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Study of the knocking propensity of 2,5-dimethylfuran–gasoline and ethanol–gasoline blends
Fuel 98 (2012) 203–212
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Study of the knocking propensity of 2,5-dimethylfuran–gasoline and ethanol–gasoline blends David A. Rothamer a,⇑, Jamie H. Jennings b,1 a b
University of Wisconsin-Madison, Department of Mechanical Engineering, 1500 Engineering Dr., RM 127, Madison, WI 53706, USA University of Wisconsin-Madison, Department of Mechanical Engineering, 1500 Engineering Dr., Madison, WI 53706, USA
a r t i c l e
i n f o
Article history: Received 11 November 2011 Received in revised form 13 March 2012 Accepted 19 March 2012 Available online 12 April 2012 Keywords: 2,5-Dimethylfuran DMF Knocking combustion Biofuel Gasoline blends
a b s t r a c t Measurements of the knocking propensity of two biofuels blended at low levels with a gasoline fuel were performed in a single-cylinder direction-injection research engine at three load conditions. Blends tested included three blends containing 2,5-dimethylfuran (DMF) blended by volume with gasoline at concentrations of 5%, 10%, and 15%. A blend of 10% DMF/10% ethanol with a balance of gasoline was also tested. The knocking propensity of the blends using the potential advanced biofuel were compared to the performance of E10 and gasoline. Knocking intensity was quantified using in-cylinder pressure measurements. A comparison of the improvement in the knock-limited spark advance (KLSA) relative to the baseline gasoline was performed for the fuels tested, where the knock limit was defined as the spark crank angle where more than 10% of the cycles knock with a knocking intensity greater than 100 kPa. All of the blends using either DMF or ethanol showed improvement in the KLSA relative to the baseline gasoline fuel. The blend with 10% ethanol and 10% DMF showed the best performance for the fuels tested and gave an improvement in the KLSA of 7 crank angle degrees at the full load condition. Ethanol showed a greater ability to reduce knocking tendency than DMF for the same volumetric blend percentage. The higher autoignition resistance of the ethanol blends is hypothesized to be due at least in part to the higher latent heat of vaporization reducing the in-cylinder temperature. Ó 2012 Elsevier Ltd. All rights reserved.
1. Introduction Research into the production of fuels derived from biological renewable feedstocks has greatly intensified over the past decade. This work was, and still is, motivated by three primary factors: global climate change, energy security, and economics. Ethanol in particular has a long history as a biofuel with potential to displace petroleum derived gasoline. Significant research effort has been devoted recently to the production of biofuels from cellulosic biomass as opposed to its production from edible sugars and starches [1,2]. Of the possible alternatives, corn-based ethanol has seen widespread introduction as a blending component with gasoline. Gasoline with 10% ethanol is commonly available in the United States and other low-level blends of ethanol and gasoline are available throughout the world. By blending ethanol into gasoline at low levels essentially no changes are needed to current vehicles and engines to use these fuels. Such a strategy of low-level blending offers a potential gateway to new biofuels for introduction into the marketplace with lower risk of problems related to compatibility of the fuel with the current vehicle and engine infrastructure. ⇑ Corresponding author. Tel.: +1 608 890 2271. 1
E-mail address: [email protected] (D.A. Rothamer). Present Address: John Deere, USA.
0016-2361/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.fuel.2012.03.049
Although, multiple regulatory hurdles would need to be cleared before a new blend component, not substantially similar to gasoline, could be introduced into the marketplace [3]. Due to the incredibly large volumes of fuel consumed world wide blending biofuels in volume fractions up to 10% offers a large marketplace for new fuels. Low-level blends of biofuels with gasoline also offer a potential advantage to the fuel producer and consumer. If blending a high-octane biofuel component with gasoline, the overall octane number of the fuel can potentially be increased. Modern engines calibrated appropriately can take advantage of the higher octane fuel by advancing spark timing and increasing engine efficiency. Due to the current trend in spark-ignition engines of moving toward turbo-charged downsized engines, to improve overall vehicle fuel economy, the advantages associated with lower knocking propensity fuels significantly increase. Turbo-charged SI engines operate rich at high loads with the spark-timing retarded to avoid knock and to keep inlet temperatures to the turbocharger within the temperature limits of the device. Operating rich results in lower combustion efficiency which could be improved if knocking combustion could be avoided at more advanced spark times. Due to potential advantages of blending high-octane biofuels with gasoline, a study was undertaken to measure the relative knocking propensity of blends of gasoline and two different
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D.A. Rothamer, J.H. Jennings / Fuel 98 (2012) 203–212 Table 1 Engine specifications. Parameter
Value
Engine configuration
Single-cylinder, Direct-injection, Dual-overhead camshafts 9.02 cm 10.89 cm 12.03 0.696 L
Bore Stroke Compression ratio Displacement volume Valve timing IVO IVC EVO EVC
360 CAD 133 CAD 153 CAD 360 CAD
high-octane biofuel blending components: 2,5-dimethylfuran (DMF), and ethanol. Both of the biofuel blending components can be produced as advanced biofuels from cellulosic materials [4–7]. Recent biofuels research utilizing catalytic processing and ionic liquids has demonstrated that DMF can be produced directly from cellulose through the platform chemical 5-hydroxymethylfurfural (HMF) [6,8,7]. The volumetric energy density of DMF (29.6 MJ/L [9]), measured by the lower heating value of the fuel, is comparable to gasoline (32 mJ/L) and significantly higher than ethanol (21.2 MJ/L [9]). This, in addition to the fact that DMF is relatively insoluble in water (solubility in water of 1.47 g/L [10]), makes DMF a potentially attractive biofuel alternative to ethanol if it can be produced at a competitive price [11]. 2. Materials and methods 2.1. Experimental setup 2.1.1. Engine The engine used in this work was a 4-stroke direct-injection spark-ignition engine with a single-cylinder Ricardo Hydra block. The pentroof cylinder head used was of a spray-guided design with a close-coupled location of the spark plug and fuel injector. A flattop piston was used for all measurements. The fuel injector utilized was an eight-hole Bosch development injector (HDEV1). The single cylinder engine has a displacement of 0.696 L representative of one cylinder of an automotive engine. Valve timing was setup to give minimal overlap at top-dead-center (TDC) of the exhaust stroke. Maximum valve lifts for the camshafts were 8.5 mm and 9.25 mm for the intake and exhaust camshafts, respectively. The specifications for the engine are summarized in Table 1. 2.1.2. Laboratory control and data acquisition A MotoHawk Control Solutions 128-pin ECU was employed to control fuel injection and spark timing, as well as, to monitor and control real-time values of intake and exhaust gas temperatures and pressures. A piezo-electric pressure transducer (Kistler 6123A) was installed in the engine head to acquire in-cylinder pressure data. The transducer has a natural frequency of approximately 100 kHz. The pressure signal was sent to a charge amplifier and acquired by a Hi-Techniques WIN600 data acquisition system. Pressure acquisition was triggered using a 0.25 CAD resolution shaft encoder coupled to the crankshaft. In-cylinder pressure measurements where acquired for 400 consecutive cycles for each operating condition. Intake air-flow was metered using calibrated critical flow orifices. Intake air temperature downstream of the metering point was controlled using a three-phase 4.5 kW Chromalox air heater followed by five tape heaters which maintained the temperatures of the pipes and intake surge tank at the desired set point temperature ±1 K using
Fig. 1. Schematic of the laboratory setup used for knocking propensity measurements.
proportional-integral–differential (PID) control. Intake and exhaust pressures were controlled to ±1 kPa. Emissions measurements were performed with a 5-gas Horiba emissions bench providing measurements of CO2, CO, hydrocarbons, NO, and O2 concentrations in the engine exhaust. Emission measurements were used to determine the air–fuel ratio for the operating conditions applying the calculation methodology of Stivender [12]. A schematic of the general experimental layout is shown in Fig. 1. 2.2. Fuel properties The fuels tested for the current study are composed of blends of a base EPA Tier II EEE gasoline (Haltermann Products) (referred to as EEE for the remainder of the paper) mixed with either DMF, ethanol, or DMF and ethanol blended at low volumetric concentrations (<15% by volume). The blends tested include DMF blends of 5%, 10%, and 15% by volume (denoted as DMF5, DMF10, and DMF15), a 10% ethanol blend (E10), and a blend of 10% ethanol and 10% DMF (E10DMF10). Other blends of potential interest, including a 20% DMF blend and a 20% ethanol blend to compare to the E10DMF10 blend, were not tested due to cost and time limitations. The DMF used for the blends was obtained from Aldrich Chemical Company and had a purity greater than 99%. Denatured anhydrous ethanol (Haltermann Products) with 0.5% water by volume and 2.00% denaturant by volume was used in the ethanol containing blends. Table 2 shows general properties for the base fuels used to make the blends. The fuels used are all relatively high-octane. The EEE gasoline fuel had a research octane number (RON) of 96.8 and a motored octane number (MON) of 89.7 for a pump octane or anti-knock index (AKI) of 92.7. The octane numbers for ethanol are higher than those of the base gasoline fuel with a RON of 108.6, a MON of 89.7, and an AKI of 99.2. DMF is reported to have a high research octane number [8,18], but what has typically been quoted in the literature is a blending research octane number (BRON) of 119 which has its origin in a European patent [19]. This has been mislabeled as the RON for the fuel in several papers [8,18,17]. The RON and MON were
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D.A. Rothamer, J.H. Jennings / Fuel 98 (2012) 203–212 Table 2 Base fuel properties.
a
Property
Gasolinea
Ethanol
DMF
RON/MON Net heat of combustion (LHV) Gravimetric [MJ/kg] Volumetric [MJ/L] Density (293 K) [kg/L] Boiling point [K] Heat of vaporization (289 K) [kJ/kg] Reid vapor pressure (310.9 K) [kPa] Chemical formula H/C [mol/mol] O/C [mol/mol] Stoichiometric A/F ratio (weight) Water solubility [g/L]
96.8/88.5
108.6/89.7 [13]
101/88 [14]
42.89 31.82 0.742 IBP 304/FBP 476 349 [16] 62.7
26.87 [9] 21.22 0.789 [15] 351.4 [9] 931.1 [15] 15.9 [15] C2H6O 3.000 0.500 8.98 >1000 (Fully Miscible) [10]
33.27 [9] 29.55 0.888 [9] 366.2 [9] 380.7 [9] 13.4 [17] C6H8O 1.333 0.167 10.75 1.47 [10]
1.879 0.000 14.57 0
All properties for gasoline are from the Haltermann Products fuel data sheet unless otherwise noted.
previously measured to have values of 101 and 88 respectively [14]. BRON is determined by measuring the octane number of a fuel blend consisting of a base fuel with a usually small blend percentage of an additive fuel. The BRON is determined from the measured octane for the blend and the measured octane for the base fuel by [20]:
BRON ¼
RON of Blend Vol% Base fuel RON of Base Fuel : Vol% of Blending Agent ð1Þ
The BRON of a fuel is not a fundamental measure of the fuel’s autoignition characteristics since it is dependent upon the base fuel used and the percentage of additive blended in [20,11]. This is why all studies which have determined BRON or blending MON (BMON) for DMF give different results for the blending octane number when blending with gasoline. Values of BRON and BMON for DMF in the literature range from 119 to 215 [19,21,11,17]. Blending octane number generally decreases with increasing octane number of the base fuel and with increasing blend percentage of the additive, with the octane number approaching the additive neat octane number as the additive concentration approaches 100% [11]. The most relevant BRON measurements to consider for the current work are the recent measurements of Christensen et al. [17] where BRON was determined for DMF and ethanol blended into the same 85 RON base gasoline blend stock. The BRON of DMF was determined to be 153 and the BRON of ethanol was found to be 134.7 for volumetric blending percentages of 13.4% and 9.7%, respectively. These results indicate that significant improvement in the autoignition resistance would be expected when blending with either DMF or ethanol, but a slightly greater improvement would be anticipated using DMF based solely on its BRON. The energy density of DMF is attractive relative to other oxygenated fuels since, on a volumetric basis, the net heat of combustion of DMF is 93% that of gasoline. In comparison, the energy density of ethanol in volumetric terms is only 67% that of gasoline. Another property which makes DMF attractive for blending with gasoline is the low water solubility of DMF. The water solubility of DMF is less than one part in one hundred, whereas ethanol is completely miscible with water in any proportions. This may make it possible to blend DMF with gasoline at the refinery instead of blending at terminals outside the refinery [22] and might make it possible to send gasoline containing DMF through pipelines. In addition to the octane number, other key properties for gasoline use in spark-ignition engines are related to the volatility of the fuel. These include the Reid Vapor Pressure (RVP), heat of vaporization, and boiling point/distillation curve. Blending of ethanol with gasoline is non-ideal, i.e., it deviates substantially from
Raoult’s law, therefore, the volatility properties of ethanol do not give a good indication of how it will behave when blended with gasoline [22]. For instance, the RVP of ethanol is substantially lower than that of gasoline, however, ethanol actually increases the RVP of ethanol/gasoline mixtures for ethanol concentrations as high as 60% ethanol by volume [22,17]. Blending ethanol at 10% volume concentration increases the RVP of the mixture by approximately 8 kPa, while blending DMF at 13.4% by volume decreases the RVP of the blend by 3 kPa [17]. Decreasing the RVP is actually advantageous to the blender since it is generally more expensive to make fuels with lower RVP. Ethanol also significantly impacts the ASTM D 86 distillation curve when blended with gasoline due to the formation of azeotropes with various hydrocarbons in gasoline [22]. This tends to flatten out the distillation until the compounds involved in the azeotrope boil off. DMF blending at low blend levels appears to have minimal impact on the distillation besides that expected from blending in a single component into gasoline with ideal behavior [17]. The heat of vaporization impact of blending ethanol and DMF with gasoline is directly related to the heat of vaporization of the neat compounds. Heat of vaporization impacts the ability to cold start spark-ignition engines. It also impacts operation for directinjection spark-ignition engines since the enthalpy needed to vaporize the fuel comes from the in-cylinder air, reducing in-cylinder temperatures thereby helping prevent knocking combustion. The heat of vaporization of ethanol is almost three times higher than that of gasoline, therefore, even blending small amounts of ethanol can significantly impact the heat of vaporization of the fuel blend. DMF has a heat of vaporization within 10% of gasoline and is not expected to significantly impact in-cylinder temperatures at low blending levels. 2.3. Engine operating conditions Measurements for the baseline EEE gasoline fuel, E10, the three DMF blends, and the E10DMF10 blend were performed at an engine speed of 2000 rpm. Spark timing sweeps were performed at three different intake pressures for each fuel. The intake manifold absolute pressures (IMAPs) used were 50 kPa, 80 kPa, and 100 kPa. Testing was focused predominantly on high-load conditions where knocking combustion is highly probable with advanced spark timings. Gasoline measurements were run at a stoichiometric air–fuel ratio of 14.6 ± 0.2 or an equivalence ratio of 1 ± 0.014. The oxygenated blends were run using a constant fuel energy for each blend. This maintains the equivalence ratios for the blends within the range for the base gasoline of 1 ± 0.014. Injection timing can significantly influence the operation of DISI engines. The intent of the current study was to focus on the
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Table 3 Operating conditions.
Table 4 Vibrational mode factors and resonant frequencies
Parameter
Value
m, n
1, 0
2, 0
0, 1
3, 0
1, 1
Engine speed Intake pressure Spark timing
2000 rpm 50 kPa, 80 kPa, 100 kPa Swept from 5 CAD to 35 CAD or slightly past knock limit 280 CAD 10 MPa 303 K 103 kPa 343 K 343 K
am,n
1.841 6.17
3.054 10.2
3.832 12.8
4.201 14.1
5.332 17.9
Injection timing (EOI) Injection pressure Intake air temperature Exhaust back pressure Coolant temperature Oil temperature
influence of the biofuel oxygenated blend components on the knocking propensity in a DISI engine architecture representative of modern engines, as opposed to a standard ASTM D 2699 RON or ASTM D 2700 MON test which utilize outdated engine technology and, therefore, are not representative of modern SI engine operation. To make the test widely applicable across engine platforms injection timing was set to an end-of-injection (EOI) of 280 CAD approximately half way through the intake stroke (throughout the paper engine times in crank angle degrees (CAD) range from 360 CAD to +360 CAD where 0 CAD is top-dead-center of the compression stroke). At this injection timing the piston is far away from the injector location. The piston speed is close to its maximum and the intake valve lift is also near its maximum, such that the in-cylinder flow momentum is high, promoting rapid mixing of the spray and minimal wall wetting. At the time of spark it is expected that the mixture is nearly homogeneous with only small inhomogeneities still present due to the direct-injection process. Early-injection homogeneous operation is the typical high-load operation mode for all DISI engines. A fuel injection pressure of 10 MPa was used for all measurements. Table 3 summarizes the detailed engine operating conditions used. Spark timing was swept from 5 CAD to 35 CAD or slightly past where significant knocking combustion was observed, which ever occured first. It was only possible to sweep the entire range for the 50 kPa intake pressure. Knocking was confirmed by audible means and by observation of pressure oscillations on the measured pressure trace. The intake air temperature was maintained above ambient to avoid day-to-day fluctuations caused by the room air temperature changes in the laboratory. Coolant and oil temperatures were maintained at 343 K. 3. Theory/calculation 3.1. Heat release analysis A standard single-zone heat-release analysis was used to investigate the relative phasing and duration of combustion. The analysis follows directly from the work of Gatowksi et al. [23]. Heat transfer was calculated using the Woschini equation [24]. Expressions for the ratio of specific heats as a function of temperature were taken from Chun and Heywood [25]. A low pass filter with a cutoff frequency of 3 kHz was applied to the individual cycle pressure data to eliminate pressure oscillations due to knocking combustion for the heat release calculations. All 400 cycles of pressure data were averaged and heat release analysis was performed on the filtered averaged pressure trace for each operating condition. 3.2. Knock intensity calculations Knocking combustion has been studied since the first observations of it in the 1880s [26]. As instrumentation for engines
fm,n [kHz]
improved, a wide variety of techniques to detect and measure knock intensity have been developed and explored [27–37]. One of the most commonly used metrics for knock intensity is the maximum amplitude of the bandpass-filtered pressure trace [29,37,38,27]. In this method the maximum amplitude of the knock-induced pressure oscillations is determined by applying a bandpass filter with suitable cutoffs to the measured pressure trace. The maximum-amplitude method shows good correlation with other knock detection methods [37,27] and provides a meaningful measure of knocking magnitude which is related to the pressure loads seen on the piston [29]. A key aspect of applying the maximum amplitude method is selection of the bandpass cutoff frequencies. The goal is to eliminate pressure variations from the pressure trace due to piston motion and normal combustion, which have low frequency content, and to isolate the pressure variations associated with knocking combustion. A bandpass filter is used as opposed to a high-pass filter to enable the elimination of high-frequency noise content from the trace. The combustion chamber has natural resonant frequencies based on the combustion chamber shape. The resonant frequencies even for pentroof cylinder heads are approximated well by analytical solutions to the wave equation for a cylindrical combustion chamber with flat ends [29,38]. The natural frequencies determined from the solution as originally published by Draper [39] are given by:
fm;n ¼
C qm;n pB
ð2Þ
where fm,n are the natural frequencies of circumferential mode number m and radial mode number n in Hz, C is the speed of sound in the combustion chamber, am,n is the vibrational mode factor, and B is the bore of the engine. A speed of sound of 950 m/s [29] was used to estimate the vibrational mode frequencies. The vibrational mode factors and estimated frequencies for the first several circumferential and radial modes are given in Table 4. The fast fourier transform (FFT) of several knocking cycles was calculated to verify that the frequencies given in Table 4 are in agreement with peaks in the calculated spectrum. Fig. 2 shows a representative example of the FFT calculated for a knocking cycle for the 100 kPa intake pressure using EEE as the fuel with a spark time of 15 CAD. As expected, the magnitude is the largest at low frequencies corresponding to the normal compression and combustion process. Two distinct peaks are also seen at higher frequencies, one at 6.4 kHz and another at 13.0 kHz. These correspond closely in frequency to estimated frequencies in Table 4 for the first circumferential mode (1, 0) and the first radial mode. Peaks at these frequencies are consistently seen for knocking cycles. Therefore, the pass band selected for the filtering was designed to allow these frequencies to be transmitted. A finite impulse response (FIR) filter was designed using the filter design panel in IGOR Pro 6.2. The FIR filter is a combination of a high-pass and low-pass filter each using 1440 coefficients. The end of the blocking band for the high-pass filter was chosen to be 4.75 kHz and the start of the transmission band was selected to be 5.25 kHz. For the low-pass filter the transmission band ends at 14.75 kHz and the blocking band starts at 15.25 kHz. The resulting passband has 50% transmission points at 5 kHz and 15 kHz and has rejection better than 1010 outside the pass band.
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Fig. 2. Cylinder pressure fast fourier transform magnitude for a knocking cycle acquired with an intake pressure of 100 kPa and a spark time of 15 CAD. Peak corresponding to the frequency of the first circumferential mode and first radial mode are labeled using the (m, n) nomenclature.
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Fig. 4. Gross indicated mean effective pressure vs spark timing (gIMEP) measured for all fuels at three intake manifold pressures of 50, 80, and 100 kPa.
4. Results
Fig. 3. Pressure traces for a knocking and non-knocking cycle filtered using the designed FIR bandpass filter. Pressure traces were taken when operating on EEE fuel with an intake pressure of 100 kPa and a spark timing of 15 CAD.
A bandpass-filtered pressure trace corresponding to the same engine cycle used in Fig. 2 for the FFT is shown in Fig. 3. A non-knocking cycle from the same data set is also shown in Fig. 3. The filtered knocking pressure trace isolates the pressure oscillations characteristic of knocking combustion. The pressure oscillations quickly build to a peak magnitude of 330 kPa and then slowly decay as expansion occurs. In comparison the non-knocking cycle only shows small oscillations of peak magnitude less than 6 kPa. The absolute peak value of the bandpass filtered pressure trace is taken to be the knock index (KI) for the cycle. The timing for the initiation of the knocking event is determined using a threshold value equal to 0.5 times the KI. This gives reliable determination of knock onset (KO) for cases with KI greater than 20 kPa. Because pressure is only measured at one location in-cylinder, the measured knock onset is only accurate within a few crank angle degrees since the knock initiation location is not known and can vary from cycle-to-cycle [40]. A large number of cycles are needed to get a statistically converged measurement of the KI distribution. In the current work 400 cycles are acquired at each operating condition limited by number of cycles which can be acquired by the data acquisition system used. This number is lower than the 1000 cycles recommended by Brunt et al. [29] so care will need to be taken when comparing knock distributions between the fuels to insure that differences in knock are compared at high levels of confidence.
The data taken for the study consists of spark timing sweeps taken at three different intake manifold absolute pressures of 50, 80, and 100 kPa. The gross indicated mean effective pressure (gIMEP) for the three intake manifold pressures over the spark timing sweeps is shown in Fig. 4. As seen in the plot, the spark time was only swept over the entire range from 5 to 35 CAD for the 50 kPa intake pressure. For the higher intake pressures knocking became too severe to advance the timing as early as 35 CAD for any of the fuels. Instead, the spark timing was advanced several CAD past the knock-limited spark advance (KLSA). Here the KLSA is defined as the spark time where 10% of the cycles had a KI greater than 100 kPa. In Fig. 4, the 5 CAD spark time for the E10 fuel gave a lower than expected gIMEP, this is due to unstable combustion at this operating condition with a coefficient of variance (COV) of gIMEP of 12%. The loads for each of the fuels using the same calculated input fuel energy are within ±45 kPa of the mean gIMEP of the fuels for all IMAPs tested (disregarding the low gIMEP case for E10 with 80 kPa IMAP). The variation in gIMEP is highest amongst the fuels at the latest spark time of 5 CAD where the COV of IMEP is highest. The lowest intake pressure shows no significant knocking combustion for even the most advanced spark time of 35 CAD with any of the fuels tested. Therefore, the focus will be on the results for the 80 kPa and 100 kPa intake pressures.
4.1. Combustion phasing comparison To ensure that comparisons of the knocking tendency for the fuels are made on a fair basis, it is necessary to examine the relative phasing of the combustion process for the different fuels and the pressure traces under similar combustion phasing. The combustion phasing is best defined by the crank angle where 50% of the total fuel energy has been converted to sensible energy. The crank angle of 50% heat release versus spark time for the 80 kPa and 100 kPa IMAPs are shown in Figs. 5a and 5b. For the 80 kPa condition the combustion phasing is approximately the same within the uncertainty of CA50 determination (±1 CAD). The exception to this is the data for the DMF10 fuel which shows later combustion phasing at all spark times by approximately 3 CAD. To account for the delayed phasing of the DMF10 data, comparisons will be performed using data with CA50 closest to the EEE fuel. This amounts to using the data with a spark timing advanced by 2.5 CAD relative to the EEE data when comparing the 80 kPa results. For spark times more advanced than 10 CAD the 100 kPa data
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Fig. 5. Crank angle of 50% heat release for (a) the 80 kPa IMAP spark sweep and (b) the 100 kPa IMAP spark sweep.
Fig. 6. In-cylinder pressure for (a) the 80 kPa IMAP at a spark time of 22.5 CAD and (b) the 100 kPa IMAP at a spark time of 15 CAD.
(Fig. 5b) exhibit combustion phasing that is the same within the measurement uncertainty. The in-cylinder pressure traces for the two higher intake pressures and all fuels tested are shown in Fig. 6 for the most advanced spark timing which EEE data was taken. For the 80 kPa intake pressure the most advanced timing tested for the EEE fuel was 22.5 CAD and for the 100 kPa intake pressure the most advanced spark time tested was 15 CAD. The pressure traces in Fig. 6a show that the fuels, when phased appropriately, have similar pressure histories and peak pressures. Peak pressures for the 80 kPa intake temperature range from 4.97 MPa to 5.34 MPa. For the 100 kPa intake pressure average peak pressures range from 5.46 MPa to 5.77 MPa for the 6 fuels tested. 4.2. Knock statistics results The similarity of the in-cylinder conditions for all fuels tested allows for comparison of the knocking tendency of the fuels. The probability density function (PDF) of the KI was calculated for each spark time and each fuel. The PDFs were then integrated to determine the cumulative distribution function (CDF) which is used to determine the point of KLSA. The impact of spark timing on the KI PDFs and CDFs is illustrated in Figs. 7a and 7b where the KI PDFs and CDFs are shown for all spark times tested with E10 at the 80 kPa intake pressure.
At the most retarded spark times no knocking is observed in the PDFs and CDFs as indicated by the sharply peaked PDFs at KI values corresponding to the noise level of the bandpass filtered pressure and the CDFs going to a value of 1 for KI greater than the noise value of approximately 5–10 kPa. No measurable knock is seen for spark times retarded from 22.5 CAD. A slight amount of low intensity knock is noticeable for the -22.5 CAD spark time. This shows up as a decrease in the peak value of the PDF and slight shifting of the CDF curve towards higher KI values. As spark time is advanced beyond 22.5 CAD, average KI increases rapidly and the PDFs of KI broaden significantly. The PDF distributions for knocking cases are skewed towards lower KI values and have a long tail extending to high values of KI. At the earliest (most advanced) spark time of 35 CAD the average KI value is 113 kPa. This is a moderately knocking case which has more than 10% of the cycles exhibiting a KI greater the 200 kPa. Similar trends have been observed previously by Leppard [27] and others regarding the shape of the KI distributions. 4.2.1. 80 kPa Knock intensity comparison The EEE fuel used as the baseline for comparisons showed the highest tendency for knocking combustion. At the 80 kPa IMAP condition a spark timing of 20 CAD corresponds to the first advanced spark time with greater than 10% of the cycles exhibiting a KI > 100 kPa, therefore, it is the approximate KLSA. This spark
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209
Fig. 7. Knock index distributions for the 80 kPa IMAP with E10: (a) probability distribution function (PDF) (b) cumulative distribution function (CDF).
Fig. 8. Cumulative distribution functions for all fuels with phasing matched to EEE fuel at spark time of (a) 20 CAD and (b) 25 CAD for an 80 kPa IMAP.
time was used to compare the KI CDFs for all 6 fuels tested. The EEE CDF in Fig. 8a shows the largest number of cycles with significant knock, whereas, all of the DMF blends appear to perform similarly with a low level of knocking slightly greater than the essentially non-knocking results seen for the blends containing ethanol. The distributions of KI are non-normally distributed so many statistical tests are not applicable to determine whether the observed differences between fuels are statistically significant. One statistical test that is relevant regardless of the distribution shape is the Kolomogorov-Smirnov (KS) test [41]. Statistical significance of the observed distribution differences were tested using the KS test at a significance level of a = 0.01, giving a confidence of 99%. The EEE knock distribution at the 20 CAD spark time is statistically different than all other fuels tested. The DMF blends at this spark time do not show significantly different behavior from each other, but the distributions are statistically different than the ethanol blends. To further investigate the relative knocking propensity of the fuel blends, the CDFs at a spark time of 25 CAD are shown in Fig. 8b. EEE was not run at this advanced of a spark time since the knock intensity was too large to safely run. All of the oxygenated fuel blends showed an increased amount of knocking at this more advanced spark time. From the CDFs it appears that the DMF fuels behave quite similarly, whereas, the E10 and E10DMF10 blends appear to have less severe knocking than the
Fig. 9. Knock-limited spark advanced determined for the 6 fuels at the 80 kPa IMAP condition.
DMF blends. However, none of the knock distributions are statistically different at the 0.01 significance level using the KS test at a spark time of 25 CAD. With that said, at more advanced spark times the KI CDFs for all fuels are significantly different. The KLSA was determined from the spark timing sweeps at the 80 kPa intake pressure. The KLSA for all six fuels is shown in Fig. 9.
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As already discussed the EEE fuel was most susceptible to knock and exhibits the least advanced KLSA of 18.4 CAD. This spark advance is still sufficient to almost reach the maximum brake torque (MBT) spark time of 20 CAD. The three DMF blends showed similar KLSA with values ranging from 24 to 25.1 CAD all within the estimated uncertainty of ±1.0 CAD. E10 and E10DMF10 blends had KLSAs of 26.5 CAD and 27.2 CAD, respectively. The two ethanol containing blends show more advanced KLSA times than the DMF5 blend, but are just within the uncertainty when comparing to the DMF10 and DMF15 blends. The improvement in KLSA relative to the base EEE fuel for the DMF and ethanol containing blends is large ranging from approximately 6 to 9 CAD.
4.2.2. 100 kPa Knock intensity comparison The 100 kPa operating condition is a more severe test of the autoignition properties of the fuel blends with higher in-cylinder pressures and higher peak temperatures. As was the case at the 80 kPa IMAP condition, the base EEE fuel was most prone to knocking combustion. The spark time tested which was closest to the estimated KLSA for EEE was 15 CAD. The CDFs of knock intensity at a spark time of 15 CAD are shown in Fig. 10a. The EEE fuel at 15 CAD spark time shows significantly higher KI than all other fuels. The EEE distribution is statistically different, at a confidence level of greater than 99%, from all other fuels at this spark time except for the DMF5 fuel blend. The DMF15, E10, and E10DMF10 blends are essentially non-knocking at this spark time. Measurements at a more advanced spark time of 21 CAD are plotted in Fig. 10b for the DMF15, E10, and E10DMF10 blends. At 21 CAD the DMF15 blend shows moderate to high knock intensities with several cycles showing KI > 500 kPa. The DMF15 blend shows a statistically significant difference in the distribution of KI based on the KS test at 99% confidence compared to the E10 and E10DMF10 blends. At a more advanced time of 22 CAD (not shown) the KI distributions for E10 and E10DMF10 blends show a statistically significant difference with the E10DMF10 having a lower average KI. The KLSA was calculated for the 100 kPa intake pressure using linear interpolation between measured data points. Results for the KLSA for all six fuels are shown in Fig. 11. All fuels except for DMF5 have an significant impact on the KLSA greater than the uncertainty limits. The DMF5 blend does advance the KLSA but the change is within the estimated ±1 CAD uncertainty bands. In comparison to the 80 kPa results there is a slightly larger difference between the DMF blends. The impact of DMF blending seems to be more linear with DMF concentration than was the case for the
Fig. 11. Knock-limited spark advanced determined for the 6 fuels at the 100 kPa IMAP condition.
80 kPa IMAP. The E10 blend results in greater knock resistance and spark advance than is capable with any of the DMF blends tested, although the E10 and DMF15 blends have uncertainty bands which overlap. Concomitant blending with 10% ethanol and 10% DMF did advance the KLSA relative to 10% blend with either of the fuels individually. The improvement in KLSA for the E10DMF10 mixture was 6.6 CAD relative to the EEE base fuel. 5. Discussion The results of the 80 kPa and 100 kPa intake pressure spark sweeps show with statistical significance that blending either DMF, ethanol, or DMF and ethanol into an EPA certification gasoline results in improvement in the autoignition resistance of the fuel. Additionally, the improvement in autoignition resistance is operating condition dependent. This is not unexpected since the effective octane index (OI) of a fuel is dependent on the pressure and temperature history [42]. What is unexpected is the apparently superior performance of ethanol relative to DMF in improving the KLSA of the base EEE fuel. Work studying the ignition properties of 2-methylfuran (MF), which has the same molecular structure as DMF minus one methyl group, showed that on an equal mole fraction basis MF increases ignition delay by a greater amount than ethanol when blended with real gasoline [43]. The experiments, performed in a rapidcompression machine at engine like conditions, indicate that both
Fig. 10. Cumulative distribution functions at a spark time of (a) 15 CAD (all fuels) and (b) 21 CAD (DMF15, E10, E10DMF only available) for a 100 kPa IMAP.
D.A. Rothamer, J.H. Jennings / Fuel 98 (2012) 203–212
MF and ethanol inhibit low temperature (cool flame) chemistry. Wu et al. [44] recently studied the combustion of DMF in a lowpressure flame using synchrotron radiation. Their results suggest that the easily abstractable hydrogens attached to the methyl groups could play a role in suppressing the low temperature chemistry of the gasoline components. Abstraction of hydrogens from the methyl groups by reactive radicals results in a stable intermediate 2-(5-methyl) furanylmethyl thereby reducing the pool of reactive species. For the current work data have been compared on an equal volume blending basis. To compare ethanol and DMF on an approximately equal mole fraction basis the E10 and DMF15 blends can be compared, keeping in mind that the mole fraction of DMF for the DMF15 blend is still approximately 15% lower than the mole fraction of ethanol in the E10 blend. Comparing the E10 and DMF15 blends it is seen that on a mole basis DMF provides close to the same impact as ethanol on the fuels autoignition properties. However, the E10 blend does slightly outperform the DMF15 blend with a statistically significant difference in the CDFs of KI on both a mole fraction and volume fraction based comparison. These results differ from the recent work of Christensen et al. [17] who found a higher BRON for DMF of 153 compared to a BRON of 134.7 for ethanol. For their study, measurements were performed using the standard ASTM D 2700 test to measure the RON of the fuel blends. The fuel used as a base fuel was an 85 RON gasoline intended for blending with ethanol. The current results suggest a lower BRON for DMF relative to ethanol. The primary difference in the tests relate to engine conditions and the base gasoline fuel used. Direct fuel injection allows the vaporization and mixture formation impacts of the fuel to play a more significant role versus just the autoignition chemistry. Therefore, the physical properties of the fuel can play a role in determining the temperature and pressure history which the fuel experiences. In the case of the E10, the higher heat of vaporization of ethanol should result in slightly cooler in-cylinder temperatures relative to the DMF containing blends potentially resulting an effectively higher BRON under direct-injection operation. The difference in base fuel used for the blending could also contribute to the difference in blending behavior observed. In particular, the impact of the blend oxygenate on the autoignition properties of the fuel blend is dependent on the composition and octane number of the base fuel [11]. The base fuel used for the current study was a relatively high octane fuel to begin with. In fact, the improvement in KLSA is impressive for the blends tested given that the base fuel had a RON of 96.8. Previous work has shown that the improvement in KLSA is linearly related to the RON of the fuel. Typical range of values for the slope of this curve is 0.5–1.0 RON/(CAD advanced)[27]. For the current data improvements in KLSA range from 6 to 9 CAD for the 80 kPa intake pressure condition corresponding to RON improvements potentially ranging from 3 to 9 octane numbers. At the 100 kPa IMAP the improvements in KLSA for the fuels tested was in the range of 2–7 CAD which corresponds to RON improvements ranging from 1 to 7 octane numbers. The potential improvements in anti-knock properties for blending either DMF or ethanol are significant even when blending into high octane fuels.
6. Conclusions The anti-knock properties of a base gasoline fuel and 5 different fuel blends were tested at three different load conditions. The blends consisted of two oxygenates, DMF and ethanol blended in at low volumetric blending fractions. Several key conclusions and finding were derived from the presented results and are summarized here.
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1. The impact of blending small amounts of oxygenates with gasoline to improve autoignition properties is dependent on several factors: chemical composition and anti-knock properties of the base gasoline, physical properties of the blend additive and influence of additive on the physical properties of the mixture, in-cylinder temperature and pressure history (e.g. engine load), and the mixture preparation process (i.e. direction-injection versus port fuel injection, injection timing, etc.). 2. The different fuel blends tested all showed statistically significant differences in the KI distribution if the spark time was advanced far enough. 3. The relative efficacy of preventing autoignition for the blends tested was in the following order: DMF5 < DMF10 < DMF15 < E10 < E10DMF10. Contrary to results from other investigators using the ASTM RON test, the current results demonstrate a higher autoignition resistance for ethanol containing blends relative to DMF blends at the same volumetric blending fraction. This is likely the result of the engine conditions tested in the current work which are representative of modern direct injection spark-ignition engines and which take advantage of the higher latent heat of vaporization of the blends containing ethanol. 4. Both DMF and ethanol showed statistically significant impacts on improving the autoignition resistance of the base highoctane gasoline fuel. Improvements in octane number ranging from approximately 1 to 9 octane numbers appear possible for volumetric blend percentages of less than 20%. The results indicate that for direct-injection operation, ethanol is potentially more effective at reducing engine knock than DMF at the same blend percentage. However, due to the attractive energy density and much lower water solubility of DMF, it is a potentially competitive blending additive assuming it is not too toxic for widespread use. The results also indicate that co-blending of ethanol and DMF does seem to give an increased benefit relative to blending with either of the fuels individually at a lower blend percentage. Acknowledgements The authors thank Prof. Ronald Raines for supplying fuel through funding from the Great Lakes Bioenergy Research Center. This work was funded in part by the DOE Great Lakes Bioenergy Research Center (DOE Office of Science BER DE-FC0207ER64494). We would also like to thank him for engaging us to look at the combustion properties of DMF as a transportation fuel. Additionally, we would like to acknowledge funding received from the Wisconsin Alumni Research Foundation which helped support this work. References [1] Huber GW, Iborra S, Corma A. Synthesis of transportation fuels from biomass: chemistry, catalysts, and engineering. Chem Rev 2006;106(9):4044–98. [2] Serrano-Ruiz JC, West RM, Dumesic JA. Catalytic conversion of renewable biomass resources to fuels and chemicals. Annu Rev Chem Biomol Eng 2010;1(1):79–100. [3] Yanowitz J, Christensen E, McCormick RL. Utilization of renewable oxygenates as gasoline blending component. Tech. rep., National Renewable Energy Laboratory; 2011. [4] Chundawat SPS, Beckham GT, Himmel ME, Dale BE. Deconstruction of lignocellulosic biomass to fuels and chemicals. Annu Rev Chem Biomol Eng 2011;2:121–45. [5] Aden A, Foust T. Technoeconomic analysis of the dilute sulfuric acid and enzymatic hydrolysis process for the conversion of corn stover to ethanol. Cellulose 2009;16(4):535–45. [6] Binder JB, Raines RT. Simple chemical transformation of lignocellulosic biomass into furans for fuels and chemicals. J Am Chem Soc 2009;131(5):1979–85. [7] Caes BR, Raines RT. Conversion of fructose into 5-(hydroxymethyl)furfural in sulfolane. Chemsuschem 2011;4(3):353–6.
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[8] Roman-Leshkov Y, Barrett CJ, Liu ZY, Dumesic JA. Production of dimethylfuran for liquid fuels from biomass-derived carbohydrates. Nature 2007;447(7147):982. U5. [9] Yaws CL. Yaws’ handbook of thermodynamic and physical properties of chemical compounds; 2003. [10] Chemidplus advanced; 2011.
[27] Leppard WR. Individual-cylinder knock occurence and intensity in multicylinder engines. SAE Technical Paper 820074; 1982. [28] Puzinauskas PV. Examination of methods used to characterize engine knock. SAE Technical Paper 920808; 1992. [29] Brunt MFJ, Pond CR, Biundo J. Gasoline engine knock analysis using cylinder pressure data. SAE Technical Paper 980896; 1998. [30] Millo F, Ferraro CV. Knock in S.I. engines: a comparison between different techniques for detection and control. SAE Technical Paper 982477; 1998. [31] Sanches AG, Pau FS, dos Santos AM, da Silva JA. Identification and quantification of knock in internal combustion engines using spectral analysis. SAE Technical Paper 2001-01-3923; 2001. [32] Worret R, Bernhardt S, Schwarz F, Spicher U. Application of different cylinder pressure based knock detection methods in spark ignition engines. SAE Technical Paper 2002-01-1668; 2002. [33] Brecq G, Bellettre J, Tazerout M. A new indicator for knock detection in gas SI engines. Int J Therm Sci 2003;42(5):523–32. [34] Daniels CF, Zhu GG, Winkelman J. Inaudible knock and partial-burn detection using in-cylinder ionization signal. SAE Technical Paper 2003-01-3149; 2003. [35] Naber JD, Blough JR, Frankowski D, Goble M, Szpytman JE. Analysis of combustion knock metrics in spark-ignition engines. SAE Technical Paper 2006-01-0400; 2006. [36] Zavala JC, Folkerts C. Knock detection and estimation based on heat release strategies. SAE Technical Paper 2011-01-1409; 2011. [37] Burgdorf K, Denbratt I. Comparison of cylinder pressure based knock detection methods. SAE Technical Paper 972932; 1997. [38] Elmqvist C, Lindstrm F, ngstrm HE, Grandin B, Kalghatgi G. Optimizing engine concepts by using a simple model for knock prediction. SAE Technical Paper 2003-01-3123; 2003. [39] Draper CS. Pressure waves accompanying detonation in the internal combustion engine. J Aeronaut Sci 1938;5(6):219–26. [40] Chun KM, Heywood JB. Characterization of knock in a spark-ignition engine. SAE Technical Paper 890156; 1989. [41] Nist/sematech e-handbook of statistical methods; 2011.
Contents lists available at SciVerse ScienceDirect
Fuel journal homepage: www.elsevier.com/locate/fuel
Study of the knocking propensity of 2,5-dimethylfuran–gasoline and ethanol–gasoline blends David A. Rothamer a,⇑, Jamie H. Jennings b,1 a b
University of Wisconsin-Madison, Department of Mechanical Engineering, 1500 Engineering Dr., RM 127, Madison, WI 53706, USA University of Wisconsin-Madison, Department of Mechanical Engineering, 1500 Engineering Dr., Madison, WI 53706, USA
a r t i c l e
i n f o
Article history: Received 11 November 2011 Received in revised form 13 March 2012 Accepted 19 March 2012 Available online 12 April 2012 Keywords: 2,5-Dimethylfuran DMF Knocking combustion Biofuel Gasoline blends
a b s t r a c t Measurements of the knocking propensity of two biofuels blended at low levels with a gasoline fuel were performed in a single-cylinder direction-injection research engine at three load conditions. Blends tested included three blends containing 2,5-dimethylfuran (DMF) blended by volume with gasoline at concentrations of 5%, 10%, and 15%. A blend of 10% DMF/10% ethanol with a balance of gasoline was also tested. The knocking propensity of the blends using the potential advanced biofuel were compared to the performance of E10 and gasoline. Knocking intensity was quantified using in-cylinder pressure measurements. A comparison of the improvement in the knock-limited spark advance (KLSA) relative to the baseline gasoline was performed for the fuels tested, where the knock limit was defined as the spark crank angle where more than 10% of the cycles knock with a knocking intensity greater than 100 kPa. All of the blends using either DMF or ethanol showed improvement in the KLSA relative to the baseline gasoline fuel. The blend with 10% ethanol and 10% DMF showed the best performance for the fuels tested and gave an improvement in the KLSA of 7 crank angle degrees at the full load condition. Ethanol showed a greater ability to reduce knocking tendency than DMF for the same volumetric blend percentage. The higher autoignition resistance of the ethanol blends is hypothesized to be due at least in part to the higher latent heat of vaporization reducing the in-cylinder temperature. Ó 2012 Elsevier Ltd. All rights reserved.
1. Introduction Research into the production of fuels derived from biological renewable feedstocks has greatly intensified over the past decade. This work was, and still is, motivated by three primary factors: global climate change, energy security, and economics. Ethanol in particular has a long history as a biofuel with potential to displace petroleum derived gasoline. Significant research effort has been devoted recently to the production of biofuels from cellulosic biomass as opposed to its production from edible sugars and starches [1,2]. Of the possible alternatives, corn-based ethanol has seen widespread introduction as a blending component with gasoline. Gasoline with 10% ethanol is commonly available in the United States and other low-level blends of ethanol and gasoline are available throughout the world. By blending ethanol into gasoline at low levels essentially no changes are needed to current vehicles and engines to use these fuels. Such a strategy of low-level blending offers a potential gateway to new biofuels for introduction into the marketplace with lower risk of problems related to compatibility of the fuel with the current vehicle and engine infrastructure. ⇑ Corresponding author. Tel.: +1 608 890 2271. 1
E-mail address: [email protected] (D.A. Rothamer). Present Address: John Deere, USA.
0016-2361/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.fuel.2012.03.049
Although, multiple regulatory hurdles would need to be cleared before a new blend component, not substantially similar to gasoline, could be introduced into the marketplace [3]. Due to the incredibly large volumes of fuel consumed world wide blending biofuels in volume fractions up to 10% offers a large marketplace for new fuels. Low-level blends of biofuels with gasoline also offer a potential advantage to the fuel producer and consumer. If blending a high-octane biofuel component with gasoline, the overall octane number of the fuel can potentially be increased. Modern engines calibrated appropriately can take advantage of the higher octane fuel by advancing spark timing and increasing engine efficiency. Due to the current trend in spark-ignition engines of moving toward turbo-charged downsized engines, to improve overall vehicle fuel economy, the advantages associated with lower knocking propensity fuels significantly increase. Turbo-charged SI engines operate rich at high loads with the spark-timing retarded to avoid knock and to keep inlet temperatures to the turbocharger within the temperature limits of the device. Operating rich results in lower combustion efficiency which could be improved if knocking combustion could be avoided at more advanced spark times. Due to potential advantages of blending high-octane biofuels with gasoline, a study was undertaken to measure the relative knocking propensity of blends of gasoline and two different
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D.A. Rothamer, J.H. Jennings / Fuel 98 (2012) 203–212 Table 1 Engine specifications. Parameter
Value
Engine configuration
Single-cylinder, Direct-injection, Dual-overhead camshafts 9.02 cm 10.89 cm 12.03 0.696 L
Bore Stroke Compression ratio Displacement volume Valve timing IVO IVC EVO EVC
360 CAD 133 CAD 153 CAD 360 CAD
high-octane biofuel blending components: 2,5-dimethylfuran (DMF), and ethanol. Both of the biofuel blending components can be produced as advanced biofuels from cellulosic materials [4–7]. Recent biofuels research utilizing catalytic processing and ionic liquids has demonstrated that DMF can be produced directly from cellulose through the platform chemical 5-hydroxymethylfurfural (HMF) [6,8,7]. The volumetric energy density of DMF (29.6 MJ/L [9]), measured by the lower heating value of the fuel, is comparable to gasoline (32 mJ/L) and significantly higher than ethanol (21.2 MJ/L [9]). This, in addition to the fact that DMF is relatively insoluble in water (solubility in water of 1.47 g/L [10]), makes DMF a potentially attractive biofuel alternative to ethanol if it can be produced at a competitive price [11]. 2. Materials and methods 2.1. Experimental setup 2.1.1. Engine The engine used in this work was a 4-stroke direct-injection spark-ignition engine with a single-cylinder Ricardo Hydra block. The pentroof cylinder head used was of a spray-guided design with a close-coupled location of the spark plug and fuel injector. A flattop piston was used for all measurements. The fuel injector utilized was an eight-hole Bosch development injector (HDEV1). The single cylinder engine has a displacement of 0.696 L representative of one cylinder of an automotive engine. Valve timing was setup to give minimal overlap at top-dead-center (TDC) of the exhaust stroke. Maximum valve lifts for the camshafts were 8.5 mm and 9.25 mm for the intake and exhaust camshafts, respectively. The specifications for the engine are summarized in Table 1. 2.1.2. Laboratory control and data acquisition A MotoHawk Control Solutions 128-pin ECU was employed to control fuel injection and spark timing, as well as, to monitor and control real-time values of intake and exhaust gas temperatures and pressures. A piezo-electric pressure transducer (Kistler 6123A) was installed in the engine head to acquire in-cylinder pressure data. The transducer has a natural frequency of approximately 100 kHz. The pressure signal was sent to a charge amplifier and acquired by a Hi-Techniques WIN600 data acquisition system. Pressure acquisition was triggered using a 0.25 CAD resolution shaft encoder coupled to the crankshaft. In-cylinder pressure measurements where acquired for 400 consecutive cycles for each operating condition. Intake air-flow was metered using calibrated critical flow orifices. Intake air temperature downstream of the metering point was controlled using a three-phase 4.5 kW Chromalox air heater followed by five tape heaters which maintained the temperatures of the pipes and intake surge tank at the desired set point temperature ±1 K using
Fig. 1. Schematic of the laboratory setup used for knocking propensity measurements.
proportional-integral–differential (PID) control. Intake and exhaust pressures were controlled to ±1 kPa. Emissions measurements were performed with a 5-gas Horiba emissions bench providing measurements of CO2, CO, hydrocarbons, NO, and O2 concentrations in the engine exhaust. Emission measurements were used to determine the air–fuel ratio for the operating conditions applying the calculation methodology of Stivender [12]. A schematic of the general experimental layout is shown in Fig. 1. 2.2. Fuel properties The fuels tested for the current study are composed of blends of a base EPA Tier II EEE gasoline (Haltermann Products) (referred to as EEE for the remainder of the paper) mixed with either DMF, ethanol, or DMF and ethanol blended at low volumetric concentrations (<15% by volume). The blends tested include DMF blends of 5%, 10%, and 15% by volume (denoted as DMF5, DMF10, and DMF15), a 10% ethanol blend (E10), and a blend of 10% ethanol and 10% DMF (E10DMF10). Other blends of potential interest, including a 20% DMF blend and a 20% ethanol blend to compare to the E10DMF10 blend, were not tested due to cost and time limitations. The DMF used for the blends was obtained from Aldrich Chemical Company and had a purity greater than 99%. Denatured anhydrous ethanol (Haltermann Products) with 0.5% water by volume and 2.00% denaturant by volume was used in the ethanol containing blends. Table 2 shows general properties for the base fuels used to make the blends. The fuels used are all relatively high-octane. The EEE gasoline fuel had a research octane number (RON) of 96.8 and a motored octane number (MON) of 89.7 for a pump octane or anti-knock index (AKI) of 92.7. The octane numbers for ethanol are higher than those of the base gasoline fuel with a RON of 108.6, a MON of 89.7, and an AKI of 99.2. DMF is reported to have a high research octane number [8,18], but what has typically been quoted in the literature is a blending research octane number (BRON) of 119 which has its origin in a European patent [19]. This has been mislabeled as the RON for the fuel in several papers [8,18,17]. The RON and MON were
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D.A. Rothamer, J.H. Jennings / Fuel 98 (2012) 203–212 Table 2 Base fuel properties.
a
Property
Gasolinea
Ethanol
DMF
RON/MON Net heat of combustion (LHV) Gravimetric [MJ/kg] Volumetric [MJ/L] Density (293 K) [kg/L] Boiling point [K] Heat of vaporization (289 K) [kJ/kg] Reid vapor pressure (310.9 K) [kPa] Chemical formula H/C [mol/mol] O/C [mol/mol] Stoichiometric A/F ratio (weight) Water solubility [g/L]
96.8/88.5
108.6/89.7 [13]
101/88 [14]
42.89 31.82 0.742 IBP 304/FBP 476 349 [16] 62.7
26.87 [9] 21.22 0.789 [15] 351.4 [9] 931.1 [15] 15.9 [15] C2H6O 3.000 0.500 8.98 >1000 (Fully Miscible) [10]
33.27 [9] 29.55 0.888 [9] 366.2 [9] 380.7 [9] 13.4 [17] C6H8O 1.333 0.167 10.75 1.47 [10]
1.879 0.000 14.57 0
All properties for gasoline are from the Haltermann Products fuel data sheet unless otherwise noted.
previously measured to have values of 101 and 88 respectively [14]. BRON is determined by measuring the octane number of a fuel blend consisting of a base fuel with a usually small blend percentage of an additive fuel. The BRON is determined from the measured octane for the blend and the measured octane for the base fuel by [20]:
BRON ¼
RON of Blend Vol% Base fuel RON of Base Fuel : Vol% of Blending Agent ð1Þ
The BRON of a fuel is not a fundamental measure of the fuel’s autoignition characteristics since it is dependent upon the base fuel used and the percentage of additive blended in [20,11]. This is why all studies which have determined BRON or blending MON (BMON) for DMF give different results for the blending octane number when blending with gasoline. Values of BRON and BMON for DMF in the literature range from 119 to 215 [19,21,11,17]. Blending octane number generally decreases with increasing octane number of the base fuel and with increasing blend percentage of the additive, with the octane number approaching the additive neat octane number as the additive concentration approaches 100% [11]. The most relevant BRON measurements to consider for the current work are the recent measurements of Christensen et al. [17] where BRON was determined for DMF and ethanol blended into the same 85 RON base gasoline blend stock. The BRON of DMF was determined to be 153 and the BRON of ethanol was found to be 134.7 for volumetric blending percentages of 13.4% and 9.7%, respectively. These results indicate that significant improvement in the autoignition resistance would be expected when blending with either DMF or ethanol, but a slightly greater improvement would be anticipated using DMF based solely on its BRON. The energy density of DMF is attractive relative to other oxygenated fuels since, on a volumetric basis, the net heat of combustion of DMF is 93% that of gasoline. In comparison, the energy density of ethanol in volumetric terms is only 67% that of gasoline. Another property which makes DMF attractive for blending with gasoline is the low water solubility of DMF. The water solubility of DMF is less than one part in one hundred, whereas ethanol is completely miscible with water in any proportions. This may make it possible to blend DMF with gasoline at the refinery instead of blending at terminals outside the refinery [22] and might make it possible to send gasoline containing DMF through pipelines. In addition to the octane number, other key properties for gasoline use in spark-ignition engines are related to the volatility of the fuel. These include the Reid Vapor Pressure (RVP), heat of vaporization, and boiling point/distillation curve. Blending of ethanol with gasoline is non-ideal, i.e., it deviates substantially from
Raoult’s law, therefore, the volatility properties of ethanol do not give a good indication of how it will behave when blended with gasoline [22]. For instance, the RVP of ethanol is substantially lower than that of gasoline, however, ethanol actually increases the RVP of ethanol/gasoline mixtures for ethanol concentrations as high as 60% ethanol by volume [22,17]. Blending ethanol at 10% volume concentration increases the RVP of the mixture by approximately 8 kPa, while blending DMF at 13.4% by volume decreases the RVP of the blend by 3 kPa [17]. Decreasing the RVP is actually advantageous to the blender since it is generally more expensive to make fuels with lower RVP. Ethanol also significantly impacts the ASTM D 86 distillation curve when blended with gasoline due to the formation of azeotropes with various hydrocarbons in gasoline [22]. This tends to flatten out the distillation until the compounds involved in the azeotrope boil off. DMF blending at low blend levels appears to have minimal impact on the distillation besides that expected from blending in a single component into gasoline with ideal behavior [17]. The heat of vaporization impact of blending ethanol and DMF with gasoline is directly related to the heat of vaporization of the neat compounds. Heat of vaporization impacts the ability to cold start spark-ignition engines. It also impacts operation for directinjection spark-ignition engines since the enthalpy needed to vaporize the fuel comes from the in-cylinder air, reducing in-cylinder temperatures thereby helping prevent knocking combustion. The heat of vaporization of ethanol is almost three times higher than that of gasoline, therefore, even blending small amounts of ethanol can significantly impact the heat of vaporization of the fuel blend. DMF has a heat of vaporization within 10% of gasoline and is not expected to significantly impact in-cylinder temperatures at low blending levels. 2.3. Engine operating conditions Measurements for the baseline EEE gasoline fuel, E10, the three DMF blends, and the E10DMF10 blend were performed at an engine speed of 2000 rpm. Spark timing sweeps were performed at three different intake pressures for each fuel. The intake manifold absolute pressures (IMAPs) used were 50 kPa, 80 kPa, and 100 kPa. Testing was focused predominantly on high-load conditions where knocking combustion is highly probable with advanced spark timings. Gasoline measurements were run at a stoichiometric air–fuel ratio of 14.6 ± 0.2 or an equivalence ratio of 1 ± 0.014. The oxygenated blends were run using a constant fuel energy for each blend. This maintains the equivalence ratios for the blends within the range for the base gasoline of 1 ± 0.014. Injection timing can significantly influence the operation of DISI engines. The intent of the current study was to focus on the
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Table 3 Operating conditions.
Table 4 Vibrational mode factors and resonant frequencies
Parameter
Value
m, n
1, 0
2, 0
0, 1
3, 0
1, 1
Engine speed Intake pressure Spark timing
2000 rpm 50 kPa, 80 kPa, 100 kPa Swept from 5 CAD to 35 CAD or slightly past knock limit 280 CAD 10 MPa 303 K 103 kPa 343 K 343 K
am,n
1.841 6.17
3.054 10.2
3.832 12.8
4.201 14.1
5.332 17.9
Injection timing (EOI) Injection pressure Intake air temperature Exhaust back pressure Coolant temperature Oil temperature
influence of the biofuel oxygenated blend components on the knocking propensity in a DISI engine architecture representative of modern engines, as opposed to a standard ASTM D 2699 RON or ASTM D 2700 MON test which utilize outdated engine technology and, therefore, are not representative of modern SI engine operation. To make the test widely applicable across engine platforms injection timing was set to an end-of-injection (EOI) of 280 CAD approximately half way through the intake stroke (throughout the paper engine times in crank angle degrees (CAD) range from 360 CAD to +360 CAD where 0 CAD is top-dead-center of the compression stroke). At this injection timing the piston is far away from the injector location. The piston speed is close to its maximum and the intake valve lift is also near its maximum, such that the in-cylinder flow momentum is high, promoting rapid mixing of the spray and minimal wall wetting. At the time of spark it is expected that the mixture is nearly homogeneous with only small inhomogeneities still present due to the direct-injection process. Early-injection homogeneous operation is the typical high-load operation mode for all DISI engines. A fuel injection pressure of 10 MPa was used for all measurements. Table 3 summarizes the detailed engine operating conditions used. Spark timing was swept from 5 CAD to 35 CAD or slightly past where significant knocking combustion was observed, which ever occured first. It was only possible to sweep the entire range for the 50 kPa intake pressure. Knocking was confirmed by audible means and by observation of pressure oscillations on the measured pressure trace. The intake air temperature was maintained above ambient to avoid day-to-day fluctuations caused by the room air temperature changes in the laboratory. Coolant and oil temperatures were maintained at 343 K. 3. Theory/calculation 3.1. Heat release analysis A standard single-zone heat-release analysis was used to investigate the relative phasing and duration of combustion. The analysis follows directly from the work of Gatowksi et al. [23]. Heat transfer was calculated using the Woschini equation [24]. Expressions for the ratio of specific heats as a function of temperature were taken from Chun and Heywood [25]. A low pass filter with a cutoff frequency of 3 kHz was applied to the individual cycle pressure data to eliminate pressure oscillations due to knocking combustion for the heat release calculations. All 400 cycles of pressure data were averaged and heat release analysis was performed on the filtered averaged pressure trace for each operating condition. 3.2. Knock intensity calculations Knocking combustion has been studied since the first observations of it in the 1880s [26]. As instrumentation for engines
fm,n [kHz]
improved, a wide variety of techniques to detect and measure knock intensity have been developed and explored [27–37]. One of the most commonly used metrics for knock intensity is the maximum amplitude of the bandpass-filtered pressure trace [29,37,38,27]. In this method the maximum amplitude of the knock-induced pressure oscillations is determined by applying a bandpass filter with suitable cutoffs to the measured pressure trace. The maximum-amplitude method shows good correlation with other knock detection methods [37,27] and provides a meaningful measure of knocking magnitude which is related to the pressure loads seen on the piston [29]. A key aspect of applying the maximum amplitude method is selection of the bandpass cutoff frequencies. The goal is to eliminate pressure variations from the pressure trace due to piston motion and normal combustion, which have low frequency content, and to isolate the pressure variations associated with knocking combustion. A bandpass filter is used as opposed to a high-pass filter to enable the elimination of high-frequency noise content from the trace. The combustion chamber has natural resonant frequencies based on the combustion chamber shape. The resonant frequencies even for pentroof cylinder heads are approximated well by analytical solutions to the wave equation for a cylindrical combustion chamber with flat ends [29,38]. The natural frequencies determined from the solution as originally published by Draper [39] are given by:
fm;n ¼
C qm;n pB
ð2Þ
where fm,n are the natural frequencies of circumferential mode number m and radial mode number n in Hz, C is the speed of sound in the combustion chamber, am,n is the vibrational mode factor, and B is the bore of the engine. A speed of sound of 950 m/s [29] was used to estimate the vibrational mode frequencies. The vibrational mode factors and estimated frequencies for the first several circumferential and radial modes are given in Table 4. The fast fourier transform (FFT) of several knocking cycles was calculated to verify that the frequencies given in Table 4 are in agreement with peaks in the calculated spectrum. Fig. 2 shows a representative example of the FFT calculated for a knocking cycle for the 100 kPa intake pressure using EEE as the fuel with a spark time of 15 CAD. As expected, the magnitude is the largest at low frequencies corresponding to the normal compression and combustion process. Two distinct peaks are also seen at higher frequencies, one at 6.4 kHz and another at 13.0 kHz. These correspond closely in frequency to estimated frequencies in Table 4 for the first circumferential mode (1, 0) and the first radial mode. Peaks at these frequencies are consistently seen for knocking cycles. Therefore, the pass band selected for the filtering was designed to allow these frequencies to be transmitted. A finite impulse response (FIR) filter was designed using the filter design panel in IGOR Pro 6.2. The FIR filter is a combination of a high-pass and low-pass filter each using 1440 coefficients. The end of the blocking band for the high-pass filter was chosen to be 4.75 kHz and the start of the transmission band was selected to be 5.25 kHz. For the low-pass filter the transmission band ends at 14.75 kHz and the blocking band starts at 15.25 kHz. The resulting passband has 50% transmission points at 5 kHz and 15 kHz and has rejection better than 1010 outside the pass band.
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Fig. 2. Cylinder pressure fast fourier transform magnitude for a knocking cycle acquired with an intake pressure of 100 kPa and a spark time of 15 CAD. Peak corresponding to the frequency of the first circumferential mode and first radial mode are labeled using the (m, n) nomenclature.
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Fig. 4. Gross indicated mean effective pressure vs spark timing (gIMEP) measured for all fuels at three intake manifold pressures of 50, 80, and 100 kPa.
4. Results
Fig. 3. Pressure traces for a knocking and non-knocking cycle filtered using the designed FIR bandpass filter. Pressure traces were taken when operating on EEE fuel with an intake pressure of 100 kPa and a spark timing of 15 CAD.
A bandpass-filtered pressure trace corresponding to the same engine cycle used in Fig. 2 for the FFT is shown in Fig. 3. A non-knocking cycle from the same data set is also shown in Fig. 3. The filtered knocking pressure trace isolates the pressure oscillations characteristic of knocking combustion. The pressure oscillations quickly build to a peak magnitude of 330 kPa and then slowly decay as expansion occurs. In comparison the non-knocking cycle only shows small oscillations of peak magnitude less than 6 kPa. The absolute peak value of the bandpass filtered pressure trace is taken to be the knock index (KI) for the cycle. The timing for the initiation of the knocking event is determined using a threshold value equal to 0.5 times the KI. This gives reliable determination of knock onset (KO) for cases with KI greater than 20 kPa. Because pressure is only measured at one location in-cylinder, the measured knock onset is only accurate within a few crank angle degrees since the knock initiation location is not known and can vary from cycle-to-cycle [40]. A large number of cycles are needed to get a statistically converged measurement of the KI distribution. In the current work 400 cycles are acquired at each operating condition limited by number of cycles which can be acquired by the data acquisition system used. This number is lower than the 1000 cycles recommended by Brunt et al. [29] so care will need to be taken when comparing knock distributions between the fuels to insure that differences in knock are compared at high levels of confidence.
The data taken for the study consists of spark timing sweeps taken at three different intake manifold absolute pressures of 50, 80, and 100 kPa. The gross indicated mean effective pressure (gIMEP) for the three intake manifold pressures over the spark timing sweeps is shown in Fig. 4. As seen in the plot, the spark time was only swept over the entire range from 5 to 35 CAD for the 50 kPa intake pressure. For the higher intake pressures knocking became too severe to advance the timing as early as 35 CAD for any of the fuels. Instead, the spark timing was advanced several CAD past the knock-limited spark advance (KLSA). Here the KLSA is defined as the spark time where 10% of the cycles had a KI greater than 100 kPa. In Fig. 4, the 5 CAD spark time for the E10 fuel gave a lower than expected gIMEP, this is due to unstable combustion at this operating condition with a coefficient of variance (COV) of gIMEP of 12%. The loads for each of the fuels using the same calculated input fuel energy are within ±45 kPa of the mean gIMEP of the fuels for all IMAPs tested (disregarding the low gIMEP case for E10 with 80 kPa IMAP). The variation in gIMEP is highest amongst the fuels at the latest spark time of 5 CAD where the COV of IMEP is highest. The lowest intake pressure shows no significant knocking combustion for even the most advanced spark time of 35 CAD with any of the fuels tested. Therefore, the focus will be on the results for the 80 kPa and 100 kPa intake pressures.
4.1. Combustion phasing comparison To ensure that comparisons of the knocking tendency for the fuels are made on a fair basis, it is necessary to examine the relative phasing of the combustion process for the different fuels and the pressure traces under similar combustion phasing. The combustion phasing is best defined by the crank angle where 50% of the total fuel energy has been converted to sensible energy. The crank angle of 50% heat release versus spark time for the 80 kPa and 100 kPa IMAPs are shown in Figs. 5a and 5b. For the 80 kPa condition the combustion phasing is approximately the same within the uncertainty of CA50 determination (±1 CAD). The exception to this is the data for the DMF10 fuel which shows later combustion phasing at all spark times by approximately 3 CAD. To account for the delayed phasing of the DMF10 data, comparisons will be performed using data with CA50 closest to the EEE fuel. This amounts to using the data with a spark timing advanced by 2.5 CAD relative to the EEE data when comparing the 80 kPa results. For spark times more advanced than 10 CAD the 100 kPa data
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Fig. 5. Crank angle of 50% heat release for (a) the 80 kPa IMAP spark sweep and (b) the 100 kPa IMAP spark sweep.
Fig. 6. In-cylinder pressure for (a) the 80 kPa IMAP at a spark time of 22.5 CAD and (b) the 100 kPa IMAP at a spark time of 15 CAD.
(Fig. 5b) exhibit combustion phasing that is the same within the measurement uncertainty. The in-cylinder pressure traces for the two higher intake pressures and all fuels tested are shown in Fig. 6 for the most advanced spark timing which EEE data was taken. For the 80 kPa intake pressure the most advanced timing tested for the EEE fuel was 22.5 CAD and for the 100 kPa intake pressure the most advanced spark time tested was 15 CAD. The pressure traces in Fig. 6a show that the fuels, when phased appropriately, have similar pressure histories and peak pressures. Peak pressures for the 80 kPa intake temperature range from 4.97 MPa to 5.34 MPa. For the 100 kPa intake pressure average peak pressures range from 5.46 MPa to 5.77 MPa for the 6 fuels tested. 4.2. Knock statistics results The similarity of the in-cylinder conditions for all fuels tested allows for comparison of the knocking tendency of the fuels. The probability density function (PDF) of the KI was calculated for each spark time and each fuel. The PDFs were then integrated to determine the cumulative distribution function (CDF) which is used to determine the point of KLSA. The impact of spark timing on the KI PDFs and CDFs is illustrated in Figs. 7a and 7b where the KI PDFs and CDFs are shown for all spark times tested with E10 at the 80 kPa intake pressure.
At the most retarded spark times no knocking is observed in the PDFs and CDFs as indicated by the sharply peaked PDFs at KI values corresponding to the noise level of the bandpass filtered pressure and the CDFs going to a value of 1 for KI greater than the noise value of approximately 5–10 kPa. No measurable knock is seen for spark times retarded from 22.5 CAD. A slight amount of low intensity knock is noticeable for the -22.5 CAD spark time. This shows up as a decrease in the peak value of the PDF and slight shifting of the CDF curve towards higher KI values. As spark time is advanced beyond 22.5 CAD, average KI increases rapidly and the PDFs of KI broaden significantly. The PDF distributions for knocking cases are skewed towards lower KI values and have a long tail extending to high values of KI. At the earliest (most advanced) spark time of 35 CAD the average KI value is 113 kPa. This is a moderately knocking case which has more than 10% of the cycles exhibiting a KI greater the 200 kPa. Similar trends have been observed previously by Leppard [27] and others regarding the shape of the KI distributions. 4.2.1. 80 kPa Knock intensity comparison The EEE fuel used as the baseline for comparisons showed the highest tendency for knocking combustion. At the 80 kPa IMAP condition a spark timing of 20 CAD corresponds to the first advanced spark time with greater than 10% of the cycles exhibiting a KI > 100 kPa, therefore, it is the approximate KLSA. This spark
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Fig. 7. Knock index distributions for the 80 kPa IMAP with E10: (a) probability distribution function (PDF) (b) cumulative distribution function (CDF).
Fig. 8. Cumulative distribution functions for all fuels with phasing matched to EEE fuel at spark time of (a) 20 CAD and (b) 25 CAD for an 80 kPa IMAP.
time was used to compare the KI CDFs for all 6 fuels tested. The EEE CDF in Fig. 8a shows the largest number of cycles with significant knock, whereas, all of the DMF blends appear to perform similarly with a low level of knocking slightly greater than the essentially non-knocking results seen for the blends containing ethanol. The distributions of KI are non-normally distributed so many statistical tests are not applicable to determine whether the observed differences between fuels are statistically significant. One statistical test that is relevant regardless of the distribution shape is the Kolomogorov-Smirnov (KS) test [41]. Statistical significance of the observed distribution differences were tested using the KS test at a significance level of a = 0.01, giving a confidence of 99%. The EEE knock distribution at the 20 CAD spark time is statistically different than all other fuels tested. The DMF blends at this spark time do not show significantly different behavior from each other, but the distributions are statistically different than the ethanol blends. To further investigate the relative knocking propensity of the fuel blends, the CDFs at a spark time of 25 CAD are shown in Fig. 8b. EEE was not run at this advanced of a spark time since the knock intensity was too large to safely run. All of the oxygenated fuel blends showed an increased amount of knocking at this more advanced spark time. From the CDFs it appears that the DMF fuels behave quite similarly, whereas, the E10 and E10DMF10 blends appear to have less severe knocking than the
Fig. 9. Knock-limited spark advanced determined for the 6 fuels at the 80 kPa IMAP condition.
DMF blends. However, none of the knock distributions are statistically different at the 0.01 significance level using the KS test at a spark time of 25 CAD. With that said, at more advanced spark times the KI CDFs for all fuels are significantly different. The KLSA was determined from the spark timing sweeps at the 80 kPa intake pressure. The KLSA for all six fuels is shown in Fig. 9.
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As already discussed the EEE fuel was most susceptible to knock and exhibits the least advanced KLSA of 18.4 CAD. This spark advance is still sufficient to almost reach the maximum brake torque (MBT) spark time of 20 CAD. The three DMF blends showed similar KLSA with values ranging from 24 to 25.1 CAD all within the estimated uncertainty of ±1.0 CAD. E10 and E10DMF10 blends had KLSAs of 26.5 CAD and 27.2 CAD, respectively. The two ethanol containing blends show more advanced KLSA times than the DMF5 blend, but are just within the uncertainty when comparing to the DMF10 and DMF15 blends. The improvement in KLSA relative to the base EEE fuel for the DMF and ethanol containing blends is large ranging from approximately 6 to 9 CAD.
4.2.2. 100 kPa Knock intensity comparison The 100 kPa operating condition is a more severe test of the autoignition properties of the fuel blends with higher in-cylinder pressures and higher peak temperatures. As was the case at the 80 kPa IMAP condition, the base EEE fuel was most prone to knocking combustion. The spark time tested which was closest to the estimated KLSA for EEE was 15 CAD. The CDFs of knock intensity at a spark time of 15 CAD are shown in Fig. 10a. The EEE fuel at 15 CAD spark time shows significantly higher KI than all other fuels. The EEE distribution is statistically different, at a confidence level of greater than 99%, from all other fuels at this spark time except for the DMF5 fuel blend. The DMF15, E10, and E10DMF10 blends are essentially non-knocking at this spark time. Measurements at a more advanced spark time of 21 CAD are plotted in Fig. 10b for the DMF15, E10, and E10DMF10 blends. At 21 CAD the DMF15 blend shows moderate to high knock intensities with several cycles showing KI > 500 kPa. The DMF15 blend shows a statistically significant difference in the distribution of KI based on the KS test at 99% confidence compared to the E10 and E10DMF10 blends. At a more advanced time of 22 CAD (not shown) the KI distributions for E10 and E10DMF10 blends show a statistically significant difference with the E10DMF10 having a lower average KI. The KLSA was calculated for the 100 kPa intake pressure using linear interpolation between measured data points. Results for the KLSA for all six fuels are shown in Fig. 11. All fuels except for DMF5 have an significant impact on the KLSA greater than the uncertainty limits. The DMF5 blend does advance the KLSA but the change is within the estimated ±1 CAD uncertainty bands. In comparison to the 80 kPa results there is a slightly larger difference between the DMF blends. The impact of DMF blending seems to be more linear with DMF concentration than was the case for the
Fig. 11. Knock-limited spark advanced determined for the 6 fuels at the 100 kPa IMAP condition.
80 kPa IMAP. The E10 blend results in greater knock resistance and spark advance than is capable with any of the DMF blends tested, although the E10 and DMF15 blends have uncertainty bands which overlap. Concomitant blending with 10% ethanol and 10% DMF did advance the KLSA relative to 10% blend with either of the fuels individually. The improvement in KLSA for the E10DMF10 mixture was 6.6 CAD relative to the EEE base fuel. 5. Discussion The results of the 80 kPa and 100 kPa intake pressure spark sweeps show with statistical significance that blending either DMF, ethanol, or DMF and ethanol into an EPA certification gasoline results in improvement in the autoignition resistance of the fuel. Additionally, the improvement in autoignition resistance is operating condition dependent. This is not unexpected since the effective octane index (OI) of a fuel is dependent on the pressure and temperature history [42]. What is unexpected is the apparently superior performance of ethanol relative to DMF in improving the KLSA of the base EEE fuel. Work studying the ignition properties of 2-methylfuran (MF), which has the same molecular structure as DMF minus one methyl group, showed that on an equal mole fraction basis MF increases ignition delay by a greater amount than ethanol when blended with real gasoline [43]. The experiments, performed in a rapidcompression machine at engine like conditions, indicate that both
Fig. 10. Cumulative distribution functions at a spark time of (a) 15 CAD (all fuels) and (b) 21 CAD (DMF15, E10, E10DMF only available) for a 100 kPa IMAP.
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MF and ethanol inhibit low temperature (cool flame) chemistry. Wu et al. [44] recently studied the combustion of DMF in a lowpressure flame using synchrotron radiation. Their results suggest that the easily abstractable hydrogens attached to the methyl groups could play a role in suppressing the low temperature chemistry of the gasoline components. Abstraction of hydrogens from the methyl groups by reactive radicals results in a stable intermediate 2-(5-methyl) furanylmethyl thereby reducing the pool of reactive species. For the current work data have been compared on an equal volume blending basis. To compare ethanol and DMF on an approximately equal mole fraction basis the E10 and DMF15 blends can be compared, keeping in mind that the mole fraction of DMF for the DMF15 blend is still approximately 15% lower than the mole fraction of ethanol in the E10 blend. Comparing the E10 and DMF15 blends it is seen that on a mole basis DMF provides close to the same impact as ethanol on the fuels autoignition properties. However, the E10 blend does slightly outperform the DMF15 blend with a statistically significant difference in the CDFs of KI on both a mole fraction and volume fraction based comparison. These results differ from the recent work of Christensen et al. [17] who found a higher BRON for DMF of 153 compared to a BRON of 134.7 for ethanol. For their study, measurements were performed using the standard ASTM D 2700 test to measure the RON of the fuel blends. The fuel used as a base fuel was an 85 RON gasoline intended for blending with ethanol. The current results suggest a lower BRON for DMF relative to ethanol. The primary difference in the tests relate to engine conditions and the base gasoline fuel used. Direct fuel injection allows the vaporization and mixture formation impacts of the fuel to play a more significant role versus just the autoignition chemistry. Therefore, the physical properties of the fuel can play a role in determining the temperature and pressure history which the fuel experiences. In the case of the E10, the higher heat of vaporization of ethanol should result in slightly cooler in-cylinder temperatures relative to the DMF containing blends potentially resulting an effectively higher BRON under direct-injection operation. The difference in base fuel used for the blending could also contribute to the difference in blending behavior observed. In particular, the impact of the blend oxygenate on the autoignition properties of the fuel blend is dependent on the composition and octane number of the base fuel [11]. The base fuel used for the current study was a relatively high octane fuel to begin with. In fact, the improvement in KLSA is impressive for the blends tested given that the base fuel had a RON of 96.8. Previous work has shown that the improvement in KLSA is linearly related to the RON of the fuel. Typical range of values for the slope of this curve is 0.5–1.0 RON/(CAD advanced)[27]. For the current data improvements in KLSA range from 6 to 9 CAD for the 80 kPa intake pressure condition corresponding to RON improvements potentially ranging from 3 to 9 octane numbers. At the 100 kPa IMAP the improvements in KLSA for the fuels tested was in the range of 2–7 CAD which corresponds to RON improvements ranging from 1 to 7 octane numbers. The potential improvements in anti-knock properties for blending either DMF or ethanol are significant even when blending into high octane fuels.
6. Conclusions The anti-knock properties of a base gasoline fuel and 5 different fuel blends were tested at three different load conditions. The blends consisted of two oxygenates, DMF and ethanol blended in at low volumetric blending fractions. Several key conclusions and finding were derived from the presented results and are summarized here.
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1. The impact of blending small amounts of oxygenates with gasoline to improve autoignition properties is dependent on several factors: chemical composition and anti-knock properties of the base gasoline, physical properties of the blend additive and influence of additive on the physical properties of the mixture, in-cylinder temperature and pressure history (e.g. engine load), and the mixture preparation process (i.e. direction-injection versus port fuel injection, injection timing, etc.). 2. The different fuel blends tested all showed statistically significant differences in the KI distribution if the spark time was advanced far enough. 3. The relative efficacy of preventing autoignition for the blends tested was in the following order: DMF5 < DMF10 < DMF15 < E10 < E10DMF10. Contrary to results from other investigators using the ASTM RON test, the current results demonstrate a higher autoignition resistance for ethanol containing blends relative to DMF blends at the same volumetric blending fraction. This is likely the result of the engine conditions tested in the current work which are representative of modern direct injection spark-ignition engines and which take advantage of the higher latent heat of vaporization of the blends containing ethanol. 4. Both DMF and ethanol showed statistically significant impacts on improving the autoignition resistance of the base highoctane gasoline fuel. Improvements in octane number ranging from approximately 1 to 9 octane numbers appear possible for volumetric blend percentages of less than 20%. The results indicate that for direct-injection operation, ethanol is potentially more effective at reducing engine knock than DMF at the same blend percentage. However, due to the attractive energy density and much lower water solubility of DMF, it is a potentially competitive blending additive assuming it is not too toxic for widespread use. The results also indicate that co-blending of ethanol and DMF does seem to give an increased benefit relative to blending with either of the fuels individually at a lower blend percentage. Acknowledgements The authors thank Prof. Ronald Raines for supplying fuel through funding from the Great Lakes Bioenergy Research Center. This work was funded in part by the DOE Great Lakes Bioenergy Research Center (DOE Office of Science BER DE-FC0207ER64494). We would also like to thank him for engaging us to look at the combustion properties of DMF as a transportation fuel. Additionally, we would like to acknowledge funding received from the Wisconsin Alumni Research Foundation which helped support this work. References [1] Huber GW, Iborra S, Corma A. Synthesis of transportation fuels from biomass: chemistry, catalysts, and engineering. Chem Rev 2006;106(9):4044–98. [2] Serrano-Ruiz JC, West RM, Dumesic JA. Catalytic conversion of renewable biomass resources to fuels and chemicals. Annu Rev Chem Biomol Eng 2010;1(1):79–100. [3] Yanowitz J, Christensen E, McCormick RL. Utilization of renewable oxygenates as gasoline blending component. Tech. rep., National Renewable Energy Laboratory; 2011. [4] Chundawat SPS, Beckham GT, Himmel ME, Dale BE. Deconstruction of lignocellulosic biomass to fuels and chemicals. Annu Rev Chem Biomol Eng 2011;2:121–45. [5] Aden A, Foust T. Technoeconomic analysis of the dilute sulfuric acid and enzymatic hydrolysis process for the conversion of corn stover to ethanol. Cellulose 2009;16(4):535–45. [6] Binder JB, Raines RT. Simple chemical transformation of lignocellulosic biomass into furans for fuels and chemicals. J Am Chem Soc 2009;131(5):1979–85. [7] Caes BR, Raines RT. Conversion of fructose into 5-(hydroxymethyl)furfural in sulfolane. Chemsuschem 2011;4(3):353–6.
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