., ref 25). The merit function, M, is generally defined as

M≡

Figure 29. NOx (on a logarithmic scale) versus filter smoke number for each of the four test fuels. Lower NOx levels correspond to lower intake-O2 mole fractions.

smaller and rises at a lower intake-O2 mole fraction. Thus, when employing low to moderate levels of oxygen dilution, a larger decrease in NOx could be obtained for B100 compared to D2 and CN80 before reaching a given FSN limit. (For example, for a FSN upper limit of 1, NOx can be reduced to ∼0.4 g/hp‚h for B100, compared to ∼0.6 g/hp‚h for CN80 and ∼0.8 g/hp‚h for D2.) The data suggest that the use of oxygenated fuels in reducing soot (PM) emissions offers significant opportunities for simultaneous PM and NOx reductions when dilute combustion strategies are used. Overlimit Function. One challenge faced during the development of advanced engines is the need to simultaneously satisfy constraints on a number of different operating parameters such as emissions, fuel consumption, peak pressure-rise rate, peak cylinder pressure, and so forth. For example, an engine that produces ultralow NOx and PM but that has excessive fuel consumption may be compliant with prevailing emissions legislation but not marketable due to its high fuel costs. A simple parameter for quantifying the extent to which a given combustion strategy is able to simultaneously achieve all relevant constraints is valuable during the screening of advanced combustion strategies. The overlimit function is introduced to serve this purpose. The overlimit function, denoted F, is defined as

F≡

∑i

(

max 0,

xi xi*

-1

)

(3)

where xi is the value of the ith constrained parameter at the given operating condition, xi* is the constraint on the ith parameter, and i is an index over all of the constraints. One constrained parameter that is likely to appear in an overlimit function is brake-specific NOx emissions. If the constraint on this parameter is xi* ) 0.2 g/(hp‚h) and the engine emits xi ) 0.4 g/(hp‚h), then the contribution to F from the NOx emissions is 1. In other words, the observed NOx emissions are 100% over the specified limit. Hence, the contribution to F from a given constrained parameter is provided in units of the corresponding limit value. From the definition (eq 3), the contribution will be zero if and only if the measured value is less than or equal to the specified limit. Therefore, it is easy to tell when all operational targets have been simultaneously achieved, since this corresponds to F ) 0. When F is nonzero, the contribution to F from each constrained parameter can be examined separately to quantify the severity of its noncompliance. It is also noted that reduction of a constrained parameter’s

C

( ) ( ) ( ) x1 x1*

k1

x2 + x2*

k2

x3 + x3*

k3

(4) + ...

where C is a scalar constant (typically C ) 1000), xi and xi* have the same definitions as in the overlimit function, and each ki is a positive scalar exponent. The merit function was initially used in this work to quantify the extent to which all of the constraints on the engine system are simultaneously satisfied, but the overlimit function was ultimately selected for several reasons. First, the overlimit function enables simple quantification and separation of the contribution to F from each individual constrained parameter xi, whereas the contribution from each xi to the merit function depends on all of the other xi and ki values. Second, if the value of a given xi exceeds its limit value, the overlimit function provides a linear measure of the degree of noncompliance in units of the limit value, while the merit function is a nonlinear measure of the degree of noncompliance. Third, further reductions after compliance has been achieved for a given xi will continue to improve the value of the merit function, which does not accurately reflect legislative pressures on emissions. One benefit of being able to separate the contributions to F from each of the constrained parameters is that the individual contributions can be plotted as engine operating parameters are adjusted (e.g., injection timing, EGR rate, or boost pressure) to succinctly show the effects on the constrained system. Figure 30 shows plots of F as a function of the intake-O2 mole fraction for each of the fuels examined in this study. In each plot, the height of each shaded region shows the contribution of that parameter to F, such that the total height of the stacked shaded regions at a given intake-O2 mole fraction is equal to the value of F at that dilution level. In Figure 30, the limit values for the various exhaust emissions were taken from U.S. EPA on-highway regulations that come into force in 2010,22 specifically, NOx* ) 0.2 g/(hp‚ h), HC* ) 0.14 g/(hp‚h), and CO* ) 2.6 g/(hp‚h). Note that indicated-specific values are used because brake-specific values were not measured with the optical engine. Since PM was not measured, smoke emissions are used, and smoke* ) 0.1 FSN is employed to roughly represent achievement of the legislated target of PM* ) 0.01 g/(hp‚h). Finally, indicated specific energy consumption (ISEC) is used as the measure of overall engine efficiency. ISEC is defined as the ratio of chemical energy supplied to the indicated work output of the engine. It is used instead of brake-specific fuel consumption because all of the fuels used in this study have different lower heating values. The limit is ISEC* ) 2.30, which corresponds to a gross indicated fuel-conversion efficiency of 43.5%. Figure 30 shows that high NOx, smoke, and HC and/or CO emissions are the primary problems at low, moderate, and high dilution levels, respectively. (Note that HC and CO emissions (25) Liu, Y.; Reitz, R. D.; Lu, F. Development and Application of a Non-Gradient Step-Controlled Search Algorithm for Engine Combustion Optimization. SAE Tech. Pap. Ser. 2006, 2006-01-0239.


Figure 30. Plots showing how the overlimit function, F, changes as a function of intake-O2 mole fraction for each of the fuels used in this work. The height of each shaded region shows the contribution of a given parameter to F, such that the total height of the stacked shaded regions at a given intake-O2 mole fraction is equal to the value of F at that dilution level. The limit values used to generate these plots correspond to untreated exhaust and are NOx* ) 0.2 g/(hp‚hr), HC* ) 0.14 g/(hp‚hr), CO* ) 2.6 g/(hp‚hr), smoke* ) 0.1 FSN, and ISEC* ) 2.30 (which corresponds to a gross indicated fuel-conversion efficiency of 43.5%).

data were not measured for DGE, but smoke data were.) Smoke emissions show a consistent decrease from D2 to CN80 to B100, and smoke levels for DGE are simply so low that they do not contribute to F. Contributions from the ISEC term are not visible for any of the fuels in Figure 30 because they are small relative to the emissions terms at even the highest dilution levels. Interestingly, Figure 30 shows that the minimum value of F is achieved at ∼16% O2 for D2, CN80, and B100 and that unfortunately this minimum value is several to many times greater than the limit values for NOx and smoke. On the other hand, if the HC and CO emissions follow similar trends for DGE as they do for the other fuels, operation with DGE at ∼12% O2 could provide the best opportunity for achieving compliance for the current set of constraints without using aftertreatment. Changing the set of limit constraints can dramatically affect the shapes and magnitudes of features in plots of F. Therefore, careful selection of the limit constraints is important, as is keeping these limits in mind when interpreting the plots of F that result. The set of limit constraints can also be modified to evaluate “what-if” scenarios, such as the effects of aftertreatment systems on F. Figure 31 shows the data from Figure 30 replotted assuming that the exhaust system is fitted with an oxidation catalyst that is 90% efficient at removing both HC and CO, and with a diesel particulate filter (DPF) that is 95% efficient at attenuating smoke. This corresponds to changing limits to HC* ) 1.4 g/(hp‚h), CO* ) 26 g/(hp‚h), and smoke* ) 2.0 FSN (the other limit values are held constant). Note that the upper bounds in the plots of F in Figure 31 have been decreased to 2.0. Figure 31 shows that, as expected, introduction of an oxidation catalyst and a DPF is effective at dramatically reducing emissions. Minima in the F curves now occur between 9 and


Figure 31. Plots of the overlimit function data from Figure 1, assuming that the exhaust system is fitted with an oxidation catalyst that is 90% efficient at removing both HC and CO, and with a diesel particulate filter that provides 95% smoke attenuation. The corresponding limit values are NOx* ) 0.2 g/(hp‚hr), HC* ) 1.4 g/(hp‚hr), CO* ) 26 g/(hp‚hr), smoke* ) 2.0 FSN, and ISEC* ) 2.30.

12% O2 for all of the fuels, which suggests the use of a different engine operating strategy than if no aftertreatment were used. As was evident from Figure 30, the beneficial effect of fuel oxygenation on smoke emissions is again clearly important for achieving compliance. With the addition of the oxidation catalyst and DPF, emissions levels are reduced far enough that the small contribution to F from the ISEC constraint is now visible at the bottom of each plot. As a final note, if it is desired that credit be given for reduction of a parameter value below its limit (e.g., for ISEC), the contribution to F from that term can simply be modified from max{0, (xi/xi*) -1} to (xi/xi*) - 1. Introduction of this modification, however, could lead to a situation where F ) 0 but certain emissions limits have not been met. One way to avoid this issue is to calculate and track two separate overlimit functions: one for emissions parameters and one for parameters that receive credit for being driven below their respective limit values. 4. Summary and Conclusions An optically accessible diesel engine was used to investigate the impact of fuel type on dilute diesel combustion and emissions. A number 2 diesel fuel, a primary reference fuel blend, and a neat, soy-derived biodiesel were evaluated using intake oxygen mole fractions ranging from 8% to 21%. In addition to conventional pressure-based and engine-out emissions measurements, optical diagnostics were used to quantify the start of fuel injection as well as in-cylinder natural luminosity levels. Data previously obtained with the fuel DGE also were presented to provide additional insight. The following observations and conclusions can be made on the basis of the results of this study: •With increasing dilution, heat-release rates decreased and combustion duration increased. These observations apply to the premixed burn portion of diesel combustion as well as to the overall combustion event.


•Fuel-conversion efficiencies for all fuels decreased at the highest dilution levels due to lower combustion efficiencies and delayed heat release. However, the drop in fuel-conversion efficiency was far less dramatic for DGE, presumably due to its high oxygen content and excellent ignition quality (i.e., high cetane number). The oxygenated B100 fuel also maintained somewhat higher fuel-conversion efficiencies at the highest dilution levels. The results suggest that oxygenated fuels with ignition quality superior to conventional diesel fuel are capable of highly dilute diesel combustion with less significant penalties to fuel consumption. •SINL data are consistent with the pressure-based data in revealing a longer combustion duration with increasing dilution. While quantitative in-cylinder soot volume fractions cannot be determined, the SINL data indicate that lower amounts of hightemperature soot are produced in-cylinder under highly dilute operating conditions. •Filter smoke numbers initially increase with increasing dilution, followed by a sharp decrease at the lowest intake-O2 mole fractions. When viewed along with the SINL data, results suggest that most of the in-cylinder soot formed at the highest intakeO2 mole fractions is oxidized before entering the exhaust system. As the intake-O2 mole fraction is reduced, the competing processes of soot formation and soot oxidation change in terms of their importance and impact on engine-out soot (filter smoke number). •For all fuels, NOx emissions decrease with increasing dilution and the associated reductions in combustion temperatures. Nevertheless, notable NOx differences were observed from fuel to fuel at the 21% and 18% O2 levels. These differences are correlated with differences in the average in-cylinder temperatures for the various fuels.

Cheng et al.

•HC and CO emissions increase sharply at the lowest intakeO2 mole fractions, coinciding with the sharp drops observed in fuel-conversion efficiency. •Overall, changes in the intake-O2 mole fraction appear to produce similar trends in emissions for each of the different fuels investigated. However, differences do exist which may provide opportunities for implementing effective fuel-specific emissions-control strategies. •A parameter called the overlimit function is introduced to succinctly quantify the “distance” separating an engine’s measured performance parameters (e.g., emissions, efficiency, and peak pressure-rise rate) from a given set of target values. Examples are given showing how the overlimit function can be used to evaluate and optimize engine systems with multiple simultaneous constraints as operating conditions are varied. Acknowledgment. The work described in this paper was supported by the U.S. Department of Energy (DOE) Office of FreedomCAR and Vehicle Technologies and by the San Francisco State University Summer Stipend Award Program. The authors thank DOE program managers Kevin Stork, Gurpreet Singh, and Stephen Goguen for their support of this work. The authors also gratefully acknowledge Dr. Glen Martin for many helpful discussions and his assistance with data acquisition and experimental hardware. The experiments were conducted at the Combustion Research Facility, Sandia National Laboratories, Livermore, California. Sandia is a multiprogram laboratory operated by Sandia Corporation, a Lockheed Martin Company, for the U.S. DOE’s National Nuclear Security Administration under contract DE-AC04-94AL85000. EF0606456

Investigation of Fuel Effects on Dilute, Mixing-Controlled Combustion

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