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High-pressure direct-injection of natural gas with entrained diesel into a compression-ignition engine Brown, Benjamin Scott 2008

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HIGH-PRESSURE DIRECT-INJECTION OF NATURAL GAS WITH ENTRAINED DIESEL INTO A COMPRESSION-IGNITION ENGINE by BENJAMIN SCOTT BROWN BASc., University of British Columbia, 2006  A THESIS SUBMITTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR THE DEGREE OF  MASTER OF APPLIED SCIENCE  in  THE FACULTY OF GRADUATE STUDIES  (Mechanical Engineering)  THE UNIVERSITY OF BRITISH COLUMBIA (Vancouver) August 2008  © Benjamin Scott Brown, 2008  Abstract The high-pressure direct-injection (HPDI) of natural gas in a compression ignition engine has the potential to reduce demand for petroleum derived fuels and significantly reduce the level of pollutants and greenhouse gases emitted from heavy duty transport vehicles. A new HPDI injector was tested where diesel is injected into a gas/diesel reservoir in the injector and the diesel and gas are then co-injected into the combustion chamber. In order to identify interactions between the diesel and gas in the reservoir, two different injector geometries were tested: prototypes A and B. Prototype B had reduced reservoir volume to increase gas velocity inside the injector. A majority of the tests were conducted in a single-cylinder test engine derived from a Cummins ISX diesel engine. As prototype A was being modified to create Prototype B this test engine was moved to a larger test cell. After updating the electrical, mechanical, and safety systems, the test engine in the new test cell was found to run repeatably; however, emissions comparisons between both test cells was not possible due to different analyzers being used. Single gas and double gas injections were conducted for both injector prototypes. The single gas injection tests found that increasing the diesel injection mass reduced the mass of gas injected. Increased diesel injection mass also shortened ignition delay, reduced unburned and partially burned fuel and increased NOx emissions. Holding the diesel injection mass constant and reducing the gas injection mass had the same effect as increasing diesel on ignition delay and gaseous emissions. If the diesel injection mass was kept constant and a second gas injection was added, the heat release due to the first injection decreased and the start of combustion was retarded. This appears to have occurred because some of the diesel was carried into the cylinder by the second injection and less diesel was available in the first injection to promote ignition. Double gas injection tests were conducted where the load, speed, and combustion timing were controlled in order to determine how injector operation affects parameters such as knock intensity, and gaseous emissions. At lower diesel injection masses, retarded combustion timing led to shorter ignition delays and less intense knock and lower unburned fuel emissions at lower loads. Longer relative times between the diesel and gas injections had a similar effect as lower diesel injection mass, especially at advanced combustion timing. For these tests Prototype B exhibited shorter ignition delays but higher knock intensities than Prototype A.  11  Table of Contents Abstract  ii  Table of Contents  iii  List of Tables  vii  List of Figures  ix  Nomenclature  xiv  Acknowledgments  xvi  Chapter 1  —  Introduction  1  1.1  Current Issues Facing Diesel Engines  1  1.2  Natural Gas Use for Heavy-Duty Engines  4  1.3  Objectives and Scope  7  1.4  Thesis Structure  8  Chapter 2— Background  9  2.1  9  2.2  Current Natural Gas Technologies 2.1.1  Stoichiometric SI Natural Gas Engines  10  2.1.2  Lean-Burn SI Natural Gas Engines  11  2.1.3  Lean-Burn Pilot-Ignited Natural Gas Engines  13  High Pressure Direct Injection 2.2.1  2.3  Ignition Delay in HPDI Engines  Co-Injection  13 14 17  2.3.1  Co-injector Operation  18  2.3.2  Previous Work at UBC on Co-injection  20  2.3.3  Patents and Studies on Gas/Liquid Co-injection  23  111  Chapter 3 3.1  3.2  —  Apparatus and Procedures  Single Cylinder Research Engine  29 29  3.1.1  Test Cell  31  3.1.2  Fuel Supply System  33  3.1.3  Air Supply System  35  3.1.4  Emissions Measurements and Calculations  37  3.1.5  Engine Speed, Temperature, and Flow Measurement  40  3.1.6  Engine Control, Monitoring, and Data Acquisition  41  3.1.7  HPDI Co-injector Operation  44  3.1.8  Injection Command Parameters  45  Cylinder Pressure Measurement and Analysis  47  3.2.1  Equipment Description  48  3.2.2  Gross Indicated Mean Effective Pressure (GIMEP) and Engine Variability  49  3.2.3  Heat Release Rate (HRR)  51  3.2.4  Ignition Delay  53  3.2.5  Knock  54  3.3  Perfonnance Comparisons for CERC and Kaiser Tests  63  3.4  Injector Characterization Flowbenches  69  Chapter 4  —  Results  70  4.1  Overview of Testing  70  4.2  Single Injection Flow Characterization  71  4.2.1  Test Series I and II: Flowbench Tests at Westport Innovations  72  iv  4.2.2  Test Series III, IV, and V: Gas/Diesel Characterization of Single Injection Tests atUBC  4.3  4.4  Test Series III and IV: Single Injection Emissions and Combustion Characteristics for Prototype A  79  Test Series VI and VII: Pilot/Main Injection Interactions  82  4.4.1  Other Factors Affecting Ignition Delay and IHR Ratio  4.4.2  Comparison Between Test Vu-A and Test Vu-B: Injector Geometry Effects on Ignition Delay and IHR Ratio, 1200 RPM  4.5  75  Test Series VIII: Emissions and Combustion for Multimode Timing Sweeps 4.5.1  Test VIII-B2: Combustion and Emission Comparisons Between Test Modes  4.5.2  Test Series VIII-A and VIII-B: Combustion Comparisons  Chapter 5  —  Conclusions  89  92 97 99 116 122  5.1  Injector Flow  123  5.2  Ignition Delay and Heat Release  124  5.3  Knock and Combustion Variability  125  5.4  Emissions  127  5.5  Conceptual Model of Co-injection  128  5.7  Co-injector Operation and Co-Injector Outlook  129  5.6  Future Work  132  References  134  Appendices  143  A.  Instrumentation List  143  B.  Results of Test Series VI and VIII Not Discussed in Body  147  V  C.  Carbon Balance and Airflow  D.  SCRE Factsheets  E.  F.  .  168 174  1.  U1-FAC-093-TEST Heather Jones  175  2.  U1-FAC-098-Test  184  3.  W1-FAC-3788-ANYS  -  —  Gord McTaggart-Cowan —  Phil Hill  190  Emissions Spreadsheets for VII and VIII Tests  197  1.  Vu-A tests at 800 RPM  198  2.  Vu-A tests at 1200 RPM  200  3.  Vu-B tests at 800 RPM  202  4.  Vu-B tests at 1200 RPM  205  5.  VIII-A tests  209  6.  VIII-B tests  211  7.  VIII—B2 Tests  215  Pressure and Heat Release Rate Curves  219  1.  Test Series VII-A-800 RPM: Pressure and HRR Curves  219  2.  Test Series VII-A-1200 RPM: Pressure and HRR Curves  239  3.  Test Series VII-B-800 RPM: Pressure and HRR Curves  256  4.  Test Series VII-B-1200 RPM: Pressure and HRR Curves  289  5.  Test Series VIII-A: Pressure and HRR Curves  330  6.  Test Series Vu-B: Pressure and HRR Curves  338  7.  Test Series VII-B2: Pressure and HRR Curves  355  vi  List of Tables Table 1.1  Exhaust emissions standards for heavy-duty engines in the United States (Dieselnet n.d.)  Table 3.1  List of components inside test cell  Table 3.2  Engine size comparison between Volvo TD100 (Christensen et al. 1998) and Cummins ISX (Duggal et al. 2004)  Table 3.3  3 32  60  Comparison of performance parameters between CERC (2008) and Kaiser (2006) tests using J36-008 injector  65  Table 3.4  Implications of 7.5% increased conversion factor  67  Table 4.1  Chronological overview of Test Series  71  Table 4.2  Controlled parameters for Test Series I and II: Westport Flowbenches BTR2 andEFSl  Table 4.3  Controlled parameters for Test Series III, IV, and V: single injection tests in SCRE  Table 4.4  72  75  Test matrix for Test Series VII: normal double, and retarded double injection operation in SCRE for both Prototype A and Prototype B. Engine Speed: 1200 RPM, manifold temperature: 70 oC, MAP—90 kPa, exhaust pressure—50 kPa, 2RIT—1.3ms  Table 4.5  85  Test matrix for test Series VIII: double injection timing sweeps for comparison of Emissions in SCRE  99  Table 5.1  Injector comparisons between the HPDI-J36 and the co-injector  130  Table A. 1  Data acquisition cards  144  Table A.2  Data acquisition hardware  144 vii  Table A.3  Pressure and temperature transducers  145  Table A.4  Gaseous emissions analyzers  146  Table B.1  Controlled Parameters for Test Series VI  147  Table B.2  Test Matrix for Test Series VIII-B at 800 RPM  150  Table B.3  ANOVA for Test Series VIII: fixed load/changing speed  159  Table B.4  ANOVA for Test Series VIII: fixed speed/changing load  166  Table C.l  Specific measurements contribution to Airflow Uncertainty (Carbon Balance)  171  Table C. 2  Specific measurements contribution to Airflow Uncertainty (Venturi)  172  Table E.1  Appendix E.1  Vu-A tests at 800 RPM  198  Table E.2  Appendix E.2  Vu-A tests at 1200 RPM  200  Table E.3  Appendix E.3  Vu-B tests at 800 RPM  202  Table E.4  Appendix E.4  VII-B tests at 1200 RPM  205  Table E.5  Appendix E.5  VIII-A tests  208  Table E.6  Appendix E.6  VIII-B tests  211  Table E.7  Appendix E.7  VIII— B2 Tests  215  —  —  —  —  —  —  —  viii  List ofFigures Figure 1.1  HPDI injector schematic  Figure 2.1  Injection delay, physical delay, and chemical delay for a typical HRR curve  6  16 Figure 2.2  Injector nozzle schematic for HPDI injector operation and HPDI co-injector operation  17  Figure 2.3  Injection sequence for normal injection operation  19  Figure 2.4  PM emissions for J36 and Co-injector Prototype A at 75% load and 1100 RPM (Jones 2006)  Figure 2.5  NOx emissions for J36 and Co-injector Prototype A at 75% load and 1100 RPM (Jones 2006)  Figure 2.6  21  Movie stills of Prototype B with 2 MPa bias, 1.0 ms diesel pulse width, 1.95 gas pulse width (Marr 2007)  Figure 2.7  21  23  Four flow regimes expected in HPDI co-injector: a) bubbly flow, b) plug flow, c) annular flow, and d) dispersed flow  26  Figure 3.1  Installed SCRE in CERC looking Northeast from the entrance  31  Figure 3.2  Test cell setup schematic  32  Figure 3.3  Diesel and natural gas process diagram with bias control loop using needle valves  34  Figure 3.4  Process diagram for combustion air, facility air, and cooling water  36  Figure 3.5  Data acquisition flow diagram  43  Figure 3.6  Geometry of an HPDI co- injector nozzle  45  ix  Figure 3.7  Commanded injection operation for the Westport controller and the FPGA controller  Figure 3.8  Figure 3.9  46  Sample indicated pressure curves for Prototype B for two different GPWs (1200 RPM, 24 MPa diesel rail pressure)  49  HRR curves for In-cylinder pressures shown in Figure 3.8  52  Figure 3.10 Comparison of ignition delay calculation methods  54  Figure 3.11 FFT of pressure data from Figure 3.8 for the high-knocking case (0.70 ms pilot GPW)  56  Figure 3.12 Knock Intensity (maximum amplitude of difference between filtered and unfiltered pressure signal) for 45 cycles, pressures from Figure 3.8  57  Figure 3.13 Knock Intensity Plotted vs. Maximum Rate of Pressure Rise (max dP/dCA). Test Series VII defined in Chapter 4  58  Figure 3.14 Distribution of maximum rate of pressure rise for Test Series VII-1200 RPM  61  Figure 3.15 Distribution of maximum rate of pressure rise for Test Series VII-800 RPM  61  Figure 3.16 Reduction of engine knock by reducing diesel injection mass (Prototype B)  62  Figure 3.17 Reduction of engine knock by reducing diesel injection mass and pilot GPW (Prototype B)  62  Figure 3.18 Comparison of in-cylinder pressure curves for SCRE setup in CERC (2008) and Kaiser (2005) for J36-008. 1200 RPM, 8 bar GIMEP, 0.40 EQR, 90 kPa MAP, 1.OmsRIT, 16.7:1 CR  63  Figure 3.19 Comparison of heat release rates for SCRE setup in CERC (2008) and Kaiser (2005) for J36-008. 1200 RPM, 8 bar GIMEP, 0.40 EQR, 90 kPa MAP, 1.0 ms RIT, 16.7:1 CR, Heated Intake Air  64  x  Figure 3.20 Comparison of in-cylinder pressure for SCRE setup in CERC (2008) and Kaiser (2005) for J36-008. Unheated Intake Air  66  Figure 3.21 Comparison of in-cylinder pressure for SCRE setup in CERC (2008) and Kaiser (2005) for J36-008. Unheated Intake Air  66  Figure 4.1  Gas injection mass as measured at Westport Innovations in BTR2  73  Figure 4.2  Gas injection mass as measured by the gas/diesel flowbench (EFS1) at Westport Innovations  Figure 4.3  Changes in CNG injection mass with increased DPW at different manifold pressures. 800 RPM, 18 MPa injection pressure, 0.75 ms GPW  Figure 4.4  74  77  Comparison of Gas injection mass of Prototype A and Prototype B measured in the SCRE  79  Figure 4.5  Ignition delay and COV GIMEP vs. gas/diesel volume ratio (Single Pulse)  80  Figure 4.6  CO and NOx vs. gas/diesel volume ratio (Single Pulse)  81  Figure 4.7  4 and nmHC vs. gas/diesel volume ratio (Single Pulse) CR  81  Figure 4.8  Representative HRR (Filtered) and IHR curves (Test Series VII-A -4, 800 RPM 22.0 mg/inj)  Figure 4.9  86  Comparison of HRR curves for different relative injection timing. Double injection tests at a low diesel injection mass (VII-A-29)  90  Figure 4.10 Unfiltered HRR curves for Prototype A at 24 MPa Diesel Rail Pressure (Vu-A 29 and VII-A-30)  92  Figure 4.11 Unfiltered HRR curves for Prototype B at 24 MPa Diesel Rail Pressure (Vu-B 29 and VuI-B-30)  93  xi  Figure 4.12 Ignition Delay comparisons between Prototype A and Prototype B at 1200 RPM, 24 MPa Diesel Rail Pressure  94  Figure 4.13 Knock intensity comparisons between Prototype A and Prototype B at 1200 RPM, 24 MPa diesel rail pressure  95  Figure 4.14 Ratio of heat released during the Pilot Combustion Event for Prototype A and Prototype B at 1200 RPM, 24 MPa diesel rail pressure  97  Figure 4.15 a) ignition delay, b) combustion duration at low load/1100 RPM  104  Figure 4.16 a) COV GIMEP, b) Knock Intensity at low load/i 100 RPM  105  Figure 4.17 a) CO, b) NOx at low load/i 100 RPM  106  Figure 4.18 a) CH4, b)tHC at low load/i 100 RPM  107  Figure 4.19 a) ignition delay and b) combustion duration at Low Speed/High Load Mode  108  Figure 4.20 a) COV GIMEP and b) Knock Intensity at Low Speed/High Load Mode  109  Figure 4.21 a) CO and b) NOx at Low Speed/High Load Mode  110  Figure 4.22 a) CH 4 and b) tHC at Low Speed/High Load Mode  111  Figure 4.23 a) ignition delay and b) combustion duration at High Speed/High Load Mode 112 Figure 4.24 a) COV GIMEP and b) Knock Intensity at High Speed/High Load Mode  113  Figure 4.25 a) CO and b) NOx at High Speed/High Load Mode  114  Figure 4.26 a) CH 4 and b) tHC at High Speed/High Load Mode  115  Figure 4.27 Ignition Delay and combustion duration for load/speed timing sweeps  119  Figure 4.28 COV GIMEP and knock intensity for load/speed timing sweeps  120  Figure 4.29 Knock intensity/ignition delay tradeoff curve  121  Figure B.i  Ignition delay for single injection vs. double injection  149  Figure B.2  Ignition Delay for Vu-A and Vu-B tests at 800 RPM  152  xii  Figure B.3  IHR ratio for Vu-A and Vu-B tests at 800 RPM  153  Figure B.4  Knock Intensity for Vu-A and Vu-B tests at 800 RPM  154  Figure B.5  Ignition delay and COV GIMEP for 13 bar GIMEP for a) 1100 RPM and b) 1400RPM  161  Figure B.6  4 and uHC emissions for 13 bar GIMEP for a)1 100 RPM and b) 1400 RPM 162 CH  Figure B.7  NOx and CO emissions for 13 bar GIMEP for a)1 100 RPM and b)1400 RPM  163  Figure B.8  Ignition delay and COV GIMEP for 6 bar GIMEP and 1100 RPM  167  Figure B.9  4 and uHC emissions for 6 bar GIMEP and 1100 RPM CH  167  Figure B.10 NOx and CO emissions for 6 bar GIMEP for 1100 RPM  167  Figure C.1  170  Comparison of Airflow Calculations in Kaiser and CERC  Figures F. 1.1 to F.7. 104: Pressure and heat release rate curves for double injection operation  224  xlii  Nom en cia ture 2GEOI 2GPW 2GSOI 2RIT A/D ANOVA ATDC BDC Bsfc BTDC CA CERC CF CE 4 CI CLD CO C02 COV CR DAQ DEOI Df DIR DPW DSOI EGR EQR ESC 13 FID FPGA GEOI GHG GIMEP GLR GLVR GPW GSOI  2nd Gas End of Injection 2nd Gas Pulse Width 2nd Gas Start of Injection 2nd RIT (end of pilot to start of main) Analog to digital Analysis of Variance After Top Dead Centre Bottom Dead Centre brake specific fuel consumption Before Top Dead Centre Crank angle Clean Energy Research Centre wet-to-dry conversion factor methane Compression Ignition Chemiluminescent Detector Carbon Monoxide Carbon Dioxide Co-efficient of Variation Compression Ratio Data Acquisition : Diesel end of injection : degrees of freedom : Diesel Return : Diesel Pulse Width : Diesel Start of Injection : Exhaust Gas Recirculation : Equivalence Ratio 13 mode European Steady Cycle Test : Flame Ionization Detector : Field-Programmable Gate-Array Gas End of Injection Greenhouse Gas Gross Indicated Mean Effective Pressure : Gas-to-liquid Ratio (mass basis) : Gas-to-liquid Volume Ratio : Gas Pulse Width : Gas Start of Injection xiv  HC HCCI HPDI HRR IHR IVC MAT MCE MS NDIR NIMEP nmHC NO NOx NV oCA OCS P PCE PIDING PM R RIT SCRE SI SOC SOL SS T TDC tHC UBC uHC V WP  : : :  Hydrocarbon Homogenous Charge Compression Ignition High-Pressure Direct-Injection Heat Release Rate Integrated Heat Release Inlet Valve Closing Manifold Air Temperature Main Combustion Event\ Mean of Squares Non-dispersive infrared Net Mean Effective Pressure non-methane Hydrocarbons Nitrogen Monoxide Oxides of Nitrogen Needle Valve degrees Crank Angle Orbital Combustion System Pressure Pilot Combustion Event Pilot Ignited Direct Injection Natural Gas Particulate Matter Gas Constant (for air) Relative injection timing (end of pre-inj ection to start of pilot) Single Cylinder Research Engine Spark-ignited Start of Combustion Solenoid Sum of Squares Temperature Top dead centre total hydrocarbons University of British Columbia Unburned Hydrocarbons Volume Westport  xv  Acknowledgements I would like first like to acknowledge the mentorship of my supervisor, Dr. Steve Rogak, for expertly provided direction in helping me develop my skills as a researcher, and patiently helping me hone and polish this thesis. This work could not have been completed without the excellent resources and experimental facilities at the University of British Columbia. I would like to thank especially Bob Parry for his endless work in maintaining the research engine and long hours in helping me with testing. Also, thanks to Gord Wright in his expertise and work in helping to re-install the electrical and control systems for the engine. Thanks to the previous researchers at UBC without whose work this thesis would not be possible. Thanks to Joey Mikawoz, Mike Marr, and Heather Jones.  In particular, thanks to Gordon McTaggart-Cowan for knowing  everything there is to know about the research engine, and for providing support and advice on several occasions. I would also like to acknowledge both the financial and technical support from Westport Innovations. Issues that I had been struggling with for days were resolved through short conversations with individuals such as Sandeep Munshi, Al Welch, and Mike Wickstone. Thank you to Dr. Phil Hill who always provided support and encouragement. I’d also like to thank Koyo Inokoshi and Mike Baker for their help with the co-injector.. My thanks to my many fellow graduate students and the help and camaraderie they provided. In particular thanks to Nick Berger, Edward Chan, Chris Laforet, Andrew Mezo, Wu Ning,  xvi  James Saunders, Malcolm Sheild, and Michael Yeung.  The collective experience and  knowledge they provided was a crucial in ensuring my research continued to progress. Lastly I would like to thank my parents for always encouraging me through my childhood and youth to strive to be the best I can and to follow my dreams. Lastly, thanks to my loving wife and eternal companion, Donna, for her unconditional love and support. Thank you, Donna, for providing an ear when it was needed, and for your loving care and support over throughout these last few years.  xvii  Chapter 1 Introduction -  Due to favorable fuel efficiency, power density, and reliability, diesel-fuelled compressionignition (CI) engines power an overwhelming majority of heavy-duty vehicle applications. Heavy-duty vehicles (gross weight> 3856 kg) are used in areas such as public transportation, commercial goods transportation, construction, and waste disposal.  Due the significant  impacts of diesel engine exhaust on air quality, as well as rising petroleum prices, there is great interest in engine emissions and fuel economy. The use of natural gas as an alternative to petroleum derived diesel is also being investigated.  1.1  Current Issues Facing CI Engines  In terms of air quality and health one of the pollutants of most concern in diesel engines are oxides of nitrogen (NOx). NOx consists mainly of two components: nitrogen oxide (NO), ) (Seinfeld and Pandis 2006). Increased levels of NOx in ambient 2 and nitrogen dioxide (NO  air cause irritation to the eyes, nose, mouth and lungs and lowers resistance to respiratory infection (US EPA 2008). However, NOx by themselves are of little concern. The “safe” levels of NOx as outlined by the US national ambient air quality standard (US NAAQS) is rarely exceeded in US and Canadian cities (Ontario MOE 2001).  Secondary reactions  involving NOx and unburned hydrocarbons (uHC), however, have contributed to increased levels of ground-level ozone and photochemical smog. For persons with existing respiratory issues, high levels of ozone have been shown to increase the hospitalization rate due to damage to the lung tissue. Atmospheric quantities of ozone as low as 80 parts per billion (80 ppb) have been shown to reduce lung function and increase susceptibility to respiratory  1  infections (US EPA 2008). In a study conducted on Canadian and international cities, 18 of 27 cities exceeded the one hour US NAAQS of 120 ppb (Ontario MOE 2001). Particulate matter (PM) is another emission from the diesel-fuelled CI engines. From diesel fuelled engines, PM consists mostly of solids with some adsorbed organic compounds ) in diameter are able to 25 (Heywood 1988, 627). PM which is less than 2.5 microns (PM enter deep into the respiratory tract, agitating the lungs or entering directly into the blood 25 have been linked to increased heart attacks, asthma stream. Short term exposures to PM attacks and acute bronchitis. Countries around the world have regulations to reduce the level of pollutants emitted from For example, Table 1.1 shows North American required  heavy-duty diesel engines.  reductions between 1988 to 2010 of uHC, NOx, and PM. The most significant pollutant reductions in the last 20 years have been the 2007-2010 emission standards. If engine manufacturers do not meet the required emissions standards then they must pay increasing non-conformance penalties which will either force the engine manufacturers to fix the non compliant engines or stop distribution (US EPA 2002a). 25 standards has been difficult With conventional diesel engines, meeting both NOx and PM due to the well-known NOx  —  PM tradeoff (Heywood 1988, 866). In order to meet the 2010  standards, exhaust aftertreatment devices will need to be installed. Ceramic particulate filters have been added to reduce the level of PM.  Three-way catalytic converters used in  stoichiometric spark-ignited (SI) engines cannot be used in lean burning diesel engines.  2  Table 1.1: g/bhp-hr  Exhaust emissions standards for heavy-duty engines in the United States,  (Dieselnet  n.d.)  Year 1988 1990 1991 1994 1998 2004 2007 2010  uHC* 1.3 1.3 1.3 1.3 1.3 0.5 0.14** 0.14  NOx 10.7 6 5 5 4 2.5 0.2** 0.2  PM 0.6 0.6 0.25 0.1 0.1 0.1 0.01 0.01  Non-methane hydrocarbons only **Half of the engine sales must meet 2010 Emissions regulations and remainder must meet 2004 standards *  Therefore, in order to reduce NOx, either a lean NOx trap or a NOx scrubber must be used. Two-way catalytic converters are used to simultaneously reduce CO and uHC emissions. However, all of these emission-control devices are expected to increase fuel consumption by 1-3%, add around $4000 to the cost of the engine, and add significant future maintenance and replacement costs to the engine (Schubert and Fable 2005; US EPA 2002b). There is increasing pressure to reduce greenhouse gas (GHG) emissions from the transportation sector.  Although CO 2 emissions from heavy-duty engines are not yet  regulated, soot (the black body carbon component may contribute to global warming) is controlled through the health-motivated standards. Methane is a powerful GHG and it is regulated in Europe in natural gas engines to 1.1 glkWh (Dieselnet n.d.). Due to interactions between supply and demand, the price of petroleum-based fuel has been steadily increasing over the last two decades (EIA 2007). The total fuel costs over the life of the heavy-duty vehicle is one of the more significant life-cycle costs of the vehicle (Schubert and Fable 2001). In the past, global fuel prices have risen in response to wars, political  3  instability, natural disasters, or trade embargoes (ETA 2007). Demand is outpacing supply, which is contributing to higher fuel prices. Governments are therefore looking at ways to distribute risk through the use of non-petroleum based fuels for medium duty and heavy duty applications. Different fuels are being investigated and subsidized by governments both to reduce greenhouse gas emissions (GHGs) and provide viable alternatives to petroleumderived diesel.  For example in the US the “Energy Policy Act” was implemented in 2005 in  order to reduce dependence on foreign oil and reduce greenhouse gases (US DOl 2005). It provides up to $32,000 in tax credits for purchasing heavy-duty natural gas, propane, or hybrid electric vehicles (US IRS 2005).  1.2  Natural Gas Use for Commercial Vehicles  Natural gas is a leading alternative fuel that can be used to simultaneously address climate change and energy demand issues from the use of petroleum-derived diesel in medium and heavy-duty vehicles. The main constituent, methane, has the lowest carbon-to-hydrogen ratio of any organic compound. Engines running primarily on natural gas have been shown to emit significantly lower GHGs provided that methane emissions are low (McTaggart Cowan 2006a). Throughout the world, sources of natural gas are more distributed than petroleum (Radler 2006) and presently natural gas prices are lower than diesel. In addition, methane can be considered a renewable resource since it can be produced through anaerobic digestion of waste.  4  One of the ways to efficiently run heavy-duty engines on natural gas is to install a high pressure direct-injection (HPDI’) system developed by Westport Innovations Inc. For an HPDI engine, a small amount of diesel is required to initiate combustion of the natural gas jet. Not only is most of the diesel replaced (—95% by energy content) with natural gas, the engine-out NOx emissions have been shown to be reduced by 40% and the PM by 70%, without compromising performance or efficiency (Dumitrescu et at. 2000; Harrington et at. 2002). Figure 1.1 shows the HPDI injector used to inject both diesel and natural gas. The inner (diesel) and outer (gas) injection systems are controlled separately by different solenoids, with the diesel fuel also acting as a hydraulic fluid in order to lift the injector needles. Gas and diesel can then be separately and independently injected through concentric injection systems. This injector is usually installed in an unmodified diesel engine. To date, Westport has successfully installed its HPDI technology in off-road mining trucks near Queensland, Australia, transport trucks in Ontario, and shipping trucks in ports throughout California. In the San Francisco Bay area for example, HPDI fuelling systems have been successfully installed in 13 refuse-hauling trucks. Over 10 million km have been logged by this fleet (Westport Innovations 2008a).  1  “HPDI” isa trademark of Westport Power Inc.  5  Diesel In Gas In  Diesel Return  Diesel Needle  Gas Needle  Gas Needle Solenoid Actuator  Figure 1.1: HPDI injector schematic  A new prototype HPDI injector has been devised where the pilot injection consists of a small amount of the pilot diesel co-injected with natural gas. In co-injection, gas-blast atomization is used to inject both diesel and natural gas through the same injection holes. Depending on how and when the diesel is introduced into the gas/diesel reservoir of the injector, the atomization process and subsequent combustion will be affected.  This process is also  affected by the number of injections per cycle. This “co-injector” design has potential to significantly improve on the simplicity and thus the cost of the HPDI injector. Since the diesel is entrained into the combustion chamber by the natural gas, the diesel injection system may not be needed. Significant material and manufacturing cost reductions may result. However, further study is required to characterize and optimize the injector.  6  1.3  Objectives and Scope  Ultimately, one needs to know whether the co-injector has similar or better performance than the current industry standard—the J36 Westport HPDI injector. However, the J36 (and associated control strategies) has been optimized over more than 15 years of research and development. This is the first thesis on this type of co-injector and it has not been previously established which parameters (ignition delay, combustion variability, knock intensity, etc.) set the boundaries of the operable region of the injector. Therefore, the goal of this thesis is to determine the parameters important to operation with only prelininary comparisons between the HPDI-J36 and the co-injector. In addition, since the gas and diesel are injected into the combustion chamber as a mixture, another objective of this research is to better understand the gas/diesel interactions. These objectives are attained mostly through experiments conducted in a heavy-duty four stroke single-cylinder research engine with undiluted charge air during single and double gas injection operation as will be explained in Section 2.3. Also introduced in 3.1.7 are the two different co-injector geometries used in this study: Prototype A and Prototype B. During this research, the test engine had to be moved. As a result, thousands of hours were spent reconnecting, redesigning, and testing the control and measurement systems.  An  additional objective of this thesis is to document the major changes to the usability and performance of the facility, particularly those that might affect comparisons between old and new measurements.  7  1.4  Thesis Structure  Chapter 1 provides background and motivation for studying HPDI co-injection and outlines the research objectives. Chapter 2 discusses prior work on two-phase injection and injection systems. Chapter 3 discusses the research engine as well as the injector flowbenches used for this research. Chapter 4 is divided into five sections which outline the testing procedure, single injection flow tests, double/single injection comparisons, and double injection emissions tests. Finally, a summary of the significant findings as well as recommendations for future work are discussed in Chapter 5.  8  Chapter 2 Background -  Numerous studies have been conducted to determine the benefits of using different natural gas engines such as stoichiometric, lean burn, or natural gas direct-injection for heavy-duty vehicles. Research on liquid/gas co-injection, however, has been limited. This chapter briefly describes the differences between types of heavy-duty natural gas engines and summarizes work that contributes to better understanding of the processes involved in two phase direct injection for internal combustion engines.  2.1  Current Natural Gas Technologies for Medium-Duty and HeavyDuty Engines  Ideally, natural gas would replace the diesel in order to avoid the extra costs of an additional fuelling system. However, in order for natural gas to auto-ignite, the temperature needs to be over 1100-1200 K which would necessitate compression ratios exceeding 23:1 (Aggarwal and Assanis 1998).  The high temperature needed would adversely affect the engine  performance and emissions. Therefore, ignition is usually assisted through the use of a spark, a hot surface or a pilot injection. Currently, stoichiometric spark-ignited (SI) engines, lean burn SI engines, and pilot-ignited lean burn engines are the most prevalent in industry and will be the only alternatives to HPDI discussed.  9  2.1.1 Stoichiometric Spark-Ignited Natural Gas Engines Stoichiometric spark-ignited natural gas engines can most economically meet the 2010 emissions standards (Chiu et al. 2007). Stoichiometric engines operate without excess fuel or oxygen, so the resulting exhaust emissions of CO and uHC are oxidized and the NOx are reduced over the three-way catalyst. The vehicle-out emissions can therefore be substantially reduced with commercially-proven exhaust treatment technologies. Premixed natural gas engines have exceptionally low emissions of PM (Faiz et al. 1996).  The carbon-based  particulate emissions from pre-mixed combustion are derived mostly from the engine oil (Faiz et at. 1996; Heywood 1988). Because emission control technology is already well developed for these engines, the life cycle costs of stoichiometric SI natural gas engines have been found to be similar to if not better than projected life-cycle costs of a similar sized diesel-fuelled engine that require exhaust gas treatments (Schubert and Fable 2005). Stoichiometric natural gas engines, however, have significant drawbacks that have prevented them from being more widely adopted. Since the air-fuel mixture needs to be kept relatively close to stoichiometric, the charge air needs to be metered as well as the fuel. At part load, the charge air is throttled, introducing significant pumping losses. Duggal et al. (2004) measured the performance of a diesel heavy-duty engine on the 13-mode European Steady Cycle (ESC 13), which can be compared with the measurements by Chiu et at. (2004) for a stoichiometric natural gas engine of similar size. Averaged over the 13 modes, the natural gas engine had 25% higher brake specific fuel consumption.  10  Current stoichiometric SI engines have significantly lower power and torque capabilities compared to a diesel fuelled engine. This is due, in part to lower volumetric efficiency due to the gaseous fuel displacing combustion air that could have been inducted into the combustion chamber. More importantly, there are design constraints that limit the maximum compression ratios to about 11:1 (Faiz et at. 1996; Chiu et at. 2004; Zhang et a!. 1998). The factors that limit the maximum power and torque are engine knock and excessive exhaust Lower maximum temperatures are  temperatures (above 973 °C) (Zhang et at. 1998).  important to reduce mechanical and thermal wear to the engine components such as the gaskets, exhaust valves, cylinder heads, and turbochargers.  Similarly, excessive and  prolonged engine knock breaks down thermal boundary layers which can cause severe damage to the piston, piston rings, etc (Taylor 1985, 39; Heywood 1988, 456). Current production stoichiometric natural gas engines are limited to about 261 kW (350 hp) and 895— 1356 nm (660—1000 lb-ft) of torque (Westport Innovations 2008b; Cummins n.d.), limiting their use to applications such as transit, and medium-duty vehicle applications. 2.1.2 Lean-Burn Spark-Ignited Natural Gas Engines The torque and efficiency limits of stoichiometric SI engines can be addressed through the use of a lean combustion. Prior to increased restrictions on pollutant emissions, most natural gas-thelled engines were lean burning engines.  In lean burning engines, the excess  combustion air lowers the exhaust temperature which reduces the engine-out NOx emissions and minimizes engine damage from thermal cycling.  Slightly higher power and torque  characteristics can be attained through the use of pre-mixed lean burning of natural gas since the maximum power and performance of the engine is not dependent on the maximum  11  cylinder temperature. Previous lean bum technologies were 10  —  20 percent more efficient  than stoichiometric engines (Faiz et al. 1996), although still lower than the efficiency of a diesel engine. Unlike stoichiometric engines, throttling is not necessary for lean burn SI engines at most operating points since these engines can operate stably at a wide range of equivalence ratios. Only at low load, where there were large cycle-to-cycle variation is throttling necessary to avoid bulk extinction during the expansion stroke. Since a leaner mixture is used, lower combustion temperatures allow higher compression ratios before the onset of knock. The biggest drawback for lean-bum engines is that the exhaust treatment technologies are relatively immature. Lean combustion reduces engine-out NOx, but not enough to meet recent heavy-duty engine NOx regulations in many applications. Excess oxygen prevents NOx from being reduced on a three-way catalyst. In addition, compared to stoichiometric natural gas engines, lean burn engines have higher levels of unburned fuel, most particularly methane. Methane emissions are important for two reasons. First, methane is a significant greenhouse gas which by mass is 21 times more 2 (Foster et al. 2007). Second, the catalytic powerful at warming the atmosphere than CO conversion efficiency for CH 4 is strongly dependent on temperature.  Low exhaust  temperatures common to lean bum engines (200 °C to 400 °C) lead to conversion efficiencies around 10  —  15% (Duggal et al. 2004).  12  2.1.3 Lean-Burn Pilot-Ignited Natural Gas Engines Lean bum pilot-ignited natural gas engines are commonly referred to as “dual-fuel engines” (Srinivasan et a!. 2006; Taylor 1985). Natural gas is introduced into the combustion chamber so that it will form a homogenous mixture with the combustion air. Instead of using a spark, a small amount of diesel (about 20% on an energy basis) is used to ignite the mixture (Srinivasan et a!. 2006).  Therefore, similar to SI engines the part load efficiency and  maximum torque are limited.  Potentially, more diesel could be used at these operating  points; however, the engine-out emissions of NOx are observed to increase with increased diesel injection mass (Srinivasan et al. 2006).  2.2  High-Pressure Direct-Injection  High-pressure direct-injection (HPDI) natural gas engines provide diesel-like performance, reliability, and efficiency for both two stroke engines (Hodgins et a!. 1996; Douville et a!. 1998; Harrington et a!. 2002) and four stroke engines (Dumitrescu et a!. 2000; Duggal et al. 2004). Since HPDI engines inject all of the fuel near the end of the compression stroke, most of the fuel bums in a turbulent diffusion flame. Therefore, higher load capacities can be achieved since HPDI engines are not limited by the onset of knock. Compared to SI natural gas engines, higher compression ratios can be used (around 15-19:1 rather than around 11:1) and thus higher thermal efficiencies are observed in HPDI engines. The load is controlled by metering the fuel only; therefore, diesel-like efficiencies at part load are achievable. While the performance and efficiency are similar to a diesel engine, the NOx, PM emissions are significantly lower. Duggal et al. (2004) summarized the performance of the HPDI  13  fuelling system in a Cummins ISX engine modified with a smaller compressor and intercooler. They reported that the installed HPDI system was able reduce NOx by 40% and the PM by 80%. Goudie et al. (2005) determined that at extremely high EGR rates (40% EGR) with an oxidation catalyst, the PM and NOx emissions were 0.36 g/bhp-hr and 0.04 g/bhp-hr respectively for an ESC 13 mode test cycle. Similar to lean bum natural gas SI engines the exhaust aftertreatment options are immature and expensive. However, engine-out NOx are much closer to the 2010 emissions standards and therefore fewer exhaust treatments may be necessary. For example, with higher levels of EGR, the NOx emissions may be met and the PM can be filtered using a particulate filter (Williams 2007). 2.2.1 Ignition Delay in HPDI Engines  For any direct-injection compression-ignition engine, there is a time lag between the introduction of fuel into the chamber and combustion. When a cool liquid jet is introduced into the turbulent high-temperature high-pressure environment, numerous things happen before it bums. Cavitation and turbulence in the liquid jet cause it to disintegrate into smaller droplets (Adomeit et al. 2002; Rotondi et al. 2001) Aerodynamic forces will also cause droplet breakup as Kelvin-Helmholtz (sinusoidal) instabilities overcome surface tension and liquid viscosity (Rotondi et al. 2001; Lörcher and Mewes 2001). Due to the higher ambient temperature the droplets heat up and begin to evaporate. The gas from the surface diffuses outward and mixes with the surroundings to form a combustible mixture. Due to the high temperature and pressure, radicals then start to form in the mixture. Exothermic reactions occur, leading to exponentially more exothermic reactions.  14  The time from the introduction of the droplet to ignition is referred to as the ignition delay. It is composed of two parts: the physical delay and chemical delay. Physical Delay is the time it takes for the fuel to establish an ignitable mixture (Teng et al. 2003) consisting of physical phenomenon such as mixing, heating, and evaporation. Chemical delay is harder to predict due to the hundreds of possible chemical reactions that can take place simultaneously. Chemical delay is often calculated from empirical Arhennius-type relations:  Tjg =  A.exp(E/RT).[Fuel]’ .[Oxygen]’  (2.1)  Where A, a, and b are constants found experimentally (e.g. Shock tube, combustion bomb). Chemical delay for diesel fuel under normal operating conditions is less than 0.67 ms (Teng et al. 2003). Compared to the physical delay, chemical delay is usually considered to be much shorter since physical delay includes the slow processes of heat and mass transfer and evaporation (Teng et al. 2003; Sazhina 1999). The chemical and physical delay cannot simply be added up as chemical reactions can take place as the droplet is evaporating and mixing. In addition to physical and chemical delay, there is a noticeable delay after the commanded pilot injection till the injector needle lifts. This is referred to as injection delay. injection delay has been found to be around 0.5  —  This  0.7 ms for the J36 series HPDI injectors  used for this work (Kostka 2008). Figure 2.1 shows the ignition delay broken down into injection delay, physical delay, and chemical delay.  15  150 —2OmgIinj-O.47ms Chemical Delay  1 00  ::  InJectio:DelaY\  50  0  Physical Delay  J’  --% Ignition  Commanded Injection  -50 -30  -20  -10  0  10  20  30  40  Crank Angle (deg) Figure 2.1: Injection delay, physical delay, and chemical delay for a typical heat releas rate (HRR) curve  2.3  Co-Injection  For the Westport HPDI injector, natural gas and diesel are separately injected into the combustion chamber as shown in Figure 2.2a. The HPDI co-injector was constructed from a J36-03 build HPDI injector by modifying the inner diesel injection system so that diesel is injected into the gas reservoir instead of directly into the combustion chamber, as shown in Figure 2.2b. This was done by plugging the needle tip and drilling holes through the gas needle. A more detailed description of the modifications needed to make Prototype A (the original co-injector concept) and Prototype B (Prototype A with an added sleeve in to change the geometry of the gas/diesel mixing chamber) can be found in Section 3.1.7.  16  Natural Gas Needle  Diesel Needle  L  Figure 2.2: Injector nozzle schematic for HPDI injector operation and HPDI co-injector operation  The gas and diesel injections are also shown in Figure 2.2. Whereas the diesel and gas are injected separately with the HPDI injector, a gas/diesel two-phase mixture is injected with the HPDI co-injector.  2.3.1 Co-injector Operation For the co-injector, three injections are needed for normal operation: the diesel pre-injection, the pilot gas injection, and the main gas injection. Figure 2.3 shows the three injections. During the pre-injection, the diesel injection needle lifts and diesel is injected into the common gas/diesel reservoir. For the J36-HPDI injector, the diesel injection is referred to as the pilot injection, since the diesel is directly injected into the combustion chamber and acts as the ignition source. For the HPDI co-injector, however, the pilot injection refers to the  mixture of diesel and natural gas and is injected as the gas needle lifts. Shortly after the pilot injection, the gas needle lifts again and the main charge of natural gas is injected.  17  Commanded Diesel Commanded Gas  1. Pre-Injection Diesel needle lifts and diesel is injected into the common gas/diesel reservoir 2. Diesel needle re-seats.  3. Pilot Injection Gas needle lifts and gas/ diesel mixture is injected into the combustion chamber 4. Gas needle re-seats and there is still diesel in the injector (hypothesis to be checked).  5. Main Injection Gas needle lifts and gas/ diesel mixture is injected into combustion chamber. 6. Gas needle reseats and there is very little diesel in the injector.  Figure 2.3: Injection sequence for normal double injection operation  18  For some cases at low load and low engine speed the co-injector may operate with a single gas injection. For this case, all of the diesel injected into the gas/diesel reservoir is injected during the pilot gas injection, unlike normal double gas injection operation where the injected diesel can potentially be divided between the two gas injections due to the gas and diesel mixing in the gas/diesel reservoir and liquid diesel sticking to the walls. Prototype A (the original co-injector concept shown in Figures 2.2b and 2.3) has the following potential advantages over the Westport HPDI injector. Firstly, since the diesel is co-injected with the natural gas, there will be a natural gas jet for every diesel jet. With previous HPDI injectors the exhaust emissions were observed to cycle every two minutes, supposedly due to the diesel holes in the gas needle changing position as the gas needle rotated. HPDI injectors therefore have more diesel holes than required to ensure stable combustion, which implies that some of the gas jets are not optimally aligned with the pilot sprays (Dumitrescu et at. 2000; Ouellette et at. 1998). Secondly, there will be overlap between the diesel and gas jets.  McTaggart-Cowan found  for the HPDI injector that when the gas was injected before or very shortly after the diesel injection significant reductions in PM were observed. He speculated that auto-ignition of a diesel-natural gas mixture may have occurred which led to substantial PM reductions which remained low independent of the EGR level (McTaggart-Cowan 2006a; McTaggart-Cowan et al. 2003). Finally and most importantly, significant system cost reductions are potentially available with future models of the HPDI co-injector.  Since diesel is injected into the gas/diesel  reservoir instead of directly into the engine, the diesel injection system could potentially be 19  simplified substantially.  The major costs of the HPDI injector include machining the  injectors to tight tolerances to prevent diesel and gas leaks into the cylinder from around the injector needles and to allow separate passages for the diesel and natural gas. The cost reduction from one Less injector needle, actuator, and injector driver per injector approaches the 50% cost reduction goal of this project, especially when the machining cost reductions are included.  The design simplification would also allow a reduction in injector size,  possibly making HPDI co-injection viable for light duty applications. 2.3.2 Previous Work at UBC on Co-injection No previous work has been published on the operation of the HPDI co-injector prior to 2007. Engine tests by Jones and McTaggart-Cowan were distributed only as internal fact-sheets with very basic descriptions of the tests conducted and analysis of the results. Some of the key findings are described below. In November 2005, Jones completed two test sets with Prototype A with single gas injections (Jones 2005a; Jones 2005b). More diesel was required to run Prototype A stably than the J36  4 and injector, especially for starting the engine. Also, the co-injector produced higher CH CO emissions at low load. One of the most striking results of the single-injection tests was the presence of “ringing” (periodic pressure fluctuations over 3 bar—discussed in Section 3.2.5) at high diesel injection masses. Although the ignition delay was only slightly longer  than for a typical HPDI injector, the energy injection rate from the co-injector was high since large quantities of fuel were injected prior to ignition. The solution proposed was to have a short “pilot” injection, followed by a “main” injection, as shown in Figure 2.3.  20  Later, Jones (2006) compared double injection tests of Prototype A with normal operation of the J36 injector at mid speed/low load, mid speed/high load, and high speed/high load which are similar to the ESC 13 test modes #7, #6, and #4 respectively (Dieselnet n.d.). Jones found at high loads that the HPDI co-injector had lower PM, less fuel consumption, and similar NOx, tHC, and CO emissions. Figures 2.4 and 2.5 show the PM and NOx for one case.  J36-0%EGR,piIotl 5mg/inj 0 121-0%EGR,pilotl 5mglinj  0.25  • J36-30%EGR,pilot=l5mg/inj • 121-30%EGR,pilot=1 5mglinj  J36-30%EGR,pilot=7mglinj  0.20  5  0.15  a. 0.05 0.00 4  5  6  8 9 10 11 12 50% IHR (deg ATDC)  7  13  14  15  16  Figure 2.4: PM emissions for J36 and Co-injector Prototype A at 75% load and 1100 RPM (Jones 2006) 10  9 8  —7  e4  3 z 2  0 4  5  6  7  8  9 10 11 12 50% IHR (deg ATDC)  13  14  15  16  Figure 2.5: NOx emissions for J36 and Co-injector Prototype A at 75% load and 1100 RPM (Jones 2006)  21  These figures show significant improvements in the PM emissions for the HPDI co-injector compared with the original HPDI injector at high engine loads without significantly increasing the NOx. Jones also found that PM emissions could be reduced in the J36 HPDI injector by reducing the diesel injection mass.  However, lower injection masses for  Prototype A could not be tested at this point as diesel injection masses lower than 12 mg/inj led to misfiring of the engine. In addition, Jones found the tHC and CO emissions were significantly higher for the co-injector at low loads.  In continuation of the work done by Jones, McTaggart-Cowan (2006b) performed single injection tests to determine the best method for reducing the high levels of uHC at low loads. McTaggart-Cowan found that the high cycle-to-cycle variation can be substantially reduced by increasing the diesel flow rate, increasing the intake air pressure, or reducing the gas injection pressure. He suggested that the higher combustion variability and higher uHC emissions were due to a lower diesel/gas volume ratio by volume injected.  He found,  however, that the transition from single injection to double injection operation was sensitive to operating condition and that the introduction of the main injection had the potential to stop combustion due to the pilot injection. Further details of the test conditions and results from McTaggart-Cowan can be found in Chapter 4, since they are closely related to tests performed for this thesis.  The full factsheets compiled by Jones (2006) and McTaggart  Cowan (2006b) can be found in Appendix D. Optical studies of the co-injector Prototypes have also been conducted on Prototype A by Mikawoz (2005) and for Prototype B by Marr (2007).  From the movies recorded of  Prototype A by Mikawoz (2005), most of the Viscor (a replacement for diesel used for 22  injector calibration) appeared to be finely atomized and injected during the beginning of the injection. The work by Marr (2007) on Prototype B supported this observation as he found that the majority of the liquid is injected near the start of the injection. Figure 2.7 shows movie stills collected by Marr (2007).  SUms  4:25 ins  7Sms  4i0  400 s  475  jc  Figure 2.6: Movie stills of Prototype B with 2 MPa bias, 1.0 ms diesel pulse width, 1.95 gas pulse width (Marr 2007)  The diesel (shown as the dark jet) is injected around 0.5 ms after the start of injection. The time of maximum diesel flux was found to be relatively independent of bias pressure, diesel pulse width, and gas pulse widths (GPW5). From these preliminary tests, Marr also found  23  that for Prototype B, GPWs less than 1.5 ms restricted the Viscor injection volumes and that the co-injector seemed to reach steacy state almost immediately.  2.3.3 Patents and Studies on Gas/Liquid Co-injection Injecting a gas/liquid mixture in internal combustion engines is not new. The first use of air to help atomize diesel was in 1893 with the original diesel engine (Stone 1999, 9). It wasn’t until 1910 that it was replaced with the high pressure liquid jet injectors used today. Currently, there are patents for improving atomization in SI engines (Kimmel and Dillon 2002) and for CI engines (Tarr et al. 1999) using gas-assist atomization for the liquid fuel. Fundamental studies conducted with these injectors are described below. Other patents have been disclosed which use natural gas as the primary fuel as well as the atomizing fluid for the igniter. For example, Hill et al. (1991) describe the use of natural gas to continuously atomize diesel using a pre-chamber. Similarly, Yang (2002) describes a dual fuel injector that uses natural gas to bring the diesel into the combustion chamber at injection pressures between 1.5  —  4.0 MPa. These devices sound similar in operation and purpose to the HPDI  co-injector (patent pending); however, peer-reviewed studies related to direct-injection natural gas engines with entrained diesel cannot be found. For SI engine applications, Orbital Engine Company produces an air-assisted direct fuel injection system referred to in the Orbital Combustion System (OCS). The OCS is used both for stratified charge and homogenous charge SI engines (Boretti et al. 2001). The OCS is similar to the HPDI Coinjector in that fuel is metered into the mixing chamber using an injector. In the case of the OCS, a conventional pencil stream port fuel injector is used 24  (Boretti et al. 2001). Cathcart and Zavier (2000) report the mass of fuel injected into the combustion chamber changes with time, and is dependent on the delay time between the pre injection event and the direct injection event with a maximum flux around 1 ms with an injection pressure of 6.5 bar and a gasoline  —  air bias of 0.7 bar. For the OCS, the gas/liquid  mass ratio (GLR) is 2 to 0.2 going from low load to high load (Houston and Cathcart 1998). Since the gas used for the HPDI co-injection is also a fuel, the GLR can be between 0.5  2.5  (McTaggart-Cowan 2006b). For low diesel injection masses the co-injector GLR can be as high as 3.5. Although the OCS has similarities to the HPDI co-injector, there are fundamental differences between the two. Compared to diesel direct injection, fuel injection into the combustion chamber begins much earlier in the compression stroke (about 80°  —  150° BTDC) in order to  allow a stratified charge to form before the spark event. Relatively low injection pressures are therefore needed for the OCS (on the order of 6.5 bar) (Borretti et at. 2007; Houston and Cathcart 1998).  Higher injection pressures are needed for direct-injection compression-  ignition engines to inject the fuel near the end of the compression stroke. Perhaps most importantly, air is used as the atomizing gas instead of natural gas. Fundamental studies have been published for a wide range of injection pressures, including diesel relevant injection pressures and conditions, using gas/liquid injection processes referred to as “effervescent atomization”. The gas injected into the liquid is done in order to reduce the injection pressure needed to produce small liquid droplets. Therefore, work with effervescent atomization has concentrated on low GLRs. Sovani et al. (2001b) found that previous studies on effervescent atomization with diesel or a diesel substitute were conducted  25  at a GLR between 0  —  0.3. Still, the fundamental studies conducted by these workers are  beneficial in attempting to understand the injection processes for the HPDI co-injector. Roesler and Lefebvre (1988), Lörcher and Mewes (2001), and Chin and Lefebvre (1993) studied the internal flows of an effervescent atomizer. There are four flow regimes reported that may be applicable to the flow in the HPDI co-injector, namely, bubbly flow, plug flow, annular flow and dispersed flow. These are shown in Figure 2.6. Chin and Lefebvre (1993) reported that as the injection pressure increased, the range of the bubbly flow regime was extended to higher GLRs. For their tests at injection pressures of 8 bar, they reported that at GLRs greater than 0.4 the liquid (water) was completely broken up by the gas (air) and was dispersed as droplets in the atomizing gas. For a non-homogenous mixture of gas and liquid in the mixing reservoir, the flow will transition from one flow regime to the next, depending on the distribution of liquid.  a  c  Figure 2.7: Four flow regimes expected in HPDI co-injector: a) bubbly flow, b) plug flow, c) annular flow, and d) dispersed flow  Liquid/gas injection has potential to reduce the droplet diameter in three ways. First, in two phase flow the speed of sound is lower (Sherstyuk 2000; Sovani et al. 2001) than for a pure  26  gas injection. This means that flow chokes at a much lower velocity, and therefore there will be a steep pressure jump across the minimum flow area.  This steep pressure drop is  beneficial in increasing atomization quality (Sovani et al. 2001). Good atomization of the fluid can result, even if there are large exit orifices, low injection rates, or low injection pressures (Sovani eta!. 2001). Second, two-phase flow can effectively reduce the size of the orifice for the liquid. This can be seen in Figure 2.6c where the liquid is pushed to the outside of the orifice wall for annular flow. Finally, as the gas expands after the orifice, the rapidly expanding gas core will break the annular flow into smaller ligaments which will then form smaller droplets (Sovani et al. 2001). In summary, gas-blast atomization is not new, and has been used before in direct injection engine applications. However, there have been no peer reviewed papers on its use as a way to deliver pilot diesel in a natural gas direct-injection engine. The preliminary work of Jones (2005a; 2005b; 2006) and McTaggart-Cowan (2006b) show that the relationship between the diesel injection mass and ignition is complex and requires further research. These studies concluded that double gas injection operation was required for most operating conditions. In comparison with the J36 over a range of operating conditions the co-injector exhibited lower PM emissions, especially when EGR was used. However, it was also found that work was needed in order to lower the CH 4 and CO emissions at low load as well as to reduce the dependence of operating condition on the diesel injection mass. The objectives first explained in Section 1.3 can now be developed into the following four research objectives in order to forward the work done by McTaggart-Cowan and Jones:  27  1. For single-injection operation, determine how the gas injection mass changes in response to changes in the commanded gas injection duration and injected diesel mass. 2. For single-injection operation, determine the effect the relative amounts of gas and diesel have on exhaust emissions, combustion variability, and ignition delay. 3. For double-injection operation, observe how a second gas injection affects the amount of diesel injected into the combustion chamber during the first injection. 4. For double-injection operation, determine how injector geometry affects the injector operation and quantify the effect injector geometry has on emissions, combustion variability, knock, and ignition delay. By addressing these four research objectives, the importance of the gas/diesel interactions on engine exhaust emissions and combustion variability can be better understood.  28  Chapter 3- Apparatus and Procedures 3.1  Single-Cylinder Research Engine (SCRE)  The test engine used for this study is derived from a 400 hp 6-cylinder Cummins ISX engine modified to operate with one firing cylinder (2.5 L displacement, 137 mm bore, 169 mm stroke, 261.5 mm connecting rod).  The other five working pistons were replaced with  drilled-through pistons. On the deactivated cylinders, the intake and exhaust valves were bolted shut, and the rocker arms were removed (McTaggart-Cowan 2006a). The SCRE can use several different pistons and air inlet systems.  Due to scheduling  constraints, several series of tests compare two injector variants for the 16.7:1 compression ratio (CR) enforcer piston. Several other test series compare the injector variants using a 15:1 CR piston insert with swirl plates at the intake valve. The SCRE was run in two different locations. The tests in 2007-2008 were conducted after the engine had been moved from Kaiser 1180 to the Clean Energy Research Centre (CERC). The test cell setup prior to 2007 is described extensively by McTaggart-Cowan et at. (2004), McTaggart-Cowan (2006a), and Jones (2004). Since the modifications to the engine may have an impact on the operation of the engine, the engine setup in CERC will be described in detail, with any important changes from the previous system noted. The engine speed is controlled by a General Electric eddy-current water-cooled dynamometer connected to the engine through a flexible spider coupling. At low loads, a Baldor 30 kW electric ‘vector’ drive motor will assist in overcoming the frictional losses of the non-firing cylinders. In Kaiser 1180 the vector drive was attached to the dynamometer using a belt 29  drive. Due to frequent belt failures, a flexible spider coupling connects the two in the CERC test cell. The engine coolant thermostat has been bolted open to allow continuous flow of coolant through the engine. The cooling water is fed to the cooling tower through a flow control valve in order to control the coolant temperature from 77 to 80 °C. 3.1.1 Test Cell  In CERC, the SCRE has been mounted in a large temperature-controlled test cell as shown in Figure 3.1. This test cell is much larger than the previous test cell in Kaiser. Figure 3.2 shows a plan view of the components inside of the SCRE test cell with a summary of the components listed in Table 3.1. The CERC installation allows the operator to monitor and operate the engine in a much safer manner than in Kaiser. A large shatterproof window allows the operator to safely monitor most of the components seated beside the Data Acquisition (DAQ)/Control Computer inside the engine control room.  During operation the test cell is ventilated at a rate of 55 air  changes per hour resulting in cell temperatures between 10 and 25°C as air is drawn across the test cell from the cell intake air duct to the air evacuation air duct (Veco 2004). The test cell CH 4 and CO detectors are integrated into the building shutdown controls.  30  B  C  D  °  t.J  —  C  ‘-  C  -  1j  .  C  h•  ;B  Ij  -  B BD  .  E  rM D rM  B  C  .  -  0’. (J  -  0  —  .  C  -  -  1 D  C  tD  —  C  C  CD0  -  0  z  O  O  s  ,  0  B  (,  .  a  0  -  — — — — —  B  -  B•  ci  C  ,J ) ) “0 00  D  :i)  -  C C C  c  =)_  .  ————  t  —  ———t—-  B  D  -  CA)  3.1.2 Fuel Supply System Figure 3.3 shows the flow diagram for the fuel supply system. The method of pressurizing and circulating the diesel is the same described by McTaggart-Cowan (2006a). The bias pressure (diesel  —  gas rail pressure) is usually set to around 0.80 MPa for the J36-HPDI  injector in order to ensure that the gas does not leak into the diesel. This is accomplished through the use of a dome-loaded self-venting regulator, PCV-NG-500, which ensures constant bias pressure between the diesel and gas. For co-injector testing, the bias pressure needs to be much greater in order to inject 10-20 mg of diesel during the 5 ms maximum needle lift duration.  In the tests conducted in 2006 and 2007 (in both test cells), the  additional bias was created through the use of a high-pressure regulator installed on the natural gas line downstream of the dome-loaded regulator.  However, the pressure  fluctuations caused by gas injection may have been causing poor performance and/or deterioration of the regulator. Therefore, for the tests conducted in 2008, the high-pressure regulator was removed and the bias pressure at the injector was controlled by lowering the diesel pressure at the dome loaded regulator through the use of needle valves, (needle valves NV-DIR-620 and NV-DIR-630 in Figure 3.3). The natural gas for CERC is supplied continuously at pressures up to 5000 psi from a dedicated gas line and is compressed by an integrated three-stage piston compressor. The Kaiser installation used separate multi-stage compression systems. In the event of an emergency or a rapid engine shutdown, two solenoid valves will shut off the gas supply to the test cell. These are shown as SOL-NG-400 and SOL-NG-401 in Figure 3.3.  33  r  SOLNG:io  m/t 1411  T ‘L  iZR0  -j  —  Facility  +L =-  Jjfntelli’-Faucet  ‘-rn  °‘  <L  I  I —  — - -  EVt;:fkv.5G  ----  Di  ACW-Di 140  2 mjcf  I [  r -,x  -  —  4  cNty Jr  Supply  O& P4tu,n  -  — —  _--—---  ___  Process Linetype Legend Nolurol Gas HP Deuel  rM  3.1.3 Air Supply System An oil-flooded screw-type compressor was used to supply the combustion air. A refrigerated dryer and low-pressure-drop filter were used to remove the water and oil from the intake air. For the tests in 2007 and 2008 automatic controls were installed for the intake air and exhaust back pressures.  The back pressure can also be controlled manually with a motorized  butterfly valve controlled through a ten-turn potentiometer.  The temperature of the  combustion air is controlled with a three-phase/20 amp/240V resistance heater to ±2 °C. A 90 L intake surge tank and an insulated 90 L insulated exhaust surge tank were used to dampen pressure fluctuations in the intake and exhaust lines. The surge tanks are located on a nearby platform. The pipe lengths and volumes between the surge tanks and the engine were similar in both installations.  The intake surge tank is vertical to allow water  condensation to drain from the system. Figure 3.4 shows the flow diagram for the air, exhaust, and cooling systems.  35  , ‘  I  -PjZJ  \ iLJ  C  ‘1  I: \;[\ I  *  -  rQ  L’N  /  A  L  L —  11  j  0  U.  :-  Iq h  t. Figure 3.4: Process diagram for combustion air, facility air, and cooling water  36  3.1.4 Emissions Measurement and Calculations , CO, CH 2 , uHC, and NON) 4 As with the previous system, the gaseous emissions (02, C0 measurements are taken downstream of the exhaust surge tank in order to ensure homogeneity in the exhaust stream. The exhaust passes through a heated line and filter to arrive at the AVL Emissions Bench, CEB II, which has limit monitoring and automatic calibration. Inside the emissions bench the exhaust is split into two branches: the wet measurements and the dry measurements. On the wet side (water not removed) the CH , 4 uHC, and NOx concentrations are measured. All other gases are measured as on the dry side. All emissions are measured according to SAE vehicle exhaust measurement standards (SAE 1993, 1995). Appendix A lists the stated accuracy and range for each analyzer. The uHC and CH 4 are measured using a Flame Ionizing Detector (FID). In the emission bench used in 2006, only the uHC was measured in this fashion. A hydrogen flame inside a constant electric field ionizes organic carbon to produce a current proportional to the amount of carbon present (Pierburg 2002a). A portion of the sample is passed through a thermo chemical converter which converts all non-methane hydrocarbons to CO 2 and water. The 4 concentration is measured through a second FID. The resulting currents are compared CH against the reference span gases of methane, and propane listed in Appendix A. During post processing, the propane-equivalent measurement of the uHC is converted to a methane equivalent measurement by dividing by 3 (the carbon number ratio for propane to methane). The NOx is measured using a chemiluminescent detector (CLD) which measures the light intensity of NO burning with ozone. To measure the N0 concentration, NO 2 is first reduced NO using a thermo-catalytic converter. During the oxidation process, light is generated  37  between 600 and 1200 nm. Low absolute pressures are used to increase the probability of producing light and reduce the cross sensitivity from other components (Pierburg 2002b). The NO is multiplied by the K-NO correction factor which is used since the amount of NO formed in combustion is dependent on the humidity of the inlet air (SAE International 1995). The remaining constituents need to be measured with the water removed. The amount of water in the exhaust (used for calculating the “wet” concentrations of  02,  , CO) is 2 C0  calculated assuming complete combustion of the fuel in air, minus the uHC, which is usually negligible. The following approach can then be used in converting the dry measurements to wet measurements (SAE 1995), starting with the stoichiometry,  CH  =  2 nO  +  )+mi-1 —> CO n(3.76N 0 2 2 ++mJH O + x0 2 2 +n(3.76 N ) (3.1) 2  In this equation, the variables y, n, and m, and x represent the atomic hydrogen-to-carbon ratio of the gas/diesel injection, the moles of oxygen in air to the engine, the moles of water in the combustion air, and the moles of excess oxygen (SAE 1995).  —  O.5y+(7.63x10 h)n-2tHC 1 (4.76+7.63x10 h)n+O.25y  In this equation, h is the specific humidity expressed in terms of gH2o/kgd 1 airS The conversion factor (CF) to convert the dry values is therefore CF=1—W  (3.3)  Oxygen concentrations are measured using the paramagnetic properties of the gas (02 becomes magnetized when under an external magnetic field). The instrument consists of an 38  oxygen free gas enclosed in a dumbbell shaped body under a non-uniform magnetic field. The oxygen will migrate towards the magnetic field at one side of the dumbbell and the resulting higher pressure will cause the dumbbell to rotate. The voltage needed to keep the dumbbell horizontal is proportional to the oxygen concentration (ABB Automation 2001). The interference factor can be calculated by Equation 3.4 (SAE 1993). Interference  =  28.8x%NOxO.O1+O.623x%CO x 2 O.O1  (3.4)  Although other gases such as CO 2 and CO are weakly paramagnetic, and NO are diamagnetic (repelled by a magnetic field), the interference for the worst case (high CO 2 low NOX) for this study was less than 0.03% (SAE 1993). CO and CO 2 are measured with Non-Dispersive Infrared absorption (NDIR) instrumentation. Non-elemental gases will absorb discrete bands of infrared energy. The frequency of light absorbed depends directly on the type of gas. A light emitter of known frequencies and amplitudes goes through the sample gas and light is absorbed. Constant pressure columns of the reference gases are located at the other end which converts light absorption into volume change of a diaphragm (ABB Automation 2000). At the beginning of each day that testing occurred, the emissions analyzers were re-calibrated using zero and span calibration gases. At the end of the day, the calibrations were checked to determine whether the calibration of the analyzers had changed. In January 2008, problems were noted in the uHC measurements that eventually led to a complete servicing of the emissions bench. It was believed that this servicing did not affect any of the tests. This was checked by repeating an entire test series in June 2008.  39  Note that the old emissions bench in Kaiser was not frequently checked for linearization, nor did it have pressure and flow checks to ensure proper operation of the analyzers. After comparing repeatability points for the Kaiser and CERC installations, it was found that the two emissions systems might be significantly different (see Section 3.3). 3.1.5 Engine Speed, Temperature, and Flow Measurement In both the CERC and Kaiser installations, Hall-Effect sensors are installed on the crank, cam, and dynamometer shaft in order to measure engine speed and position. The crank and cam sensor signals are conditioned and amplified at the sensor and sent to the controller. The dynamometer shaft sensor signal is sent to the Digalog Dynamometer Controller where it is conditioned, amplified, and used for engine speed control. The fuel, intake and exhaust pressures are measured with strain gauge diaphragm pressure transducers. These transducers were re-calibrated when the engine was moved to the new test cell. Temperatures are measured with K-type thermocouples. Appendix B gives the instrument list and the expected accuracy of each updated from what was reported by McTaggart-Cowan (2006a). The diesel fuel is kept in a small recirculation tank which was refilled as needed. The diesel flow rate is calculated gravimetrically by determining the change of the diesel mass in the recirculation tank over a sample time of 120 s or more (pail-and-scale). The natural gas mass flow rate is measured using a Micromotion Coriolis effect mass flow meter. A UBC-built venturi meter is used for measuring airflow.  With the current calibration,  however, there is an offset in the carbon mass balance. This offset was assumed to be due to  40  a lower air-flow rate than expected. This may be caused by unresolved leaks in the intake air system as well as errors in the calibration of the venturi. Therefore, as described in Appendix C, the air flow rate was determined through the use of the carbon balance. 3.1.6 Engine Control, Monitoring, and Data Acquisition  In the test cell at CERC, most of the control and monitoring of the SCRE takes place in  engine control room. A field-programmable gate-array (FPGA) is installed in the control computer with the ability to send and receive both analog and digital signals (NI 7831 -R). A simplified information flow diagram is shown in Figure 3.5. The review of the shutdown logic and the operating procedures was a significant part of this thesis work over the winter of 2007/2008. Detailed process and instrument drawings and fault scenarios were prepared for approximately 6 hours of review meetings with Westport technical staff.  The final  operating procedures and Labview control logic are the result of this process. Details are given in the electronic appendix of this thesis (...rogak!sbrown/Thesis/Brown_Thesis.zip). The information flow diagram shows data flow through the sensors, connection panels and multiplexers to the control computer. The multiplexers combine, amplif’ (for thermocouple measurements), and condition several analog signals for transmission through a single cable. The signals are converted to digital signals through a 12 bit Analog-to-Digital (A/D) card in the computer (PCI-MIO-16E-1).  The SCXI 1001 chassis collects either “slow-speed”  temperature and voltage signals at about 1 Hz, or “high-speed” voltage signals every V 2 degree crank angle (°CA).  41  Included on the FPGA board are -1OV to 1OV digital-to-analog converters. Since the FPGA operates at a clock speed of up to 40 MHz, it can be used for the high-speed control of fuel injection, a function previously taken care of by a Westport controller board. Similarly, the intake air selection, intake air pressure, exhaust back pressure, and coolant temperature are controlled by digital and analog control signals from the FPGA. The remaining controls needed for engine operation are controlled manually through the control panel or regulators and valves in the test cell. The intake air heater, motorized back pressure valve, and engine speed and load are controlled at the control panel. Only control of the diesel pressure, diesel-gas bias pressure and intake venturi pressure require the regulator/valve to be manually opened or closed. The control panel, the Labview control program and FPGA control program include safety logic. The integrated safety system is capable of monitoring the temperatures and pressures important to the functionality and safety of the test cell, warning the operator of unsafe conditions, and shutting down the engine as shutdown limits are reached. The shut down levels are set by the user. During a shutdown, the control computer decides that a shutdown is necessary and sends a ‘software shutdown’ signal to the FPGA. The FPGA then decides that a shutdown is necessary and sets all of the actuators to their default positions and sends a ‘shutdown output’ signal to shut down the remaining actuators. Similarly, a shutdown signal can be invoked through the FPGA board, the control panel, loss of power, or cooling water.  42  XI  U, U,  0 0  = 0 C-)  S  0 0  0  Ir  U,  =  U)  0  C,)  43  3.1.7 HPDI Co-Injector Operation The co-injector injects diesel into the common gas/diesel reservoir through 7 holes of As the diesel needle lifts, diesel is injected into the gas/diesel  0.17 mm diameter.  reservoir where it mixes with the gas. The amount of diesel injected depends on the bias pressure between the diesel and gas rail pressures and the injection duration. The gas diesel mixture is then injected into the combustion chamber during the pilot injection. The pilot injection usually lasts between 0.47 and 0.7 ms. Finally, in 0.3 to 1 ms after the pilot injection the main injection occurs. For high load applications, most of the gas is injected during the main injection with injection durations ranging from 0.8  —  1.1 ms.  Based on measured diesel flow rates and the commanded diesel needle opening time, the velocity of the fluid entering the gas/diesel reservoir ranges from 10 to 80 m/s. Based on the measured diesel  —  gas bias pressure and Bernoulli’s equation, the maximum velocity  should range from 45 to 88 m/s.  Both estimation methods are crude and serve only to  show that the diesel may move a significant distance inside the injector before the gas needle opens. While the engine was being moved, the HPDI co-injector prototype was also being modified. In an attempt to improve the design of Prototype A (the original HPDI co injector concept), the internal geometry was changed by adding a sleeve to create Prototype B. Figure 3.6 shows the modification that was made. The injector sleeve reduces the inner reservoir volume by 33%. It decreases the minimum annular area in the injector from 30 2 mm to 10 mm , resulting in three times higher fluid velocity. In an 2 attempt to keep the gas/diesel mixture near the injector tip, the annular area expands to 30 2 near the diesel holes. This volume is about 35 mm mm , enough for 30 mg of diesel. 3 44  Injector Sleeve to make Prototype B Diesel Gas Holes Pilot Plug  Figure 3.6: Geometry of a HPDI co-injector nozzle  3.1.8 Injection Command Parameters  The Westport Controller used in Kaiser was replaced with injection control using the FPGA board. For the Westport (WP) Controller, the commanded injections were based on timing of the commanded diesel end-of-injection (DEOI) to top dead centre (TDC), and the gas timings relative to end of the previous injection (Figure 3.7) TDC was calculated based on the 2 missing teeth in the crank signal at 6O0 after top dead centre (ATDC). The end of the pilot injection was then timed to end at a specified time before  45  TDC. The commanded first gas injection was then commanded to occur at a specified time after the end of the diesel pre-injection, referred to as the relative injection timing (RIT). Similarly, the commanded second gas was injected a short time after the end of the first injection, referred to as the second RIT (2RIT).  -60 0ATDC Missing CrankTooth  WP2GPW WPGPW (ms) (ms) •———-——bI* WP2RIT WPRITI (ms) (ms)  TDC  FPGA 2GPW (ms)  I  I  FPGA GPW (ms)  I  FPGA2GSOI(deg>  PGA GPW (ms) 4  I  °  FPGA GSOI (deg)  j  FPGA DSOI (deg)  1 WP DPW (ms)  I WP DEOI (ms) FPGA °TDC WP  tTDC  =  60 (deg):  100001RPM (ms)  a  Legend: FPGA Controller (CERC) Westport Controller (Kaiser) Black Line Control Parameter in timE Grey Line Control Parameter In deg  —  —  —  Figure 3.7: Commanded injection operation for the Westport Controller and FPGA controller The FPGA control logic was based on absolute injection angles instead of the timing of the commanded pulse widths; therefore, the absolute injection angle is controlled. Whereas the original controller calculated the time until TDC from the two missing teeth 46  on the engine crank signal at 6O0 ATDC, the FPGA system reset the counters at 6O0 ATDC and then compared the crank angle to the commanded gas/diesel crank angle. An optical encoder attached to the flywheel provides ¼ degree resolution as a comparator. Because this comparison is reset at 6O0 ATDC, commanded injections that overlap this point will not operate properly. The diesel start of injection (DSOI), gas start of injection (GSOI), and second gas start of injection (2GSOI) are specified in crank angle degrees after top dead centre (ATDC) of the power stroke. The diesel pulse width (DPW), gas pulse width (GPW), and second gas pulse width (2GPW) are specified in milliseconds (ms). The differences between the Kaiser and CERC control systems make exact comparisons of repeatability points difficult. For the emission tests discussed in Section 4.5, the required injection angle to obtain a specified RIT for the first commanded gas injection is calculated and input into the Labview control program. Several methods were used to check the alignment of the rotational encoders and injection command timing, as described in Wi -FAC-3788-ANYS (Appendix D).  3.2  Cylinder Pressure Measurements and Analysis  The cylinder pressure is used to determine heat release rates, indicated power, ignition delay, combustion variability, and knock intensity. These are the main parameters used to characterize combustion, so they warrant careful discussion.  47  3.2.1 Equipment Description An AVE water-cooled QC33C piezoelectric transducer measures the in-cylinder pressure and a charge amplifier converts the signal from the transducer to a voltage. Because the piezoelectric transducer measures the change in pressure, the absolute pressure in the cylinder is modified so that it matches the pressure measured in the intake manifold near the time the intake valve closes (-1 80°ATDC). The intake manifold pressure is measured every V 2 degree crank angle by a high speed strain-gauge pressure transducer. This procedure was used also in the Kaiser installation. Piezoelectric pressure transducers should have factory-calibrated charge/pressure conversion factors, so that knowing the charge amplifier characteristics, no further calibration is needed. The practice in Kaiser (McTaggart-Cowan, 2008) was to choose the bar/volts factor in order to reconcile motoring curve behaviour with calibration experiments in a small constant volume chamber. Whether or not this procedure was optimal did not affect previous results, which were all done for a consistent pressure measurement procedure. In January of 2008, it was necessary to replace the old charge amplifier (Kistler 503) with a model 504D Kistler charge amplifier. The overall bar/volt conversion was set to 2.850 bar/V to reconcile pressure traces from Kaiser and CERC for the same operating condition (previously set to 3.866 bar/V in Kaiser). Section 3.3 discusses this further. In April 2008, the charge amplifier was replaced with the AVL model F1exIFEM. The QC33C piezoelectric pressure transducer was also replaced with a new QC33C pressure  48  transducer. This time the pressure/voltage conversion was set to 2.000 bar/V, consistent with the factory settings.  The bar/V conversion factor was confirmed in the small  constant volume chamber. The impact of these changes on data comparisons is discussed in Section 3.3. Figure 3.8 shows a two representative indicated pressure vs. crank angle plots obtained in this study.  In Section 3.2.5 the pressure fluctuations are analyzed.  120 100 z 0 0  —20 mg/inj -0.70 ms, 20 mg/inj -0.47 ms  80 60 40  ‘20 0 -60 -50 -40 -30 -20 -10 0 10 20 30 40 50 60 Crank Angle [deg]  Figure 3.8: Sample indicated pressure curves for prototype B for two different pilot GPWs (1200 RPM, 24 MPa diesel rail pressure)  3.2.2 Indicated Mean Effective Pressure (IMEP) and Engine Variability  The indicated pressure curves can be used to determine the amount of work output from the engine. The net indicated mean effective pressure (NIMEP) is a measure of the indicated work output per unit of swept volume and can be expressed as the cyclic integral of work (PdV) for each of the four strokes (Sonntag et al. 2003).  PdV NIEP  PdV =  =  swept volume  comp,expansion,exhausl,intake  (3.5)  swept volume  49  Small pressure differences at low cylinder pressures during the intake and exhaust strokes led to higher uncertainties in the pressure measurements during these strokes, which reduced the confidence in the NIMEP. For this reason, the gross indicated mean effective pressure (GIMEP), which takes into account only the compression and expansion stroke was used. GIMEP has been previously used in the SCRE (McTaggart-Cowan 2006; McTaggart-Cowan et al. 2006; Jones 2004, 2006) as well as by other workers (Boretti et al.2007; Cathcart and Zavier 2000; Cairns et a!. 2006) as a measure for defining the engine output. —180  (3.6)  fPdV GIMEP =  180  swept volume  Since the pressure is measured every V 2 °CA, the integral becomes a summation from bottom dead centre (BDC) of the compression stroke to BDC of the expansion stroke. Due to the encoder offset, the pressure is not recorded at BDC and therefore the volume 0 at BDC is computed from the swept and clearance volumes, as shown in the following V equation, assuming the pressure at BDC is the same as the first measured pressure, P . 0 (P GIMEP=  +  0 P  (v,  —  (Vswept  +  Vciearance))+  Pk + k=2  Vswept  2  k-l  (Vk  —  ) 1 Vk (3.7)  The coefficient of variation (COV; standard deviation/mean) of the mean effective pressure has been used widely for determination of the cyclic variability of the engine (Duggal et al. 2004; Cathcart and Zavier 2000; Boretti et a!. 2007; McTaggart-Cowan et al. 2006; Zhang et a!. 1998; McTaggart-Cowan 2006). The maximum acceptable COV  50  IMEP is usually between 3-6% (Zhang et at. 1998; Cathcart and Zavier 2000; Duggal et al. 2004).  3.2.3 Heat Release Rate  The in-cylinder pressure as a function of the engine crank angle can be used to determine the heat release rate (HRR). The heat release rate is an approximation of the amount of heat that would need to be added (due to the release of chemical energy) to the combustion cylinder to observe the measured in-cylinder pressure (Stone 1999, 547).  HRR during the power stroke is based on an air standard cycle which has the following assumptions (not considering the pumping work) (Sonntag et al. 2003, 410): •  A fixed mass of air is the working fluid through the entire cycle, and the air is always an ideal gas. Thus there are no inlet process and exhaust process.  •  The combustion process is replaced by a process transferring heat from an external source.  •  Air has a constant specific heat.  With these in mind, the HRR curve is usually only computed from -180 ° to 180 °ATDC. Using the First Law of Thermodynamics for a closed volume (after the inlet valve closes and before the exhaust valve opens), the net heat release can be written (Stone 1999, 548):  dQ net dO  where  ‘  =  YdV ‘y—l dO  +  _i_v.P ‘y—l  dO  (3.8)  is the constant specific heat ratio for the exhaust gas mixture (set at 1.30 for the  diesel process) (Heywood 1988, 510).  51  The crevice regions of the combustion chamber at TDC are non-negligible and can make up a few percent of the clearance volume. The gas the crevice is cooler and denser and has properties much different from those in the rest of the cylinder (Heywood 1988, 509). In some previous work, multi-zone HRR models which take into account heat transfer to the walls and combustion processes have been used (Hill and Douville 1997, Hill and McTaggart-Cowan 2005). For this study, the single zone heat release rate model was used since the more complicated versions of heat release are still approximations. Semiquantitative comparisons of ignition delay, 50% IHR, and knock can still be drawn from the simpler HRR model. Figure 3.9 shows the unfiltered HRR curves derived from the pressure data of Figure 3.8.  500 400 300 200 100  -100 -200 -20  -10  0  10  20  30  40  Crank Angle [deg]  Figure 3.9: HRR curves for in-cylinder pressures shown in Figure 3.8 Note that in Figure 3.9 that the pressure fluctuations of Figure 3.8 are exacerbated on the HRR curves. In order to better calculate performance metrics such as ignition delay, the high frequency pressure fluctuations were filtered out using a low-pass Gaussian digital  52  filter. As discussed in Section 3.2.5, the cutoff frequency was chosen so that it would filter out any frequencies higher than 3.5 kHz.  3.2.4 Ignition Delay  As discussed in Chapter 2, the ignition delay is measured from the start of commanded injection to the start of combustion (SOC). The start of combustion is difficult to define (Stone 1999, 549). In previous work the start of combustion was defined as the 5% IHR (Asad and Zhang 2008, McTaggart-Cowan 2006b), rapid pressure rise (Donghui et at. 2004), the point where the indicated pressure trace separates from the motoring pressure trace (Srinivasan et at. 2006, Dumitrescu et al. 2000). Stone recommends using the minimum of the IHR, or the first non-zero HRR (Stone 1999, 549) which is what was used by McTaggart-Cowan (2006a) and McTaggart Cowan et at. (2006). For this study, the start of combustion (shown in Figure 3.9) was found as the intercept of the HRR curve of the first combustion event with the zero HRR /deg and 20 3 axis. This intercept was calculated by finding the slope between 30 kJ/m /deg. 3 kJ/m Figure 3.10 shows the ignition delay as defined by the 5% IHR plotted against the ignition delay defined by the first non-zero HRR for Test Series VIII-B2 tests (tests are described later in Section 4.5). Note that both calculation methods are correlated, with shorter ignition delays calculated from the 5% IHR typically longer by about 1 ms; larger ignition delays are in agreement.  53  5.0 y=O.8021x+0.9124 4.0  0.00  1.00  2.00 3.00 Ignition Delay from HRR (ms)  4.00  5.00  Figure 3.10: Comparison of ignition delay calculation methods  3.2.5 Knock  Heywood (1988, 505) describes combustion in a compression-ignition engine as a rapid premixed burning phase followed by a controlled burning phase. During the testing process, there were specific operating conditions where significant fluctuations were observed on the indicated pressure trace. The observed pressure fluctuations are due in part to the rapid energy release and subsequent pressure rise, shock wave and wave reflection off of the chamber walls during the premixed burning phase (Heywood 1988, 454; Taylor 1985, 95).  The amount of engine noise and engine knock with the co  injector was greater in magnitude than that of a J36-HPDI injector where diesel and gas are injected separately. Although the exact cause of engine knock has not yet been ascertained, a possible cause could be a high initial injection rate of natural gas and diesel causing a larger proportion of fuel to burn during the premixed phase.  54  Knock intensity can be measured in two ways. First, the knock intensity can be attained by determining the maximum amplitude of the fluctuating portion of the indicated pressure curve (Heywood 1988, 455). Vibrations that are picked up by the pressure transducer are a combination of the structural vibration as a result of a rapid pressure rise and pressure waves in the cylinder (Christensen et al. 1998). The pressure fluctuations are due to the gas in the cylinder resonating at the first transverse mode acoustical frequency (Heywood 1988, 455).  The first transverse mode acoustical frequency is  defined as  (3.9)  where c is the speed of sound of the gas, a is a geometric constant (taken as 1.84 for this case), and D is the cylinder diameter (Eng 2002). The speed of sound will be dependent on the gas temperature. Given the pressure and volume data, the in-cylinder temperature can be approximated by the ideal gas law relation:  =2T 2 T 1  (3.10)  Where the T 1 is the intake manifold temperature, V 1 and Pi are the volume and pressure when the intake valve closes, and T ,V 2 , and P 2 2 are the temperature, volume and pressure right before ignition. The speed of sound of the compressed gas can then be calculated assuming a constant specific heat ratio by the relation c  =  where R is the gas  constant for air (287 JIkgK). For the test case, the first transverse mode frequency should  55  therefore be about 3.7 kHz, which is expected since the first mode frequency for engine cylinders is usually between 3  —  10  kHz (Heywood 1988).  Ideally, a bandpass filter set to the first transverse acoustical frequency would be put on the signal output with the maximum amplitude from each cycle representing the knock intensity (Heywood 1988, 455).  For this study, since there was no filter installed, the  frequency is found through a Fourier transformation of the raw in-cylinder pressure data to the frequency domain. Figure 3.11 shows that there was an experienced frequency between 3 and 4 kHz on a frequency domain plot.  ‘I’  Unfiltered Data Filtered Data  30 20  0  1000  2000  4000 3000 Frequency (Hz)  5000  6000  7000  Figure 3.11: FFT of pressure data from Figure 3.8 for the high-knocking case (0.70 ms pilot GPW)  The knock intensity can be obtained by finding the maximum difference between the filtered and unfiltered indicated pressure curve for each of the 45 recorded cycles. Figure 3.12 shows the maximum amplitude of the fluctuations occurring for the high and low  56  knocking cases of Figure 3.8. The boundary between “normal” and “heavy” knock is not well defined. For this study, knock intensities over 300 kPa are considered heavy knock, consistent with a description by Heywood (1988, 455). Note that the knock intensity for high diesel, long pilot GPW duration is well above 300 kPa (3 bar), signif’ing that intense knocking is occurring.  900 800 -  700  0 -  >  5 500 J)  o 300 0  200 100 0 0  5  10  15  20  25  30  35  40  45  Cycle Number Figure 3.12: Knock intensity (maximum amplitude of difference between filtered and unfiltered pressure signal) for 45 cycles, pressures from Figure 3.8  Another way to report the level of knocking is to report the maximum rate of pressure rise (max dP/dCA). Shiga et al. (1988) related the maximum rate of pressure rise of a diesel engine to the knock intensity. They found that knock intensity increased with the square of the maximum rate of pressure rise. Usually, a maximum rate of pressure rise between 6-10 bar/deg is considered the stress limit for diesel engines (Christensen et al.  1998; Obert 1973, 589).  57  Figure 3.13 shows the relationship between knock intensity and the maximum rate of pressure rise.  From these tests, there is appears to be a correlation between knock  intensity and maximum rate of pressure rise.  It appears that a 3 bar knock intensity  corresponds to a maximum rate of pressure rise of 6  -  12 bar/deg, consistent with the  recommendations for maximum rates of pressure rise made by Christensen et al. (1998) and Obert (1973).  8  x  7  x  (0  0  x x  O  Co 0  —3  x  .  C.,  02 C  x Vu-A oVII-B  0  -r  0  2  4  6  8  10  12  14  16  18  20  22  24  26  Max dPIdCA (barldeg) Figure 3.13: Knock intensity plotted vs. maximum rate of pressure rise (max dP/dCA). Test Series VII defined in Chapter 4.  The conventional detection method of engine knock, however, does not include information about the amount of energy released or the cylinder pressures and temperatures at the time of knock occurrence.  Engine geometry and the pressure at  which knock occurs are just as important as the knock intensity (Taylor 1985, 39; Fitton and Nates 1996). Specific cases of intense and prolonged engine knock without engine  58  damage has been previously reported (Christensen et al. 1998; Vressner et al. 2003, Taylor 1985, 39). Damage due to knock can occur as the thermal boundary layer around the piston and cylinder walls breaks down due to the pressure waves inside the cylinder (Stone 1999, 74). Aluminum pistons are especially susceptible to melting or holing during engine knock due to the lower melting temperature of aluminium. Piston rings and lands can be broken due to the pressure waves. When a piston ring fails, it will likely score the internal combustion engine’s piston liner. The high frequency mechanical vibrations can cause extra fatigue on engine components shortening their lives. Damage in spark ignited engines due to knock is much more prevalent, since it usually occurs near stoichiometric conditions. A large amount of energy is therefore released. Note that for all of the cases considered, the maximum cylinder pressure for the SCRE is always below 150 bar and the exhaust temperature is below 600 °C. There are no peer reviewed studies on damage due to knock in a pilot-ignited directinjection natural gas engine. The closest studies found for similar sized engines were with natural gas homogenous charge compression ignition (HCCI). Christensen et al. (1998) operated a similar sized engine, fuel, and boost pressure with significant knock occurring.  This engine was operated as a homogenous charge  compression ignition engine for short durations at 40 bar/deg and continuously at 15-20 bar/deg without any noticeable wear to the engine. Table 3.2 shows compares the geometries and temperatures of the engines. Note that for this study, the manifold temperatures were quite a bit lower. 59  Table 3.2: Engine Size Comparison Betveen Volvo TD100 (Christensen et al. 1998) and Cummins ISX (Duggal et al. 2004)  Volvo TD100 1.6 Displacement (L) 0.14 Stroke (m) 0.12 Bore(m) Connecting Rod (m) 0.26 Compression ratio 17.5 Boost Pressure (bar) 1 Manifold Temperature °C 100 -170 Fuel Tested Natural Gas Equivalence Ratio 0.33 0.4 —  Cummins ISX 2.5 0.17 0.137 0.26 16.7 1 30 Diesel/Natural Gas 0.3-0.55  It is important to note that the engine was running on a homogenous charge of natural gas at manifold temperatures higher than the controlled temperature of the SCRE in order to promote autoignition of the natural gas. This would lead to in-cylinder gas temperatures at TDC (without combustion) 350  —  450 K higher than the Cummins ISX. Christensen et  al. (1998) found the peak cylinder temperatures to be up to 2000 K. Vressner et al. (2003) attribute the lack of damage to the low temperatures near the wall due to lean combustion.  These peak temperatures would be similar in the SCRE at lower  equivalence ratios. Figures 3.14 and 3.15 show the proportion of Test Series VII cases (described in Chapter 4) that exceed the 10 bar/deg limit by Obert (1973) and the 20 bar/deg limit by Christensen et al.(1998). Less than 1% of all of the tests run exceeded 20 bar/deg. Both the 16.7:1 and 15:1 pistons were inspected after over 20 hours of operation with the co-injector, and there was no sign of damage to the pistons or rings. The valve train did not appear to be affected either. However, test conditions that exhibit heavy knock should  60  be avoided unless there are specific attributes of the co-injector that need to be observed at high diesel/high gas pilot injections.  80% 75% 70% 65% 60% 55% >< 50% 45% C.) 40% 35% • 30% 25% 20% 15% 10% 5% 0%  —S—A- 1200 RPM -a 8-1200RPM  U,’  21% 1% 6% 15  10  5  0  X  =  20  25  30  dP/dCA max (bar/deg)  Figure 3.14: Distribution of maximum rate of pressure rise for Test Series VII-1200 RPM 80% 75% 70% 65% 60% 55% >< 50% 45% o 40% 35% • 30% 25% 20% 15% 10% 5% 0% 0  5  10  X  =  15  20  25  30  dP/dCA max (bar/deg)  Figure 3.15: Distribution of maximum rate of pressure rise for Test Series VII-800 RPM  Lower diesel injection masses or shorter pilot gas injection durations can reduce the severity of diesel knock.  Figures 3.9, 3.16, and 3.17 show that the fluctuations can  effectively be reduced by reducing the pilot injection duration, reducing the diesel injection mass, or both. Lower diesel injection masses will prevent many cases which 61  would otherwise exhibit strong diesel knock. Chapter 4 results include more detail on the conditions that affect knock.  500 400 300 200 .  100  I  -100 -200 -20  -10  0  10  20  30  40  Crank Angle [deg] Figure 3.16: Reduction of Engine Knock by Reducing Diesel Injection Mass (Prototype B) 500 400  a) 200 E ioo ‘  -100 -200 -20  -10  0  10  20  30  40  Crank Angle [deg] Figure 3.17: Reduction of Knock Intensity by Reducing Diesel Injection Mass and Pilot GPW (Prototype B)  62  3.3  Performance Comparisons for CERC and Kaiser Tests  In order to determine the long-term repeatability of the engine, a repeatability test was conducted in the engine to begin each day at oil temperatures of 95-100 °C. This was especially important with the Westport J36-008 HPDI injector, which has been used in benchmark testing since 2005. The indicated pressure and HRR curves shown in Figures 3.18 and 3.19 compare the recorded combustion for repeatability testing in CERC and in Kaiser with the J36-008 injector. Unfortunately, the tests conducted in February-March 2008 were conducted with heated intake air to 70°C. These points are therefore compared against bracketed temperatures (57°C and 81 °C) collected in May 2006. As mentioned in Section 3.2.1, the in-cylinder pressure measurements in 2005 were calibrated against a small volume pressure chamber. This method could not be used in the new test cell due to some leakage between the small volume pressure chamber and the transducer. Therefore, to continue on with the testing, the voltage-to-pressure conversion factor for the tests in 2008 was set so that the GIMEP was similar to that in 2005. Figures 3.18 and 3.19 show that for this condition, the peak pressure location and magnitude, and the HRR curves were comparable. These similarities indicate that even though GIMEP is similar between the 2005 and 2008 tests by design, engine operation is similar between both test cell locations.  63  120 Average - Jan to Mar 2008-70 oC May 2005-55 oC ..  /.‘  Crank Angle (deg AT DC) Figure 3.18: Comparison of in-cylinder pressure curves for SCRE setup in CERC(2008) and Kaiser(2005) for J36-008. 1200 RPM, 8 bar GIMEP, 0.40 EQR, 90 kPa MAP, 1.0 ms RIT, 16.7:1 CR 200  Average - Jan to Mar 2008- 70 oC May 2005-55 oC May 2005-81 oC  160 120’ 80 40  ,1  0•  —  -  --e.  -  -40 -30  -20  -10  0  10  20  30  Crank Angle (deg AT DC) Figure 3.19: Comparison of Heat Release Rates for SCRE setup in CERC(2008) and Kaiser(2005) for J36-008. 1200 RPM, 8 bar GIMEP, 0.40 EQR, 90 kPa MAP, 1.0 ms RIT, 16.7:1 CR, Heated Intake Air  In June, 2008, further repeatability tests were conducted with the J36-008 with unheated air. These results were compared with the tests conducted in 2006 by Jones. Table 3.3  64  compares the CERC in June of 2008 measurements to those made at similar operating conditions in Kaiser in January of 2006. As seen from the table, the intake manifold temperature was slightly lower in the CERC setup, which would be expected from the larger, more effectively ventilated test cell. Similarly, the electrical systems in the CERC test cell have been laid out to reduce the electrical noise in the signals. This could potentially reduce the calculated knock intensity. Even with the controlled parameters in the CERC setup set to ±5% of the Kaiser setup, there were still significant differences in the emissions measurements. Most notably were the CO measurements where the Kaiser emissions were more than 100% higher than the CERC measurements.  Similarly, the CH 4 emissions were lower in the CERC setup and  the uHC were higher, which resulted in the non-methane hydrocarbons (nmHC) to be more than 100% lower in the Kaiser setup.  65  Table 3.3: Comparison of Performance Parameters Between CERC (2008) and Kaiser (2006) tests using J36-008 injector 3  Number of Tests  % Change  3  (2006 2008) /2008 x 100 -  Engine Speed (RPM) Gas Rail Pressure (MPa) Air Flow (kglhr) Exhaust Back Pressure (kPa) Diesel Injection Duration (ms) Equivalence Ratio GIMEP (bar) 50% IHR  -  w  Manifold air temperature ‘  u  .  .  Ui  .  Manifold air pressure (kPag) Disel injection mass (mg/mi) CNG injection mass mg/in.) Peak Pressure (bar) Knock Intensity (kPa) CO (g/kg of fuel) NOx (g/kg of fuel) CH4 (g/kg of fuel) HC (g/kg of fuel) C02 (kg/kg of fuel) 02 (kg/kg of fuel) Carbon balance ratio Nitrogen balance ratio Hydrogen_balance_ratio Oxygen balance ratio  2008 tests 1206 +1- 4 20.9 +1- 0.1 143.0 +1- 0.4  2006 tests 1210 +1- 1 21.4 +1- 0.2 142.8 +1- 0.9  79 +1- 7  76 +1- 7  0.65  0.3% 2.6% -0.2% -2.7%  0.65  0.394 +1- 0.007 8.33 +1- 0.27 10.4 +1- 0.3  0.408 +1- 0.004 8.48 +1- 0.03 10.0 +1- 0.1  3.7% 1.8% -4.0%  20.7 +1- 0.5  25.0 +1- 1.0  20%  60.9 +1- 0.1  59.4 +1- 2.8  -2.4%  8.2 +1- 0.9  8.2 +1- 1.1  0.7%  84.3 +1- 1.7  87.1 +1- 1.1  3.4%  86.5 +1- 1.8 108.2 +1- 6.4 2.21 32.51 2.33 4.30 2.63 5.10 1.051 1.00 1.00 0.984  +1+1+1+1+1+1+/-  0.10 0.68 0.27 0.45 0.01 0.15 0.004  +1- 0.001  92.1 145.5 4.51 26.13 3.13 2.94 2.44 4.99 0.977 1.00 1.00 0.991  +/-  0.63 0.41 0.15 0.02 0.04 0.04 0.019  6.5% 34.5% 104% -20% 34% -32% -7.1% -2.1% -7.1%  +/-  0.004  0.7%  +1- 0.8 +1- 7.4 +1÷1+1+1+/+/-  For these tests the AVE charge amplifier was used with a pressure/voltage conversion  factor of 2.000 bar/V. The indicated pressure and HRR in Figures 3.20and 3.21 compare the recorded combustion for similar operating points in CERC and in Kaiser for repeatability testing. Note that the measured in-cylinder pressures for the June-CERC tests were lower, probably due to pressure calibration errors in the Kaiser setup. It is likely  that the  uncertainty  associated with the  calibration  of the pressure  transducer/charge amplifier in the small constant-volume chamber was large.  66  100 90 80 70 D 60 50 jj 40 c 30 20 é 10 0 -40  -30  -20  -10  0  10  20  30  40  Crank Angle (deg ATDC) Figure 3.20: Comparison of In-Cylinder Pressure for SCRE setup in CERC(2008) and Kaiser(2005) for J36-008. Unheated Intake Air 180 —a--- Pilot 2006 -*—Gas 2006 *• Pilot 2008 ‘-s--Gas 2008 —Average 2008 Average 2006  160 140 120  100 .  80  60 4Q 20  0 -40  -30  -20  -10  0  10  20  30  40  Crank Angle (deg ATDC) Figure 3.21: Comparison of In-Cylinder Pressure for SCRE setup in CERC(2008) and Kaiser(2005) for J36-008. Unheated Intake Air  These repeatability tests show that emission measurements from the two test cells cannot be compared quantitatively. Furthermore, the in-cylinder pressure is slightly different between the CERC tests after April 2008 and the other tests.  Table 3.4 shows the  implications of a larger or smaller conversion factor on the cylinder pressure measurements and analysis.  67  Table 3.4: Implications of 7.5% increased conversion factor Conversion Factor GIMEP COVGIMEP PcyI max COV PcyI max CA@ PcyI max COV CA @ PcyI max HRR max dPIdCA max Knock Intensity HR max Start of Combustion 2% IHR 5% IHR 0% IHR 50% IHR 90% IHR Combustion Duration (90% IHR 5% HR) -  Units barN bar % bar % oCA % kJ/m3/oCA bar/oCA bar 3 kj/m °CA °CA °CA °CA °CA °CA  A 2.151 8.8 1.8 92.1 0.7 13.3 4.6 174.7 4.5 1.2 1594.4 3.96 -6.6 2.4 5.4 10.5 23.0  B 2.000 8.2 1.8 85.8 0.7 13.3 4.6 162.6 4.2 1.1 1496.9 4.05 -6.6 2.9 5.4 10.6 24.5  %change (A-B)IB 7.5% 7.5% 0.0% 7.4% 0.1% 0.0% 0.0% 7.5% 7.5% 7.5% 6.5% -2.2% 0.0% -16.9% 0.0% -0.8% -6.1%  °CA  20.5  21.5  -4.7%  Parameters dependent on the pressure such as the GIMEP, maximum cylinder pressure, maximum rate of pressure rise, knock intensity, etc. change at the same rate as the pressure conversion factor. The COV GIMEP, the CA 50% IHR, the start of combustion, and the combustion duration are affected to a lesser degree. The parameters that are affected less represent robust parameters for analysis. Although the IHR appears to be affected by the choice of conversion factor, the ratio between two IHR (discussed in Section 4.4) also represents a robust measurement between Prototype A and Prototype B.  68  3.4  Injector Characterization Flowbenches  Two injector characterization flowbenches at Westport were used in this study. The BTR2 is used to check the quality of the all Westport injectors. This rig provides a static back pressure of 80 bar with a gas injection pressure of 16 MPa. Nitrogen is used as a substitute for CNG for these tests. Viscor® calibrating fluid is used instead of diesel because it has density and viscosity similar to diesel while being less of a fire hazard. For HPDI injectors, the liquid and gas injections can be tested separately; however with the co-injector prototype, only the gas flow response to commanded injection duration was tested. The EFS1 injector characterization flowbench has the ability to test up to 6 injectors simultaneously to determine injector-to-injector flow differences from a common rail. Injectors are installed in a modified engine head and the injectors can be tested against a static pressure of 60 bar simulating the engine cylinder pressure near the end of the compression stroke. For single injector testing, five blanks were installed in the EFS 1. This EFS 1 uses natural gas (rather than nitrogen), but the diesel was again replaced by Viscor.  In order to determine the gas flow response to changes in  commanded pulse width as well as changes in diesel injection mass, the mass flow rate of the diesel was measured gravimetrically (pail and scale), while a coriolis flow meter was used to measure the mass flow rate of the gas.  69  Chapter 4 -Results  4.1  Overview of Testing  Tests for this study were conducted so that the flow and combustion characteristics of two injector geometries could be compared. Table 4.1 shows the 8 different test series (I  —  VIII)  used in this study as well as the engine location and engine setup for each test. Three of these test series (I, VII, and VIII) were repeated with Prototype A and Prototype B. Flow characterization (I and II) of the prototypes was conducted at Westport Innovations (Section 4.2.1 and Section 4.2.2). All other test series (III  —  VIII) were conducted at UBC  with different combinations of test cell location, piston geometry and injector geometry as shown in Table 4.1. Series III, IV, and V were conducted in the engine using one injection per cycle, making it possible to determine the gas and diesel flow characteristics of each injector. Two series (VI and VII) compared single and double injection operation to determine the influence of the second injection on the first (pilot) injection. Finally, Series VIII examined the combustion variability, knock and emissions for a few standardized operating modes.  70  Table 4.1: Chronological Overview of Test Series 41: I  CD CCD  0_  o  •eCD  ()  -  0  4  I  J’  i-t•c3  CD  CD  I  I-A 111-A IV-A VI-A Vu-A VII-B VIII B -  11-B I-B V-B  SCRE, Kaiser BTR2, Westport SCRE, Kaiser SCRE, Kaiser SCRE, Kaiser SCRE, Kaiser SCRE, CERC SCRE, CERC EFS 1, Westport BTR2, Westport SCRE, CERC  VIII-B2  SCRE, CERC  09-01-06  13-01-06  HJ  09-03-06  09-03-06  KI  CD  6  15  11  0  11  0  10  16.7 Kist. 503  2 1  -  -  43  15  Kist. 503  1  0  10  15  Kist. 503  1  18  0  18  15  Kist. 503  2  16  14  30  15  Kist. 503  1,2  40  37  77  15  Kist. 503  1,2  9  15  24  16.7  Kist. 504D  2  23  0  23  -  -  11  0  11  -  -  11  0  11  16.7  3  90  16.7  —  21-03-06  22-03-06  GMC 43 —  22-08-06  22-08-06  SB  22-08-06  22-08-06  SB  13-09-06  15-09-06  SB  — — —  17-12-07  15-01-08  SB  22-02-08  26-02-08  SB  29-02-08  11-03-08  KI  27-03-08  27-03-08  KI  13-06-08  13-06-08  SB  — — — — —  09-06-08  13-06-08  SB  HJ = Heather Jones, id = lcoyo InoKosril, b =  4.2  9  —  CD  0  ..  VIll-A  H)  30  scott brown, (Mc  AVL F1exIFEM  AVL F1exIFEM  1 1 1 2  = (.orcI MCI aggart-Uowan  Single Injection Flow Characterization  Five different tests were conducted with single injection operation: two  at Westport  Innovations on flow benches to characterize the gas flow (Test Series I and II) and three at UBC in the SCRE (Test Series III, IV, and V). The results from these tests are described hereafter.  71  4.2.1 Test Series I and II: Flowbench Tests at Westport Innovations The BTR2 injector flow characterization rig was used for Test Series I for both injector prototypes. Table 4.2 shows the range of variables tested for both test series. Table 4.2: Controlled Parameters for Test Series I and II: Westport Flowbenches BTR2 and EFS1  Simultated Engine Speed (RPM) Gas Rail Pressure (Mpa) Diesel Gas Bias Pressure (MPa) Back Pressure (bar) RIT (ms) Gas Liquid GPW (ms) DPW(ms) #Tests/Prototype -  Test Series I 1800 16 0.5 0.8 80 n/a Nitrogen Viscor 0.5 3 0 llforA&B -  -  Test Series II 1200 23 1.2 60 I Natural Gas Viscor 0.45-0.7 0-2 23forB  Test Series I provides important information about the gas flow over a wide range of GPWs; however, this flow test does not provide adequate resolution in the area of interest (0.5  —  0.7  ms GPW), nor can it be used to determine the gas flow response to different diesel injection masses. In Series I (Figure 4.1) gas characterization tests were conducted in the BTR2 for both Prototypes A and B testing the gas injection separate from the diesel injection. Prototype B (with the sleeve) exhibited 8  —  26% lower mass flow rates than Prototype A,  depending on injection durations. At a gas pulse width (GPW) of 0.7 ms, the gas mass flow rate is 26% lower for Prototype B than it is for Prototype A. This reduction could be due to higher friction losses in the gas/diesel reservoir which would result in lower injection flows. The thick dashed lines in Figure 4.1 represent the acceptable injector-to-injector variability of Westport J-36 injectors. Both A and B were within these limits.  72  180 160 140  igg 40 20 0 0.5  1.5  1.0  2.0  PW (ms) Figure 4.1: Gas injection mass as measured at Westport Innovations with only a gas injection  The EFS1 flowbench was used for Test Series II, in order to test the diesel and gas injections together. These were tested over a range of pilot GPWs and diesel injection masses common for normal operation. A full factorial test was conducted over 6 gas pulse widths (0.45 to 0.7 ms in 0.05 ms increments) and 4 diesel pulse widths (0 ms, 1 ms, 1.5 ms, and 2 ms) resulting in 23 data points (0.45 ms GPW w/ 0 ms DPW not tested).  The diesel flow rate was  calculated from the change of diesel mass for test durations between 10 and 20 minutes. The natural gas flow rates were averaged over 100 seconds at a collection rate of 1 Hz. For the EFS1 flowbench, only Prototype B was tested. Figure 4.2 shows the results of the GPW sweeps with different diesel fuelling amounts.  73  40 B Oms DPW (1.5 mgfir) —o-—B-1msDPW(6.5mgiIr) ......B15msDPW(95mgñnj) •.x..B2msDPW(12.5mg[q) ——  .  35  I  •  -  .rj  0:  0.4  0.45  0.5  0.55  0.6  0.65  0.7  0.75  Pilot Gas Pulse Width (ms) Figure 4.2: Gas injection mass as measured by the gas/diesel flowbench (EFS1) at Westport Innovations  For injection durations shorter than 0.5 ms, the gas injection mass drops off very quickly.  For GPWs from 0.5  —  0.65 ms, the injection mass is, surprisingly, almost independent of the  gas injection duration. The observed plateau in this test is likely not perfectly flat as shown but appears flat due to a quantization error. At pilot GPWs longer than 0.65 ms, the gas injection mass again increases. Because only one test was run at each condition, flow measurements were taken from SCRE experiments to check the trends.  4.2.2 Test Series III, IV, and V: Gas/Diesel Characterization of Single Injection Tests at UBC Series III was performed by McTaggart-Cowan (2006) to determine the minimum diesel flow rate for low-load single-injection operation.. It is beneficial to study single injections to elucidate the behaviour of the pilot injection of normal double-injection operation. With  74  single injection operation, it is possible to estimate the gas and diesel masses for each injection from the averaged gas and diesel injection rate. With double injections, the distribution of masses among the two pulses is indeterminate. The single injection tests are shown in Table 4.3.  Table 4.3: Controlled parameters for Test Series III and IV: single injection tests in SCRE  Prototype A Ilib  lila  Test Series  Compression  Prot. B Ilic  IV  V  15  15  15  15  16.7  16.5  22.5  27.5  22.4  22.3  800  800  800  800  800  18.5  24.7  29.5  24  24  -10  -9.5  -7  -9  -8  0.7 unheated  0.7  0.7  0.7  1  unheated  70  56  Ratio Gas Rail  Pressure (MPa) Engine Speed (RPM) Diesel Rail Pressure  (MPa) Pilot SOT (deg ATDC) RIT (ms) MAT (°C)  —  unheated  (Si  (D  —  ()  -  Test Point  °  0)  -s-i p  p  P  PiIotGPW (ms) Diesel Injection  01  (TI  0’  Mass  0  (mg/mi) Intake Manifold Pressure (kPa)  0  0  ..  0  0  0  0  0)0)  CJ1  )  0  0  ö  01 I  * I  Bc 5• *  (Si  0 01  I  I  0  00  LjL.j  (.71  3 ()1  D  CD  0) 01  0) C31  0) 01  0) 01  —  —  i-,.  .-,.  3 CD  D  001  4  0)  k)  (31  01  Cii  -  -  —  0) (71 —  0 (71  0 01  0 (31  0 01  0 (71  .  Cii (Ti  0) Cii —  o (31  Cii  01 .-.-  01  o 01  75  Test Series III was conducted at three different injection pressures whereas Series IV was conducted only at moderate injection pressures. Note that as the gas pressure increased, the pilot SOT was set to occur later and the pilot GPW is set to be shorter. This was done in an attempt to offset the effect of higher injection rates at higher injection pressures. For these tests, the injection timings were chosen so that it would simulate the pilot injection for normal injection operation (McTaggart-Cowan 2006b). For Test Series III the intake air was unheated, resulting in temperature fluctuations in the intake manifold ranging from 20 to 30 °C. The intake air for Test Series IV, however, was heated to 70 °C. Test Series V is a set of single-injection tests that was conducted with Prototype B, intended only to characterize flow as a function of commanded pulse width. The compression ratio was not the same for Test Series V as it was for Series III and IV, but the gas and diesel flow rates should be similar, at similar cylinder pressures. Assuming constant specific heat during compression of an ideal gas, the cylinder pressure, Ptdc, and the in-cylinder temperature, Ttd can be estimated using the compression ratio, CR, and the polytropic constant, n (Sonntag et al. 2003, 278). For these tests n was set to 1.35, which near both to the polytropic constant suggested by Heywood (1984, 385) and the constant found from the measured pressure rise during the compression stroke.  tdo  TtdC  (cR)  (4.1)  1 TbdC (cR)  (4.2)  bdc  =  x  76  Dropping the MAP from 60 kPa (for Test Series IV case) to around 40 kPa keeps the peak cylinder pressures nearly constant. Similarly, a MAT of 56°C (329 K) will lead to similar in-cylinder temperatures close to that of the lower CR tests with a MAT of 70 °C. For different injection pressures, pilot GPWs, injection masses and manifold pressures, the fuel specific emissions, combustion variability, and ignition delay were calculated. The diesel injection mass or pilot GPW “min*” refers to the minimum amount of diesel or gas necessary to maintain stable combustion and may change from test point to test point (as seen in Figure 4.3). The intake manifold pressure was changed in 20 kPa increments. The importance of matching the in-cylinder pressures for injector comparisons can be seen in Figure 4.3. In this data from Test Series III conducted by McTaggart-Cowan and re-analyzed for this study, the amount of gas injected changes with manifold air pressure.  20  •-  12  •D  10  Representative Error Bar • A-5 kPa Manifold Pressure D A-25 kPa Manifold Pressure —o-— A-40 kPa Manifold Pressure t A-60 kPa Manifold Pressure  8 .E  6  z  4 2 0 0  I  I  1  2  3  I  I  4  5  6  Diesel Pulse Width (ms) Figure 4.3: Changes in CNG injection mass with increased diesel injection mass at different  manifold pressures (Test Series lIla). 800 RPM, 16.5 MPa gas injection pressure, 0.75 ms GPW 77  Note that the minimum diesel pulse width is shorter for higher manifold air pressures. At the time of commanded injection, the diesel fuel used to hydraulically hold the injector needle closed is drained from the injector. Higher cylinder pressures may cause the opening force to overcome the closing force sooner so that the injector needle lifts sooner, leading to an earlier injection of the gas/diesel mixture. The actual injection duration (as opposed to the commanded duration) is increased (Jones 2005b; McTaggart Cowan 2006b). Also seen in this figure is the effect of diesel pulse width on the amount of gas injected. The gas injection mass decreases as the amount of diesel injected increases. Figure 4.4 compares the gas injection mass rates of prototypes A (Series IV) and B (Series V). The gas flows are lower for Prototype B, consistent with the Westport BTR2 tests (Figure 4.1).  Additional tests with Prototype B and the 16.7:1 piston and 60 kPa MAP  produced flows that were higher than the Protype B flows of Figure 4.4 but lower than the Prototype A flows of Figure 4.4.  78  50.0  .E  •  45 0 -  40.0  -  -x-  IV A-2.2 DPW (16 mg/inj diesel) IV A-3.4 DPW (20 mg/in] diesel) V-B 2.2 DPW (14 mg/in] diesel) V-B-2.2 DPW -  -  -  o  E  -  -  I Representative Error Bar 25.0  C..)  5.0 0.0  0.40  0.45  I  I  I  I  I  0.50  0.55  0.60  0.65  0.70  0.75  0.80  Pilot Gas Pulse Width (ms)  Figure 4.4: Comparison of Gas injection mass of Prototype A and Prototype B measured in the SCRE Similar to BTR2 flowtests shown in Figure 4.2, the dependence of gas injection mass on pilot GPW is non-linear. For Prototypes A and B, the CNG injection mass decreased with increasing diesel injection mass for all GPWs. This behaviour makes sense, but contrasts with some of the Westport tests (see Figure 4.2).  4.3  Test Series III and IV: Single Injection Emissions and Combustion Characteristics for Prototype A  The series III and IV tests were conducted for a wide range of conditions and were not originally conceived as a systematic study of single-injection operation.  Nevertheless,  certain patterns emerge that will be useful in understanding the series VII and VIII double-  injection experiments.  Figures 4.5-4.7 show the ignition delay, COV GIMEP, and fuel  79  specific emissions plotted as a function of the gas/diesel volume ratio (GDVR), which is defined as  GDVR  = Pd,esel  (4 3)  <  Pgas  mdiese,  Where p is the fluid in-cylinder density at the time of injection and m is the mass injected. The natural gas density inside the combustion chamber is approximated by the peak cylinder pressure, and the start of combustion is approximated by the 5% IHR, consistent with the work done by McTaggart-Cowan (2006b). While one can’t expect this ratio to characterize all aspects of the combustion, it is clear from the figures that the emissions and ignition delay converge towards low GDVRs, for a wide range of conditions.  At high GDVRs, the  combustion is apparently more sensitive to other factors that would require further study.  3.5  35  ±Represer,tative Error Bar.  3  . .. .  2.5 E  Q  •. q 0  1.5  16 mg/mi diesel  -  25 0 LI  ()  x  X  -  XIV 22 mg/mi diesel  4  •  Ill 0 IV  30  .  .  F-fH Representative  Error Bar  20 15 10  0.5  0 lv  -  16 mg/inj diesel  xlv 22 mg/inj diesel -  0  10  20 Gas!DieseI Volume Ratio  30  40  :  0  10  20  30  40  GasIDiesel Volume Ratio  Figure 4.5: Ignition delay and COV GIMEP vs. gas/diesel volume ratio (Single Pulse)  80  250  100-.111 0  200  .111 oPI-l6mg/injdiesel XIV 22 mg/mi diesel  90  IV  16 mgñnj diesel  -  80  xlv -22 mglinj diesel  .  -  >-  70  -.  -  150  Representative Error Bar  0,  100  :  •  C.,  .  •  •  •,< ••  X  50  0  .•..Q LI  0  Representative ErrorBar  -  •  .  x400  --4  Z ( 3 )  00 . •  •  20 10  10  0  • . • ••  .<  20  40  30  0  ,•  20  10  GaslDiesel Volume Ratio  •:  ••  30  40  GaslDiesel Volume Ratio  Figure 4.6: CO and NOx vs. gas/diesel volume ratio (Single Pulse) 60  160 . Ill  140 120  .  100  --‘  80  .111  a IV  -  xIV  -  o IV  50  16 mglinj dieseir 22 mg/in] diesel  -  16 mg!inj diesel  XIV -22 mg/mi diesel ‘40  •  Represehtative ErrorBar  —4--,  30  ----*-.--  Representative Error Bar  0  ----.-  .•--  %cI 4 •< .•  .•  10  ••  <:  X  20  30  •• .  .  .  0.e  GasIDlesel Volume Ratio  .  C.,  0  •.  •  • •  40  0  •  00•  20  10  30  40  Gas(DIesel Volume Ratio  Figure 4.7: CH 4 and nmHC vs. gas/diesel volume ratio (Single Pulse)  The results of Test Series III and IV show that the observed correlations between emissions and GDVR might be related to the ignition delay (Figure 4.5). The smaller diesel quantities in Test Series III (i.e. larger gas/diesel volume ratios) resulted in the diesel being more dispersed throughout the gas so that the diesel would take more time to form an ignitable mixture with the air in the combustion chamber. Since the start of combustion was retarded and more variable this resulted in higher uHC. Similarly, increased gas injection mass in Test Series IV would also lead to less likelihood of having an ignitable mixture shortly after  81  the pilot injection. For single injection operation, the combustion stability did not appear to have a significant influence on emissions.  4.4  Test Series VI and VII: Pilot/Main Injection Interactions  Except at extremely low loads, the HPDI co-injector operates with double injections (both a diesel/gas pilot injection and a main gas injection). If the co-injector is to operate with lower diesel quantities, multiple injections are required at higher loads and speeds. Single injection operation at high speed would require high diesel mass injection rates to increase the likelihood of the injected diesel mixing with the combustion air to an ignitable mixture. However, as discussed in Section 2.3.2, lower diesel injection masses are desired in order to maintain an acceptable knock intensity level. Lower diesel injection rates require shorter gas pulse widths in order to maintain low gas/diesel volume ratios and thus acceptable combustion variability and uHC emissions. A main gas injection is therefore required after the primary pilot injection for high loads and/or engine speeds. Also, for diesel fuelled engines, significant NOx emission reductions for similar PM emissions can be achieved with multiple injection operation (Nehmer and Reitz 1994; Ghaffarpour et at. 2006). The same amount of energy is being released over a longer period, resulting in cooler cylinder temperatures and thus lower NOx. PM emissions would also be also lower as the soot producing regions at the jet tips are broken down and restarted with the second injection (Nehmer and Reitz 1994; Ghaffarpour et al. 2006). Double gas injection operation for the HPDI co-injector could potentially have similar effects.  82  Two different tests (VI and VII) were conducted with double-injection operation. Series VI examined the effect of the second injection on the first injection for a wide range of operating conditions. The results were qualitatively very similar to those of Series VII. However, Series VI had fewer repeats at each condition, so the trends were less clear than in Series VII. Therefore, Series VI results were moved to Appendix B. Test Series VII was conducted in order to determine the significance of the interactions between the pilot injection and the main injection. For this test series, three test modes were conducted for each test. First, the engine was run normally with a double gas injection. At this mode, the main injection followed shortly after the pilot injection in order to ensure stable engine speed and low uHC emissions.  For the second mode, the pilot injection  pressure, timing and duration were held constant.  The high speed data was recorded  immediately after removing the main injection. The thermal mass of the piston and cylinder allowed the wall temperatures, and therefore the diesel evaporation rates from the walls was expected to be similar. This procedure was repeated for the third mode, except instead of removing the main injection, the main injection was retarded to past 10 degrees after top dead centre (10° ATDC). The in-cylinder pressure and temperatures were controlled through control of the intake air pressure and temperature, as well as the back pressure. Back pressures higher than 30 kPa above the intake would cause the exhaust gas residuals to exceed a mass fraction of 0.03 (McTaggart-Cowan 2003). For VII tests, the back pressure was set under the intake pressure to minimize the amount of residuals. If there were no injection-to-injection variations and the pilot injection was truly independent of the main injection, then the heat release duration,  83  timing, and magnitude during the pilot combustion event would also be the same regardless of whether there was a main injection present. Table 4.4 summarizes the different points that were tested for Test Series VII at 1200 RPM. Similarly, Test Series VII at 800 RPM (Test points 1-24) is also discussed in Appendix B. The sample times, operating parameters and performance measurements for each test can be found in Appendix E. For these tests, the diesel flow rates were controlled at two different levels: low flow and high flow. Low diesel flow rates were controlled to around 10 mg/inj and high diesel flow rates around 20 mg/inj. However, the exact control of the diesel flow rate was found to be time consuming. It appeared that fluctuations in the bias pressure and/or extra noise on the scale may have been a contributing factor. Therefore for VII Series tests, some control over the diesel injection mass was sacrificed for longer sample times in order to ensure accurate mass flow measurements. These tests were recorded from 180s to 300s. In addition, at moderate speeds, the pilot injection by itself was not sufficient to run the engine. In both locations for the SCRE, the engine speed was reduced by up to 12% at 1200 RPM when the main injection was removed. The implications of this are discussed in Section 4.4.1.  84  N  I..  o.  ,  rM  z —  .  (1)0  # of Repeats Prototype B  .  # of Repeats Prototype A  .  Pilot Start of N Injection(degATDC)  .  N  —  Lf  N  -  -  N N N  -  N  -  N N  —  -  —  N  .  -  •  N  —  -  tkO  00  ©  N  >  -  00  N ci  -  00  -  C C C  -Z -Z  -  NN  00 00  j-  00 a C  C”  •  —  C C  N  tO  -  —  -  -  c’ ‘t j-  —  O  -Z .-Z -Z  -  -  N N N N N  I  00  Ir  i-  -  -  t(  -  N  tf  -  -  -  -  -  tIj  NNN  tIO  P 00 00  ‘1-  C  N N N  —  t  Cl  C  -  ‘  >  —  J  -  (-‘a  -  N  —  tJ  NN  —  00 00 t  C  -—  N 00 C  ri  -  .—  .D  ———————  q  00 00  —  Bias Pressure t II I! L( I t( (Diesel-Gas) ‘i ri ri c’i ri ri ci ei MPa  (ms)  .  iese T injecilon Mass (mg/inj) DieseiRail r155U1  •  esL r Olfli  ç) p  (MPa)  I..  “  ,  ‘P ,,  CI)  -  C C  a) 0 a)  0  N  ti)  0 Q  0  rJ  -c0 0 ID a)  0 a) a)  00  For Series VII tests the diesel introduced during the pre-injection was assumed to be injected into the combustion chamber both during the pilot injection as well as the main injection. Sample Integrated Heat Release (IHR) and the Heat Release Rate (HRR) curves for the three operating modes are shown in Figure 4.8.  3250  250 200  2750  150 2250  100 0  50  1750 1250 O  x  -50 X -100  750  Engine Speed: 800 RPM Diesel Press.: 18 MPa, Bias: 2.5 MPa DSOI: -10°ATDC, DPW: , GSOI: 10°ATDC , GPW:  -150 -200  250  -250  -250 -10  -5  5  10  15  20  25  30  Crank Angle (deg)  Figure 4.8: Representative HRR (Filtered) and IHR curves (Test Series Vu-A- 4, 800 RPM, 22.0 mg/mi)  These curves are typical for the Series VII experiments. Note that for normal operation, two distinct combustion events were observed: the pilot combustion event (PCE) starting about  -  2° ADTC and the main combustion event (MCE) which starts about 4° ADTC. The MCE was greater in magnitude and longer in duration than the PCE. At lower diesel injection masses a lower magnitude PCE was observed. Two combustion performance metrics used to quantitatively characterize the combustion performance are also displayed in Figure 4.8: the ignition delay and the IHR ratio. The ignition delay has been used previously and is defined in Section 3.1.6. Ignition was earlier for the PCE of the single injection mode  (PCESINGLE)  compared to  PCEDOUBLE.  The 86  integrated heat release for single injection  (IHRSINGLE) and IHRDOUBLE  were found at the end  /deg). Note in Figure 4.8 that 3 of combustion of the PCE (crank angle where HRR<1O kJ/m the  IHRSINGLE  was greater in magnitude than  IHRDOUBLE for the  PCE, indicating that more  heat was released during the PCE of the single injection mode. The IHR ratio is defined as IHRDOUBLE/IHRSINGLE.  As seen later in this section, for Series VII tests, the ignition delay was affected significantly by the diesel injection mass. The ignition delay for these tests is indicative of the ignitability of the fuel injected. Since the diesel is the fuel used for ignition is already finely atomized, shorter ignition delays may indicate that more diesel is available to mix with the combustion air early in the combustion cycle. The other factors affecting ignition delay for these test series are discussed briefly in Section 4.4.1. The IHR ratio is a comparison of both the energy of fuel injected and the combustibility of the mixture through the compression and power stroke. At lower engine speeds, higher diesel injection masses, larger pilot GPWs, or higher pressures, higher IHR ratios were observed (see Appendix E/Appendix F). Ignition delay and IHR ratio may be correlated since a longer ignition delay may lead to over-leaning of the fuel before ignition. Also, longer ignition delays push the PCE further into the expansion stroke increasing the likelihood of bulk extinction. Therefore the longer ignition delays for double-injection operation may lead to a lower-magnitude PCE. IHR ratios lower than one can mean a number of things. First, it may indicate that for double injection operation that less energy is released per unit of fuel injected for the same amount (or increased levels) of fuel injected during the pilot injection. It was previously thought that  87  the change in timing and magnitude of the HRR curves due to main injection addition was due to flame quenching and/or ignition delay from the addition of a cool turbulent jet during the main injection (Jones 2006). Note in Figure 4.9 the  PCEDOUBLE  is nearly the same for the  retarded and normal double injection. If there were adverse interactions between the PCE and the main injection, a significant difference would be observed with the PCE during normal double and retarded double injection operation.  Low IHR ratios may also indicate  that for double injection operation the same amount of energy is released per unit of fuel, but less fuel is injected. Without an accurate combined flow/combustion model, however, it is almost impossible to separate the importance of each factor. The observed differences in the HRR traces might be explained by gas/diesel interactions inside the injector shown schematically in Section 2.3.1 (Figure 2.3). It is likely that the high diesel velocity (10-80 mIs) during the pre-injection distributes the diesel through the injection reservoir both as a thin film on the reservoir walls as well as droplets mixed with the gas. As the injector opens during the pilot injection, the gas/diesel mixture will be injected into the combustion chamber. Since the diesel pre-injection essentially occurs a full cycle before the main gas injection it is assumed that the removal of the main gas injection will not affect the mass of diesel injected. A portion of the diesel will be retained in the reservoir after the pilot injection, dependent on reservoir geometry and the distribution of the diesel in the reservoir. Marr (2007) observed that the injector achieved steady state almost instantaneously which was also observed when the main gas injection was removed. The conceptual model is effective in explaining some of the observed differences between single and double injection operation. For single-injection operation, the retained diesel will  88  be injected with the following pilot injection and will therefore be a factor in reducing the ignition delay, since the added diesel will increase the likelihood of the diesel mixing with the combustion air. For double-injection operation, this diesel will be injected during the main injection where ignition has already occurred. Since more diesel is injected into the combustion chamber during the pilot gas injection, the IHR will be higher for the PCE which will reduce the IHR ratio.  4.4.1 Other Factors Affecting Ignition Delay and IHR ratio For both prototypes the relative injection timing (RIT) between the end of the pre-injection to the beginning of the pilot injection was changed while the start of the pilot injection remained relatively the same. Prior to these tests, Jones (2006) found that the RIT had very little effect on the HRR or emissions. These tests, however, were not conducted over a wide range of RITs or diesel injection masses. Figure 4.9 shows the HRR curves for different RITs for Prototype A at a low diesel injection mass (between 10  —  15 mg/inj). It appears that  the timing of the diesel injection into the gas/diesel reservoir makes a difference to the HRR curve. As seen in the HRR curves in Appendix F.2, at longer pilot GPWs and higher diesel injection masses, the difference is less evident. The RITs that result in the shortest ignition delay seem to be negative RITs and those close to 0. It is clear that although RIT may not affect the shape of the HRR (in some cases), it can affect the start of combustion. This is consistent with the conclusions of Jones (2006) as shown in Section 4.5.2. Also of significance in Figure 4.9 are the diesel to pilot RITs of 1.10 ms, 1.25 ms, and 1.45 ms. These test points resulted in significantly higher CH 4 and uHC emissions. The diesel  89  injection mass for these cases was found to be significantly lower (around 10 mg/inj as opposed to 15 mg/inj as found in Appendix E.2). The mechanism of how the RIT affects diesel injection mass and combustion is unclear. From Figure 4.9 it appears that the timing of the injection is important, which may indicate an optimal distribution of the diesel within the gas/diesel reservoir. Whether or not the RIT affects the needle opening time is also unclear. Additional testing is required to determine more conclusively the relationship between RIT and ignition.  260 220 -  180  140 •— 100  ‘60 20 -20  -10  -5  0  5  10 15 crank angle (deg)  20  25  30  Figure 4.9: Comparison of HRR curves for different relative injection timing. Double injection tests at a low diesel injection mass (VII-A-29)  Pilot injection pressure, pilot timing, intake and exhaust air pressures, intake manifold temperature, and engine speed are relevant factors that affect ignition delay (Heywood 1988,  546).  Although these parameters are controlled, as discussed in Section 3.3.4,  perfect  90  control of these parameters is nearly impossible considering the variation in load applied during single and double injection operation and the difficulty in control of these parameters in the SCRE. In the absence of the main injection at 1200 RPM, the engine speed may fall by as much as 12% (150 rpm), due to the difficulty the dynamometer and electric drive motor have of instantaneously reacting to the change in fuelling. A reduction in engine speed would be expected to show lower IHR ratios since at lower engine speeds there is more time before the expansion stroke for combustion to occur, increasing the magnitude of PCESINGLE. At lower engine speeds, less swirl changes the fuel evaporation rates as well as the mixing processes. In addition, lower peak compression temperatures will result from more heat lost per stroke (Heywood 1988, 546). At reduced engine speeds, longer ignition delays are therefore expected. In comparing double to single injection in Test Series Vu-A there is a possibility that lower engine speeds would result in shorter ignition delays. As mentioned in Section 3.2.1, the start of the pilot injection for Test Series Vu-A occurs before TDC based on a measured time rather than measured crank angle; therefore, the actual crank angle the pilot injection begins will be closer to TDC at lower speeds. It is likely that the pilot injections closer to TDC will reduce the ignition delay, since the initial temperatures and pressures inside the combustion chamber are higher.  The expected change in ignition delay due to higher in-cylinder  pressures, however, would only partially explain the shorter ignition delays observed for single injection operation. Also, this change is only applicable to the high speed tests from Test Series Vu-A since no significant speed change was observed at 800 RPM and the  91  injector control for Test Series Vu-B ensured the pilot injection occurred at the same crank angle when the main injection was removed.  4.4.2 Comparison Between Vu-A and Vu-B: Injector Geometry Effects on Ignition Delay and IHR ratio, 1200 RPM  Figure 4.10 shows representative HRR curves for Prototype A at 1200 RPM, whereas Figure 4.11 shows the same for Prototype B. The vertical lines represent the start of commanded diesel pre-injection, pilot injection, and main injection (around -30, -17, -5 deg BTDC). Note that the main gas injection is later for higher pilot GPWs since the 2RIT is held constant.  300 250 .  200 150 100  50 0 -50 -35  -25  -15  -5  5  15  25  35  Crank Angle [deg] Figure 4.10: Unfiltered HRR curves for Prototype A at 24 MPa Diesel Rail Pressure (VuI-A-29  and VuI-A-30)  For Prototype A, longer GPWs resulted in an advanced start of combustion, whereas for Prototype B, the start of combustion was not dependent on the GPW. In addition for some of the low diesel injection masses and with short pilot GPWs, no significant PCE was observed for Prototype A. On the other hand a distinct PCE was observed for Prototype B for nearly  92  the same conditions. The absolute start of combustion was also observed to be sooner for Prototype B.  200 150  -35  -25  -15  15 -5 5 Crank Angle [deg]  25  35  Figure 4.11: Unfiltered HRR curves for Prototype B at 24 MPa Diesel Rail Pressure (VII-B-29  and VII-B-30)  In order to determine whether the trends observed in Figures 4.10 and 4.11 hold true for different diesel injection masses, Figure 4.12 shows the measured ignition delay for both Prototype A and Prototype B at 1200 RPM and 24 MPa diesel rail pressure for different diesel injection masses. Since no difference was observed in ignition delay between the high and low bias cases, these are plotted on the same figure. Note that the test points with no measurable PCE have been identified with a  “+“,  since the ignition delay is also dependent  on the main injection timing for these cases. For these points, ignition delays greater than 3 ms were typical for Prototype A. For Prototype B, there was always a PCE present, even at low diesel injection masses.  93  4.00 3.50 Cl)  >. 2.50 ci)  x  C  . 1.50 C 9) 1.00 0.50 0.00 0.00  o  x  A  •  A  . o B-0.47ms • A-0.47 ms +NOPCE 5.00  10.00  B-0.7 ms A-0.7 ms XNOMODEb A  15.00  20.00  25.00  30.00  Diesel Injection Mass (mglinj) Figure 4.12: Ignition Delay comparisons between Prototype A and Prototype B at 1200 RPM, 24 MPa diesel rail pressure  Two observations can be made about the difference in ignition delay between Prototype A  and Prototype B. First, the ignition delay for Prototype B is consistently shorter than the ignition delay for Prototype A, especially at lower diesel injection masses.  At larger  injection masses, it is unclear whether there is a difference in ignition delay between prototypes. Ignition delay for Prototype B is less dependent on the diesel injection mass and therefore, as the ignition delay increases for Prototype A, the ignition delay for prototype B stays around 2 ms. This indicates that for Prototype B, the fuel mixture is more ignitable. Second, the minimum diesel needed for stable operation was observed to be significantly lower for Prototype B. For Prototype A, diesel injection masses under 12 mg/inj resulted in high COy GIMEP and methane emissions indicative of total or near-total mis-firing of the  94  engine. Similar engine variability was observed in Prototype B around 8 mg/inj. The fact that there was always an observed PCE for Prototype B, even if it was very small, may have had an influence on lower attainable diesel injection masses. Figure 4.13 shows the knock intensity (defined in Section 3.2.4) plotted against diesel injection mass for both Prototype A and B. Although it appears that Prototype B has higher knock intensity than Prototype B, the difference is far less evident than the difference in ignition delay.  10.00  ‘  9.00 8 00 7.00 6.00 5.O0 4.00 3.00 2.00 1.00 0.00 0.00  o B-0.47ms • A-0.47 ms +NOPCE  -  B-0.7 ms A A-0.7  ms  XNOMODEb  A  A  x  A  X I  I  I  I  I  5.00  10.00  15.00  20.00  25.00  30.00  Diesel Injection Mass (mg/inj) Figure 4.13: Knock Intensity comparisons between Prototype A and Prototype B at 1200 RPM, 24 MPa diesel rail pressure  For Prototype A there were many cases where a significant PCE was only observed with the main injection removed.  For these cases, an IHR ratio of zero was assigned and plotted as  “NO PCE”. This does not mean that no fuel was injected during the pilot injection, rather  95  both diesel and gas were injected but not at a sufficient quantity to initiate combustion. Conversely, for some of the test cases for Test Series VII with Prototype B, there was no PCE present when the second injection was removed.  These test points are indicated as  “NO MODE b” tests in the Figures 4.12 to 4.14 (plotted with an “x”). At 1200 RPM, these test points were most often observed at low bias cases. Without injector visualization at low diesel-to-gas bias pressures, determining the source of injector variability is difficult. Figure 4.14 shows the IHR ratio for Test Series VII for both Prototype A and Prototype B. Note that PCESINGLE was greater in magnitude for most cases than  PCEDOUBLE.  For the same  injector, the IHR ratio was closer to one at higher diesel injection masses and at longer pilot GPWs. At very low diesel injection masses, the IHR ratio is reduced as the ignition degraded and there was no observable PCE for all or some of the 45 cycles of recorded high-speed data. As the diesel mass increased, a lack of a significant PCE was less of an issue, but there still might have been diesel retained in the injector. If a specific amount of diesel was retained due to areas of low velocity or recirculation in the injector then the higher diesel injection masses would result in the observed higher IHR ratios since the fraction of diesel retained would be relatively less important. Similarly, longer pilot GPW durations would have higher IHR ratios since there would be more time to clear out the diesel and energy wise the retained diesel would have less of an impact. There may also be a maximum amount of diesel which can be injected during the pilot injection (for a specific GPW). In this case, the IHR ratio would decrease as the diesel mass reaches its maximum. More measurements would be needed to determine the relative importance of each model.  96  1.60 1.40 1.20 0  1.00  o B-0.47ms . A-0.47 ms + NO PCE  B-0.7 ms A A-0.7 ms x NO COMBUSTION  x  0.80 .  0  0.60  A  A  A  .  0.40 A  0.20 0.00 0.00  . J. .11. T1!T  5.00  10.00  .1 TT  15.00  20.00  25.00  30.00  Diesel Injection Mass (mg/mi) Figure 4.14: Ratio of heat released during the Pilot Combustion Event for Prototype A and Prototype B at 1200 RPM, 24 MPa diesel rail pressure  4.5  Test Series VIII: Emissions and Combustion for Multimode Timing Sweeps  The effects of double injection operation on combustion variability, emissions, ignition delay, and knock intensity were tested in the SCRE in Test Series VIII-A, VIII-B, and VIII B2 (see Table 4.10). These tests were conducted at three of the European Stationary Cycle (ESC 13) test modes (#7, #6, and #4) which are 30% load/i 100 RPM, 75% load/i 100 RPM, and 75% load/1400 RPM respectively. Test Series VIII-A was conducted by Jones (2006). Although the bias pressure was slightly lower for VIII-B and VIII-B2, the difference did not affect the operation of the injector since the diesel injection mass was held constant. The  97  exhaust manifold pressure would affect the residual fraction of exhaust gas retained in the cylinder and was therefore fixed to around 10 kPa (exhaust  —  intake pressure) for all tests.  As discussed in Section 3.2 the pre-injection could not overlap -60° ATDC in the CERC location since the comparators used for injection control are reset at this point. This is only important for the diesel pre-injections, since the fuel injection into the combustion chamber would not occur so early in the compression stroke. For Prototype B, the importance of pre injection timing on emissions was investigated by changing the RIT. Table 4.4 shows the four timing sweep test modes conducted for the three load/speed combinations. The baseline test mode consisted of nine test points (three timings for the three loadlspeed combinations) at a specified RIT, diesel injection mass, and pilot GPW. The RIT and diesel injection mass were then changed separately for an additional two test modes. Finally, the pilot GPW was changed from 0.7 ms to 0.6 ms for low speed/low load timing sweep. Shown also in Table 4.4 are the two tests conducted with Prototype B (B and B2) and the test conducted with Prototype A (A). The measured values for the controlled parameters, combustion parameters, power specific emissions, and injection timing are tabulated in Appendix E and the indicated pressure and heat release rate curves are shown in Appendix F.  98  Table 4.5: Test matrix for Test Series VIII: double injection timing Sweeps for comparison of emissions in SCRE Test Point  4 5 6 7 8 9 10 11 12 13 14 15 16 17 18  22 23 24  28 29 30  4.5.1  50% IHR (deg ATDC)  5 10 15 1 15 5 10 15  RIT  Engine Speed  GIMEP  (ms)  (RPM)  (bar)  1 1 1 1 1  1100 1100 1100  13 13 13  10 15 5 10 15  -7.3 -7.3 -7.3 -7.3 -7.3 -73 -7.3 -7.3 -7.3  5 10 15  1 I 1  5 10 15  3  1  1 I 1  1 1100 1100 1100 1100 1100 100 1400 1400 1400 11 11 11 1100 1100 1100 1400 1400 1400 1100 1100 1100  6 6 6 13 13 13 13 13 13 6 6 6 13 13 13 13 13 13 6 6 6  Series VIII-B and VIII-B2:  EQR  0.55 0.55 0.55 .55 0.55 0.55 0.3 0.3 0.3 0.55 0.55 0.55 0.55 0.55 0.55 0.3 0.3 0.3 0.55 0.55 0.55 0.55 0.55 0.55 0.3 0.3 0.3  Diesel Injection Mass  GPW  (mg/inj)  (ms)  A  B  B2  15 15 15  0.7 0.7 0.7  2 6 2  1 2  4 4 4  15 15 15 15 15 15 15 15 15 15 15 12 12 12 12 12 12 12 12 12 15 15 15  Repetitions  0.7 0.7 0.7 2  0.7 0.7 0.7 0.7 0.7 0.7 0.7 0.7 0.7 0.7 0.7 0.7 0.6 0.6 0.6  2 6 6 5  3 3 3  2 1 1 2 1  3 3 3 3 3 2 3 3 3 3 3 3 3 3 3 3 3 3 3 3  Effect of Operating Mode and Injection  Parameters  In Section 4.4, the influence of the diesel pre-injection timing on combustion characteristics was briefly discussed. The influence of the PIT, diesel injection mass, and pilot GPW on emissions were not discussed for Test Series VII due to the intrinsic changes made to the load and equivalence ratio.  99  Figures 4.15 to 4.18 compare the combustion characteristics and power specific emissions for the four different test modes at low load/i 100 RPM for the Test Series VIII-B and VIII-B2. Each figure compares the timing sweeps for the four test modes (baseline, -7.3 ms RIT, 12 mg/mi diesel, and 0.6 ms GPW). Similarly, Figures 4.19 to 4.22 and Figures 4.23 to 4.26 are comparisons at high load/1100 RPM and high load/1400 RPM respectively. Since the VIII tests were conducted at constant equivalence ratio and load, the response to different changes such as changes to the RIT, diesel injection mass, and pilot gas duration can be tested. Figures 4.15 to 4.26 show that the ignition delay is dependent on the diesel injection mass, that the effect of the diesel injection mass and relative injection timing are similar in many cases, that at low load the CH , tHC and CO emissions correlate strongly with the ignition 4 delay, the NOx emissions were independent of diesel injection mass and relative injection timing, and that at higher loads the higher load test points have similar ignition delays but higher knock intensities. First, the ignition delay gets shorter as the amount of diesel injected is increased. From Figure 4.15, the ignition delay was observed to be 3 ms for the low diesel injection mass (points 19-21) as opposed to 2 ms for the baseline (points 1-3). This trend was also observed in the Series VII tests (Figure 4.12). High load/i 100 RPM (Figure 4.19), and high load/1400 RPM (Figure 4.23) also showed longer ignition delays for lower diesel injection masses, especially at advanced combustion timing. For a purely diesel fuelled engine, the amount of diesel injected (holding load constant) has little effect on ignition delay (Heywood 1988, 546). For the HPDI co-injector, the amount of diesel injected early in the injection cycle may play a critical role in the observed shorter ignition delay times for higher diesel injection masses. If more diesel were injected earlier during the injection process then this finely  100  atomized diesel would mix with the combustion air earlier leading to earlier combustion. As with the Series VII tests (Figure 4.12) the pilot GPW made little difference to ignition delay. Second, increased RIT (as observed with a negative RIT) and reduced diesel injection mass had a similar effect on performance. From Figure 4.15, both increased RIT (VIII-B2 10-12) and lower diesel injection mass (VIII-B2 19 -21) had ignition delays longer than the baseline case (VIII-B2 1—3). This was more evident at advanced combustion timing. At high loads, the lack of a significant pilot combustion event at advanced combustion timing led to longer ignition delays and higher knock intensity, as some fuel which was injected during the pilot gas injection would potentially still be at the right combustible mixture at the time of ignition. Similar operation between longer RITs and lower diesel injection masses could both be related to the amount of diesel injected during the pilot gas injection. Longer dwell times between the diesel pre-injection and pilot gas injection mean that the diesel may be more distributed throughout the gas diesel reservoir, reducing the proportion that is available for ignition. Figures 4.16, Figure 4.20, and Figure 4.24 show that reduced diesel injection mass and increased RIT (negative RIT) had the same effect of reducing the knock intensity, especially at lower engine speeds. Third, the CO, CH 4 and tHC emissions at low load/i 100 RPM followed the ignition delay trends. Test modes with longer ignition delay resulted in higher CO, CH 4 and tHC emissions. It appears that at lower loads, longer ignition delay leads to overleaning of the fuel mixture which leads to higher CH 4 and tHC emissions. At higher loads and higher engine speeds, there was little or no difference in these gaseous emissions, consistent with the tests by Jones (2006) which found the RIT made little  101  4 and tHC emissions are difference to the power specific emissions. At higher loads the CH near the limits of detection (150 ppm and 230 ppm respectively) such that determining differences in emissions between test modes would be difficult. Fourth, for Test Series VIII-B2, the advanced combustion timing was observed to increase the NOx emissions. Also, cases where there was no PCE (such as test points 22-24 in Figure 4.21 and test points 16-18 and 25-27 in Figure 4.25) NOx emissions were higher. NOx formation for non-EGR cases at constant speed and load should only be affected by the injection parameters (Heywood 1988, 863). NOx emissions were not affected by the RIT between the diesel and gas injections, diesel injection mass, or pilot GPW, indicative that the spray characteristics were not affected sufficiently to observe a difference in NOx emissions. Finally, comparing the tests from low load/i 100 RPM and high load/ii00 RPM, the ignition delay is similar (Figure 4.15 vs. 4.19), but the knock intensity is greater (Figure 4.i6 vs. Figure 4.20). Since the RIT is constant at I ms between tests, the pilot start of injection is advanced by about 2.5 degrees for the high load case. Based on the dependency of ignition delay on combustion timing, ignition delay should be longer for the higher load case due to an advanced pilot injection of 2  —  3 degrees. The similarity between the ignition delay and  increase in knock intensity between low load and high load cases may be due to higher in cylinder temperatures at higher loads.  Hot walls and residuals would increase diesel  evaporation rates and chemical reaction rates during both the ignition and premixed burn phases of combustion.  This might explain why knock intensities increase even though  ignition delay has not been extended due to the earlier injection.  102  Series VIII-B tests are also compared to VIII-B2 tests in Figures 4.15 to 4.26. Due to the variance in VIII-B tests, the NOx, CO, CH , and tHC emissions are for many modes similar 4 to Test Series VIII-B2. Ignition was found to be similar for most cases; however, Test Series VIII-B2 showed longer ignition delays at advanced injection timing. In addition, at low load, the CO, Cl 4 and uHC are predictably lower for VIII-B tests. The differences could be related more to variation in the controlled operating conditions rather than any variations in the injector. Since the airflow for Series VIII-B tests was based on an assumption that the airflow included unresolved air leaks (see Section 3.1.5) the intake pressure was slightly higher for VIII-B tests, resulting in earlier injector needle opening times. Since the air leaks have been resolved, engine control has been reasonably controllable over a wide range of injection timings and engine speeds.  103  4.0  25  3.0  -2O  a)  E C  —  0  2.5  r  a)  15  2.0  C  .2 10  E .2 1.5 10  .0  0.5  E  0.0  0 0  10  5  15  20  0  5  50% HRR (deg ATDC)  20  25 I  3.5  o  VIII-B2  I  o  E. 0  sVHk  _  .2 j 15 O C o 10  >2.5 2.0  U)  15  50% HRR (deg ATDC)  4.0  —  10  0  U)  1.0  .0  E5 C)  0.5 0.0 0  5  10  15  0  20  0  50% HRR (deg ATDC)  10  5  15  20  50% HRR (deg ATDC) 4.0  2b  E  C)  °  3.5  a)  :g. 20  3.0  ••U)  C 0  >2.5  15 o 10  2.0 0 5 gi  o  U)  )10 •  E  .0  —  o  0.5  0  —  0.0  0 0  5  10  15  20  0  5  50% HRR (deg ATDC)  15  20  15  20  25  4.0 0)  35  a)  :  3.0 ‘.  10 50% HRR (deg ATDC)  20  C  2.5  15  2.0 io  0  Ca  C 9•0  0.5  0  0  0.0  0 0  5  10  15  50% HRR (deg ATDC)  20  0  5  10 50% HRR (deg ATDC)  Figure 4.15: Ignition delay and combustion duration for 1100 RPM and 6 bar GIMEP  104  -  E  3.0  4.0 3.5 3.0 2.5 2.0 0 1.5 1.0 0.5 0.0  2.5 2.0 1.5  i.o 0 0.5 0  10  5  15  0.0  20  0  5  50% HRR (deg ATDC)  10  20  15  50% HRR (deg ATDC)  3.0  4.5  4.0  0  VIII-B2  3.5  •  VIIIB  2.5  c3.0 LU  •  2.5  E — c,  C.)  0  o  0  1.5  2.0 Q 015  I  1.0  C lr  0.5  0.5 0.0  0.0 0  5  10  15  20  0  5  50% HRR (deg ATDC)  —  —  E  —  4.5 4.0 3.5 3.0 2.5 2.0 0 5 . 1 Q 1.0 0.5 0.0  0  0  1.5  i.o 0.5 0.0 10  5  15  20  0  5  10  15  20  15  20  50% HRR (deg ATDC)  4.5 4.0 3.5 3.0 2.5  3.0 0 ..  2.5  2.0 1.5  >20 o_  —  20  2.0 C..  50% HRR (deg ATDC)  E  15  3.0  0  It..  10  50% HRR (deg ATDC)  1.0 0.5 0.0  o_ ‘V  0.5 0.0 0  5  10  15  50% HRR (deg ATDC)  20  0  5  10  50% HRR (deg ATDC)  Figure 4.16: COV GIMEP and knock intensity for 1100 RPM and 6 bar GIMEP  105  12  18 16  0  10  14 8  12 0  0  0  8  E  04 0  36  0  4  2  2 0  0  0  10  5  15  0  20  5  10  15  20  50% HRR (deg ATDC)  50% HRR (deg ATDC)  12 18  0  0 —  -  — —  cl  E  VIII B2  10  x  B  8  6  04 z  4  2  2  0  0 0  5  10  15  5  20  10  20  50% HRR (deg ATDC)  50% HRR (deg ATDC)  12  18  0  10  16 —  15  0  14  co 04  8 0 06  Z  2  4  E  0  2  0  0 0  5  10  15  20  5  10  15  20  50% HRR (deg ATDC)  50% HRR (deg ATDC) 12  18  00  10  16  6  io e—  0  8 0  0  04 Z  2 4 0  2  0  0 0  5  10  15  20  5  10  15  20  50% HRR (deg ATDC)  50% HRR (deg ATDC)  Figure 4.17: CO and NOx for 1100 RPM and 6 bar GIMEP  106  20  20  15  .15  •c  -  .910  >10 0  0  •  0  E  0  0  10  5  15  0  20  0  5  50% HRR (deg ATDC)  10  15  20  50% HRR (deg ATDC)  20 20  II  ..-..15  0  VIII-B2  S  VIII-B  I  -c  II  15  I  I  I I  ?10 0  x 05  o 0  0  0 5  0  10  15  0  20  0  50% HRR (deg ATDC)  5  10  15  20  50% HRR (deg ATDC)  20  20  0 15  15 0 -  0 05  5  0  E 0  0 0  5  10  15  0  20  5  50% HRR (deg ATDC)  I.  10  15  20  50% HRR (deg ATOC)  20  20  15  15  910  110  0  05 5  E  0 0  5  10  15  20  50% HRR (deg ATDC)  0 0  5  10  15  20  50% HRR (deg ATDC)  Figure 4.18: CH4 and tHC for 1100 RPM and 6 bar GIMEP  107  50 4.0  $  3.0 2.5 2.0 gl.5  S  o.s 0.0  —  0  a  VIII-B  0  VIII-B2  5  45 40 35 30 25 20 15 10 0 5 10  15  0  20  5  50% HR (deg ATDC)  z  E° >25 Cu. 2.0 1.5 )1•0 0.5 0.0  35 30 25 20 15 10 05  0 0  5  10  15  0 0  20  5  0  E  S  15  20  50 ) 45 g. 40 35 30 25  4.0 3.5  Cl  10  50% HR (deg ATDC)  50% HR (deg ATDC)  _  20  In-’  S 5  C.’  15  50  4.0 3.5  f-I  10  50% IHR (deg ATDC)  2.0 gi.  15  0.5  0  0  5  0.0  0  5  10  15  50% IHR (deg ATDC)  20  0  5  10  15  20  50% HR (deg ATDC)  Figure 4.19: Ignition delay and combustion duration for 1100 RPM and 13 bar GIMEP  108  1.6 1.4  6.0  VIII-B VIIIB2  0  1.2  0  -  0  0.8  E C.)  3.0  0.6 0.4 0.2 0.0  0 ‘  1.0 0.0  10 15 5 50%IHR (deg ATDC)  0  1.6 1.4  2.0  C  20  0  5  10  15  20  15  20  10 15 50% HR (deg ATDC)  20  50% HR (deg ATDC)  5.0 0  1.2  0  uJ10 0.8 —  I)  3.0  0 (1)  0.6  £2.0  04  E  0 0 C  0.2  1.0  0.0 0  5  10  15  0.0  20  0  50%IHR (deg ATDC)  10 50% IHR (deg ATDC)  1.6 F—  5  5.0  1.4 1.2 a  0  0  0 .  c 0.8 0.6  0  0  £2.0  0.4  0 C  1.0  0.2  0.0 —  3.0  C  0  5  10 50%IHR (deg ATDC)  15  20  0.0 0  5  Figure 4.20: COV GIMEP and knock intensity for 1100 RPM and 13 bar GIMEP  109  18 16 14 12 10  2.0  1.5  o  VIII-B  I  VIII-B2  I  I  ‘  I  C)  1.0 0  0.5  4 2 0  V  E  0.0 0  5  10  15  0  20  5  18 16 14 12  2.0 —‘1.5  0  4 z 2 0  0.0 0  5  10  15  20  0  5  50% IHR (deg ATDC)  18 16 14 12 10  1.5  15  20  0 0  C)  0  0)1.0 0 00.5  10  50% HR (deg ATDC)  2.0  2 0  0.0  9 —  20  0  0 00.5  a?E  15  0  1.0  e._’r,)  10  50% HR (deg ATDC)  50% IHR (deg ATDC)  0  5  10  15  20  50% HR (deg ATDC)  0  5  10  15  20  50% IHR (deg ATDC)  Figure 4.21: CO and NOx for 1100 RPM and 13 bar GIMEP  110  1.4  0.8  0.7  s viii-B  0.6  4  c  I  0.5  0  1.2  I  1.0  VIII-B2  0.8  0.4 —  x  1)  0.4  °0.2  E  0.2  0.1  0.0  0.0 0  5  10  15  0  20  5  50% IHR (deg ATDC)  0.8  1.4  0.7  1.2  0.6  1.0  0.5 0.3 0.2  E  —  0.4  0.1  0.2  0.0  0.0 0  5  10  15  20  0  5  10 15 50% HR (deg ATDC)  20  0  5  10 15 50% IHR (deg ATDC)  20  50% IHR (deg ATDC)  0.8  1.4  0.7  1.2  0.6 (  1.0  ‘0.5  0.8  .0.4  E  20  0.6  I)  Qt  15  O.8  0.4 —  10 50% HR (deg ATDC)  0.6  0.3 0 0.2  0.4  0.1  0.2  0.0  0.0 0  5  10  15  20  50% HR (deg ATDC)  Figure 4.22: CH4 and tHC for 1100 RPM and 13 bar GIMEP  111  4.5  50  4.0  45 -40  ..35  35 3.0 >.. 2.5 Cl)  E  30 25  g  2.0 $_l.5 2,1.0  o  E  •  VIII-B  0  VIIIB2  20 15  .  1  1  05 0.0  0 0  5  10  15  5  20  10  15  20  15  20  15  20  50% HR (deg ATDC)  50% HR (deg ATDC)  4.5  50  4.0  .; 45  z  3.5 E  -ct E  25 a)  Cl  Cl)  2.0  N  $_1.5  35 •i 30 C 25 20 15 10 05  05 0.0  0 0  10  5  15  20  0  5  50% HRR (deg ATDC)  10 50% HRR (deg ATDC)  4.5 4.0  ‘1) 40  3.5  35 30 25  •N  2.0  20  1.5  15  io  .OIU  2,.  E —  E o5  0.5  0.0  0 0  5  10 50% HR  (deg  15 ATDC)  20  0  5  10 50% THR  (deg  ATDC)  Figure 4.23: Ignition delay and combustion duration for 1400 RPM and 13 bar GIMEP  112  1.8  5.0 4.5 4.0 ‘—35 3.0 2.5  1.6  1.4  E  a  VIII-B  0  VIII-B2  I I  I  1.2  .‘  0 ‘  E E  0  >0.8 80.6 04  1.5 C 1.0 0.5 0.0 C)  0  0.2 0.0 0  5  10  15  0  20  5  50% IHR (deg ATDC)  —  ef  bO  E  —  1.8 1.6 1.4 0 1.2 Ui 1.o 0 >0.8 30.6 04 0.2 0.0  E  20  15  20  8.0  a  ,,7.O ‘a a  S  a  .‘  5.0  C ci)  4.0  3.0 02.0 C 1.0 0.0 0  C,)  15  10 50% IHR (deg ATDC)  1.8 1.6 1.4 1.2 1.0 > 0.8 80.6 0.4 0.2 0.0  5  10 15 50% HRR (deg ATDC)  0  20  5  10  50% HRR (deg ATDC)  5.0 4.5 ‘4.0  0  .0  3.0 a) 5 . 2 2.0 o 1.5 0 C1.0 0.5 0.0 C  0  0  5  10 50% HR (deg ATDC)  15  20  0  5  10  15  20  50% HR (deg ATDC)  Figure 4.24: COV GIMEP and knock intensity for 1400 RPM and 13 bar GIMEP  113  4.0 F  3.5 3.0  •  VIII-B  o  VIII-B2  20 18 16 14 12 10 x8 0  2.5 IcUE  .9 r  o  1.5 1.0  4 2 0  0.5  E  0.0  tn  0  5  10  15  20  0  F  I  S  4.0 3.5 3.0 2.5 , 2.0 o° 1.5 1.0 0.5 0.0  ?i  0  5  10  5  10  15  20  50% HR (deg ATDC)  50% IHR (deg ATDC)  15  20  20 18 16 2 14 12 ?10 x8 0 z6 4 2 0  50% HRR (deg ATDC)  0  5  10  15  20  50% HRR (deg ATDC)  —  .  20 18 16 14 12  4.0 3.5 3.0  F  r-.  io  0 1.5 1.0 0.5 0.0  x  0  6 Z  0  5  10  15  20  50% HR (deg ATDC)  4 2 0 0  10 5 15 50% HR (deg ATDC)  20  Figure 4.25: CO and NOx for 1400 RPM and 13 bar GIMEP  114  2.5  4.0  3.5 2.0  3.0  1.5  25  .g 1.0  2.O 1.5  E  1.o  S  0.5  “  0.5  0.0  0.0 0  10  5  15  20  I  0  5  50% HR (deg ATDC)  10  15  20  50% HR (deg ATDC)  0.8  1.2  s  0.7  1.0  0.8  •  S q)  (_  0.4  0.6  O.3 0 0.2  0.4  0.1  0.2  ,  E  ‘  -  0.0  0.0 0  5  10  15  20  0  5  50% HRR (deg ATDC)  10  15  20  50% HRR (deg ATDC)  2.5  4.0  3.5 2.0 C’)  —.  3.0  25  1.5 -  2.0 1.0  1.5  0.5  E —  *—  i.o s  0.5  0.0  0.0 0  5  10  15  20  50% HR (deg ATDC)  0  5  10  15  20  50% IHR (deg ATDC)  Figure 4.26: CH4 and tHC for 1400 RPM and 13 bar GIMEP  115  4.5.2 Series VIII-A and VIII-B Combustion Comparisons As mentioned in Section 3.3, comparisons of the emissions between Prototype A and Prototype B are problematic, since each injector prototype was tested in a different location with different gaseous emissions analyzers. A discussion of the emission comparisons between Prototype A and Prototype B is discussed in Appendix B. The combustion parameters obtained from the high speed in-cylinder pressure data, however, should be easily compared. Comparisons of ignition delay, combustion duration, COV GIMEP, and knock intensity between VIII-A, VIII-B, and VIII-B2 tests are shown in Figures 4.27 an 4.28 for the three different load/speed combinations. Figure 4.27 shows that at low load/i 100 RPM (0.6 ms pilot GPW), high load/i 100 RPM, and high load/1400 RPM, Prototype B has a shorter ignition delay and longer combustion duration.  Since the intake pressure and back pressure are similar for these cases, the  difference in ignition delay is not related to the residual, but is indicative of the improved performance of Prototype B due to the inserted sleeve.  The diesel injection mass was  comparable between Prototype A and Prototype B, although the diesel flow rates were much more variant for VIII-A tests. The relative injection timing (RIT) was 1.0 ms compared to 0.3 ms for VIII-B2 tests. Since shorter RITs advances the start of combustion, the difference in ignition delay for VIII-B2 tests is actually more significant. The primary purpose of the diesel is to promote ignition. Therefore the shorter measured ignition delay is indicative of better performance with the modified injector geometry, since  116  the diesel is being used more efficiently (less diesel is needed for the same operating point). In addition, at higher loads the added sleeve effectively increases the allowable fuelling rates (to allow stable combustion with acceptable knock) which extends the operating range of the injector, potentially allowing lower engine emissions through new operating strategies. Figure 4.28 shows the combustion variability and knock intensity comparisons between Prototype A and Prototype B. At higher loads the difference in combustion variability and knock intensity between prototypes is less evident. At low load Prototype B has lower combustion variability but higher knock intensity. A similar knock intensity and combustion variability could be attained for Prototype A by increasing the amount of diesel injected; therefore, at low loads the range of operation of the injector seems be similar between injector prototypes, but shifted to lower diesel injection masses for Prototype B. Figure 4.29 compares the ignition delay against knock intensity for all of the VIII tests. This figure shows the tradeoff between ignition delay and knock intensity over a wide range of loads, speeds, and combustion timings. At ignition delays shorter than 2 ms the knock intensity increased substantially. The knock intensity for Prototype A appeared to increase at longer ignition delays (about 2.3 to 2.1 ms) compared to Prototype B (2.0  —  1.8 ms). Both Prototype A and B exhibited a small “tail”  that didn’t follow the ignition delay/knock intensity tradeoff curve. For Prototype B these were test points at lower diesel injection masses and retarded combustion timing (VIII-B2 24) which also had some of the shortest ignition delays. For Prototype A these points were mid-load with 50% IHR at 1 0°ATDC where the lowest knock intensity levels were recorded.  117  At ignition delays greater than 3.0 ms the knock increased as ignition delay was extended. These test points occurred mostly at advanced combustion timing which had small pilot combustion events. The characteristics of the ignition delay/knock tradeoff curve show the same things that were observed in previous tests; namely, shorter ignition delays lead to higher knock intensity, when there is no pilot combustion event present the knock intensity increases, and lower diesel injection masses lead to lower magnitude knock intensity.  118  I  0  0  0  rj•j  0  0  —  0  C)’  0  p  0 0  0  C)’  -a  p 0  0  C,,  0  a  r.J  0  C)’  0  a. p  CD  C)’  0  a. CD (a > -1 ci 0  I  0  01  0 0  0  X  X  L  0  co  x80  .L......L..L.....  0 010 C)’ 0 0 0 CYT 0 CJ1  Combustion Duration (deg)  8 x 0  c 0  0  (71  .9  ID  >  CD Co  0  C,,  0  ci  -i  >  0  o  0  C,,  C  0  C)’  0  0  p 0  0  01  -.  Combustion Duration (deg)  x  x  -  0C0C0Cfl0C)’0  a- p.  CD Co  I  0 0  C)’  0  0.x  Ignition Delay (ms) 0QC)C)4. 0010010010010  Ignition Delay (ms) 00C)C)4. bO01QC)10•Olb  p 0  VIII-B 4- 6 High Load/i 100 RPM 0.7msGPW  VIII-B 7- 9 High LoadJl400 RPM 0.7 ms GPW  0  01 0  C 0  C  p  I’.)  0  0  1..  C  01 C  -1  > ci C-)  L’3  0 0  0  1..  a- p. CD (0  I  C)’  0  I  0. CD  x  0  01  0 C  C)  0  C)’  0  -  x  xG9  C)’  -  00  0  I’3  C)’  r)  Combustion Duration (deg)  cX  ppaar\)1\) 001001CC)’  Ignition Delay (ms)  VIII-B 28 -30 Low Load/I 100 RPM 0.6msGPW  L’J C  0  O 0  00  -I  -  ii P 0  P.)  0  9’  0. CD  -&  01 0  P 0  I  0  01  0 0  0. CD  01 0  0  0  P.3  0  9.’  -  0  01 0  0 0  o  0  -  0)  0  coX  P.)  xcb 0  .  P.3 01  010)  Knock Intensity (bar)  ox  ()  Q2< (W  0  .-  COy GIMEP (%) r’.) oi o 01 0  00  VIII-B 7- 9 High Load/1400 RPM 0.7msGPW  0  0  C.)  >  0. CD CO  01 0  0 0  F’.)  0  91...  -  0 0  -  01 0  0 0  0  P.)  0 0  ‘—0  -  >  CD CD  o  c01  0 0  0 0 0 0  F’.)  .ox  01  -  P.30)  •  010)  I  Xo  )<  .  01  OP3  Knock Intensity (bar)  cD 01  CCV GIMEP (%)  VIII-B 4- 6 High Load/I 100 RPM 0.7 ms GPW  0 0  0  P.)  0 0  ‘0  D09’  0  O  P.)  0 0  p..  I  0101 O•  0. CD  I  01 0  0 0  .-  01  0  F’.)  0 x9  -  C.)  .  01  010)  Knock Intensity (bar)  OOx  @X3X  x 0  .  0  CCV GIMEP (%) r’j  Op 01 O  VIII-B 28-30 Low Load/i 100 RPM 0.6 ms GPW  8.0  70  I  L  XVIII-A .VllI-B  I  0  6.0 a a  5.0 U  e.  0 U  0  4.0 &  0  XX X  C.)  o  00  30  a  a  0  a  X  0  a 0 00  oaXX X ao°%°  2.0  • U  1.0  U  0  0  0  0.0 0.00  00  0  0  0  X X  <  ca  0  U  Xo  X>*1?<  I  1.00  2.00  3.00  4.00  5.00  Ignition Delay (ms) Figure 4.29: Knock Intensity/Ignition Delay Tradeoff Curve  121  Chapter 5: Conclusions and Recommendations  The objective of this research was to understand the interactions between the diesel and natural gas in an injector prototype and how these interactions affected the combustion performance and emissions of the engine. The fuel injector used was a high-pressure directinjection natural gas injector where the pilot diesel was first mixed with the natural gas inside the injector and then co-injected with the gas into the combustion chamber. The combustion performance of the injector was addressed through studies where the injector geometry and injector operation were varied. The geometry of the injector was modified by inserting a sleeve into the common gas/diesel reservoir. While the injector was being modified, much work was also done in moving the single cylinder research engine (SCRE) from one location to another and comparing its operation. It was concluded that comparison of the emissions between the two test cells would be difficult due to different emission benches being used. Operation of the engine and the in-cylinder pressure, however, were observed to be similar in both test cells. This chapter summarizes the general observations and conclusions made from the each of the tests conducted and recommends future work with HPDI co-injection. Similar to previous work this study concluded that for most operating conditions two gas injections were needed: the pilot gas injection and the main gas injection. Also, consistent with previous tests the relative amounts of gas and diesel injected during the pilot injection were important to engine performance. It was found that the main injection reduced the effectiveness of the pilot injection by scavenging diesel from the gas/diesel mixture in the injection reservoir, 122  lengthening ignition delay and requiring more diesel for stable operation. An added injector sleeve which increased fluid velocity in the gas/diesel reservoir and attempted to segregate the gas and diesel was found to reduce the amount of diesel needed for stable operation and reduce the ignition delay. In addition, it was determined the maximum allowable amount of diesel in the pilot injection was limited by engine knock (rapid energy release which causes high frequency in-cylinder pressure fluctuations).. 5.1  Injector Flow  From single-injection tests in the SCRE, the in-cylinder pressure was observed to have a significant effect on the gas injection rate. As the manifold air pressure (and thus the incylinder pressure at the time of injection) increased, the gas injection mass was also observed to increase. This had been noted previously by both McTaggart-Cowan (2006) and Jones (2006). Higher cylinder pressures may lift the injector needle earlier and hold open the needle longer thus increasing the gas injection mass. For a given injection pressure and gas pulse width, both the test engine and the Westport flow rigs showed that a 25% increase in diesel injection mass resulted in a 10  —  15 % reduction in gas injection mass.  Also observed  was that the gas injection response to commanded pulse width was non-linear. Between 0.5 and 0.6 ms commanded gas pulse width (GPW) durations, the change in gas injection rate was less steep. It is unclear whether this observation is caused by force balance on the injector needle or whether it was exclusive to the co-injector. For both the Westport flow rigs and the SCRE, the difference in injector flow was measured for both co-injector prototypes. The sleeved injector (Prototype B) exhibited 8-25% lower gas injection masses than the unsleeved injector (Prototype A) for similar rail pressures,  123  cylinder pressures, and injection durations. Still, the measured gas flow rates for Prototype B were within an acceptable range of operation.  5.2  Ignition Delay and Heat Release Rate  For single injection operation, the ignition delay was shortened as more diesel or less gas was added. Ignition delay was strongly correlated with the ratio of gas to diesel on a volume basis at the time of injection. Over a wide range of equivalence ratios this relationship was found to be true whether the gas pulse width duration was held constant and the diesel injection mass was changed, or vice versa. During normal injection operation, two gas injections were used resulting in a bi-modal heat release rate (HRR) curve comprised of the pilot combustion event and the main combustion event. Assuming the pilot gas injection and subsequent combustion were independent of the main injection then the HRR curve for single injection operation should have accurately represented the pilot combustion event for double injection operation. However, when the main injection was removed (keeping the pilot injection timing and duration unchanged), ignition delay was shorter and the magnitude of the heat released during the pilot combustion event was larger for both co-injector geometries. The difference was more apparent at lower diesel injection masses, and shorter injection durations. Injection pressure and bias pressure had a minor effect on the change in ignition delay. The ignition delay was found to be most dependent on the diesel injection mass. Higher diesel injection masses led to shorter ignition delay times.  Interestingly, increasing the  relative injection timing (RIT) between the diesel injection (diesel injected into the gas/diesel  124  reservoir) and the gas injection (gas/diesel mixture injected into combustion chamber) had the same effect in many cases as lowering the diesel injection mass, especially at advanced combustion timing. This may be an indication of diesel distribution in the spray being dependent on the distribution of the diesel in the gas/diesel reservoir. The added sleeve made a significant difference to the ignition delay and heat released during the pilot combustion event. Significantly shorter ignition delays were observed with the added sleeve consistently over different speeds and operating conditions. The difference in ignition delay was most evident at lower loads.  In addition, for some cases with the  unmodified co-injector (Prototype A), no significant pilot combustion event was observed until after the main injection was removed. With Prototype B, no such observations were made. Due to the increased pilot combustion heat release, the added sleeve also significantly reduced the amount of diesel needed for stable combustion. Up to 20% less diesel was needed for the modified co-injector to run the engine without misfiring.  5.3  Knock and Combustion Variability  For single injection operation (and the pilot injection for double injection operation) the minimum diesel and gas injection masses were limited by combustion variability as measured by the coefficient of variation of the gross indicated mean effective pressure (COV GIMEP). For single injection operation, the combustion variability increased as the diesel injection mass was reduced. The combustion variability could not be reduced by reducing the gas pulse width.  125  For double injection operation the combustion variability increased as the load was lessened or the combustion timing was retarded. At all other points tested the COV GIMEP remained relatively constant over all combustion timings for both injector prototypes. This indicated that when sufficient diesel was present for combustion, the sleeve did not positively or negatively affect combustion variability. For single injection operation (and the pilot injection for double injection operation) the maximum diesel and gas injection masses were limited by the onset of heavy “knock”. The indicated pressure curves (for both single and double injection operation) exhibited pressure fluctuations around 3-4 kHz which was found to be the first transverse mode acoustical frequency of the cylinder. The relationship between knock intensity and ignition delay is complicated. For lower diesel injection masses (12 mg/inj) and high diesel injection masses (15 mg/inj) with longer RITs, knock intensity increased at longer ignition delays. Since knock intensity increased with increased pre-mixed combustion, the longer ignition delays lead to higher knock intensity levels. However, for higher diesel injection masses injected into the gas/diesel reservoir just prior to the gas injection, the knock intensity was reduced at later combustion timing. In-cylinder temperature may have also been a factor in increased knock intensity. At higher loads (with accompanying higher cylinder and exhaust temperatures) the knock intensity was observed to increase, even though the ignition delay was held relatively constant. The higher temperatures may have caused faster reaction rates which would lead to higher rates of pressure rise.  126  At higher diesel injection masses, the knock intensity for the sleeved injector (Prototype B) was slightly higher, especially at lower engine speeds. However, at lower diesel injection masses and double injection tests with no significant pilot combustion event for Prototype A, the knock intensity was observed to be greater for Prototype A due to additional premixed combustion.  For a given speed and load if the range of diesel injection masses were  bracketed on one side by a significant pilot combustion event and on the other by knock intensity, then the injector sleeve moved this bracket towards lower diesel injection masses.  5.4  Emissions  For single injection operation, the fuel specific emissions of CO, and CH 4 from the engine could be reduced by either shortening the gas pulse width or increasing the amount of diesel injected, effectively lowering the ratio between the volume of natural gas and liquid diesel at cylinder pressures. This correlation was attributed to increased gas volumes adversely lowering the likelihood of the diesel mixing with the air to an ignitable state. Strong negative correlations were also observed between NOx emissions and the gas/diesel volume ratio. For longer ignition delays, the injected fuel mixes past combustibility before ignition occurs which increases the amount of unburned and partially burned fuel emitted. For single injection operation and for double injection operation with a short second injection (low load cases), a large portion of unburned and partially burned fuel is not re-ingested by the flame. These emissions represent a substantial portion of the CH 4 and uHC emissions for low load and single injection operation. At higher loads (longer second injection) much of these emissions are re-ingested into the flame, which significantly lowers the uHC emissions.  127  Although the uHC and CH 4 emissions may be related to the ignition delay at higher loads, the emissions bench could not detect differences between the test modes. Due to improvements made to the research engine, emissions between Prototype A and Prototype B could not be compared since the analyzers used to measure emissions in both cases were different.  However, since the ignition delay was significantly shorter for  Prototype B, one would expect the CH , uHC and CO emissions to be similarly lower at low 4 load for Prototype B with little change in the NOx emissions.  5.5  Conceptual Model of Co-injection  A conceptual model based on the observations about injector flow, combustion characteristics, and emissions is as follows:  diesel fuel is injected into the gas/diesel  reservoir at high velocities such that during the pre-injection the diesel is distributed through the injection reservoir both as a thin film on the reservoir walls as well as droplets mixed with the gas. As the injector opens during the pilot injection, the gas/diesel mixture will be injected into the combustion chamber. Because diesel is injected with the gas, increased diesel injection mass will displace the natural gas. A significant portion of the diesel will be retained in the reservoir after the pilot injection, depending on the reservoir geometry and the distribution of the diesel in the reservoir. For single-injection operation, the retained diesel will be injected with the following pilot injection and will therefore be a factor in reducing the ignition delay and increasing the magnitude of heat released.  For double-injection operation, this diesel will be injected  during the main injection, and is therefore unavailable as an ignition promoter.  128  In Prototype B, the added sleeve reduced the volume of the gas/diesel reservoir, resulting in higher fluid velocities inside the injector. These higher velocities could have sheared the diesel off of the walls more efficiently and swept the diesel out of the injector more quickly. In addition, the sleeve may help contain a higher concentration of diesel near the injector tip so that the highest concentration of diesel is injected near the beginning of the injection event. The finely atomized diesel introduced earlier in the injection event would have more time to mix to an ignitable mixture with the air, reducing the ignition delay and increasing the proportion of heat released during the pilot combustion event. Knock intensity may also increase with increased diesel concentrations, since higher concentrations of diesel may lead to more ignition sites for a faster burn.  5.6  Co-injector Operation and Co-injector Outlook  Since this is the first thesis on the HPDI co-injector, comparisons between the co-injector prototypes and the industry standard HPDI J36 are of interest.  Table 5.1 outlines the  similarities and differences in operation and performance between the injectors. Overall, both Prototype A and Prototype B operated surprisingly well considering that very little has been done to optimize the geometry of the injector for mixed diesel/gas operation. At high load the engine out gaseous emissions were similar to the J36 injector with lower PM emissions which is probably due to better diesel atomization (Jones 2006).  129  Table 5.1: Injector comparisons between the IIPDI-J36 and the co-injector  Injector Operation, Performance Ability to reproduce a given operating condition with a fixed set of operating and injector parameters  At high load emissions will depend mainly on the main gas timing, equivalence ratio, and oxygen concentration. At low loads, the pilot injection has a larger effect and emissions can be quite different.  .  .  2  Effect on emissions (general)  3  Effect on PM  4  Effect on NOx  5 6  Effect on uHC, CH4 Combustion variability (CCV  7  Knock intensity  8  Sensitivity to diesel quantity  9  Sensitivy to engine speed  ..  .  ..  Generally lower PM for the co-injector. For the J36 PM emissions depend strongly on the diesel pilot_injection. LovrNOxat low load. At high load similar NOx emissions. .  .  .  .  .  .  Sensitivity to cylinder pressure 11  Engine Startup  12  Transient engine control  .  .  and Emissions Comparisons between J36 and Co-injector Reproducibility similar to J36 IF fuel and cylinder pressures are identical, higher test-to-test variability at lower diesel injection masses. Additional variability due to extra control over gas to diesel bias pressure and relative effect of cylinder pressure.  .  Higher CH4 emissions at low load. At high load similar CH4 emissions Slightly higher combustion variability. Variable, from levels similar to J36, to above 10 bar. Knock is controlled by the pilot gas injection duration and diesel injection masses Higher sensitivity. Limited at low diesel quantites due to combustion variability/unburned fuel. Limited at high fuel quantities due to knock. High sensitMty. More diesel is needed at higher engine speeds for stable operation High sensitivity. Lower cylinder pressures (either lower boost or advanced injection timing) cause gas needle to lift later. Higher diesel quantities (20- 30 mg/inj vs. 10- 15 mg/inj for J36), and earlier injection timing (-14 deg ATDC vs. -8 deg ATDC for J36) needed to start. Unknown. Transient control currently could be limited by software. Transient control problems were identified with Co-injector A but not extensively described. No transient control problems have yet been identified for Prototype_B.  There were, however, issues observed with the repeatability and combustion variability using the co-injector. Similar operating conditions are produced with the J36-HPDI injector for a given set of pulse width durations, injection pressures, and cylinder pressures. Repeatability with the co-injector is dependent on parameters such as combustion timing, diesel injection quantities and diesel injection timing relative to the gas injection. The operation of the J36 injector is less sensitive to differences in cylinder pressures and diesel quantities. Most of these operational difficulties may be related to the ability to control the quantities of diesel and gas injected during the pilot injection. Unlike the J36-HPDI injector, co-injector  130  operation is highly sensitive to the amount of diesel injected during the pilot gas injection. Combustion variability increases at low diesel injection quantities and high knock increases at high diesel injection quantities. The added sleeve appears to widen the window of acceptable operation, especially at higher loads. Using shorter pilot gas injection durations is also an effective strategy in reducing sensitivity to diesel quantity; however, the current injector design limits the minimum achievable gas pulse width to 0.46 ms with recommended gas pulse widths above 0.6 ms.  Shorter gas pulse widths are an issue with the current  injector since cylinder pressure has a greater effect on gas injections especially at lower gas injection durations, making early combustion timing (before 5 °ATDC) troublesome. Issues with engine startup (Jones 2005a, 2005b; McTaggart-Cowan 2006b) and transient operation (Jones 2005a) have been identified in previous works. However, these points do not currently seem to be an issue. Higher diesel quantities (‘-20 mg/inj) are needed to start the engine naturally aspirated. Sensitivity to changes in engine operating speed have not been observed with the current engine setup. It is unclear whether this is a result of the changes in the injection control or to the injector geometry. For both Prototype B and future single-actuator injectors the injector design could be better optimized to increase repeatability and reduce the effect of cylinder pressure. Optimization of the injector should concentrate on better gas needle response at earlier injection timings and lower manifold pressures, more repeatable injector operation at shorter gas pulse widths, and internal injector geometries that prevent diesel from mixing excessively with the gas before the pilot gas injection.  131  5.7  Future Work  Future work on HPDI co-injectors (either the current Prototype B or future variants with a single actuator) could be broken down into the following categories: injector modeling, injector visualization, and engine tests. Even though the conceptual model of the gas/diesel interactions adequately described the observations of improved combustion performance with the modified co-injector, it does not describe all of the observations made. The model does not adequately describe the effect of the relative injection timing (RIT) on ignition delay, knock intensity, or diesel injection mass. For longer RITs, diesel will be injected into the gas/diesel reservoir earlier and the gas/diesel bias pressure may also be changing which would affect the diesel injection rate. Whether this allows the diesel to be more dispersed in the natural gas, whether a large concentration of diesel still exists in the reservoir, and whether the diesel has absorbed a sufficient quantity of natural gas to cause flash atomization is unknown. In addition, during the injection of the gas diesel mixture, it is unclear whether the dispersed diesel effectively lowers the critical velocity of the fluid at the choking point or whether the diesel reduces the natural gas by replacement.  Scaled models of a transparent injector would be a problem because it is  unclear which of these phenomena is important. A mathematical model that addresses all of these phenomena would be beneficial in explaining the difference in combustion. In addition to a more comprehensive model, the work started by Mikawoz (2005) and Marr (2006) in the injector visualization chamber should be continued. If similar operating points were conducted for Prototype A by Mikawoz and for Prototype B by Marr, these visualizations could be compared to determine whether there was any observable difference  132  between Prototype A and Prototype B. Further study should be done with Prototype B in order to quantify the effect of diesel injection mass and gas pulse width on the gas/diesel spray during double gas injection operation. Improvements could be made to the single cylinder research engine in order to improve experiment quality and streamline testing time. The diesel flow rate which should be based only on the bias pressure (diesel  —  gas rail pressure) and the pre-injection pulse width was  observed to be erratic for the same pulse width, both test to test and repetitions. Since engine performance with the co-injector is closely related to the amount of diesel injected, this significantly affected the repeatability of the injector. The source of the erratic diesel flow rate is unknown. As the single cylinder engine, diesel and natural gas supply systems, and ancillary sensors and analyzers are optimized, similar tests could be conducted in order to determine the source of these uncertainties. 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Experimental Study on Flashing Atomization of Methane/Liquid Fuel Binary Mixtures. Energy and Fuels, 19 (2005): 2050-2055. Zhang, J., D. Jiang, Z. Huang, X. Wang, and Q. Wei. 2006. Performance and Emissions of Direct Injection Diesel Engine Fuelled with Diesel Fuel Containing Dissolved Methane. Energy and Fuels, (2006) 20, 504-511.  142  APPENDICES  Appendix A- Instrumentation List This section describes the equipment used for controlling the SCRE and for collecting the important pressures and temperatures. The capabilities of the data acquisition hardware is given in Table A. 1 and A.2. The range and accuracy of the temperature, pressure, and flow sensors shown in Table A.3  —  A.4 are given as well as the range and accuracy of the gaseous emissions analyzers.  The following tables show the instruments used for the CERC setup. The instrumentation list for the previous setup (Kaiser) has previously been described by McTaggart-Cowan (2006a).  143  FPGA Board  National Instruments  NI 7831-R  FPGA  -  -  -  12 Slot Modular Chassis DAQ Card Thermocouple (2 hz_filter) DAQ Card Voltage (200 hz filter) DAQ Card Voltage (unfiltered)  Device  SCXI1100  SCXI 11 02B  SCXI 1102  SCXI1001 IA6FOFO  Model Number  Table A. 2: Data aquition hardware  750 kS/s  1.25 MS/s  Sampling RatelChannel  SCXI 11 02B SCXI1100  National Instruments  SCXI 1102  SCXI 1101  DAQ Flow Diagram  National Instruments  National Instruments  National Instruments  Manufacturer  -  -  16 Channels 12 bit (-10 to 1OV, 0 -_I OV) 8 Channels 16 bit (-10 to by)  Analog Input  Data Acquisition  Data Acquisition  12 Bit AID Converter  National Instruments  PCI-Mb16E-1  DAQ Card  DAQ Flow Diagram  Manufacturer  Model Number  .  Device  Table A. 1: Data acquisition cards  10 V ±1OV  —  +  —  +0.1 V  Range  Resolution  160 Lines  8 Channels, 0 ..5 V  Digital lnputlOutput  l2bit(4.88mV)  12 bit (4.88 mV)  l2bit(.049mV) -1 .2 deg C  -  8 Channels 16 bit (-by to IOV)  2 Channels -12 bit (-lOVto IOV)  Analog Output  -  Crank Angle  Charge Amplifier 503/1033 XH25D-ss-720ABCZ/AA042876  QC33C I M184  In-Cylinder Pressure  Kistler BEI Optical Shaft Encoder  AVL Piezoelectric  PCB Piezoresistive  15-1CO2EZ1V5 GBAR/405  Intake Manifold Pressure  NGK Spark Plugs Artech Industries  LZAO3-E 1  Torque  Wideband Oxygen Sensor (U EGO)  Abs.  (most):KMQXL-125U6  All T  Micromotion Coriolis force Sensor: Al-Scale UBC Design & Constructed Subsonic Venturi Omega Diaphragm PX2300-2D1_/ Omega k-type Thermocouple Setra Strain Gage 209  CMFOIOP323 NC I 397665; Transmitter: RFT9739 I 7027956 Gravimetric Scale  1705557  Manufacturer  SPD-DYN-110  PT-ENG-100  PR-INT-135  TQ-DYN-110  O2**  -  PR**  TC**  VEN-INT-l00  SCL-TNK-l00  FLM-NG-500  DAQ Flow Diagram  Sensor Description Model Number  Duff. Pressure  Air Flow  Liquid Fuel  Gaseous Fuel Flow  Device  Table A 3:Pressure and temperature transducers  -  -  0-200 bar  -  0 6 bar  —200°C to 1250°C From0-2toO 5000 psi -15.3 to 15.6 % Excess Oxygen  0-2 psi  0 500 kg/hr  0 -6 kg  0- 15 kg/hr  Range  ±0.5 deg  Linearity ±0.2%; sensitivity 28.41 pC/bar  ±0.05% FS  ±0.25% FS  ±0.75% rdg  ±1% FS  Estimated ±2% FS  ±0.1 g  ±0.5%FS (for> 0.8 kglhr)  Accuracy  Measuring Principle  Species Manufacturer Model Range INT Urasl4 0-5.0% C02 ABB EGA NDIR vol Urasl4 0-15% C02 ABB EGA NDIR vol Magnos 106 0-22% 02 ABB EGA Paramagnetic vol CLD 4000 0-2600 NOx PIERBURG hhd Chemiluminescent ppm CLD 4000 0-2600 NO PIERBURG hhd Chemiluminescent ppm Urasl4 0-2300 CO ABB EGA NDIR ppm FID 4000 0-3900 CH4 PIERBURG hhd FID ppm FID 4000 0-1500 uHC PIERBURG hhd FID ppm General Specifications Repeatability <1% FS Noise (Peak-Peak) <2% FS Drift: <2% FS/8h Linearity <2% of point between 15% and 100% of measuring range <1% FS  Table A. 4: Gaseous emissions analyzers  Appendix B: Results of Test Series VI and VIII-A not Discussed in Body B. 1  Test Series VI: Pilot/Main Injection Interactions  On the same day as the Series IV tests (single gas injection in Section 4.2.1 and 4.2.2), double gas injection tests were conducted  (Test Series VI).  Table B.1 shows the  controlled parameters with the main gas injection commanded to start 1.3 ms after the end of the pilot injection. Again for these tests, min* refers to the minimum duration pilot GPW that can be used for stable operation which can be seen on Figure B. 1. Table B.1: Controlled Parameters and Test Matrix for Test Series VI: Double Injection Tests in SCRE Effect of Diesel and Gas Injection Mass -  Test Series Gas Rail Pressure (MPa) Diesel Rail Pressure (MPa) Engine Speed (RPM) Pilot SOl (deg ATDC) RIT (ms) MAT (°C) Test Point 1-4 Pre-injection DPW (ms) 2.2 Pilot GPW (ms) 0.7 Main GPW (ms) 0.8-0.45  VI 22.4 24 800 -9 0.7 70 15-19 3.4 0.75min* EQR = 0.4  20-23 2.2 0.8min*  EQR  0.4  25-28 1.9 0.65min* EQR = 0.4  These tests were conducted at constant equivalence ratio of 0.4. For Test Series VI, two assumptions were made in order to compare the double injection operation to single injection operation. First, the CNG injection mass during the pilot injection was assumed to be independent of the CNG injected during the main injection. It was assumed that  147  since there was about 150 ms (2 engine revolutions) between the end of the main injection and the beginning of the next pilot injection that the main injection could not affect the CNG pressure at the injector tip. Second, all the diesel was assumed to be injected into the combustion chamber during the pilot injection event and the diesel mass was dependent on the pre-injection DPW only. If all the diesel was introduced into the combustion chamber during the pilot injection then the combustion characteristics of the pilot combustion event should be similar with or without a main injection.  However, Figure 4.12 shows the contrary. Comparing  single injection operation to double injection operation for the same pre-injection DPW, the ignition delay (the time between the start of the commanded Pilot injection to the start of combustion) is consistently longer for double injection operation at lower pilot GPWs. Note for double injection operation that at pilot GPWs below 0.45 ms, a significant increase in ignition delay was observed, indicating that there was not enough fuel injected during the pilot injection to initiate combustion.  148  3 —0—2.2 ms —0—3.4 ms ——2.2 ms —3.4 ms  2.5 U)  DPWDPWDPWDPW-  Single Injection Single Injection Double Injection Double Injection  >  ci) 0  2  C  10.5  -  00.4  i  I  I  I  I  I  0.45  0.5  0.55  0.6  0.65  0.7  0.75  0.8  0.85  Pilot GPW Figure B.1: Ignition delay for Single Injection vs. Double injection  Note that some of the variation between Prototype A and Prototype B could be due to test-to-test diesel mass fluctuations, since a higher gas/diesel volume ratio would result in longer ignition delays. Likewise, the longer ignition delay observed during double injection operation could be related to the gas/diesel injection interactions. B.2  Comparison Between Vu-A and WI-B: Injector Geometry Effects on Ignition Delay and IBR ratio, 800 RPM  Table B.2 summarizes the different points that were tested for Test Series VI at 800 RPM. The non-shaded regions represent regions where tests were not conducted. Note that for this test series, the tests conducted for Prototype A were much less broad. For Prototype A the relative injection timing (RIT) between the end of the pre-injection to the beginning of the pilot injection was changed while the start of the pilot injection remained relatively the same. Even when the diesel pre-injection occurred after the main 149  C C)  a>  C a> a) a>  C C a> I  C  z C) C)  C C C>  a>  aS a>  C C C) a)  C,)  b13  C 0 C’,  0 C  I..  —  a?  7 c’  In  a?  I.  —  z. .4-  r•  =  of Repeat kototype B of Repeat rototype A re-injection tc >ilot RIT foi E’rototype A (ms) >ilot Start ol njection (de 4 T1’iC” “‘‘“)  -  3ias Pressurt (Diesel Gas; YlPa  ‘  i1ot Gas Puls  ‘  m°/in b’ J)  ..  )iesel Injection Mass  t. .)‘rir..i vviutiijfl15j  .-I CI).  C)  C a?  a)  Cl)  0  rest Point  2 )iesel !• ressure (Mpa)  Cl) a?’  c)  C)  C a)  I  -  if;’  00 L 00 00  N  00  cc  N  ca  -  cc  -  N  0  cc  -  0  cc  N ‘.0  N -  C  N ‘.0  -  C  La  ‘.r  N  N ‘.0 ..L  .  N ‘.0 .L  N  00  0  -  t  1r  C  00  N  00  .  N  N  C  00  L  N  •—  00  .  L  00  N  N  N  C  N  L .L  N  C  -  N 00  N L  N -  C’  N  •—  c  c  1f  N  C  —  t i- 00  •—  N  ‘r 00 00 00 00 ‘.0 ‘.0 ‘.0 ‘.0  N  0  00 00  ‘.  ‘f  N  C  00  I.r  N  11Th  N  00 00  ‘Th  a>  I  0 C) a)  C  0 C  0  C’,  C C  a>  4-’  a)  I  C  r12  C C  a>  a>  C,)  C  In the new test cell setup (CERC), the engine speed was surprisingly difficult to control at engine speed of 800 RPM. At engine speeds lower than 850 RPM the dynamometer would intermittently cause the engine to stop. The issue was traced to the Hall Effect sensor gap on the dynamometer shaft or loose wiring between the dynamometer and the Digalog dynamometer controller.  Therefore, all of the low speed engine tests for  Prototype B for Test Series VII were conducted at 850  —  900 RPM.  Differences between Prototype A and B at 800 RPM were much less evident compared to 1200 RPM. This is due mostly to the range of pressures and flow rates tested without a significant number of repeats. Still, the comparisons were important for understanding the reasons significant differences were observed at higher engine speeds. Figures B.2 shows the measured ignition delay between Prototype A and Prototype B at 18, 24, and 28 MPa diesel rail pressure respectively.  At low injection pressures,  differences in ignition delay were not observed. It was observed that at many points there was no PCE observed, resulting in longer ignition delay durations. At 24 MPa, however, Prototype B again shows shorter ignition delays. The uncertainty on the ignition delay is less than 0.1 ms. Similar to moderate pressures at 1200 RPM, there is little distinction between the ignition delay of the two injectors at high diesel fuelling rates. Finally, shorter ignition delays were again observed for Prototype B at 28 MPa. Similar to 1200 RPM, the minimum diesel fuelling rate for stable combustion was lower for Prototype B than for Prototype A. The average minimum diesel fuelling rate was around 12-15 mglinj for Prototype A and 5-10 mglinj for Prototype B.  151  3.50 3.00  -  -  0.50-  •A-I8MPa LØB-18MPa I  -  0.00  0.0  5.0  I  •A-24MPa oB-24MPa  AA-28MPa iB-28MPa  I  10.0  I  15.0  I  I  I  20.0  I  I  25.0  I  I  30.0  Diesel Injection Mass (mg/inj) Figure B.2: Ignition Delay for Vu-A and Vu-B tests at 800 RPM The trends observed for the ll{R ratios between Prototype A and Prototype B at 800 RPM were not as evident as the VII-1200 RPM tests. As seen in Figure B.3, the IHR ratio varied widely. Still, there appears to be a shift to higher IHR ratios at higher diesel fuelling rates for both prototypes. Also, Prototype B appears to have a higher IHR ratio at moderate and high gas pressures. At lower injection pressures, the IHR ratio appeared to be slightly lower for Prototype B compared to Prototype A.  152  2.00 1.50  -  -  o  0  1.000  .2  0.500.00  -0.50 -1.00  -  -  O.  o I  ‘‘I’’  b’  • A -18 MPa oB-I8MPa  ‘  I  I  I’’’  0  • A- 24 MPa oB-24MPa  A -28 MPa B-28MPa A  -  0.0  5.0  10.0  15.0  20.0  25.0  30.0  Diesel Injection Mass (mg/mi) Figure B.3: IHR ratio for Vu-A and Vu-B tests at 800 RPM Figure B.4 shows that the knock intensity between Prototypes A and B are similar. As  with the ignition delay and IHR ratio, however, it was difficult to compare injector performance due to the lack of resolution with the Vu-A tests. Knock intensity was observed to be slightly higher for the Vu-B cases at 24 MPa rail pressure than both 28 MPa and 18 MPa test cases. This was different that what was observed for Prototype A, where higher knock intensities were observed at higher diesel rail pressures.  153  6.0 B-18MPa •A-I8MPa oB-24MPa •A-24MPa tB-28MPa AB-28MPa  5.0-  4.0 3.0-  0  o  °t<>  -  0Li  0•  L<>  1.00.0  -  oC>Q•t • 0  11111111  0.00  5.00  10.00  15.00  •  6>  0  I  I  20.00  25.00  30.00  Diesel Injection Mass (mg/mi) Figure B.4: Knock Intensity for Vu-A and Vu-B tests at 800 RPM  At both 800 RPM and 1200 RPM, diesel flow rate was shown to have the most significant influence on the pilot injection for all operating conditions. The lower the diesel flow rate, the greater the influence the second injection had on the first due to the second pulse scavenging some of the diesel from the nozzle plenum which had not been injected in the first pulse.  154  B.3  Test Series VIII: Double Injection Emissions and Combustion Characteristics  The objective for Test Series VIII was to compare ignition, combustion stability, and gaseous emissions between Prototype A and Prototype B for double injection operation. To compare emissions, timing sweeps at constant equivalence ratios and pilot fuelling rates were done. These tests were previously done for Prototype A by Jones (2006) at a CR of 16.7:1. Section 3.3.5 discusses the parameters that were held constant for this test series. G.3.1 Analysis of Variance (ANOVA) In order to determine the statistical significance of the injector type on the emissions and combustion stability, an Analysis of Variance (ANOVA) was performed. For ANOVA, the measured response (ignition delay, combustion stability, NOx emissions, etc) is related to the controlled parameters through the use of a block model. The block model used for these examples is presented in Equation B.1 (Hicks 1982, 252). In ANOVA, the hypothesis is that the treatment options  tyk  are insignificant so that the each measured  response Y,, will consist only of its population mean  =  /1 + Vik + 6 mQjk)  t,  and a random error Em(yk).  (B. 1)  The measured responses for this study were the ignition delay, the co-efficient of variation of the gross indicated mean effective pressure (COV GIMEP), and the power specific emission levels of CO, NOx, CU , uHC. The gross power and gross IMEP are 4 used since the friction losses in the SCRE are not the same as in a heavy-duty engine with 155  six working cylinders. COy IMEP has been used in previous studies as a measure of combustion stability. The treatments subscripts  i,  tyk  are the injector type (Ii), the 50% IHR (Hi), and the speed  (Sk).  The  j, and k represent the different injector types, combustion timing, and speeds  tested. Each possible combination of I, H, and S represent an observation cell which is repeated m times. These treatment terms plus the interaction terms are presented in Equation B.2.  =  p + Rm  +  I.  +  1+ H  + I,H + 1 Sk + HJLk + I,HJSk +  m(yk) 8  (B.2)  Equation B.2 includes interaction terms which may contribute to observed differences between injectors. Second order interaction terms (IH, ISk, etc) are usually negligible and third order interaction terms (I HJSk) are extremely rare. However, both second and 1 third order interactions will be included in this analysis since significant second order interaction terms make interpreting the main effects more difficult (Hicks 1982, 94). ANOVA is used since the sum of squares, SStotai, can be broken down into the sum of squares of the different treatments, interactions and error as shown in Equation B .3 SSIOIaI = SSJ+SSH+SSS+SSJXH + &ixs + SSHXS + SSIXHXS + £Serror  (B.3)  Each sum of squares is seen to be independent of the others and thus a Chi-squared distribution if divided by its degree-of-freedom (dO (Hicks 1982, 41); therefore, an F test can be employed. G.3.2 Assumptions and Corrections for ANOVA  156  Although the complete factorial was completed for the test data, there were cases of missing repetitions, mostly for Prototype A. Equal number points are needed for a full factorial analysis in order to retain orthogonality (Hicks 1982, 73). For some observation cells only a single measurement was taken. In these cases, Hicks suggests replacing the missing observations with those that make the sum of the squares of the errors a minimum (Hicks 1982, 74). Each missing term can be solved for separately solving for SSerror  over a wide range of that missing term and finding the minimum.  Also, it assumed that the treatment parameters are discrete. For the injector type this is evident since either Prototype A or B will be installed. The 50% IHR, the engine speed and the load, however are continuous and can vary widely in the SCRE. In this case, it is assumed that the change in the measured response due to the small change in H or L is much smaller than the random fluctuations,  Em().  Third, the Null Hypothesis of ANOVA is that all of the measurements are taken from a normal population with population mean  i  and variance a 2 (Devoire 2004, 689) The  sample means for a specific combination of the treatment options is allowed to vary for each case, but the variance is assumed to be the same for all of the tests (Hicks 1982, 59). For these tests, the difference in variance was quickly checked for all of the measured responses through the use of the D 4 factor as described below (Hicks 1982, 60). The D 4 factor test is usually used in quality control to check homogeneity of variance by determining for a given measured response whether all of the measured ranges of an observation cell (maximum Y  —  minimum Y) are less than D 4  where D 4 is  3.276 for a sample size of 2 and R is the average of all the measured ranges (Hicks 1982,  157  60).  It was observed that for most of the cases, all of the ranges were within the  threshold. The points outside the threshold were the CO, tHC, CH , and COy GIMEP 4 for the low load, low speed, 50% IHR at 15°ATDC.  Due to late cycle bulk flame  extinction, the absolute magnitude of these measured responses was changing rapidly at a 50% IHR of 15°ATDC. Therefore, a I degree error in setting the timing contributed to  the larger than normal variance measured for this observation cell. Also, the measurements need to be repeatable and random (Hicks 1982, 59). This means that for a given I, H, and S, the measured response for any given set of treatment options needs to be pulled from the same normal distribution. The order of the observations in a block of tests should be completely random over all of the treatment parameters. As with previous experiments in the SCRE, the tests are randomized as much as possible without significantly increasing the test time (McTaggart-Cowan 2006a; Jones 2004). In this case, the three 50% IHR timings  (100,  5°, and  150  ATDC) were conducted sequentially.  Finally, the most important violation on randomization was testing Prototype A and B in two different test locations (Kaiser and CERC). Thus, there is no way to be sure that the differences attributed to the injector have no contribution from the engine installation. Even though the variance in all of the tests was not completely equal, the treatment parameters were not discrete, and the observation cells were not in a completely random order, it was still useful to perform ANOVA in order to determine the significance of the treatment parameters.  The Null Hypothesis is rejected at an a. level of 0.01 (1 in 100  chance of identifying a significant effect when one is not present), as opposed to an a. level of 0.05 used in previous work with the SCRE by McTaggart-Cowan (2006a). The  158  treatment factors and interactions that exhibit significant differences were examined further in an attempt to understand the reasons for the differences. G.3.3 ANOVA Results Table 4.1 gives the ANOVA for the three measured emissions and engine performance metric for fixed load/changing speed. As discussed in Section 3.3.6, an equivalent table for fixed speed/changing load is presented in Appendix B. The treatment variables and the interaction terms of the treatment variables in Equation 4.2 have been transposed into the table with the second column displaying the degrees of freedom (di) which is i-i, j  —  1, etc. for the injector type, 50% IHR, etc. Although 3 tests were conducted for most of  Test Series VIII-B, only the first two repetitions were used since most of the data points for Test Series Vu-A had 2 or 1 repetitions. Table B.3: ANOVA for Test Series VIII: fixed load/changing speed df  Injector (I) 50%IHR (H) Speed (5) IxH IxS HxS IxHxS Error  1  CO EO•  2  • 3.6E-05.  1 2 1 2 2 11  2 6E-11  0.38 0.08 0.91  •  Nox  uHC  4 CH  3.8E-04  :•1.3E-Ô7  0.04  Ignition Delay (ins) 3.7E-04  0.01  0.05  .3.5E.O5•  0.77  9.413-04  L4EO6  0.39 0.99 0 24 0.58  0.88 0.16  0.08 0.40 0.98 0 77 0.60  1.OE-.  •  I .00V2.7E-04  I I  0.01 0.03 0.57  0.09 0.75 0 11 0.18  ••  I I  0.31  Coy GIMEP  0.60  The shaded regions represent those areas where the Null hypothesis (the measured response is independent of all treatments) is rejected at an x level of 0.01 (1 in 100 chance of identifying a significant effect when one is not present). The actual probability  159  is shown for these cases. Note that although the combustion timing and engine speed were observed to have significant impacts on the exhaust emissions and ignition delay, their significance will not be discussed. As seen in Table B.3, the injector geometry was observed to have the greatest significance on ignition delay, CO, NOx, and uHC emissions. These differences will be discussed in the following section. There were no factors that significantly affected COV GIMEP, possibly since there was sufficient diesel used to avoid long ignition delays discussed in the earlier test series (VII). Figures B.5  -  B.7 show the ignition delay, combustion stability, and gaseous emissions  plotted as a function of combustion timing. The difference in ignition delay due to injector geometry was found to be statistically significant at moderate engine speeds. Figure B.5 compares the ignition delay between Prototype A and Prototype B at 1100 RPM and 1400 RPM. At 1100 rpm, Prototype B exhibits ignition delays close to 1.6 ms,  whereas Prototype A exhibits an ignition delay nearly a millisecond longer. At this speed, the combustion timing has little effect on ignition delay. At 1400 RPM there is little difference in ignition delay between Prototype A and B. At 1400 RPM there is a strong relationship between ignition delay and combustion timing with later timings exhibiting shorter ignition delays for both Prototype A and B. The amount of diesel injected at 1400 rpm may not be enough since the pilot combustion was not as significant at higher engine speeds. Figure B.5 also shows the difference in combustion stability between prototypes. As previously mentioned, the engine stability was similar for both prototypes due to sufficient diesel being used to ensure stable combustion.  160  -a  C.)  o  >  C,  a. uJ  _.  0  —  a,  0  02  0.4  0.6  0.8  12  0  0  0  5  1100 RPM  0  5 10  15  15  20  4  50% HRR(deg)  10  o CoinjectorB  • CoinjectorA  50% HRR(deg)  1.5  0.6  0.4  0.5  0  0.1  02  0 0.3 o  C, >  0.. Ui  0.7  0.8  0.9  0  0.5  91  C  C  0.5  C 0  0  3.5  4  4.5  a, 2.5  .  C  1.5  m  >  C’,  >1  b) —  1100 RPM  • CoinjectorA o Coinjector B  0  2.5  3  0  0  5  5  10  50% HRR (deg)  1400 RPM  o Coinector B  • CoirijectorA  50% HRR (deg)  10  1400 RPM  0  0  • CoinjectorA o CoinjectorB  Figure B.5: Ignition delay and COV GIMEP for 13 bar GIMEP for a) 1100 RPM and b) 1400 RPM  a)  15  15  20  20  t’3  0.3  0.4  0.5  0.6  -  0  0  5 10 15 50% HRR (deg)  15  I  1100 RPM  50%H ( 9 deg)  • Coinjector A o Coinjector B  5  CoinjectorA 0 CoinjectorB  •  1100 RPM  20  20  b)  0.5  0  0.1  0.2  0.1 0  0.9 0.8 2 0.7 0.6 5 0.5 0.4 0 0.3  zD  ‘-  20.3 )  0.4  c  0.6  0  0  Figure B.6: CH. 4 and uHC emissions for 13 bar GIMEP for a)1100 RPM and b) 1400 RPM.  0.2 0.1 0  0.5  -c  0.9 0.8 0.7  0  02 I 0 0.1  )  a)  5  I  50% HRR (deg)  10  I  0  15  I  15  I  1400 RPM  50% HRR (deg)  10  • CoinjectorA 0 Coinjector B  5  I  • CoinjectorA 0 Coinjector B  1400 RPM  20  20  z  0.6  0  0.2  30.4  5  p0.8  1.2  0  2  ‘C 03  8  9  10  0  0  5  1100 RPM  0  5  1100 RPM  0  15  0 Coinjector B  50% HRR(deg)  10  15  • CoinjectorA  50% HRR(deg)  10  • CoinjectorA 0 Coinjector B  20  20  b)  ‘C  0 C)  0  z  0  .  0  0.5  1.5  2  2.5  0  2  4  8  10  12  14  0  0  Figure B.7: NOx and CO emissions for 13 bar GIMEP for a)llOO RPM and b)1400 RPM  a)  5  1400 RPM  5  1400 RPM  50%  10  HRR(deg)  15  15  • CoinjectorA o Coinjector B  50% HRR (deg)  10  • CoinjectorA o CoinjectorB  20  20  Figures B.6 shows the CH4 emissions and uHC emissions for the two different engine speeds. Although there were observed differences in engine stability and CH4 emissions between injectors, the difference was not statistically significant. The uHC however, measured significantly higher. The relatively higher uHC emissions was not expected for Prototype B since previous studies with the original HPDI injector have found that over 80% of the uFIC emissions are unburned CH 4 (Dumitrescu et at. 2000; Duggal et at. 2004). More likely, there was a linearization error in either the CH4 or uHC emissions in one of the test cell setups. The high uHC emissions are also suspect since the main causes for HC emissions in diesel engines fail to fully explain the observed differences in uHC emissions for the co injectors.  At moderate loads, there are three main mechanisms for hydrocarbon  emissions in diesel engines: over-leaning due to long ignition delay times, under-mixing from low velocity fuel vapour introduced late in the combustion process from the injector sac volume, and late cycle bulk quenching (Heywood,  .  Hydrocarbon emissions due to  over-leaning are correlated with ignition delay and should be lower for Prototype B. Similarly, the amount of low velocity fuel entering the combustion chamber late in the cycle should be nearly the same since the sac volume of the injector was not modified. Finally, the uHC emissions are seen to drop for later timings, as observed in Figure 4. If late cycle bulk quenching were important, higher uHC emissions should be observed for later injections. Figures B.7 shows the NOx emissions and CO emissions for the two different engine speeds. NOx emissions are higher and CO emissions are lower at moderate speeds due to a shorter pilot GPW for Prototype A. 164  At moderate speeds and high speeds the CO emissions are consistently lower for Prototype B. NOx emissions appear to be similar for both Prototype A and Prototype B, except for at advanced combustion timing at high engine speeds where NOx emissions are higher for Prototype B.  It is important to understand the difference in ignition delay and combustion stability between Prototype A and Prototype B since these metrics may have an influence on some of the observed differences in emissions. Over-mixing of the fuel before ignition was not observed to be an issue since the shorter ignition delay times for Prototype B at 1100 RPM did not result in significantly Lower CH 4 and uHC emissions. Shorter ignition delays for Prototype B indicated that there was more diesel in the pilot injection; therefore, more heat was released early in the combustion cycle lowering CO emissions and increasing NOx emissions. In the equation B.2, Sk, which represents the speed in the constant speed/changing load case can be interchanged with the engine load (Lk) in the constant speed/changing load case. This analysis was not included in the body of the thesis because different pilot injection durations were used between injectors, making it much more difficult to distinguish between the effects of the injector type and the effects of a shorter pilot GPW. Unfortunately, this negates any comparisons between Prototype A and Prototype B that could be done at low load for CO and NOx. The comparisons of COV GIMEP and ignition delay, however, should be fine since both were found to be independent of GPW (see Figures 4.9 and Figure 4.15 for COV GIMEP and ignition delay respectively). 165  Table 13.4 shows the ANOVA results for the constant speed (1100 RPM), changing load (0.4 & 0.55 EQR) tests. Again, the injector geometry was observed to have a significant effect on the ignition delay, and power specific CO and NOx emissions. In addition, for the CO and NOx emissions there were load x injector interactions. Table B.4: ANOVA for Test Series VIII: fixed speed/changing load Ignition  df  Co  NOx  tHC  4 CH  Delay n’s  1jl0  ‘ IxH IxL HxL IxHxL Error  Figures B.8  2 1 2  0.58 0.61 •8EO5 .:L3E-03. :O4  0.40  0.64  0.02  0.95 0.90 .3:.6O4 0.95  0.49 0.15 L1E04. 0.50  GIMEP 0:02  0.06 0.17 0.90 0.09  0.41 0.02 0.03 0.16  iii —  B.10 show the ignition delay, COV GIMEP, and power specific emissions  at 30% Load/i 100 RPM.  As with higher speeds and loads, the ignition delay is  significantly shorter for Prototype B. The COV GIMEP appears to be slightly higher for Prototype B, and the CH 4 a little lower. Longer pilot GPWs for Prototype B would allow for more early-cycle heat release which would lead to higher early-cycle cylinder temperatures. This could explain the higher NOx and lower CO emissions observed for Prototype B.  166  3.5  4.5  • Con.ctA  o Cn.ctB  • CoinjectorA 0 Coinjector B  4  3  3.5  2.5  cry  1.5  p 0  gO  02 15  C,  I  2  /  a.  2  o  0.5  0.5 0  0 0  5  10  15  20  0  5  50% HRR (deg)  10  15  20  50% HRR (deg)  Figure B.8: Ignition delay and COV GIMEP for 6 bar GIMEP and 1100 RPM 6  0.40 EQR, 1100 RPM  • C0jfliCtOIA o CoectorBo  5  2  £4  10 9 8  • CH4-CoinjectorA 0 CH4-CoinjectorB uHC-CoinjectorA uHoiecinJectorB  .9 C,  =4  ‘is 01  0 0  5  10  15  20  0  50%HRR(deg)  0  5  10  15  20  15  20  50% HRR (deg)  Figure B.9: CH4 and uHC emissions for 6 bar GIMEP and 1100 RPM 8  12  • OCoinjector Co:e:or,,,/  10  7 6  .9 Z  • CoinjectorA 0 CoinjectorB  2  2  0  0 0  5  10  50% HRR(deg)  15  20  0  10  5  50% HRR(deg)  Figure B.10: NOx and CO emissions for 6 bar GIMEP for 1100 RPM  167  Appendix C: Carbon Balance and Airflow In a perfect world, this section shouldn’t exist. However, because of random and systematic error, mis-calibrated measurement devices, and human error, the measured and calculated values contain uncertainties. Usually, the airflow rate is calculated through the use of a UBC built venturi. However, in this test engine, there have been problems getting a proper mass balance of Carbon. In addition, with the airflow measurement from the venturi, the volumetric efficiency was found to be greater than 1 for the SCRE, whereas it should be around 0.8  —  0.9.  The error in the Carbon balance indicates that there are one or more systematic errors from devices used to calculate the airflow. A systematic error is defined as an error that is independent of the number of measurements. Assuming that the emission bench analysers respond in a relatively linear fashion, systematic uncertainty should be minimized with these sensors as they are calibrated daily. The linearity C02 and 02 passed linearization checks. While the diesel flow rate has large random fluctuations associated with it, there should be no systematic errors (unless of course there were a leak somewhere). This leaves the natural gas coriolis flow meter (systematic uncertainty may be due to residual strain in the strain gauge) or the airflow venturi (systematic error may be due to calibration). For these tests, the error was assumed to come from the air flow reading. The other measurements that contribute to the mass balance of Carbon have been checked and so far, no systematic errors (offsets) have been found.  168  Therefore, in order to better approximate the airflow, an airflow was chosen so that the Carbon balance would be close to 1. Functionally, this is ignoring any measurements of airflow from the venturi and using the other measurements plus the First Law of Thermodynamics to solve for the airflow. Not only does this provide a more accurate measure of the airflow (assuming that there are no systematic errors in the other measurements), it also provides a more precise approximation of the airflow rate. This can be shown through a study of the propagation of errors in the system. Instead of using the error propagation equation to determine the r.m.s of the airflow rate, Monte Carlo simulation is used. Since there are continual improvements to reduce both the random and systematic error in the system, the program can be quickly modified to reflect those changes. Measured data was taken both from the old engine setup and the new engine setup. Assuming that the errors for each of the measurements were independent and Gaussian, a 10,000 Monte-Carlo simulation was run. Two simulations were run. The first was the airflow computed from the pressure drop through the venturi. For each run, a normally distributed measurement for the air line pressure, temperature, and venturi pressure drop were used to calculate air flow rate using the existing calculations for flow rate. Errors in the measurement of the venturi areas were not included at this point. The calculated airflow based on the Carbon balance was done by taking normally distributed Gaussian distributions for the intake (airflow, intake C02, CNG, and Diesel) and the exhaust (02, C02, CO, NOx, and tHC) to calculate the C balance. The standard deviations were assumed to be the variation of the measurement over the sample time. The airflow was then  169  changed by multiplying it by a correction factor until the C balance was equal to 1 ± 1E-6. The resulting histograms can be seen in Figures C.la and C.lb.  Airflow (kg/hr)  Airflow (kg/fir)  Figure C. 1: Comparison of Airflow Calculations in a) Kaiser, and b) CERC  Two important observations should be made about Figures C.la and C.lb. First, that there is a systematic error in one or more of the measurements observed as a shift in the mean calculated air flow rates. Air leakage in the intake air system or piston blow-by may cause air flow rates as measured by the venturi to be higher than expected. Similarly, inaccurate measurements for the diesel flow rate, poor linearization of the 02 emissions could cause high or low air flow readings using the Carbon ratio. Second, there is still improvements that can be made in the CERC setup to reduce error, as seen by comparing the rms values between Figures C.la and C.lb, shown in Table C.1. For both test locations, most of the error comes from the natural gas flow measurement, the CO 2 and the Diesel. However, in CERC, there are significant contributions from the uHC and the CO. A smaller bottle of span gas should help for the CO measurement.  170  Table C.1: Specific measurements contribution to Airflow Uncertainty (Carbon Balance)  C02 Error (kglhr) 02, CO, NOx Error (kg/hr) Co uHC Error (kg/hr) CNG Error (kg/hr) Diesel Error (kg/hr) Total Error (kglhr)  Kaiser 0.4 0.0 0.0 0.0 2.1 0.2 1.7  CERC 0.8 0.0 0.4 0.2 1.0 0.2 1.4  The measurement of the diesel mass has accumulated errors coming from two points. First errors are introduced due to the fluctuations in the actual mass of the diesel mass measured in the scale. This is caused by the re-circulating diesel. Diesel pressure fluctuations will cause flow fluctuations into the measuring tank. Electrical noise and vibration may also be a factor. The second source of error is the mode of digitizing the diesel mass. The 4—20 mA signal from the scale is first converted into a 1  —  5 V signal and then to a 12 bit number on a  scale from 0-1 OV. For the scale maximum range of 4 kg this would result in a resolution of 2g per bit. Similarly, Table C.2 shows the contributions of the specific measurements for the Airflow. Note that the Venturi pressure, and the airline pressure have the largest contributions for both sets. Not shown here are the contributions of uncertainty in the flow areas or Cv, which is used in the calculations. Depending on the uncertainty, these factors can have significant effects (up to 1.5 kg/hr error).  171  Table C. 2: Specific measurements contribution to Airflow Ijncertaint (Venturi  Venturi dP Error (kg/hr) Airline T Error (kg/hr) Airline P Error (kglhr) Total  Kaiser CERC 0.8 1.0 0.0 0.2 0.6 0.7 1.0 1.2  The pressure variations at the intake pressure are slightly larger for the new system, due partly to the fact that the air pressure is being regulated. Hysteresis in the pressure regulator introduces some random error. From this analysis, the use of the Carbon Balance as an additional measure can give accurate approximations of a specific value, if there is a systematic error present. For example, for this study, it was used to measure the airflow rate.  172  Appendix D: Factsheets The Factsheets are as follows: •  U1-FAC-093-TEST Heather Jones  •  U1-FAC-098-Test  •  W1-FAC-3788-ANYS  -  —  Gord MeTaggart-Cowan —  Phil Hill  173  Alternative Fuels Group Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, BC. V6T 1Z4  Engine Testing Results from First 121-Coinjector Prototype J36 Comparison to 121 Injector  ,5’.1?.E Project —  id 3  Round of Testing  Objectives  1. Compare the emissions using the 121 injector versus a J36 injector over a range of operating conditions 2. Study the effect of pilot injection mass on emissions and combustion stability with the J36 and 121 injector. 3. Establish that the timing of the pilot pulse for the 121 injector does not need to be strictly controlled. Test Matrix  Basically three main speed/load conditions were tested: 30% and 1100rpm, 75% load and 1100rpm, and 75% load and 1500rpm. Each of these main conditions were tested with 30% EGR and without EGR. At each condition a timing sweep was done by setting the power at mid-timing (50% IHR at 10 degrees ATDC) and then the fuel flow rate was held constant during the sweep. Hence the GIMEP changed slightly during the timing sweep but the fuel and air flows stayed relatively constant. All of the tests were completed with a fixed pilot fuelling of 15mg/injection. In addition, at 1100rpm and 75% load the pilot quantity was increased to 20mg/injection and a timing sweep was done for both injectors and the J36 was also tested with a lower pilot quantity of 7-8mg/injection (low pilot fuelling not possible with the 121 injector due to combustion instability). At 1100rpm and 75% load, the relative timing of the pilot pulse (PSEP) was varied with respect to the gas fuelling so that the pilot pulse was well ahead of the first gas pulse until it was after the second gas pulse. Table I shows the testing matrix and Table 2 shows the controlled parameters during testing. Table 1: Test matrix GIMEP = Gbar,  GIMEP = l3bar,  GIMEP = l3bar,  1100rpm  1100rpm  1500rpm  50% IHR @ 5, 10, 15 deg. 0% EGR  **  Pilot quantity 15,  50% IHR @ 5, 10, 15  15 degrees  2Omg/inj*  degrees  50% IHR @ 5, 10, 15 degrees  30% EGR *  50% IHR @ 5, 10,  PSEP (121 only) -5, -1, 0.3, I ms 50% IHR@5, 10, l5deg. Pilot quantity 15, 50% IHR PSEP (121 only) -5, -1, 0.3, degrees I ms’  @ 5,  10, 15  Timing sweep done with each quantity Done only at mid-timing (50%IHR at lOdeg.)  The gas injection pressure was fixed to 21MPa for both injectors. The 121 injector had a diesel rail pressure of approximately 23.6MPa during testing. A bias of 2.6MPa worked well in the past so this was fixed. The exhaust back pressure was fixed to approximately lOkPa over the intake pressure so that the residual fraction in the cylinder and the exhaust temperature remained relatively constant during each timing sweep and from injector to injector. The 121 injector was run only with pulsed gas injection. The first gas pulse width was fixed to 0.6ms at AUTHOR:  DOCUMENT NUMBER: KEYWORDS:  Heather Jones U1-FAC-093-ThST 121, J36, coinjector  DATE  Alternative Fuels Group Department of Mechanical Engineering, University of British Columbia 2054 —6250 Applied Science Lane, Vancouver, B.C. V6T 1Z4  .  S’(’RE Project  low load and O.7ms at high load and the second gas pulse was used to control the power output. The second gas pulse was timed to occur I .5ms after the end of the first gas pulse (shown in the last set of tests to be a good setting, refer to U 1 .-FAC-092-TEST). The overall equivalence ratio was fixed during this testing. The oxygen in the recirculated exhaust gas was included in this calculation so that when power is fixed, the intake manifold pressure will be higher to achieve the desired oxygen level (similar to supplemental EGR). Table 2: Fixed Parameters GRP 21 MPa 121 injector Bias  2.6 MPa  Exhaust BP (without EGR)  lOkPa over intake  Overall equivalence ratio (based upon oxygen/fuel) CNG flow  0.3 at low load, 0.55 at high load  Pilot timing (PSEP)  0.3ms  Pilot fuelling  15 mg/mi  121 first gas pulse width  0.6ms (low load), 0.7ms (high load)  121  nd 2  Gas pulse timing  Fixed during timing sweep  start 1 .5ms after end of first gas pulse  Results 1. EMISSIONS GIMEP  =  6bar. 1100rpm  Figures 1.1, 1.2, 1.3, and 1.4 show the power specific hydrocarbon, nitrogen oxide, particulate matter, and carbon monoxide emissions respectively at low load and 1100rpm. Operation with the 121 injector produces much higher hydrocarbon and carbon monoxide emissions than the J36 at this low load condition. However, the NO and particulate matter emissions are significantly lower with the 121 injector. It was not possible to get to the earliest timing of a 50% IHR at 5 degrees with the 121 injector due to combustion instability at this low load condition. The reason for this is unknown.  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  Heather Jones U1-FAC-093-TEST 121, 336, coinjector  DATE  I  I  20-01-2006  I  Alternative Fuels Group  $CliE Project  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T 1Z4  10  12  0J36 -0%EGR • J36 - 30%EGR 0121 0%EGR 30%EGR  10  9  ç,,  8  —  —  —  —  -  7  —L  —  •I21 -  —  -  6 5  —  —  —  —  —  —  —  0J36-0%EGR • J36 - 30%EGR 0121 - 0%EGR --‘ -30%EGR  —  —  —  :  ,-..  4  ----  C.)  —-----  -fzz:  2  0 4  5  6  7  8  12 9 10 11 50% IHR (deg ATDC)  13  14  15  0.10  —  2 —  —  —  0 4  5  6  7  0.08 0.07  •121-30%EGR-—-—-—-—  18 —  7  —  —  0.06  16  0.05  0.03  —  —  0 J36 - 0%EGR  + J36 30%EGR  0121-0%EGR  •-30%EGR  —  12  —  ———  —  —  ——.———  F  ç 10  .b —  —  6  -  —  .  08 C)  z;  0.01  16  -  —  ..‘  -  -—---------  -  -  —  0.02  15  %,  14  --7-  14  -  —  / -  13  —  —  -  0.04  10 11 12 9 50% IHR (deg ATDC)  Figure 1.2: NOx emissions at low load and mid-speed  -  0.  8  20  0 J36 0%EGR • J36 30%EGR 0121-0%EGR  0.09  —  —  3  16  Figure 1.1: tHC emissions at low load and mid-speed  —  4  —  -  -  -  2  0.00  0 4  5  6  7  10 11 12 8 9 50% IHR (deg ATDC)  13  14  15  Figure 1.3: PM emissions at low load and mid-speed  GIMEP  =  16  4  5  6  7  8 9 10 11 12 50% IHR (deg ATDC)  13  14  15  16  Figure 1.4: CO emissions at low load and mid-speed  l3bar. 1100rpm  Figures 1.5, 1.6, 1.7, and 1.8 show the power specific hydrocarbon, nitrogen oxide, particulate matter, and carbon monoxide emissions respectively at 75% load and 1100rpm. Generally, the two injectors are very comparable at this operating condition. The 121 injector gives slightly higher hydrocarbon emissions and lower particulate matter emissions that the J36 injector but the nitrogen oxide and carbon monoxide emissions are very close.  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  Heather Jones UI -FAC-093 -TEST 121, J36, coinjector  I  Alternative Fuels Group  S(’liE Project  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T 1Z4  0 J36-0%EGR,pilot=1 5mglinj C l21-0%EGR,pilot=l5mg/inj 0 J360%EGR,pilot=2OmgIinj C l21-0%EGR,pilot=2OmgIinj J36-0%EGR,8IiI2L  1.2  • J36-30%EGR,pilot=1 5mg/inj • l21-30%EGR,pilot= l5mg/inj x J36-30%EGR,pilot=2OmgIinj X 121-30%EGR,pilot=2omgRnj  10  8  1.0  >  0.8  S20)  • J36-30%EGR,pilot=1 5mglinj 0 J36-0%EGR,pilot=1 5mglinj • l21-30%EGR,pilot= l5mglinj 0 l2l-0%EGR,pilotl5mglinj 0 J36-0%EGR,pilot=2omglinj X J36-30%EGR,pilot=2omglinj C 121-0%EGR,pilot=2Omg/inj X 121-30%EGR,pilot=2omglinj J36-0%EGRpilot=8mgiinj —----------j  -—-------  —  —  -  —  —..—  0.6 F  C-) 0.4  —_  _fl  0 5  6  7  8  —  2  0.0 4  —  —  -—--—--1===__  0.2  —..  9 10 11 12 50% IHR (deg ATDC)  13  14  15  16  Figure 1.5: tHC emissions at 75% load and mid-speed  4  5  6  7  8  9 10 11 12 50% IHR (dog ATDC)  13  14  15  16  Figure 1.6: NOx emissions at 75% load and mid-speed  The effect of pilot fuelling amount on emissions can be seen in these figures as well. The J36 injector suffers higher particulate matter and carbon monoxide emissions with more diesel pilot added. Interestingly, more pilot fuelling with the 121 injector does not cause higher carbon monoxide emissions, only higher particulate emissions. It was not possible to decrease the pilot fuelling to 8mg/injection with the 121 injector due to combustion instability. The better atomization in the 121 injector really shows here with the lower particulate matter emissions with this injector. 0 J36 0%EGR, pilot=l5mg/inj • J36 30%EGR, pilot=l5mgfinj 0 J36-0%EGR,pilotl 5mg/inj • J36-30%EGR,pilot=1 5mg/mi 0 121 0%EGR, pilotl5mglinj •12l 30%EGR, pilot= 15mg/mi C 12l-0%EGR,pilot=1 5mg/mnj • 121-30%EGR,pilot=1 5mg/mnj -  -  -  C J36-0%EGR,pilot2omg/mnj C] 121-0%EGR,pilot=2Omg/inj J36-0%EGR,pilot=Bmglmnj  0.15  -  D J36 0%EGR, pilot=2omgnnj 0 121 0%EGR, pilot=2OmgIinj J36 0%EGR, pilot=8mgiinj  x J36-30%EGR,pilot=2Omg/mnj  -  X 121-30%EGR,pilot=2Omg/mnj  -  6  -  -  -  rL--  5 4  0.10  x J36 30%EGR, piIot20mgfinj x 121 30%EGR, pilot=2Omgfinj  3 0.05  2  0.00  0 4  5  6  7  8 9 10 11 12 50% IHR (deg ATDC)  13  14  15  Figure 1.7: PM emissions at 75% load and mid-speed GIMEP  =  16  4  5  6  7  8  9 10 11 12 50% IHR (dog ATDC)  13  14  15  16  Figure 1.8: CO emissions at 75% load and mid-speed  l3bar, 1500mm  Figures 1.9, 1.10, 1.11, and 1.12 show the power specific hydrocarbon, nitrogen oxide, particulate matter, and carbon monoxide emissions respectively at 75% load and 1500rpm. Generally, the hydrocarbon emissions are higher with the 121 injector and the particulate matter emissions lower. At this operating condition the effect of combustion timing has a much more dramatic effect with the 121 injector than the J36 injector. Hydrocarbon AUTHOR: DOCUMENT NUMBER: KEYWORDS:  Heather Jones U1-FAC-093-TEST 121, J36, coinjector  DATE  •1  Alternative Fuels Group  .  Project  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T 1Z4  emissions at 30% EGR are lower at early timing with the 121 injector and then much higher at late timings where as with the J36, hydrocarbon emissions are relatively constant over the timing sweep. Similarly, N0 emissions with the 121 injector are more strongly affected by timing; early timing gives much higher N0 than the J36 and later timing gives slightly lower N0. 2.5  10  J36-0%EGR,pilot=1 5mg/inj • J36-30%EGRpilot=1 5mg/inj 0121-0%EGR,pilot=l5mg/inj  2.0  8  J36-30%EGR,pilot7mg/inj  .—“  I!  .-  --------  =  x  0121-0%EGR,pilot=l5mg/inj  — —  1.5  =  J36-0%EGR,pilot=l5mgIinj • J36-30%EGR,pilot=l5mgIinj  9  •l2l-30%EGR,pilot=l5mg/inj -  \.  —--1———  —  a— 0.5  7  I 121-30%EGR,pilot=l5mglinj  —  6  --çJ3630%EGRPilot=7mhnJ  5 4  —  -  3  -  -  1ç  2  -  1  -  -  —  —  —  —  -  0.0 4  5  6  7  10 11 12 8 9 50% IHR (deg ATDC)  13  14  16  15  -  4  5  6  7  8  9  10  —  z  11  =  12  —  *  13  —  14  15  16  50% IHR (deg ATDC)  Figure 1.9: tHC emissions at 75% load and high speed  Figure 1.10: N0 emissions at 75% load and high speed  Carbon monoxide emissions also follow a much different pattern with the 121 injector compared with the J36. At early timing the carbon monoxide emissions are approximately 3 times lower with the 121 and then quickly rise to levels higher than with the J36 after mid-timing. • J36-30%EGR,pilotl 5mg/inj • 121-30%EGR,pilot=l5mg/inj  0 J36-0%EGR,pilot=1 5mg[inj 0 121-O%EGR,pilot=1 5mglinj 0.25  8  J36-30%EGR,pilot=7mglinj  • J36-30%EGR,pilot=lSmg/inj  • 121-30%EGR,pilot=l5mg/inj  7  0.20  5  0 J36-0%EGR,pilot=l5mg/inj C l21-0%EGR,pilot=l5mgñnj J36-30%EGR,pilot=7mgñnj  6  ;:::::::  0.15  a) C)  2  0.05  J;Z_  0.00 4  5  6  7  8  9  10  11  12 50% IHR (deg ATDC)  13  14  15  Figure 1.11: PM emissions at 75% load and high speed  16  0 4  5  6  7  8  9  10  11  12  13  14  15  50% IHR (deg ATDC)  Figure 1.12: Co emissions at 75% load and high speed  Particulate matter emissions consistently lower with the 121 injector. At high load with EGR, particulate matter emissions significantly increase with the J36 injector. Lower pilot fuelling is definitely key to decreasing these particulate emissions in the J36 injector. Figure 1.11 shows that a decrease in diesel pilot fuelling of 50% decreases particulate emissions by more than 50%.  AUTHOR:  DOCUMENT NUMBER: KEYWORDS:  Heather Jones U1-FAC-093-TEST 121, J36, coinjector  DATE  •1  Alternative Fuels Group  .  S’J?.E IrOjeCt  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T 1Z4  2. FUEL CONSUMPTION The fuel consumption is shown in Figure 2.1. At high load the 2 injectors have very similar fuel consumption and in fact at high load the 121 injector may have slightly better efficiency. However, at low load with EGR the 121 injector has very poor efficiency. It causes a fuel consumption of approximately 5-7% higher at early and mid timings and over 10% higher at late timing. The cause of the extremely poor operation of the 121 injector at low load with EGR is unknown but it seems to be related to the cylinder pressure. The cylinder pressure was much higher with EGR during these tests since we fixed power output and the overall equivalence ratio (similar to supplemental EGR).  :: :zz  ‘  190  —  210  --  ,  —  -  180  -  190  -  -  —  —  —  170  -  170  —  EEE  180  J36-O%EGR •J36 30%EGR 0121 0%EGR 30%EGR  C.)  -  LI. U)  C,  -  —  —  ..—  0 0 0  -  -  -  0• 0  • J36 30%EGR •121 30%EGR  -  200  —  0  J36 O%EGR 0121 O%EGR 0  •121  160 5  6  7  8  160  -  10 11 12 13 9 50% IHR (dog ATDC)  14  15  4  16  5  6  7  8  9  10  11  12  13  14  15  16  50% IHR (deg ATDC)  (b)  (a) 210  J36 O%EGR 0121 O%EGR  •J36 30%EGR •121 30%EGR  -  -  -  200 0  < 0  a, a,  —  190 —  —  —  —  -  —  180  -  —  —  — —  —  —  .—  U,  C.) U U)  170  C,  160 4  5  6  7  8  9  10  11  12  13  14  15  16  50% IHR (deg ATDC)  (c) Figure 2.1: Fuel consumption at(a) low load, 1100rpm, (b) 75% load, 1100rpm, and  (C)  75% load, 1500rpm  3. PILOT TIMING -121 INJECTOR The hypothesis was that a pilot pulse before the first gas pulse would be equivalent to a pilot pulse occurring after the second gas pulse in terms of operation. This is because the diesel is injected out of the gas sac so until the gas needle opens there is no diesel injection into the cylinder. To test this out the end of the pilot pulse was varied from 1 and 0.3 ms before the start of the first gas pulse to 1 and 5 ms after the start of the first gas pulse. AUTHOR: DOCUMENT NUMBER: KEYWORDS:  Heather Jones U1-FAC-093-TEST 121, J36, coinjector  DATE  J  I  Alternative Fuels Group  .  AS’Cl?li’ Project  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T lZ4  We will call the time between the end of the pilot pulse to the start of the first gas pulse the “pilot separation” (PSEP). Table 3.1 shows the power specific emissions for each of the variations in PSEP. The top of the column contains a picture showing where the pilot pulse is located in relation to the gas pulses. The pilot timing varied from before the gas pulse to after the second gas pulse. Due to limitations of the controller, the pilot pulse could not be moved further back than —5ms. So to get the pilot pulse to occur after the second gas pulse, the timing between the two gas pulses had to be shortened (case ‘d’) otherwise the time between the gas pulses was fixed at 1.5ms. Table 3.1: Emissions as a result of changing the pilot separation (PSEP) and the start of the first gas pulse.  (a)  (b)  I 0.87 0.41 0.008  (d)  (e)  I fllfl RHFI nI1  0.3 0.63 5.7 0.38 0.002  5.6  time between the end of pilot pulse  (C)  [IRH ‘inn PSEP (ms) Co (glGikwh) NOx (gIGikWh) tHC (glGikWh) PM (glGikWh)  —  5*  —1 0.94 6.3 0.43 0.003  5 1.29 8.1 0.51 0.002  0.77 5.8 0.43 0.004  Gas pulses Pilot pulse * 2nd gas pulse was moved closer to the first so that the pilot pulse occurred after the second gas pulse as we could not set a pilot pulse of less than -5ms on the controller Generally, all pilot timings give approximately the same emissions except for case ‘e’ where the pilot is injected nd while the gas needle is open on the 2 gas pulse. Figure 3.1 shows the corresponding pressure traces and heat release curves (averaged from 45 cycles). Basically all of the conditions run in the engine but case ‘e’ produces a sharp heat release likely because the pilot is igniting a pre-mixed gas mixture similar to HCCI combustion. Basically, as long as the pilot fuel is injected while a gas needle is closed the injector will behave the same. 200 180 —(a) PSEP=lms I  160  41D  —(b) PSEP=0.3ms —(C) PSEP=-lms  —(a)P=1ns —  —(d) PSEP=-5ms —(e) PSEP=-5ms  ‘3D  140  —)P=O.  (cI) PSSD=.51s  —  —(c)P=inE  (e) P=-&rs  120 (a 3I  ao  100 80  0  60 C)  1:  40 20 0 -30  -20  -10  0  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  10 20 CA (deg.)  30  40  Heather Jones U1-FAC-093-TEST 121, J36, coinjector  50  60  _  -10  0  10  2)  3)  40  5)  )  C(deg)  DATE  I  I  20-01-2006  I  •  Alternative Fuels Group  S’R1 Project  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T 1Z4  Figure 3.1: Cylinder pressure and heat release as the PSEP is varied (timing between end of pilot pulse and start of gas pulse) 4. COMBUSTION STABILITY Figures 4.1 through 4.6 show the coefficients of variance of GIMEP and of maximum cylinder pressure for each of the operating conditions. The COV Of Pmax at the late timing in Figure 4.2 must be disregarded as it is the maximum pressure is due to compression at this condition and not due to combustion. The combustion stability is comparable between the two injectors with the exception of two conditions; low load with 30% EGR and with high pilot fuelling. At low load with 30% EGR the CCV of GIMEP and Pm is up around 3.5% under the worst case (early timing). It was at this condition that high hydrocarbon emissions were also found. Figure 4.4 shows the COV of Pmax up around 3% with high pilot fuelling.  Figure 4.1: CCV of GIMEP at low load and 1100rpm 5.0  3.5 3.0 2.5  ‘C  > 0 C.)  2.0 1.5 1.0  o J36-0%EGR • J36-30%EGR 4.0 o 12l-O%EGR • 121-30%EGR 3.5 121-0%EGR,high pilot 3.0 121-30%EGR,high pilot 4.5  • J36-30%EGR o 121-0%EGR • 121-30%EGR 121-O%EGR, high pilot + 121-30%EGR,high pilot  4.0  C,  5.0  0 J36-O%EGR  4.5  Q LU  Figure 4.2: CCV of Pmax at low load and 1100rpm  —  2.0 1.5 1.0  -  0.5  -  0.0  0.5 0.0  4  6  8  10  12  14  50% IHR (deg. ATDC)  Figure 4.3: CCV of GIMEP at 75% load and 1100rpm  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  16  4  6  8  10  12  121, J36, coinjector  16  Figure 4.4: CCV of Pmax at 75% load and 1100rpm  DATE  Heather Jones U1-FAC-093-TEST  14  50% IHR (deg. ATDC)  I  Alternative Fuels Group  AcCI?E.’ Project  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T lZ4  5.0  3.5  4.0 —.  3...,  3.0  3.C  2.5  2.5  2.0  2.0  1.5  1.5  1.0  1.0  .i  0.5  K J36-0%EGR • J36-30%EGR O 121-O%EGR • 121-30%EGR  4.5  • J36-30%EGR o 12I-O%EGR • 121-30%EGR  4.0 —  5.0  K> J36-0%EGR  4.5  0.. 0.0  0.0 4  6  8  10  12  14  4  16  6  10  8  14  12  16  50% IHR (deg. ATDC)  50% IHR (deg. ATDC)  Figure 4.6: CCV of Pmax at 75% load and 1500rpm  Figure 4.5: CCV of GIMEP at 75% load and 1500rpm  5. EFFECT OF PILOT FUELLING AMOUNT ON COMBUSTION Figure 5.1 shows the effect of pilot fuelling on the J36 injector at 75% load, 1100rpm and no EGR (averages of 45 cycles). Basically as the pilot is increased the initial heat release moves earlier and releases more heat. However, the ignition delay is slightly longer with more pilot fuelling. 180  140 -7mgñnj  —7rrg/ir  160  l5mglinj  —15lTir  g14o  —  20iTirj  E 120  I: €60  40  40 20  20  0 -30  -20  -10  0  20 10 CA (deg.)  30  40  50  60  0 -30  -20  -10  0  10  20  30  40  50  60  CA (deg.)  Figure 5.1: Pressure trace and heat release rate of J36 injector without EGR, 1100rpm, GIMEP=l3bar Figure 5.2 shows the pressure trace and heat release of the 121 injector under the same operating condition (45cycle average). The 121 injector needs more pilot fuelling than the J36 so 7mg/injection was not possible. The injector runs well with 15mg/injection but as the pilot fuelling is increased to 20mg/injection there is some kind of “ringing” within the cylinder, this is evident in the large fluctuations seen in the pressure trace and heat release curve. It is unclear why this happens.  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  Heather Jones U1-FAC-093-TEST 121, J36, coinjector  DATE  •1  Alternative Fuels Group  5r(YljE Project  Department of Mechanical Engineering, University of British Columbia 2054—6250 Applied Science Lane, Vancouver, B.C. V6T 1Z4  140  : 80  —l5mglinj  ,  500 450  400 a) V ,-350 .  ////‘\  •  -,300 g250  .2200 4)  V  4o  150 100  20  50 0  0 -30  -20  -10  0  20 10 CA (deg.)  30  40  50  60  Li  -30  —15mgnj  -20  -10  0  10 20 CA (deg.)  30  40  5060  Figure 5.2: Pressure trace and heat release rate of 121 injector without EGR, 1100rpm, GIMEP=l3bar  Conclusions 1. The 121 injector gives emissions levels comparable to that of a J36 under most of the tested operating conditions with the exception of low load where the hydrocarbon and carbon monoxide emission were excessive. 2. The 121 injector in general always gives lower particulate matter. This is likely due to the better diesel atomization with this injector. 3. Timing of the pilot injection is not very sensitive. As long as the gas needle is closed when the pilot is injected into the gas sac, the emissions are very similar. 4. Higher and lower pilot fuelling amounts prove to be troublesome with the 121 injector. High pilot fuelling produces a “ringing” in the cylinder and the injector does not run with low pilot fuelling (less than about 12mg/injection).  Recommendations Basically the 121 injector has proven to have promise as a potentially low cost alternative to the J36 injector. Much more work needs to be done to redesign the injector so that the performance is better under all operating conditions. It is recommended that work be continued on the development of this injector.  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  Heather Jones U1-FAC-093-TEST 121, J36, coinjector  I  I  ILI 20-01-2006  Alternative Fuels Group Department of Mechanical Engineering, University of British Columbia 2324 Main Mall, Vancouver, B.C. V6T 1Z4  S”€1?E’ Project  Key Question: Can the 121 injector run at low load, and is there a minimum diesel pilot flow-rate under these conditions?  Method: The prototype 121 injector was run at UBC’s SCRE facility. The operating condition chosen was to test the low-load operation of the injector. To provide a baseline operating condition, the engine was operated at 800 RPM, 8.5 bar GIMEP with an intake manifold pressure of 65 kpa (g). A single pilot injection preceded a 2-stage gas injection process. The durations of the diesel pulse and the first gas pulse were semi-arbitrarily set for the lowest first gas pulse which retained stable operation. Timing was set for the mid-point of the heat-release rate (50%]HR) at 10°ATDC. For the low-load tests, the 2’ injection event was terminated, while the timing and duration of the jst gas pulse was held constant. The diesel end-of-injection timing was held constant, but the duration was adjusted to provide the desired quantity of diesel. The manifold pressure was then reduced in 2OkPa increments from 65 to 5 kPa (g). At each manifold pressure, three pilot flows were tested —30 mg/inj, 20 mglinj, and a ‘minimum’ which was selected as being the lowest pilot flow at which there was no evidence of the engine misfiring (for a number of conditions, this minimum was at, or above, 20 mg/inj). This third parameter was somewhat subjective, and the stability at this condition varied with the different test conditions. This procedure was carried out at 16.5, 22.5, and 27.5 MPa gas rail pressure (18.5, 24.5, and 29.5 diesel rail pressure). The testing was not randomized, with the 22.5 MPa testing carried out on 2 1/03/06 and the other two on 22/03/06. The manifold pressures were tested sequentially at each injection pressure. Uncertainties relating to this testing include the standard uncertainties relating to testing on the SCR.E, as well as: i) testing durations (3-4 minutes) were the minimum which have been shown to provide stable diesel mass flow measurements. Errors in this flow rate, in particular at low pilot flow conditions, were substantial ii) operation of the engine at low-load tends to result in many instruments (including the gas and air flow-rates) being closer to their limits-of-detection, and as a result the uncertainty in their readings tend to increase. The baseline injection parameters used in this testing are given in the table below: Gas Rail Pressure 16.5 22.5 27.5 Diesel Rail Pressure 18.5 24.7 29.5 Gas SOl (°CA) -10 -9.5 -7 0.75 Gas PW (baseline, ms) 0.7 0.6 Pilot EOI (baseline, °CA) -13 -13 -10 Parameters not included in the table, but held constant for all tests included the manifold air temperature (--28°C), the end of pilot-first gas pulse separation (0.7 ms), EGR level (0).  Discussion: The first objective of this testing was to determine whether the 121 injector was capable of running stably at low loads. In particular, concern had been raised based on previous testing that the injector would not function at near-atmospheric conditions. The coefficient of variation (COV) of the GIMEP under minimum and high diesel pilot flows are shown in Figures 1&2. Also shown in Figure 1 is the COV of GIMEP for the low injection pressure at a diesel pilot of—20 mg/inj, which is roughly equivalent to the ‘low’ pilot flows at each of the higher injection pressures. These results indicate that while high variability in the combustion may occur at low pilot flows, increasing the pilot flow will substantially reduce this variability. The pilot flow rates corresponding to these low flows are shown in Figure 3. As can be seen, the high variability at the lowest injection pressure is attributed to the low pilot flow rate. By increasing the flow (to approximately 20 mg/inj), a substantial reduction in combustion variability is achieved. In general, observation of the plots suggests that lower manifold pressures result in higher combustion instability for a given diesel pilot flow. That lower pilot flows were achievable with the low injection pressure case may be due to the lower gas flow at this condition, as shown in Figure 4. This suggests that an important parameter may be the ratio of diesel pilot to gas (in the first pulse). However, further testing is required to investigate this hypothesis in more detail. While emissions measurements were not a major objective of this work, including the HC and NOx emissions provides further insight into the combustion stability. In general, high hydrocarbons (in this case) can be attributed to high combustion variability, whereas high NO,, will indicate more stable, earlier, and more rapid combustion. Figures 5&6 show the HC emissions, with 5 for the ‘low’ pilot flow conditions and 6 for the ‘high’ pilot flow case. The equivalent NO,, emissions are shown in figures 7&8. The results agree with the previous assessment that the higher pilot flow results in much hotter, more stable combustion. This leads to high NO,, but low HC emissions.  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  G.P. McTaggart-Cowan  U 1 -FAC-09 8-Test 121 injector; low load  I  06-03-23  Alternative Fuels Group Department of Mechanical Engineering, University of British Columbia 2324 Main Mall, Vancouver, B.C. V6T 1Z4  .Sr(lj.E Project 1  The in-cylinder pressure traces tend to support these results. The pressure trace and heat-release rate for the low and high pilot flows for the low and high injection pressures at the low and high manifold pressures are shown. At the low injection pressure, the high pilot flow induces such a rapid heat-release (approaching detonation conditions) that ‘ringing’ of the in-cylinder pressure measurement is observed. This effect has been observed previously with this injector, and appears to occur for those cases with very high rate-ofincrease of the in-cylinder pressure during the initial combustion event. Whether the ringing is actual pressure waves in the combustion chamber or is a mechanical or electrical effect in the pressure transducer is unknown. Similar ringing is observed at the other high-pilot conditions except for the low manifold pressure, high rail pressure case. It would appear that in this case, the initial rate of heat release is somewhat lower and as a result the pressure rise is not as rapid. At the lower diesel flow rates, the ignition process appears to more closely resemble that of conventional HPDI combustion. Under certain conditions, there even appears to be an early first-stage combustion, followed by the main combustion event (for example, in the high injection pressure, low manifold pressure case). However, the duration between the initial and main heat releases are relatively short. Even at this condition (where ringing is not observed) the higher pilot flow can be seen to substantially increase the combustion rate. These results suggest that, at low load, the higher pilot flow is substantially increasing the initial heat-release rate of the premixed combustion phase. Due to the low load condition, the combustion is occurring primarily in the premixed phase. For the lower diesel quantities, the ignition delay is substantially increased and the peak heat-release rate is reduced. As the diesel flow gets very low (as shown in Figure 9), the overall combustion rate is greatly impaired. This is most likely due to the relatively small quantity of diesel (relative to the natural gas mass). It is likely that this small diesel quantity is more dispersed within the natural gas, impairing the ignition process. With the longer delay, the ignition also becomes more variable, resulting in some cycles with very long ignition delays (the limiting value which is approached is cycles where no ignition occurs: however, for the test points here, the conditions were selected to attempt to avoid such misfiring cycles). The effect of the relative amounts of diesel and gas (in the first injection) are shown in Figures 10-13. The effects of both the mass and the volume ratio are shown, with the COy IIvIEP, tHC, CO and NOx emissions as outcomes. The diesel volume was calculated assuming incompressible fluid at a density of 848 kg/rn . The gas density was estimated using the ideal gas law at the peak cylinder 3 pressure and at ambient temperature. While this calculation is questionable (the injection process occurs before peak pressure; the gas will certainly be at a higher temperature than ambient when injected), however it is representative of the volume of gas injected per cycle. The relative volume of gas is shown to have a very significant influence on combustion stability and emissions. As expected, increases in the volume ratio (more gas to diesel) resulted in higher combustion variability, higher unburned fuel and CO emissions, and lower NOx emissions. The role of the mass ratio can be seen to be substantially less significant, with no clear trends in NOx or combustion stability, and only very rou h trends in CO and unburned fuel. The correlation coefficients are given in the table below: tHC Parameter COV CO NOx GIMEP 0.669 0.824 0.877 -0.770 p(Volume Ratio) 0.406 0.585 0.63 -0.486 p(Mass Ratio) It should also be noted that the cross-correlation (Volume ratio mass ratio) is also strong, as would be expected, with an p-value of 0.87. Given this strong cross-correlation, it is very significant that all the outputs are much more strongly correlated with the volume ratio than with the mass ratio, indicating that it is the relative volumes of the two fuels which are most significant. —  Conclusions: 1) The 121 injector was shown to run successfully at low load conditions down to ambient manifold pressures. The engine was also started under normal (naturally-aspirated) conditions without undue problems. 2) A lower diesel mass limit of around 10-20 mglinj was identified for most operating conditions. This depended on the amount of gas being injected, the manifold pressure, and the injection pressure. In general: a) the lower the manifold pressure, the more diesel was required b) the higher the injection pressure, the more diesel was required c) the more gas was injected (in the first pulse) the more diesel was required 5t 1 3) The relative volumes of the diesel and gas ( pulse) injections had a strong influence on combustion stability and emissions, with larger volumes of gas reducing stability and increasing HC and CO emissions. 4) This suggests that it may be possible to minimize diesel consumption by reducing the natural gas in the first gas pulse. Further power nd may be developed by increasing the duration of the 2 pulse. 5) Transition from single gas pulse to double gas pulse operation proved to be sensitive to operating condition, with the potential for even very late 2” injections (as much as 3 ms after the first pulse) still being sufficient to stop the combustion event. This is thought to be a result of injector dynamics. AUTHOR: DOCUMENT NUMBER: KEYWORDS:  G.P. McTaggart-Cowan  Ui -FAC-098-Test 121 injector; low load  DATE  .i  Alternative Fuels Group  .  S’(RE Project  Department of Mechanical Engineering, University of British Columbia 2324 Main Mall, Vancouver, B.C. V6T 1Z4  Recommendations: 1) No attempt was made to optimize the injection process for low-load operation. Adjustments to the pilot-gas separation time, the absolute timing of the injection, or the diesel-gas rail bias could have substantial impacts on the overall combustion system, and hence require further investigation. 2) The response of the prototype injector to the specified commands was not always well understood. Further testing of the injector, either in the UBC spray rig or on the Westport rate tube, could provide more information regarding the injector’s actual performance. 3) Pressure pulsations in the gas rail were observed to be significant. It would be interesting to study the effect of the double-pulse injection behaviour on the rail pressure with a high-pressure, high-speed transducer. This could provide important information for both the 121 and conventional HPDI programs. Figures: -44-- 16.5 MPa  35 30  —-a.-- 16.5 (20 mg/mi)  6  —a— 22.5 MPa 27.5 MPa  5  -44-- 16.5 MPa________________  —a--— 22.5 MPa \\  ....--.  “low” diesel  .--.---  30 mg/inj  27.5MPa  25 0  w  C-)  I I,  20  ——  w  ,1  :  II  \  2  a_____EJ  x  1  5  4  0  0 80  40  60  20  80  0  60  Fig. 1: COV GIMEP at various injection pressures over a range of manifold pressures at ‘low’ diesel pilot flow (<2Omg/inj) —44-— 16.5 FPa  22.5 NFa GPW 22.5 MPa GPW .-.+.- 3OMPa —v-— —.—  35 30  E  25 E 20  g 20 a)  E  x.  0  15  :E  15  —— —  10  —44-— 16.5 MPa  10 —  .  0  Fig. 2: COV GIMEP at various injection pressures over a range of manifold pressures at ‘high’ diesel pilot flow (30mg/inj)  0.7 0.6  = =  20  Intake Manifold P (kPag)  Intake Manifold P (kPag)  30  40  —D-—  —  5  5  22.5 MPa GPW lPa GPW  = =  0.7 0.6  0  0 80  60  40  20  0  Intake Manifold P (kPag)  Fig. 3: “Minimum” diesel injection mass for the various injection and manifold pressure conditions. “Minimum” semi-arbitrary selection as point at which engine was not completely misfiring  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  80  60  40  20  Fig. 4: Gas injection mass corresponding to diesel flows indicated in figure 3.  G.P. McTaggart-Cowan  DATE  U 1 -FAC-098-Test 121 injector; low load  0  Intake Manifold P (kPag)  06-03-23  I  I  Alternative Fuels Group  AS(lili .I’roject  Department of Mechanical Engineering, University of British Columbia 2324 Main Mall, Vancouver, B.C. V6T 1Z4 —4<-— 16.5 MPa ... — 16.5 (20 mglinj) —9-— 22.5 FVFa 27.5 Fa — ——  250 200  “low” diesel  a)  \\  0)  150  0)  0)  0  C-) x  MPaI  —a-— 22.5 MPa• 27.5 MPaI  16  •--•--•  a)  —4<-— 16.5  18  30 mg/mi  .....  14 12 10 8 6  4 2 0 80  —  a) .  0) ,  0)  40 35  50 40  “low” diesel  0)  30  0)  15 10  --  30  --)<--16.5MPa —9-—  22.5 MPa___________________  --.4-.  27.5 MPa  10  Z  --..-----  5  -x 9 _—----.-  0  0 80  40  60  20  0  80  Intake Manifold P (kPag)  Fig. 7: NOx emissions for ‘minimum’ diesel pilot flows 30 mg/mi diesel I 20 mg/mi diesel PgasrajI= 16.5 MPa jIoieseIJ P5kPa 60 50  60  40  20  0  Intake Manifold P (kPag)  Fig. 8: NOx emissions for ‘high’ diesel pilot flow 350  70  Cu  ____-(  35  15  •  5  —  x 25 0 z 20  \  3Omg/inj  >‘-.-----  45 a)  25 0 z 20  0  Fig. 6: HC emissions for ‘high’ diesel pilot flow  —4<-—16.5MPa —-a---- 16.5 (20 mg/inj) —9— 22.5 MPa -.-#.. 27.5MPa  ri  20  Intake Manifold P (kPag)  Fig. 5: HC emissions for ‘minimum’ diesel pilot flows 50 45  40  60  Intake Manifold P (kPag)  300 250  0  40  200  D C’) 0  30  150  a-  20  100  10  50  a)  0  0 -60-3003060  Crank Angle (0CA)  -50 Crank Angle (0CA)  Fig. 9: P and FIRR for low manifold pressure, low rail pressure.  AUTHOR: DOCUMENT NUMBER: KEYWORDS:  G.P. McTaggart-Cowan  DATE  Ui -FAC-09 8-Test 121 injector; low load  06-03-23  I  I  Alternative Fuels Group  Project  Department of Mechanical Engineering, University of British Columbia 2324 Main Mall, Vancouver, B.C. V6T 1Z4  100  400  P=65kPa Paas = 16.5 MPa  90  —30 mglinj diesel I  350  20m1Inidiesj  ra  80  300 C.)  •  J  70 0  —30 mglinj diesel x 20 mglinj diesel low diesel  250  fj  200  50 U,  40 0  .  30  z  20 10  150  Un  Fh  aL____  0 -60  -30  30  0  ;-  60  Crank Angle (0CA)  Il_ Crank Angle (OCA)  Fig. 10: P and HRR for high manifold pressure, low rail pressure. 60  :::E  P=5kPa Pgas rail = 27.5 MPa  50  100  40  I  0  30 U) In  a. 20 10  0 F  1X  20  1  lx  0 0 -  -60  0  -30  30  60  -i6”o  10  20  -20  Crank Angle (oCA)  Crank Angle (0CA)  Fig. 11: P and HRR for low manifold pressure, high rail pressure. 700  120  30 mglinj diesel I  P=65 kPa Pgas ra 27.5 MPa  100  600  =  30mgdieseI x 2Omg[injdieselj  500 C.) 0  400  C.)  E  300 200 100  -60  -30  0  30  60  Crank Angle (oCA)  -100 Crank Angle (0CA)  Fig. 12: P and HRR for high manifold pressure, high rail pressure. AUTHOR: DOCUMENT NUMBER: KEYWORDS:  G.P. McTaggart-Cowan  Ui -FAC-098-Test 121 injector; low load  DATE  1  I  06-03-23  Alternative Fuels Group  SCRE l’rojeCt  Department of Mechanical Engineering, University of British Columbia 2324 Main Mall, Vancouver, B.C. V6T 1Z4  35  35  30  30  ?  25 aUi  20  w 20  C 15 > 0 C>  25  a . .  10  5  o  15 > 0 o 10  .  •  .  . .  . .4  ..  5  •.:i  0  .  •  •  1)4•  *4  0 0  40  20  60  0  0.5  1  Volume ratio (VcngNdies)  1.5  2.5  2  Mass ratio (mcng/mdies)  Fig. 13: COV GIMEP, for all points vs. volume and mass pilot-gas ratios. 250 200 ‘  150  ‘  •g’  .  0)  x  .  .;  •1)  ..  •  100  150  B.  •  •  .•  50  •4  ...,,: 0  4  . 4.  .  •  .  •  c3H00  •  0 40  20  *.  0  60  4  *4  .  • •  0.5  1  Volume ratio (llcngNdies)  1.5  2  2.5  2  2.5  2  2.5  Mass ratio (mcng/mdies)  Fig. 14: tHC emissions, for all points vs. volume and mass pilot-gas ratios. 250  250  200  200 •  •  4  4  4, ‘  150 .  0  •  ‘V.,  + ••. •• ..  50  •  •.  0 100 0  .  .  ••• •:  0  .  .  0  .  4 •  50  • •  • .. • • . •+.  4  .4.  4  0  0 0  40  20  60  0  0.5  1  Vobirne ratio (VcngNdies)  1.5  Mass ratio (mcnglmdies)  Fig. 15: CO emissions, for all points vs. volume and mass pilot-gas ratios. 50 45 40 _35 D ;30  50  I  .  45 40 35 30 25  •  1 •  . . .  25 ‘a,  .  ..  10 5  •  20  -  Z15  .  5* .• •  .4 •  4  20  .  40  . •  10 5 0  I,  0  .4  60  .  •‘  4  4.  0  0.5  Volume ratio IcngNdies)  ., •  .  1  •  1.5  Mass ratio (mcnglmdies)  Fig. 16: NO emissions, for all points vs. volume and mass pilot-gas ratios. AUTHOR: DOCUMENT NUMBER: KEYWORDS:  G.P. McTaggart-Cowan  Ui -FAC-09 8-Test 121 injector; low load  DATE  I  Appendix B.3  —  SCRE Timing Factsheet  This appendix (pp. 190— 196) has not been included because of copyright restrictions, It contained the following information: •  Documentation on the calibration of the optical shaft sensor to TDC  •  What crank angle offset should be used for the SCRE in calculations with the indicated pressure curve  This factsheet is available upon request. Hill, P.G. SCRE Timing Checks, W1-FAC-3788ANYS, Westport Innovations Factsheet. December 2007.  190  [Original document missing pages 191-196]  Appendix E: Emissions Spreadsheets The Emissions Spreadsheets are organized as follows: •  Appendix E. 1  •  Appendix E.2  •  Appendix E.3  •  Appendix E.4  •  Appendix E.5  •  Appendix E.6  •  Appendix E.7  —  —  —  —  —  —  —  Vu-A tests at 800 RPM Vu-A tests at 1200 RPM Vu-B tests at 800 RPM Vu-B tests at 1200 RPM VIII-A tests VIII-B tests VIII-B2 tests  On the emissions bench the top section has the test series number, test name, and date and time. The date and time format is in the same format that can be used to find the raw “slow” data files. For example, test series Vu-A- 1 has one test name “31-16-10-47”. The raw data for this file can be found in the electronic appendix (...rogak/sbrownlThesis/Brown_Thesisl) under the filename “VII-A-800/slowO7-09- 13-14.24.1 4.csv”. For VII series tests, the test files are organized first by injection pressure, then by pilot gas pulse width duration, then by diesel injection mass. The pressure traces and heat release rate plots for specific test points for Figures 4.8 to 4.14 can be found most easily through the diesel fuelling rate. These heat release figures with the accompanying “b” and “c” test modes can be found in Appendix F.1 to Appendix F.4. For test series VIII tests, the emissions spreadsheets are organized by mode number.  197  00  C  N  ç  <  I  C  N  <  70a2  47a 40-16-2070a  11.13.47  14.52.54  14.41.55  14.33.35  .—  > -  -  00 N  N  (_) 0  00  N 00  C  .  C’ C j-  ,‘  Lfl  ‘  .j.  V  \C  %C  c  .  00  c .-  -  I  -  -  -  E’-  L)  ‘  ‘i)  N  “I  —  ‘..O N  D.  -  CS  -  Lf  -  4  N  .  -  N  N  .  C  (1)  )  -  )  c•9  )  )  L)  -  07-09-13 14.24.14  N  345GRIT070913  37-16-1047  >  <  —  a  Test Name  a  34-16-10-  > —  )  ;  F  00  q -,  ‘— \_  ‘—,  OL  :r  PLl8lILO  po6coI  \_  .  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(bar) CA@Pk.press.(bar) Gross IMEP (bar) EQR 5%IHR(deg) 10%IHR(deg) 50%IHR(deg) 90%IHR(deg) COVGIMEP DSOI (deg) DEOI (deg) GSOI (deg) GEOI (deg) 2GSOI (deg) 2GEOI(deg) Comments  C  —  -  I  -  •  Vi  I -  ‘Jis  2.87 2.95 2.70 2.28 1.73 1.58 1.95 2.2 1.7 1.7 1.3 2.3 3.2 1.2 1211 1097 1041 1489 970 837 833 1.1 1.0 4.0 -0.2 0.6 1.1 0.5 872 883 878 877 879 866 881 60 60 61 61 60 60 60 93.6 93.8 95.6 92.7 93.5 93.9 92.8 17.4 17.2 17.5 17.3 17.4 17.3 17.4 18.3 18.3 18.3 18.3 18.3 18.3 18.3 57.7 55.9 51.2 49.0 58.4 50.8 57.0 134 135 133 135 133 133 135 117 116 117 116 116 117 117 0.22 0.64 0.52 0.41 0.33 0.53 0.41 1985 1268 2011 1577 831 2397 1561 0.97 1.00 1.77 1.64 1.30 2.19 1.42 46 217 267 308 150 126 35 2.62 2.53 3.45 3.14 3.21 4.37 2.84 562 500 812 626 547 992 485 16.54 16.66 14.88 15.40 15.54 13.32 15.93 174 126 252 255 147 133 173 122 123 77 64 61 57 86 210 206 270 250 248 320 232 77.0 74.2 84.7 81.3 77.5 88.4 77.2 14.1 14.5 14.7 13.6 13.6 7.8 13.0 4.36 4.29 6.46 5.84 5.63 8.17 5.17 0.22 0.21 0.31 0.28 0.26 0.37 0.24 9.1 9.6 8.6 7.1 2.1 2.1 4.6 9.6 10.1 9.1 8.1 2.6 2.6 6.1 11.4 10.6 11.3 11.9 10.5 10.3 10.4 14.1 15.1 15.6 15.6 18.1 17.6 16.6 4.2 10.9 4.1 10.0 8.7 9.9 7.7 -31.0 -29.0 -41.0 -32.0 -31.0 -41.0 -22.0 -15.1 -15.3 -14.8 -15.2 -15.2 -15.0 -15.7 -8.0 -8.0 -8.0 -8.0 -8.0 -8.0 -8.0 -5.5 -4.3 -4.4 -4.3 -5.5 -5.5 -5.3 -2.8 -2.5 -1.4 -1.8 -1.3 -1.3 -1.3 4.9 3.6 3.8 4.6 5.1 5.0 5.1  C  I  •  0  C  I  .  —  C I  C  C  0 I  —  •  Vi  J t.J  )  -.  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N  -  t.  ‘  N  N  00 \D  00  •  N  N  —  I  I  I  I  00  -  ooc9©ocN  2  OOcN 00 00 N ‘r .— cr N I I N  —  00000oC  ——  ‘.-.-  cl  Noo-NO  ,  9’  ‘.-.  ‘—.  ‘©‘©I©’c.)  :I I.I_II_I_I_I_I>  ‘—  1)  v  -  -  Lr-NNDNNOO.  .NON©  N  :  N 00 C”! t  © N — •.r-N° N ‘r  -  .0  N  “ ON”! N ‘.r a N N  ——.—  ooNN-So  2d  NC’  —  13.05.10©NN—--N  07-12-19o-ooooN  29-10-70 07-12-17 16.45.15 a 29-1070a 29-20,Oa 29-107OLBa15.11.25 29-10-  70a-  29-1047a  s  “—  O7-12-19  47a  Test  !  29-20-  > i c i n co 291047071247 al —  > c  5 ‘  C,)  I  I  I  I  I Name I  ()I G)I  IEI  0  00  C’.  “C 00  00  N N  C 00  1) I z  ?  43ia  14.13.25 1810  7°; 18-107OLBa 18-107OLBa  --  zaftera  18-10-  JZe  N a’. QO  N N  08-01-14 e 07-12-18 07-12-18  C  ,—.  a’ —  ‘1- a’ a’ c’i N N C  00  N N N a’.  Ce  V .  -  V  —  N  ‘  1) <  —  C  N  ‘•‘  i-  a’.  ,—‘  o  ..  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UI  )  1.61 1.79 2.20 160 1624 1599 2.21 1.15 1.13 0.79 0.61 1.00 1197 1209 1221 66 65 65 95.7 92.3 93.1 22.8 23.2 23.3 23.9 24.1 24.3 54.0 53.3 60.4 161 157 159 174 179 176 0.63 0.29 0.32 17.62 8.12 8.85 4.24 3.41 4.40 54 71 52 5.96 4.56 6.37 1281 969 1241 10.47 12.80 9.83 98 129 99 52 71 59 450 363 473 98.7 92.1 101.1 14.2 13.5 2.7 11.09 8.79 11.35 0.41 0.19 0.23 -3.4 3.6 -4.4 3.6 5.6 -3.4 10.8 10.3 10.4 21.1 18.6 24.1 4.2 2.0 3.3 -58.0 60.0 60.0 -22.1 77.4 78.3 -17.0 -15.9 -17.0 -13.6 -12.5 -11.9 -5.5 -4.7 -5.5 2.8 2.5 3.7  2.42 1443 0.73 0.27 1204 63 93.1 23.3 24.5 55.3 168 187 0.33 9.04 3.35 595 4.35 785 13.22 475 236 357 72.0 8.2 7.75 0.52 10.6 11.6 16.4 24.6 6.6 -42.0 -21.1 -10.0 -4.9 -2.0 4.5  0 .  Ui  .)  t’)  -  --  .  0Qo-  OOO  2.07 120 1.11 0.33 1210 64 93.4 23.9 24.7 57.8 156 174 0.29 8.09 4.07 45 5.63 917 11.00 103 61 439 81.9 17.8 10.29 -0.28 -1.4 8.6 14.8 24.6 3.8 -35.5 -27.9 -17.0 -13.6 -1.5 6.8  O  0 --  1.49 219 1.63 0.62 1209 59 94.2 23.8 24.6 53.5 161 179 1.37 15.00 3.11 36 4.83 953 12.62 118 71 367 91.3 14.0 9.01 0.23 -3.9 4.1 10.6 19.6 3.7 20.0 41.0 -17.0 -13.6 -6.0 2.3  0 .I UI  C .) UI  t  0  9)  9) o 0 ‘.CO C C t’JCc,i  -‘ .  —  c©  -.  0  -  —  0 0 UI t.) .)  -  0  c9C C C  c.-  c UIoo  1.88 2386 1.40 1.00 1206 65 93.6 23.3 24.2 64.8 157 174 0.41 11.24 5.63 195 7.48 1350 7.57 68 39 549 106.3 12.7 13.39 0.30 0.1 2.6 10.4 24.6 4.0 -42.0 -26.1 -17.0 -11.9 -4.3 4.1  1.66 1.49 1997 1579 1.87 1.70 0.95 1.00 1220 1214 64 65 93.3 93.4 23.6 23.1 24.4 24.0 59.1 61.3 158 176 175 195 0.43 0.48 11.83 13.29 4.45 3.39 89 171 6.04 4.73 1132 889 10.25 12.65 118 175 68 90 457 376 97.1 87.4 12.4 12.4 11.18 8.72 0.42 0.28 -1.4 2.6 0.1 3.6 10.5 10.9 22.6 20.1 3.0 4.8 -35.5 18.0 -27.8 39.9 -17.0 -10.0 -11.9 -4.9 -4.5 -2.0 3.9 4.6  L  .  —  0  —  C  — ..  L%) t’J  o.  ..  .) t’.)  C C —  Ui  ..D L’J  --  c  1.65 1.48 1.44 2438 1819 613 2.82 2.61 5.51 0.91 1.10 1.00 1201 1211 1216 65 65 65 96.4 92.8 94.2 22.9 23.2 23.2 23.8 24.1 24.1 58.1 53.3 55.8 160 159 160 177 177 177 0.52 0.61 0.72 14.51 16.84 19.84 5.71 3.51 3.81 49 201 267 7.63 4.63 5.45 1284 1015 1263 7.41 12.66 11.40 129 51 192 31 104 75 408 555 363 107.6 94.0 100.7 10.7 13.9 5.8 10.24 13.72 8.92 0.32 0.46 0.47 -3.4 2.1 -3.9 -2.4 5.1 -2.9 10.4 10.6 8.6 26.1 18.1 20.1 4.2 2.8 1.5 -55.0 60.0 -45.0 -26.2 77.9 -15.8 -17.0 -15.9 -15.9 -12.0 -10.8 -10.8 -4.3 -5.0 -5.0 4.0 2.6 2.7  (109,i 10)  206  Figure # T est s eries -  32 35 38 41 44 47 50 62 65 53 56 59 Vu-B- Vu-B- VII-B Vu-B- Vu-B- Vu-B- Vu-B- Vu-B- Vu-B- Vu-B- Vu-B- Vu-B 33 33 34 34 34 35 35 35 36 36 36 t3 00  -‘  i’.) 00  -  t’.) 00  -  .  C 0)  . ...j  .  0)  0)  00  00  00  0  -  0  0)  0)  0)  C  —  :  0  C  -.  00 t3 0  00 t’. 0  . -  . -..  .  0)  0)  0)  C  C  —  00  0  00 t.) 0  0)  0)  0)  00  00  0  C  .-  —  —  C  0  C  —  C t’J tJ • I -  C C •  ‘.000  -oo  —  C  -‘  C  .  1-., + LIaLe,  rime  ::::..  •---:-  IgnitionDelay(ms) IHR(kJ/m3) Knock (bar) IHR Ratio Engine Spd.(rpm) MAT(°C) MAP (kPag) CNGPress.(MPa) Diesel Press. (MPa) ExhaustPress.(kPa) Corr.Airflow(kglhr)  Airflow (kglhr) Diesel flow (kg/hr) Dieselinj.(mg/inj) CNGflow(kg/hr) CO (ppm-dry) C02 (%-dry) NOx (ppm-dry) 02(%-dry) CH4(ppm-dry,C1) tHC (ppm, Cl) ExhaustT.(°C) Pk. press. (bar) CA@Pk.press.(bar) GrossIMEP(bar) EQR 5% IHR (deg) 10% IHR (deg) 50%IHR(deg) 90%IHR(deg) COy GIMEP DSOI (deg) DEOI (deg) GSOI (deg) GEOI(deg) 2GSOI (deg) 2GEOI (deg) Comments  t’.)  (.)  ) .j I  ..  I  .oo  Coo  1.75 1486 1.69 0.63 1194 65 94.1 29.0 29.9 46.8 157 174 0.34 9.48 3.06 209 4.28 832 13.37 148 77 338 90.2 13.2 8.19 0.38 -1.4 0.6 10.4 17.6 2.5 -32.0 -27.2 -17.0 -13.6 -1.3 3.8  1.70 1506 1.75 0.60 1197 65 93.3 29.4 30.0 47.5 156 174 0.41 11.37 3.08 194 4.35 844 13.28 144 74 342 90.5 13.1 8.35 0.49 -1.4 0.1 10.4 17.6 1.8 -33.0 -28.0 -17.0 -13.6 -1.3 3.8  C  C -..  I  1.58 952 2.98 0.55 1201 66 94.9 28.5 29.5 52.2 163 181 0.76 21.15 1.90 345 3.05 433 15.67 170 108 259 75.5 14.0 5.27 0.16 6.1 6.6 12.7 19.1 5.6 -29.0 -18.6 -8.0 -4.6 0.0 5.8  C  I  k) I  ..  c-.)I  (.)  k) I  —  Ooo  Coo  vi.o  viOo  1.76 1888 1.41 0.81 1199 65 95.1 28.6 29.8 50.2 158 176 0.44 12.29 4.13 289 5.54 1126 11.12 178 90 417 99.8 13.2 10.60 0.41 -0.9 1.1 10.2 19.1 2.4 -32.0 -27.3 -17.0 -12.0 -0.3 4.8  1.79 1972 1.42 0.87 1201 65 94.1 29.0 29.9 51.2 156 173 0.48 13.34 4.23 278 5.79 1151 10.69 178 89 433 100.1 13.2 11.11 0.44 -0.9 1.1 10.4 20.1 2.0 -38.0 -33.0 -17.0 -12.0 0.3 5.3  1.71 1584 1.79 0.82 1207 66 92.4 28.6 29.4 68.9 159 177 0.72 19.92 3.30 446 4.74 893 12.55 299 161 385 87.1 15.6 8.63 0.44 6.1 7.1 12.5 19.1 4.6 -28.0 -17.9 -8.0 -2.9 1.0 6.8  1.64 1533 1.91 0.74 1192 65 95.2 29.2 29.9 49.6 158 175 0.56 15.70 3.20 149 4.55 882 12.94 138 72 356 90.6 13.4 8.52 0.39 -1.9 -0.9 10.6 18.1 3.1 -40.0 -33.3 -17.0 -13.6 -0.8 4.3  c k)  •  I  —  Ooo  .  • -.  t’J I  Ji\Q  1.56 1.53 472 1097 4.67 4.92 0.82 0.96 1196 1200 65 64 94.3 94.5 29.2 28.7 29.8 29.5 47.5 55.2 157 161 175 179 0.66 0.57 18.44 15.87 3.21 2.01 74 105 4.72 3.52 612 990 12.71 14.94 125 97 66 71 362 284 93.5 85.3 2.0 7.8 8.84 6.18 0.65 -0.02 -4.9 4.1 -3.4 4.6 10.7 11.6 18.1 20.6 2.5 6.5 -38.0 -38.0 -27.2 -18.6 -17.0 -8.0 -13.6 -4.6 -0.5 0.0 4.5 5.8  c  1.71 2116 2.20 0.98 1200 65 94.7 28.9 29.8 54.7 156 174 0.55 15.17 4.68 107 6.44 1232 9.58 127 67 473 107.6 5.6 11.88 0.69 -2.9 -1.4 10.2 22.1 4.3 -40.0 -33.2 -17.0 -12.0 1.0 6.0  C I  vi  k) I  C’.o  1.57 1.53 951 1649 3.79 4.61 0.95 0.84 1203 1204 65 62 94.7 95.0 29.1 28.4 29.7 29.5 51.1 59.3 164 157 175 182 0.57 0.72 15.91 19.85 4.53 3.23 53 66 5.02 6.41 1464 919 9.72 12.27 112 76 63 60 384 465 115.2 95.9 2.5 8.8 11.68 8.89 0.66 0.50 4.1 -4.9 -3.4 4.6 13.2 10.9 24.1 23.1 2.4 8.5 -38.0 -38.0 -27.2 -18.5 -17.0 -8.0 -11.9 -2.9 1.8 4.0 6.8 9.8 207  00  N  I  z  a) a) a)  F-  14.12.27  !  a)  o 292047a080114 14.12.40  Test Name  j.  ‘  00  .—  00  C  -‘-‘  .  N C  -  -  INN 00NQ  -  c  ‘©  ‘  ‘  E’  -  -  -  .  -  ‘-.  :00  .  F—  .  I  00  i ‘:‘ Qa)a)  .,  ‘  ‘...  9-o--  )  -d  g  a  C a)  00  c’1  N  C’  ON  00  tt  V  •  —  -  -  -  L(  -  1 -  —  -  -  -  —  C’  --  121-3-06-01-10C’’° 10.10.04 — 10 -  -  ‘  --  -  06-01-09c°° C’C 121-3-7 15.49.23 C$ N  -  I2I306-01-10N° 11.40.46 20  -  ‘r N —  -  k(N  -  C N OON N.  irN  -  00 N OON N  ,,  -  -  —  C’ N C’ C’  C’  C’ N m N N N N C’NC’ C’ C’ O C’ C’ C’  —  —  ‘r 00 ON  -  C’ N CON N  0 N  N N  00 N  •  N ,-  0  N  00  ‘.‘  N  —  C’  I  N  o  I  I  C I”  N  —  c ‘.c’  C’C’OONQ.,C’NC’N -C’C’ “? .C’ C’ — — C’ C’ C’ 0 I I — N  N C’ N C’ N 00 N N .  12I3-06-01-10 COO Ir)’_ 11.30.19000 0. cr 00 N — — .19 121-3- 06-01-10 11.08.15 18 -  —  I2I306-01-10c 10.54.58 N 17  -  tr C’ fl C’ N.4C 00°° N ,j. CO N N. N  ‘  E S.-,  N o•ON 0 N CON N  -‘  -  Cf  ‘r) C’  J-  r  t-  C’ C’ fl C’ C’ ‘1- N r  00  C’  N ‘i N —  .  .‘r  C’  ooN  00 N.N  .‘n  ‘•  --O  o_  N  oo  c’,-.’  C’ C’ 00  c  flooN  \D N C’ N C i- N NN  0’. N 00 “0 C’  0  I8cCCc) ()QL)  0  I,11  I  I  I  I  -  N N C  N  4  C’ .  a  —‘  -  I  I  I  —  —.4  ‘  (•f  N C’ 0 -  N C’ ©  N  c  cr N C  ci  N —:  rj  N C’  ri  N C  C’  00  N 0  -  C’CC’c.)  —  N  4  1r  &,N I2I306-01-1000 N0NN NON flN 10.33.31—’oo C’C’DCC’C’ -4N 12 — — 0  Test Name 0  .—  -‘  C’ N 0 06-01-09 o .-)rN 4 I 15.39.41_No 2 I 36 oN  121-3-06-01-100 10.24.33 Nii  v- o c 06-01-09 N 235 15.31.40Noo N  06-01-09 121-3-8 15.56.48  06-01-10%c’° 1 C’ • I2I39O95251 .-•C’0N — 00 N O N t C’ C’ D C’ C’ C’ -  .1 -  1  I  U,  V V  E-  I  C N  00  C)  F  00  E-:  --  121 3 2  06-01-09 121-3-1 144626  N  12.11.20 ri  00  N  0  ‘i-  0  N  -  ci  ..  o  N  o  .  00 O  t  d  00  Lf  d  00  00  d  ©  N  N N  0 00 00  I  .  .  N  •  .  0  C  12.38.210d000_  I2I306-O1-1OoCN;000ooN 15 16  13 14  Figure # T est s eries 3 Z_ )  11 12 1 2 3 4 5 6 7 8 9 10 VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- V1II-B- VIII-B- VIII-B- VIII-B- VIII-B 13 14 10 10 11 11 12 12 4 5 6 13 rnI  ti I  T1  I  M..  en  .  + L,aLe, rime  t-.  t’J  J  -3  Cj  CJ  C i.))  c’ci  .DUI  1.56 1.59 1055 1046 2.4 1.6 15.5 16.5 1102 1103 50 49 72.5 77.7 22 3 21 9 24.2 24.1 71.7 58.5 126 131 139 144 127 141 1.78 1.75 4.59 2.58 0.48 0.48 7.46 6.99 1.49 1.59 1.25 1.00 1.95 1.61 286 286 99.4 90.7 9.8 7.1 599 592 02 28 -4.1 0.9 -3.1 1.9 60 103 11.5 17.5  1.53 1065 2.0 17.5 1100 50 72.0 22 3 24.2 58.0 126 138 130 1.82 3.62 0.48 6.17 1.46 1.38 2.13 298 90.4 7.1 605 030 0.9 1.4 112 18.5  1.56 948 1.2 17.5 1104 49 77.8 21 9 23.9 71.3 131 144 1 1.70 7.81 0.49 4.29 1.82 4.66 7.12 282 78.7 8.9 530 027 5.4 6.4 161 23.0  -  .J  IgnitionDelay(ms) 1.55 IHR(kJ/m3) 1018 Knock (bar) 2.7 Comb. Dur. (deg) 16.0 Engine Spd (rpm) 1104 MAT(°C) 49 MAP (kPag) 77.6 CNG Press (MPa) 21 8 Diesel Press. (MPa) 23.9 ExhaustPress.(kPa) 72.3 Corr.Airflow(kg/hr) 130 Airflow (kg/hr) 143 Dieselmj (mg/rnj) 134 CNG flow (kg/hr) 1.74 CO (g/kW-hr) 4.37 C02 (kg/kW-hr) 0.49 NOx (g/kW-hr) 7.48 02 (kg/kW-hr) 1.64 CH4 (g/kW-hr) 1.24 tHC (g/kW-hr,C1) 1.96 ExhaustT.(°C) 280 Pk.press.(bar) 99.1 CA@Pk.press.(bar) 4.2 GrossIMi 572 EQR 028 5%IHR(deg) -4.1 10% IHR (deg) -3.1 50%IHR(deg) 65 90%IHR(deg) 12.0  .CO .  Oocii  C ..OO  -  )  —  -.  C  C  i  rn  C OO C  —  C  I  O C ...OO —  c L’J  OUI  C -OO C 00 tJ LJ  C ooOO C C -.  -  t’J  i  i  Ji  -  )  —  C ooO — C  C  C  ©  O  —O  wO  ‘J C  C C  i’J  O\  .)  )  ‘J  C  (Ji  j  C  COO L1 C  V  j  UiJ  cij  oo  tJ  tJ  1082 50 73.1 22 0 24.2 55.3 125 138 149 1.56 4.94 0.50 5.02 1.75 2.40 3.73 278 77.6 10.5 536 027 5.9 6.4 151 22.0  1.78 2204 3.4 29.5 1098 47 112.6 21 0 23.1 107.1 151 166 144 4.27 1.11 0.45 7.59 0.55 0.41 0.64 466 143.8 9.2 1294 053 -7.6 -6.6 54 22.0  1.69 2239 4.8 32.5 1096 47 112.5 20 8 23.1 107.9 151 166 155 4.26 0.56 0.46 5.76 0.56 0.36 0.57 475 123.4 11.0 1289 053 -5.1 -4.6 103 27.5  1.57 2302 5.7 30.5 1096 47 112.7 20 9 23.1 108.3 152 167 150 4.28 0.34 0.47 4.83 0.57 0.32 0.52 485 119.0 4.7 1259 053 0.4 0.4 153 31.0  1.35 1.51 1.50 1466 2143 1461 3.1 2.8 3.4 20.0 26.5 25.0 1114 1094 1113 49 48 50 87.3 112.6 86.9 20 5 21 9 20 3 22.1 24.0 22.1 71.4 100.4 72.7 135 152 135 148 148 167 170 114 172 2.60 4.08 2.68 1.84 0.56 1.00 0.50 0.45 0.53 10.07 8.87 7.26 1.07 0.59 1.05 0.73 0.43 0.63 1.24 0.67 1.04 352 446 377 118.0 146.4 100.8  -.1çj  1.57 959 1.5 16.0  .  ©j  c  t)  9.4  9.1  2.8  839 039 -8.1 -6.1 51 12.0  1262 049 -8.1 -6.1 48 18.5  834 040 -4.6 -3.1 105 20.5  COy GIMEP  1.9  1.7  2.3  1.9  4.2  2.6  0.7  0.7  0.7  1.3  0.7  1.5  DSOI (ms) DPW (ms) RIT (ms GPW (ms? 2RIT(?nS 2GPW(ms) Comments  3.0 2.1 -7 6 0.7 1 46 0.54  2.7 2.3 -7 6 0.7 140 0.57  3.8 2.1 -7 6 0.7 133 0.54  3.6 2.3 -7 6 0.7 139 0.56  4.4 2.1 -7 6 0.7 150 0.54  4.3 2.3 -7 6 0.7 134 0.56  -6.4 2.1 1 0 0.7 141 1.20  -5.8 2.0 1 0 0.7 162 1.14  -4.8 2.0 1 0 0.7 136 1.17  2.9 3.0 -92 0.7 150 0.86  2.7 2.0 -8 0 0.7 141 1.04  3.8 3.0 -92 0.7 149 0.86 VOID 211  Figure#  14 15 16 17 18 20 21 22 23 24 13 19 VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIII-B- VIH-B 16 14 15 15 16 16 16 16 16 17 17 17  T est s cries-  I—  ..d  )  I  -  Lu —-  •  tJ  C  I—  Ls i  -  ‘d  —  I  ‘  CD  (.)i  c  (Ji  C  •  •  C -  ‘-  c.ui  •  c•  (11  —  i  I  I  .  I  C o• c i Lti  I  •  —  I  ::..:  ri IT LJaLe,  rime  L.O  1-00  woo  ‘,o  000  J0O  O0  OOO  %000  OO  Oo  LuOO  L’J C  C C  L.’.) C J  W C ‘.. .)  —  C  .) C  .) 1 t  C 0 t  0 C 0 t.)  0 ‘.0 t’...)  t’J 0 0  Cj  Jj  I-t!J  chit’J  cj  -j  L  o  0 0 W t’..) .1t)  3.85 2.18 4.51 2137 2171 2220 6.1 1.9 2.5 10.0 33.5 14.5 1513 1508 1505 51 51 53 139.2 148.1 144.5 21.1 20.6 21.4 23.3 22.1 23.6 110.6 118.5 166.3 197 203 200 216 223 220 17.8 14.0 14.8 5.52 5.66 5.64 0.97 2.05 3.69 0.45 0.49 0.43 12.56 5.54 7.90 0.54 0.60 0.55 0.42 0.54 1.48 0.67 0.87 2.18 479 510 525 160.0 122.8 128.1 9.6 12.5 15.2 1247 1280 1285 054 052 053 -1.1 3.4 -6.6 0.4 -0.1 4.9 55 100 106 27.0 8.9 18.0 0.7 1.1 1.7 1.0 3.1 2.2 1.8 2.8 1.4 -71 -91 -80 0.7 0.7 0.7 152 147 176 1.19 1.05 1.11  2.87 2203 1.9 23.5 1510 51 147.1 21.5 23.6 127.7 205 225 17.9 5.67 1.70 0.44 6.16 0.56 0.56 0.83 503 127.4 13.7 1288 053 -0.6 3.9 101 23.0 0.8 2.3 1.9 -76 0.7 145 1.08  i’Jtj  IgnitionDelay(ms) IHR(kJ/m3) Knock (bar) Comb. Dur. (deg) Engine Spd (rpm) MAT (°C) MAP (kPag) CNGPress.(MPa) Diesel Press. (MPa) ExhaustPress.(kPa) Corr.Airflow(kg/hr) Airflow (kg/hr) Diesel inj. (mglinj) CNGflow(kg/hr) CO (g/kW-hr) C02 (kglkW-hr) NOx(g/kW-hr) 02 (kg/kW-hr) CH4 (g/kW-hr) tHC (g/kW-hr,C1) ExhaustT.(°C) Pk.press.(bar) CA@Pk. press. (bar) GrossIM,r) EQ 5%IHR(deg) 10%IHR(deg) 50%IHR(dg 90%IHR(deg) COVGIMEP DSOI (ms) DPW (ms2 RIT(m* GPW (ms) 2R1T(ms) 2GPW(ms) Comments  1.43 1.43 2235 1467 2.9 2.5 24.0 31.5 1094 1111 48 50 112.4 87.0 21.-7”20.4 23.9 22.1 98.2 72.3 152 135 167 148 18.4 15.5 4.18 2.61 0.32 1.18 0.46 0.53 6.39 5.71 0.57 1.11 0.36 0.63 0.58 1.14 466 377 122.6 95.1 11.2 6.1 1280 811 053 039 -5.6 0.4 -4.6 1.4 102 154 26.0 24.5 0.8 1.5 3.4 4.6 2.0 3.0 B0 92 0.7 0.7 141 148 1.04 0.86 VOID  ....  1.38 3.84 2341 2132 2.5 4.5 31.0 10.5 1094 1505 48 52 112.6 148.4 22.0 20.6 23.9 22.1 100.2 113.2 152 203 167 223 11.8 16.3 4.33 5.34 0.30 1.18 0.47 0.47 5.27 13.18 0.55 0.67 0.32 0.54 0.52 0.92 484 454 118.2 156.8 4.8 10.1 1290 1236 052 -1.1 -1.1 -0.1 0.4 52 57 30.0 9.4 0.7 1.0 4.2 2.7 2.0 2.3 -80 -92 0.7 0.7 141 151 1.04 1.08 VOID  Lu  C tJ  0  3.65 3.93 3.54 2143 2166 2091 5.4 5.0 4.5 11.5 10.0 10.5 1513 1507 1511 51 52 53 144.3 148.1 148.0 20.7 21.3 21.4 22.1 23.6 23.6 112.8 130.7 131.5 199 205 205 219 225 225 14.3 15.2 18.2 5.37 5.37 5.28 1.04 1.04 1.00 0.47 0.43 0.44 13.50 12.37 11.52 0.62 0.61 0.62 0.54 0.45 0.53 0.89 0.73 0.84 464 460 461 161.4 156.1 158.7 9.1 10.9 9.4 1252 1246 1226 051 050 050 -2.6 -0.1 -1.6 -0.6 1.4 -0.1 48 66 52 8.9 9.9 8.9 0.9 0.9 0.8 2.1 2.3 1.7 2.8 1.5 1.9 -91 -80 -76 0.7 0.7 0.7 154 145 140 1.05 1.14 1.08  —  4.01 2193 6.7 10.0 1515 53 147.8 21.6 23.6 133.8 204 224 15.3 5.58 0.75 0.44 14.33 0.58 0.37 0.60 470 164.3 9.9 1259 051 -0.6 0.9 55 9.4 0.8 1.6 1.6 -76 0.7 139 1.15  1t-)  212  ‘fl  0 Cr)  00  N  N N  ‘fi N  41:  A-lB  A-iD  14.01.36  O8-0226r’1 N  U1-LflS- O80222oN 10-55- 183040 -  r  —  ,-  C-1E  ‘fl N cr)N  08-02-26N— 00 19.30.18 c  16.24.08  —  080225ooo C 00 15.38.47 08-02-25N’ N 12.30.33  N  Lfl  ‘t  Cl  N  08-02-26a’— 00  —  080222c.0 15-55- 20.07.12 N N 1-  B-iC  c-ic  B-iF  -  C 1F  O2-U-O8O222© QD( 0555 183620 N 00 —  ._  W ,-  >  I  Test Name  B-1E  0  r00  N  (“  00  00  N  —  ,.....  c.i  ,.  N N  00  c...  00  N  ci  N  V)  D  -  —  ‘fl  cS  -  d  -  -  ‘i  Cl ‘r  ‘fl  0 tfl  -  Cl  N  N  N fl Ifi 00  ‘D ‘fi  ND00  N  t ‘fi  Q ‘fi N 0  C’ ‘fl 00  .c’n  —--°°  .._  N  © N  0 N  c’ e  ‘..D .,  00  0  .N.000N  c  C’---° C  ‘fl ‘fl  oo  N  N 00  N  0 C’ N  .  0  flN  —  —  ON .  N  ‘fl  •  ._-  Z.  L)  ‘N  -  (l  ‘  ‘  0 —  ‘fi  ‘ —  N 0  -  .  N  N  c  d  ‘  ‘  —  . c  •  .  ‘fi O © c N  .  © - -  ©  ©.NN•00N © , I  I  I  0  0 0”  —o  000)  0”  ————>  ‘—  —‘-  rM  iU)  ‘flfl NN°°00oNC-4 0 00C00NIfl00 -  I >$ ‘‘‘—L)  08-02-25ooN N 16.16.40  .  080226oN C-iD2 N 190951 ——  0)  U  F-  -  N  C  00  N  14:  > F-  08-02-25 —  N “.  N  ‘IN  —  •0000  —_-4—-4  C  ’IN 00 ON  17.03.31  ,_  \O ‘IN  C ‘IN  CONON\O ONCe cN  —  C -  (.‘IC  .  C  CNC  .  00 ,  C  .-  .  C,’  ‘IN  “  ON  N  ONC  •ONOO  tr’IN  ‘IN 08-02-26 ‘.c ON ON N ON 1N ‘ C’r 4 .fC’ r8 O 5 . 7 l c .O -  -  .  08-02-25N— ,INCCl 14.07.17 ‘.  cN  ONN  ON  C’  ON  ON  ‘  00 ‘IN C’1  C’1  ON C’ C’ 00  00  00  N N  -  “  —  C’.  Cl  C’  ‘IN  c  ON  —  C  N .  ‘r  N  .  ‘  C’I  —  ‘.  ‘.o  rr  N  ‘.  c  C  .  CR  C’  0C  ‘IN  C  N O O N 00 \O ON N ‘I 00 tIN C  C ‘IN C”  tIN  C tIN C C  -  .  ‘IN ON ‘IN ‘IN 00 tIN  ‘IN  C  0  00  —  ‘IN  ‘.0  ‘.0 ‘IN ‘IN  C  —‘  ‘IN  —  ‘IN  C  N  ‘.0  ‘IN  -  —  C  ON  00 ON IN  ON C’ m 00  .  ‘.  ‘IN  00  .tIN  C C oo CR  no-’..CR9 I t NQ0LIN — C C tIN C C C —  ‘n ‘.0 ‘IN  I  C  9)  ON  ON  ‘fl  c.,  .  ON  —  ‘IN  C  ON  CR CR  C’4  C  C  cc  N  I  CR  I  N  ‘IN C N  ON  I  CR  N  N  d  ON C N  I  ‘N 080225NC • .NNNONON ON°’I .CONo o  19.47.07  13.47.19  REP-08- 11.11.26 02-28  1555  02-26 LOW Co C  F-  (‘  10-55A-iF A-iC  08-02-26 C”  ;_.,  -  o 6’  ‘  ,  -  .‘INQ’.0tIoo’IN  VV ,  —  —  —  C  —  C,  N  ‘  ‘•••  -  EC.’) -  ,—  A-1E  )  2  z..88zc.  CCCL) -INO’  ‘.D tIN ‘.0 ON ON ‘.0 . N ‘.0 ‘IN ‘IN N 00 — N — ‘‘ ON ON tIN ‘IN N 00 ON ‘. N ONN ‘IN ‘INONON’IrN CON r —_ CC00CCC——  Test Name  .  °  08-02-25 A4A 13.57.54  03-0815-5513  19.24.15  121706 B1A2  C’ .,_  >.  -  .,-  1c9  ‘  CI) t)  r  IF-  C.’)  U  N -l  00  N  —  —  VIII-3D  VIII-2B VIII-2  VIII-2A VIII1D2 MIII  ‘r’. I..  YT  V11i115  VIII-1  12.28.00 08-06-11 15.27.00 08-06-11 09.31.41  08-06-11 15.11.08 080641 09.20.26  1 09.25.32  c  o  —  ,.  N  ,,  fl  .  —  o —  ,_  ON  ——  ‘.0  —  —  g-  ‘. ‘-ON-  ——  cr  •C’  —  00’ 0 .©N —  .  — ,  000 .0 N  15.00.33---  .  ON  12.01.07—’--’  000o0  12.40.51—’—  ’,N 080642 00  12.03.28  00 C’  VIII-3A 08-06-09  VIII-3  VIII-3B  —  r’  -  ,  —  r’ —  —  —  r  ‘  I  ‘ U CID  Name  iest  U  08-06-1  ‘  N  (•9 ‘o r  ‘D  N 0 0  0  —  0  —  —  00 0’,  j-  ON  -  N  00 V  \D  —  Cia;cr  N  ‘  0  ON ON C N 0 0 N ON ‘ N r) ON ON dr00  N  N  ‘.0  ‘.0  —  —  —  “0 ‘-  —  N ON  00  —.  ‘,  —  N  000  ‘  —  —  o o  V  00  —  ‘.0 0  —  —  .  ON —  —  N •j-  ‘  N  )  ‘  -  ‘r  ON  O’ ON ©  Lfl  —  c  ,‘  —  —  ‘  ‘—  .—  -  c  ‘r 0  —  N  0  —  n  ,-  —  ‘fl  ON  N N ON 0 ‘.0 ‘.0 0 ‘.0 C’N• 0 •  .  ‘.0 0 ‘.0 ‘.0  ci  —  -  -  N  -  ci  C’C’C’L) -O  C’.  ‘  •  N 0  N  “0 0 N  —  “0  ,-  —  0 N  -  -  .—  .—  —  ‘  N C’ N  —  0  c#  N C  00 ‘.0 0 N  cl  cl  d  00  NC’’.flNOON0NON’.O  00  (Y  ‘fl C  C’. C’  .  •j0N N 00 ‘. 0 • — —  m N  N  NOOOOOOOO0NO  ‘.0  N  N N ‘r  N  00  00  ‘fi ‘.0  N N  N C’. t  .-  N N ‘r  00  0C’.ONC’ONON”  .—  .  .:  0 00  cn —  -  N  NNON 0 — — N 0 00 — N N N 00  —  ‘.0  ‘fl  —  N N  N  -  c  ,  N’  C4  000 ‘r 000 00 c 00 N N 0 ‘.0 N N N CC’ N N 00 ‘. ON ‘.0 •o— .0o0N0 00N_r, VHI-1A 08-06-09ON0N00 N a 00 N 0 00 — N N ‘.0 12.14.42 —  c  I  •  F-’  L  —  1-  14  13  Figure# .  Test Series # -  < —  (  . .•. •.  .:.  (D  0  00  00  iime . V  V.  Ignition Delay(ms) IHR(kJ/m3) Knock (bar) Comb Dur. (deg) Engine Spd.(rpm) MAT (°C) MAP (kPag) CNGPress.(MPa) Diesel Press. (MPa) ExhaustPress. (kPa) Corr.Airflow(kg/hr) Airflow (kg/hr) Diesel inj. (mglinj) CNG flow (kg/hr) CO (g/kW-hr) C02 (kg/kW-hr) NOx(g/kW-hr) 02 (kg/k W-hr) CH4 (gIkW-hr) tHC (g/kW-hr,C1) Exhaust T.(°C) Pk.press. (bar) CA@Pk.press.(bar) 43joss1MEP (bar) •• :EQR 5%IHR(deg) 10%IHR(deg) 50%IHR(deg) 90%IHR(deg) COy GIMEP DSOI (ms) DPW(ms) RIT(ms) GPW (ms) 2RIT(ms) 2GPW (ms) Comments  18  .< — —  — —  20  19  21  22  24  23  00 .  0 O I  0 — 0  0 0 O I 0 — 0 Ui  — •..  <.:< - —  I  —  —  —  0 000 Li.) 0 •-.. O  00 00 Ui 0  0 00 0  —0 -00 Li.) 0  •  •  :  V—  Ui  .  -‘  V.  17  VIll-B2 VIll-B2 Vffl-B2 VIll-B2 VflI-B2 VIII-B2 VIII-B2 VIll-B2- VIII-B2 VIll-B2 VUJ-B2 VIH-B2 4 4 4 5 6 6 5 5 6 6 5  -  1 + L,aLe,  16  15  0  O\90  0 0 Ui O • I 0 — 0  1.51 1.77 1.72 2192 2193 2265 1.8 2.0 2.1 34.0 35.0 33.5 1101 1085 1079 48 50 52 87.1 71.4 74.5 21.1 21.3 21.0 23.4 23.6 23.4 101.5 104.3 98.7 159 143 144 159 143 144 15.1 16.0 15.8 4.22 4.16 4.23 1.37 1.07 0.93 0.47 0.46 0.47 8.27 8.47 8.07 0.59 0.48 0.47 0.44 0.43 0.41 0.44 0.43 0.41 450 486 488 135.3 131.5 133.8 9.3 9.3 9.2 12 95 13 1 13 30 0.54 0.55 0.55 -8.6 -7.1 -7.6 -5.1 -4.1 -4.6 5.5 5.2 5.2 25.5 28.0 26.0 1.0 0.8 0.7 -6.2 -6.2 -6.2 1.5 1.6 1.5 10 10 P0 0.7 0.7 0.7 1.641.59 1.59 1.27 1.24 1.33  -0  0  00  00  \C0  —0  -0  0  0  .. 0 Li.) Q • I 0 — 0  ‘.‘i  i—  .i • I 0 — 0  1.94 2148 1.9 32.0 1097 53 68.3 21.0 23.4 84.3 142 142 15.2 4.12 0.84 0.47 11.04 0.48 0.42 0.42 480 128.6 9.2 12 76 0.55 -7.6 -4.6 4.9 24.5 0.9 -6.3 1.5 10 0.7 1.64 1.28  •  -I  0 — C  çi00  0 Ui • I 0 — 0 —  1.59 1.56 1.58 546 550 565 2.8 2.4 2.3 36.0 38.0 36.0 1099 1084 1078 48 50 52 87.0 71.6 73.6 21.1 21.3 20.9 23.4 23.7 23.4 99.3 102.7 95.1 159 143 144 143 144 159 14.9 15.7 16.1 4.27 4.17 4.21 0.79 0.62 0.57 0.47 0.47 0.48 5.92 6.31 5.81 0.59 0.48 0.48 0.39 0.38 0.37 0.39 0.38 0.37 463 497 497 115.7 110.3 111.8 12.4 12.4 12.6 12 90 12 92 13 02 0.31 0.56k 0.55 -6.1 -5.6 -5.6 -4.1 -3.6 -3.6 10.4 10.5 10.3 30.0 32.5 30.5 1.0 0.8 0.5 -5.5 -5.4 -5.5 1.5 1.6 1.5 10 !10 10 0.7 0.7 0.7 1.65 1.59 1.59 1.23 1.22 1.28  0 ,-00 0 0 Ui •  0 0  O’ I —  1.67 540 2.4 37.0 1094 54 70.8 21.0 23.4 86.0 144 144 15.2 4.22 0.51 0.48 7.93 0.48 0.36 0.36 496 110.2 12.4 12 84 0.55 -6.1 -4.1  31.0 0.9 -5.5 1.5 10 0.7 1.66 1.26  0  0  I —  1.52 582 4.5 35.0 1097 48 87.1 2L1::’ 23.4 103.6 158 158 15.2 4.31 0.39 0.49 5.04 0.59 0.35 0.35 480 105.3 3.4 12 56 0.50 -1.1 -0.6 15.8 34.0 1.2 -4.5 1.5 10 0.7 1.40 1.26  Ui  0 C  O, I —  1.55 621 2.5 35.0 1083 50 70.8 21.3 23.6 81.4 143 143 15.7 4.29 0.39 0.48 5.26 0.47 0.32 0.32 500 96.6 4.4 12 98 0.50 -1.1 -0.1 15.3 34.0 0.7 -4.7 1.6 10 0.7 1.60 1.23  O O  0 0  I —  Li) •  0 0  I — ‘j  1.52 1.56 578 582 3.1 2.7 35.0 36.0 J076 1093 52 53 74.8 73.1 21.0 21.0 23.5 23.3 90.1 92.1 145 146 145 146 15.5 14.9 4.25 4.33 0.33 0.43 0.48 0.48 4.92 6.48 0.49 0.47 0.32 0.31 0.32 0.31 504 513 98.0 97.5 3.4 3.8 12 96 13 01 0.55 0.56 -1.6 -2.6 -1.1 -0.1 15.4 15.2 33.5 33.5 0.6 0.9 -4.7 -4.8 1.5 1.5 10 10 0.7 0.7 1.60 1.66 1.27 1.27  216  25  Figure # .  Test Series # -  Corr.Airflow(kgIhr)  Airflow (kg/hr) Diesel inj. (mg/inj) CNGflow(kg/hr) CO (g/kW-hr) CO2 (kg/kW-hr) NOx(g/kW-hr) 02 (kg/kW-hr) CH4 (g/kW-hr) tHC (gIkW-hr,C1) Exhaust T.(°C) Pk.press.(bar) CA@Pk.press.(bar) Gross IMEP (bar) EQR 5%IHR(deg) 10%IHR(deg) 50%IHR(deg) 90%IHR(deg) COy GIMEP DSOI (ms) DPW(ms) RJT(ms) GPW (ms) ‘U2RjT(ms) 2GPW(ms) Comments  29  —  —  —  — —  -..  I  00  -  1  00  C  C  C  C  —  C  C  C  .4 C  C C  —  C  C C  C C  0  ..  .  C C  IgnitionDelay(ms) IHR(kJ/m3) Knock (bar) Comb Dur. (deg) Engine Spd.(rpm) MAT (°C) MAP (kPag) CNGPress.(MPa) Diesel Press. (MPa) ExhaustPress.(kPa)  —  -.a  -  r L,ale,1m ime  28  —  ( G  27  30  32  31  34  33  35  36  :;:<  < -  26  Vffl-B2 VIll-B2 VIll-B2 VLII-B2 VIll-B2 VHI-B2 VIU-B2 VTII-B2 V1II-B2 VIII-B2 VIII-B2 VIfl-B2 7 8 9 10 8 8 9 10 7 9 10  I  3.28 2237 2.3 28.5 1405 53 112.2 21.2 23.3 125.4 201 201 14.5 5.43 2.20 0.45 8.01 0.58 0.52 0.52 492 132.9 11.7 13.13 0.55 -3.1 2.0 79: 25.5 0.8 -7.0 1.9 10 0.7 1.91 1.28  •  C C  2.75 2210 2.5 29.0 1401 54 96.4 21.1 23.3 113.1 185 185 14.5 5.37 2.00 0.45 8.97 0.49 0.49 0.49 514 135.8 9.0 13.08 0.55 -4.6 -0.6 52 24.5 0.9 -6.7 1.7 10 0.7 1.54 1.36  -  -.  O —  c 2.61 1.60 2227 2211 2.2 1.3 31.0 37.0 1402 1405 56 53 97.9 112.2 20.9 21.1 23.3 23.2 111.8 123.4 185 202 185 202 14.5 14.7’.. 5.47 5.44 1.43 2.46 0.46 0.46 8.45 5.32 0.48 0.59 0.48 0.51 0.48 0.51 521 505 135.5 118.8 9.0 12.1 13.20 12.92 0.55 0.55 -4.6 -5.6 -0.6 -0.1 52 98 26.5 31.5 0.6 0.7 -6.5 -6.4 1.6 1.9 10 10 0.7 0.7 1.42 1.67 1.47 1.28  t’J C’ C —  C  — —  -  -..  C C  C  ‘J C •  I  C C  C  I  —  C  —  .) C  (J  -‘  C’. •  C C  0’. I  1.99 1.89 1.79 1.69 2177 2243 514 521 1.4 2.2 1.3 1.7 38.0 39.5 42.5 42.5 1402 1400 1406 1400 55 54 56 54 96.2 98.4 111.4 96.3 21.1 20.8 21.2 21.0 20.4 23.2 23.2 23.3 111.6 115.7 120.1 114.1 185 187 201 185 185 187 185 201 15.2 14.7 14.5 14.0 5.43 5.57 5.49 5.46 2.44 1.93 1.53 1.12 0.47 0.47 0.47 0.48 5.00 4.60 3.73 3.63 0.50 0.48 0.59 0.50 0.43 0.39 0.45 0.34 0.43 0.39 0.45 0.34 540 523 557 553 109.9 111.7 100.5 96.5 12.9 12.1 2.9 5.1 12.75 13.O5 2.73 12.54 0.55 0.56 0.50 0.50 -5.1 -5.1 -5.1 -4.6 -0.6 0.5 -2.1 -2.6 106 106 145 161 33.0 34.5 37.5 38.0 0.7 0.8 1.1 0.9 -6.1 -5.4 -5.9 -5.8 1.7 1.6 1.9 1.7 10 10 10 10 0.7 0.7 0.7 0.7 1.59 1.45 1.67 1.60 1.34 1.46 1.32 1.25 -‘  a’. • C C  C  I  —  1.75 497 1.9 42.5 1399 57 98.2 20.8 23.2 112.6 185 185 14.8 5.56 1.09 0.49 3.48 0.48 0.33 0.33 564 95.5 5.1 12.73 0.56 -5.6 -3.1 151 37.0 0.7 -5.5 1.6 10 0.7 1.66 1.38  —  00 •  0  -‘  C  I  —  —  I  —  —  — — —  ‘.D  C Q  C  I  ..O  00  -.  —  I  I  — —  w  C  C  C  C C  C C  •  •  0’. I  C  2.64 1100 1.6 18.0 1102 51 48.7 21.0 23.4 61.9 127 127 15.4 1.78 9.08 0.48 8.60 1.46 2.37 2.37 292 91.4 9.2 6.07 0.30 -6.6 -2.6 53 11.5 2.3 2.3 1.6 -73 0.7 1.66 0.74  0 0’.  C  I  2.71 1082 1.4 17.0 1102 53 44.4 21.0 23.4 56.1 123 123 16.0 1.78 11.48 0.48 8.78 1.44 3.10 3.10 294 88.6 9.6 5.98 0.31 -4.6 -0.1 S5 12.5 2.4 2.2 1.5 -73 0.7 1.71 0.76  -.  C  .i. 0.  3.59 1039 1.0 13.5 1107 54 45.4 21.2 23.5 56.9 125 125 14.8 1.76 15.86 0.48 11.02 1.54 4.83 4.83 283 84.7 9.8 5.72 0.30 -0.1 1.4 58 13.5 3.6 2.0 1.6 -73 0.7 1.51 0.80  217  00  In  C  CO  44:  w  —  -  —  F  —  —  ._.  )  ..-.  00  VIII14B  13.16.48 c’i  13.21.04  VIII 11B  VIII 12A  VIII12B  08-06-11  08-06-12 12.56.42  08-06-09 12.58.01  13.12.02  —  .-  —  ‘•  in  —  N  In  —  C9 00  —  11.15.32 c’i  —  VIII14A VIII13B  -  VIIIhA 08-06-09  VIII 13A  VIII-  12.41.01  2  hA  Test Name  N  I-  C-.. .-  .-  U  c  N N N 00  N N  Lf  Inc  D N N ‘n  N N  c  —  —  —  -  -  .  .  —  -  —  ‘  -  —  —  ó  —  —  N  -  C N  .-  C N  C O in N N O C m N N N N N —  —  —  ‘  ‘  ‘‘  .-‘  ‘‘  —  E  C  oo  °. n  N C Lfl  .  N C 00  O  c  C C N  -  .J. —  .-  N  m  C  C  C  •  i  d  C N  %D C  C  ‘  -  C C (fl . C \D  00 C’I  2  N  .-  .  N  N  N N © - N  ‘-  C C C t  N —  C  N C D - N  - CL  —  —  —  .‘—  ,;;  N  ‘  I  ti  in  —  N  N  C N  —  ‘t  —  —  N In  ONcc .-  L  __  --  •  •  C  C  C  C  N  -  C  00 N N C — 00 Q N m C C 00 — C C’. 00 O c dNNóNC.hi9Cod ° 4:NNCN-C  -  i N C •C .-  ‘  ,-  .  C C C N  ‘  z. -—-9---  C  00  00  N  I’.)  C)  a)  41: I  — ri  — —  g  C’  — —  R  C  6 . 4 l 4 . l 8  08-06-11 10.45.46  —  - -  —  L  -  cl  —  N  0  cl  N  —  -  00  00  — —  cr c  —  - -  e  —  (‘4  -  —  -  •  —  —  .‘n  — .  ‘r  N00°fl  .  ‘,  — — C).‘i O0Q0°°CC C) — — — — — —  r  ‘t-  c  N C’ r4 Q  N  ‘i;’  c- c  ,‘  oo  ‘‘  —  -  —  -  ‘‘  —  —  ‘  c  j-  —  —  C  00 m o C) C)  ..  C  C  ,—‘  r-  ‘  O —  ,—‘  —.—.‘  C)  ‘n o  C)C)-N NeooC)C)O©OC)—C)  —  C.)  00C)r-  c  °  ;_-  —  c  13.32.19’—  ‘—,  ,—..  —  08-06-11 11.29.06— 13.34.11 )  0000  n o ioo a o’ o 08-06-11 _ 0 r123028 V 2 .C) HI-19 -  VIII18B  VIII18A VIII 17B VIII-17  VIII 17A VIII16B  VIII-16 VIII15B VIII-15 VIII15A  Test Name  08-06-11  22 VIII-18 08-06-11  -  —  > DO  —  —  -  ir >  —  -  a)rJ  ,4-  a)rd a)  I  Ifl  [  r  .  .  —  —  ‘  __ -  -  a)a)a)  —  Ci  —  Cd  c  I  .  —  00  -  -  -  Lf  I  I  -  N  —  N  ©  .4  I  —  (‘4  C”' N  C N  0  \0  00  N  ‘.0 ‘.0  ‘.0  1I  — C\I  >  c9  —  e 125205 VHI-23 .N  kfl  N N  .tr  C  N r  C  C  C C  C  “  C  —  -  -  -  ‘  C.  -  -  —  N C  N  C C N  C  —  —  10.5O.48—  C  fl  z  —  VIII22B  N  .‘  .  °°—‘  o  Co  N .. m’•m° C C C  C  °  CCCç)  Cl)  08O612;;NNNoooocN ©: 15.29.36— N  O  i  .  0  °  VIII22B 080611 12.58.05 -. C  C.  080613 C 10.18.02---  -  VIII21B  c  DNt  08-06-11= — .13.37.07—  -  ©  C.  VIII21A  1 VIII-22 -  c’  VIII20B -  —  )__-  ‘—  C.CNfflCC  08-06-11 C. NC13.2O.59—-  u.  —--  10.04.38—-  ,-‘  .  VIII19C  .  1i.9.46—  Test Name  VIII19A  VIII20A  -  08-06-11 ee 0 C /III2112374O —  c’  —  r -  —  > c .  .  —  II)  E  ci, )  -.  )  r  —  C  C”  73  Figure # T est s eries -  74  75  76  77  78  79  80  81  82  84  83  Vffl-B2 VIll-B2 Vffl-B2 VIll-B2 VIII-B2 VII1-B2 VIII-B2 V1II-B2 VIII-B2. VIII-B2 VIll-B2 V1fl-B2 23 23 24 24 24 25 25 25 26 26 27 26  z -  -  <<  — tJ  tiJ  t’3  .  t’.) C  0 Oe c) C  .  •  I-.  uiO  r + IT LIaLeI  rime  ‘J  IgnitionDelay(ms) I}IR(kJ/m3) Knock(bar) Comb Dur. (deg) Engine Spd.(rpm) MAT(°C) MAP (kPag) CNG Press. (MPa) Diesel Press. (MPa) Exhaust Press. (kPa) Corr.Airflow(kg/hr)  Airflow (kg/hr) Diesel inj. (mg/inj) CNGflow(kg/hr) CO (g/kW-hr) CO2 (kg/kW-hr) NOx (g/kW-hr) 02 (kg/kW-hr) CH4 (g/kW-hr) tHC (g/kW-hr,C1) Exhaust T.(°C) Pk.press.(bar) CA@Pk.press.(bar) Gross IMEP(bar) EQR 5%IHR(deg) 10%IHR(deg) 50%IHR(deg) 90%IHR(deg) COy GIMEP DSOI (ms) DPW (ms) RIT(ms) GPW (ms) 2RIT(mS) 2GPW (ms) Comments  1  2.92 2224 1.8 26.5 1096 54 70.5 21.1 23.5 84.6 143 143 12.0 4.28 0.62 0.47 9.62 0.49 0.39 0.39 486 113.0 13.5 12.78 0.55 0.4 5.4 10.3 27.0 0.8 -5.4 0.8 1.0 0.7 209 1.28  o’ I —  —  ‘ow  .  1.24 2243 1.6 28.5 1105 52 71.0 20.9 23.3 84.0 146 146 11.6 4.36 0.81 0.47 8.65 0.48 0.38 0.38 491 112.2 13.2 12.92 0.55 -0.1 4.4 10.3 28.5 0.8 -5.0 0.9 1.0 0.7 153 1.34  0 C  as I c’ .  dl  —  1.94 2293 1.3 29.0 1098 52 69.8 20.8 23.4 29.3 146 146 12.4 4.38 0.73 0.48 5.64 0.50 0.34 0.34 474 92.4 17.6 12.66 0.55 3.4 8.9 15.2 32.5 1.0 -4.2 0.8 1.0 0.7 156 1.36  .  C  u9O (..)  •  C  C  C C  -‘  —  v. as • I  -  C  C  —  CO  t’. C  C  —  0  0  as I  u. as I  C  •  —  a I  ..  •  -.  -  -  -.  .  0  c as • I  C  oo  —  0  .i.9C  0  LJ C  —  .  as as • I  as as I  C  —  0  C  C9C C  W C  .  w as • I  —  as I  -.  -  oo  ooj  v  Wç,j  Lj  1.15 1.37 3.45 3.69 3.91 2242 2242 2223 2173 2213 1.3 1.4 3.4 4.1 3.8 28.5 29.0 24.5 22.5 24.5 1095 1100 1399 1397 1400 54 52 55 57 58 70.5 70.9 91.7 94.8 96.0 21.1 20.9 20.9 20.9 20.9 23.5 23.3 23.2 23.4 23.2 83.1 84.9 113.4 121.4 108.1 144 144 182 181 184 144 144 182 181 184 11.6 11.9 13.3 12.2 12.6 4.30 4.31 5.43 5.47 5.33 0.60 0.64 0.99 1.11 1.23 0.49 0.48 0.46 0.46 0.45 7.23 7.23 13.26 14.77 15.34 0.50 0.50 0.46 0.49 0.48 0.35 0.35 0.46 0.44 0.45 0.35 0.35 0.46 0.44 0.45 503 501 517 510 502 91.8 91.6 139.7 144.1 146.1 17.6 17.7 8.9 9.3 9.3 12.42 12.41 13.12 12.86 13.12 0.55 0.55 0.56 0.54 0.55 2.4 1.9 -1.6 -1.6 -1.1 7.4 7.4 0.4 0.4 0.9 15.1 15.2 5.1 5.3 5.4 31.0 31.0 23.0 21.0 23.5 0.8 0.6 0.9 0.7 0.6 -4.2 -4.3 -6.6 -6.4 -6.8 0.9 0.8 1.4 1.2 1.4 1.0 1.0 1.0 1.0 1.0 0.7 0.7 0.7 0.7 0.7 167 159 158 159 174 1.24 1.29 1.45 1.40 1.30  2.77 2221 1.5 34.0 1398 55 91.5 20.8 22.2 106.3 182 182 13.4 5.46 1.69 0.47 7.36 0.47 0.42 0.42 543 113.5 12.7 12.94 0.56 -2.1 3.4 9.9 32.0 0.7 -5.9 1.4 1.0 0.7 158 1.42  3.20 2.80 2207 2193 1.6 1.6 28.5 34.0 1396 1399 58 56 94.1 96.4 21.0 20.9 23.4 23.2 120.6 109.1 181 185 181 185 12.4 12.6 5.51 5.56 1.69 2.42 0.47 0.47 8.43 7.53 0.47 0.49 0.40 0.42 0.40 0.42 545 536 114.7 114.4 13.6 13.2 12.91 12.93 0.56 0.55 0.9 -1.1 4.4 3.9 10.2 10.2 29.5 33.0 0.8 0.8 -5.8 -5.9 1.2 1.4 1.0 1.0 0.7 0.7 172 163 1.36 1.30  2.25 399 1.1 40.5 1394 56 98.5 20.9 23.2 115.3 187 187 13.3 5.65 1.34 0.49 4.93 0.48 0.33 0.33 562 95.3 16.0 12.89 0.56 -3.6 1.9 15.0 37.0 0.7 -5.5 1.4 1.0 0.7 177 1.42  c  .  221  Fiure# Test Series .  -  86 87 88 89 90 85 91 92 93 94 95 96 VIll-B2 Vffl-B2 Vffl-B2 VIll-B2 VHI-B2 Vll1-B2 VIII-B2 VIll-B2 VIII-B2- Vffl-B2 Vffl-B2 REP-08 27 27 29 28 28 28 29 29 30 30 30 06-09  t3 00 O i-+,--’• LPaLe,  rime  IgnitionDelay(ms) IHR(kJ/m3) Knock (bar) Comb Dur. (deg) Engine Spd.(rpm) MAT(°C) MAP (kPag) CNGPress (MPa) DieselPress.(MPa) ExhaustPress.(kPa) Corr.Airulow(kg/hr)  Airflow (kg/hr) Dieselinj (mg/mj) CNG flow (kg/hr) CO (g/kW-hr) C02 (kg/kW-hr) NOx(g/kW-hr) 02 (kg!kW-hr) CH4 (g/kW-hr) tHC (g/kW-hr,C1) Exhaust T.(°C) Pk. press. (bar) CA@Pk. press. (bar) GrossIMEP(bar) EQR 5%IHR(deg) 10%IHR(deg) 50%IHR(deg) 90%IHR(deg) COVGIMEP DSOI(ms) DPW(ms) RJT(ms) GPW (ms) 2RJT(nIS 2GPW (ms) Comments  0  -0 .j.Q  0’. 0’. L-  —.  .-  ‘.0  0’ 0 t  2.14 1.98 387 410 1.1 1.4 37.5 41.0 1395 1399 58 57 94.2 95.4 210 209 23.4 23.2 119.4 109.0 184 182 182 184 123 121 5.54 5.50 1.23 1.23 0.48 0.48 5.01 4.97 0.48 0.50 0.34 0.33 0.34 0.33 566 551 92.7 91.9 16.4 16.4 1259 1252 056, 055 -1.6 -2.6 3.4 3.9 151 153 36.0 38.5 0.6 0.6 -5.0 -5.2 1.2 1.4 10 10 0.7 0.7 154 163 1.35 1.24  0  j0 C’.  o::  0  0  .  0  00  1.93 1080 2.1 16.5 1106 56 44.4 212 23.6 64.0 124 124 148 1.79 5.58 0.49 10.80 1.43 1.62 1.62 302 90.1 9.1 594 0’3 -4.1 -3.6 55 12.5 1.2 -5.5 1.5 10 0.6 132 0.85  .  0  .00  0  ..00 çi 1 M 0’.  00 C.  0 0’.  .-  -‘  — -  1.73 426 2.8 19.5 1091 53 48.1 211 23.6 64.5 126 126 155 1.77 5.62 0.50 5.71 1.51 2.55 2.55 299 80.8 3.9 591 030< -3.1 -2.1 99 16.5 1.1 -5.2 1.7 10 0.6 174 0.72  1.83 1038 1.7 18.5 1105 56 45.1 214 23.6 63.6 124 124 137 1.80 7.05 0.49 7.92 1.48 3.44 3.44 307 76.9 13.2 584 O0 -0.6 0.4 102 18.0 1.4 -4.9 1.5 10 0.6 147 0.76  c  o  2.02 1101 1.6 17.5 1093 53 47.7 210 23.6 63.8 125 125 15.8 1.69 5.99 0.47 6.97 1.49 1.66 1.66 287 92.2 8.9 603 029 -6.1 -4.6 53 11.5 0.8 -5.9 1.7 10 0.6 173 0.74  0  ‘.0  0  2.04 1108 1.7 18.0 1103 55 48.9 209 23.5 65.1 128 128 146 1.87 4.44 0.49 11.31 1.43 1.12 1.12 299 93.2 8.5 617. 30 -5.6 -4.6 48 12.5 1.4 -5.7 1.4 10 0.6 153 0.90  —0  0  00  0  0  —0  00  00  ji00  ‘.)00  -00  o  —0 0 0’.  wo  —c  --  ..o 0’. o.L  00  0’.  V ‘.o  1.85 369 2.2 20.5 1102 55 49.6 210 23.5 66.1 128 128 135 1.87 5.33 0.49 8.26 1.48 2.28 2.28 305 79.4 12.6 604 030 -1.6 -1.1 101 19.0 1.0 -4.8 1.4 10 0.6 154 0.82  1.69 444 2.0 21.0 1090 53 49.9 211 23.6 66.8 127 127 155 1.84 5.80 0.49 4.41 1.47 4.30 4.30 311 78.3 5.4 609 030 0.4 0.9 151 21.5 1.3 -4.7 1.7 10 0.6 197 0.76  1.80 336 1.8 21.0 1101 54 49.0 210 23.5 63.3 128 128 144 1.86 7.03 0.50 6.12 1.57 5.39 5.39 310 70.8 6.4 576 030 2.4 2.9 152 23.5 1.4 -4.2 1.4 10 0.6 169 0.80  1.79 1499 2.0 27.5 1210 50 58.8 210 23.3 84.7 142 142 138 3.03 2.19 0.48 5.93 0.86 0.64 0.64 409 90.5 12.9 864 041 -3.1 -1.1 99 24.5 1.8 -5.2 1.5 10 0.7 141 0.98  0’.  Vi  —  .j  1.73 343 1.6 19.5 1102 56 44.7 214 23.6 64.0 124 124 136 1.84 8.47 0.50 6.03 1.53 6.90 6.90 317 67.3 6.6 566 031 2.9 4.4 151 22.5 1.3 -4.3 1.5 10 0.6 156 0.79  Vi 0’. ciiL  C’.  222  C -4  C -4  C -4  -4  C -4  C C -4  00 ON  N ON  C’I  —  .  W — —  VIII-27 VIII19B  ‘ 00 .  00  N  4 c-  -  C  O  00 C  cr  “  ‘fl 00  c c’1 ‘n  i  r’ C  c”  Cr4  00  00  N  cr  -  —  D  —  —  N  -  0000  ON  C  ON  C’  .  -  N  N  •Co0  .  fl  00  ‘r  r  00  .j  C  Ir  -  ON.  C  -  —  C  ‘  ON .  Cr  C  I  .  C  ‘t \D C C o t— Ndnc I C I Cr  N  N  .00Cr  Cr C 00 N 00 Q C 0000 0000tkfl -. fl ‘- C C C C ON  -4  C’  ‘  --  CS  Cr  0  N  -  N  -  -  -  -  00  -  -  C N C C C  -  C’  ON ON ON  N  00  -  -  -  N  ‘-.,  c  ON  ON  ‘  -  -  C’  --  00 C --  Cl  00 C  00  ‘-  —  ..c  -  Cl  -.  C-0.4  0 -Iic)I  I  4)  I  1)  Cl  U  4....4  C’  U  00  ICICICIç)  V  Cl  ‘  ,  ,—.  tfl C N  c,  .  .  ,  C  00CCe° C00CCC0CCd  .  —4  0 CC  09.58.00— ‘  08-O6-12 C 12.13.00 -  Vill-1D  09.21.47c-  ON  06-13  —  .  09.52.04 ‘-4  N  —  ‘‘Cr  ‘  10.03.04— U  E  —  C’  06-12  08-06-11 —VIII-11 114400 —  > —  —  >  92  -  00  —  9 0.4’  00  —  06-10  lest  Name I  = REP-08 08-06-11 0.4’ p06-1109.01.13 00  — ,  92  I  c..J  U  U .4-i  UI  0  Appendix F. 1- Test Series Vu-A -800 RPM Pressure and HRR Curves  224  VII-A-7 O1-20-47a  ca  Knock:2.O bar IHR: 1613 kJ/m 3 V  a) 0  E  100  0  a) (a  a,  a) 0 (a a) a)  0  ca0 C  :2O Crank Angle (deg ATDC)  VII-A-7 02-20-47b  ca a) U)  100  I \  (a a)  x  Crank Angle (deg ATDC) C) a)  a) Ca  C  a)  a) a)  a) (a 0  a) D 0 0 a)  100  1  Knock:3.5 bar IHR: 315 kJ/m 3 300  -)  a)  (a  ;200 2040 Crank Angle (deg ATDC)  Vll-A-7 02-20-47b2  (a  1  E  0  C  200  a)  I  200  100 )H  C -20  0 40 20 Crank Angle (deg ATDC)  2  Knock:3.4 bar 3 IHR: 320 kJ/rn  0) a)  •0 30C E  100  200  0 a) C >  C-)  (a 0 C  C -20  41 0 20 40 Crank Angle (deg ATDC)  VII-A-7 03-20-47c  40 20 Crank Angle (deg ATDC)  Knock:3. 1 bar IHR: 254 id/rn 3  C)  300 0 0 a) 0  E  100  200  a) •0 C  a) (a 0 V C  :20  0 20 40 Crank Angle (deg ATDC)  (a a)  I  100  C-20  -  A  20 40 Crank Angle (deg ATDC)  4  Figures F.1.1 to F.1.4 : Diesel flowrate: 17.2 mg/inj IHRrano: 0.80 Knock Ratio: 0.92 Ignition Offset: 2.41 deg  225  V11A8 04-20-70a  Knock:4.9 bar IHR: 1619 kJ/m 3 E  100  -  200  0  I  I-  cc  CD  50/  100 ci)  1  cc 0  V  -20  0 20 40 Crank Angle (deg ATDC)  -20  V11A8 05-20-70b  c  20 40 Crank Angle (deg ATDC)  5  Knock:5.4 bar IHR: 608 kJ/m 3 300  100 —  J%\  -  200 cc  50/  100  o  ci)  ci)  o  —  v -20  cc 0 -20  0 20 40 Crank Angle (deg ATDC)  V11A8 06-20-70c  20 40 Crank Angle (deg ATDC)  6  Knock:4.4 bar IHR: 554 kJ/m 3 300 E -,  ‘200  0  50 0  ci)  cc oO -20  ‘4  cc  0 20 40 Crank Angle (deg ATDC)  0--20  LU  20 40 Crank Angle (deg ATDC)  Figures F.1.5 to F.1.7 : Diesel flowrate: 18.6 mg/inj IHRratio : 0.91 Knock Ratio: 0.81 Ignition Offset: 1.73 deg  226  VII-A-5 07-12-47a  e. D Co (I)  c)  E  100  aV  Knock: 1.3 bar IHR: 1604 kJ/m 3  1)  50  a)  (i)  Cu a)  a>  C -20  0 20 40 Crank Angle (deg ATDC)  a) V  50/  a) Co Cu a)  V a)  0 20 40 Crank Angle (deg ATDC)  Cu a)  0) a) •0  8  Knock:3.7 bar 11IR: 259 kJ/m 3 300  100  0 -20  —  nn  1 0 20 40 Crank Angle (deg ATDC)  9  Knock:2.2 bar IITR: 141 kJ/m 3  11  _,  a)  50  )i  0 20 40 Crank Angle (deg ATDC)  II  100  a  /NJN  0  •5-.  •0 a)  a: -20  20 0 40 Crank Angle (deg ATDC)  200  I’  cr a) U) Cu Cl)  C  .—  — a:  0 -20  VII-A-5 09-12-47c  -  C -20  a) 200 a:  0  Co 0  a)  \  -,  1  0  .w  E  100  a CD 0  100  — a:  VII-A-5 08-1 2-47b Co 0  1 I  —  •0 0  200  a)  100  0— -20  20 40 Crank Angle (cleg ATDC)  10  Figures F.1.8 to F.1.10 : Diesel flowrate: 16.2 mg/inj IHR ratio : 0.55 Knock Ratio: 0.58 Ignition Offset: 4.22 deg  227  VII-A-6 10-12-70a  Ca  V  c) Cl)  E  100  ci)  0 a) V C  50  ci)  Knock:2.2 bar 111R: 1640 kJ/m 3  :: 100  ci)  ,1  ci) C.)  V C  C  -20  20 40 0 Crank Angle (deg ATDC)  VII-A-7 1O-12-70a2 a)  a) I  o -20  0  11  Knock:2.5 bar ]HR: 1594 kJ/m 3  a),,’,’ V “vi. C.)  ci)  50  11)  200  100  ;  a) a)  0 -20  \*-  E  100  .2  *  0 20 40 Crank Angle (deg ATDC)  0 20 40 Crank Angle (cieg ATDC)  VII-A-7 l1-12-70b  a)  I  -o  /  Crank Angle (deg ATDC) 0) a) 0  12  Knock:4.0 bar I[-IR: 515 kJ/m 3 300  2  [00  -) ___,  \  50  ci)  a)  200  100  a) V a) c’  0 -20  0 20 40 Crank Angle (deg ATDC)  a)  I  n -20  Vll-A-7 12-12-70c  Ca  100 200  0 a) V C  50 a)  V  0 V C  100  a)  ci)  0 -20  0 20 40 Crank Angle (deg ATDC)  13  Knock:1.7 bar IHR: 346 kJ/m 3 300  z U)  20 40 Crank Angle (deg ATDC)  a)  I  -  —  I1 f\J 0 20 40 Crank Angle (deg ATDC)  14  Figures F.1.12 to F.1.14 : Diesel flowrate: 14.9 mg/inj ratio 0.67 Knock Ratio: 0.42 Ignition Offset: 2.91 deg 1  228  V11A1 1 13-28-20-47a  e.  Knock:3.1 bar IHR: 1540 kJIm 3 E  10o  $1  -  a-  200  50/  ci, 0 inn c a, ci)  a,  (>0 -20  20 40 0 Crank Angle (deg ATDC)  -20  V11A1 1 14-28-20-47b  fl H  4à 0’ 20 Crank Angle (deg ATDC)  15  Knock:4.8 bar IHR: 532 kJ/m 3 300 E -  i,200  a:100 a,  50  100 .2  C)  o  0 -20  0—-20  20 40 0 Crank Angle (deg ATDC)  VIIA 11 15-28-20-47c  50/  16  1  E  / ‘%%%,  100  a)  0 -20  20 40 Crank Angle (deg ATDC)  -,  o  0  4’ Knock:2.4 bar IIIR: 409 kJ/m 3  ioo  a-  A  c  0 20 40 Crank Angle (deg ATDC)  0—-20  A  N  0 ‘20 40 Crank Angle (deg ATDC)  17  Figures F.1.15 to F.1.17 : Diesel flowrate: 14.7 mglinj IHR ratio : 0.77 Knock Ratio: 0.51 Ignition Offset: 2.45 deg  229  VIIA 11 16-28-20-47a  Knock:4.4 bar IHR: 1622 kJ/m 3 300  100 200  •  50/  ioo ,%—, .  cx 0 -20 —  •  0”” -20  20 0 40 Crank Angle (deg ATDC)  V11A1 1 17-28-20-47b  0 40 20 Crank Angle (deg ATOC)  18  Knock:5.5 bar IHR: 585 kJ/m 3 300  100 200  ft  50Z  C -20 —  •  20 0 40 Crank Angle (deg ATDC)  0 -20  ——  V11A1 1 18-28-20-47c  0 20 40 Crank Angle (deg ATDC)  19  Knock:4.1 bar LHR: 538 3 kJ/m 300  100  -  200 50  0  -20 —  %%%%  0 20 40 Crank Angle (deg ATDC)  0 F° 0 -20  0 20 40 Crank Angle (deg ATDC)  20  Figures F. 1.18 to F. 1.20 : Diesel flowrate: 20.3 mg/inj IHR ratio : 0.92 Knock Ratio: 0.75 Ignition Offset: 1.13 deg  230  V11A12 19-28-20-70a  —  8  Knock:2.8 bar 3 11IR: 1747 kJ/m -300 E  100  -  0  a)  50  \\N  a) C.)  C -20  100 —  a)  20 40 0 Crank Angle (deg ATDC)  926 Crank Angle (deg ATDC)  V11A12 20-28-20-70b  8 1oo  \  i  J  a .5’ o a)  200  Knock:4.9 bar 3 1HR: 1077 kJ/m 300 E  \\  --  200  50’ •_%.__  a) •5  20 40 Crank Angle (deg ATDC)  -20  V11A12 21-28-20-70c  8  40 20 Crank Angle (deg ATDC)  22  Knock:2.2 bar IHR: 945 kJ/m 3 300 2  ioo  -,  200  a  ,Z% 50 a)  100  V  a) o  21  C -20  0 20 40 Crank Angle (cleg ATDC)  C -20  1  I  -J20  ‘--  Crank Angle (deg ATDC)  40  23  Figures F.1.21 to F.1.23 : Diesel flowrate: 17.2 mg/inj IHR ratio : 0.88 Knock Ratio: 0.45 Ignition Offset: 1.39 deg .  231  V11A12 22-28-20-70a  Knock:2.8 bar LHR: 1640 kJ/m 3 300 E  100  -  200  0  i5 50  100  0  a)  I  a)  (‘0 -20  0 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  V11A12 23-28-20-70b  24  Knock:2.8 bar IHR: 953 kJ/m 3 300  A  100 —  /  I-  a)  _/  \  200  fl  c  50Z  100  a)  -  oO -20  0-•-•• -20 20 40 Crank Angle (deg ATDC)  20 0 40 Crank Angle (deg ATDC)  V11A 12 24-28-20-70c  25  Knock: 1.0 bar IITR: 692 Id/rn 3 300 E  ioo  -,  a)  200  z  50/  100  o  -o a) 0  0 -20  0 20 40 Crank Angle (deg ATDC)  0—: -20  ,c 0 20 40 Crank Angle (deg ATDC)  26  Figures F.1.24 to F.1.26 : Diesel flowrate: 14.5 mg/inj IHRratio : 0.73 Knock Ratio: 0.34 Ignition Offset: 2.84 deg  232  VII.A4 25-16-20-70a  Knock:3.0 bar IHR: 1574 kJ/m 3 300 E  ci)  c3  100  200 50’  c_)  ci)  -D  -  -  cc  00  -20  0 20 40 Crank Angle (deg ATDC)  -20  V11A4 26-16-20-70b  5O/  0 20 40 Crank Angle (deg ATDC)  27  Knock:2.0 bar IHR: 308 kJ/m 3 .300 E  100 ci)  Ai  -,  200  7  cc cj  0  0  cc 0 -20  0’-20  0 20 40 Crank Angle (deg ATDC)  V11A4 27-16-20-70c  e.  28  Knock:3.0 bar 11IR: 248 kJ/m 3 300 E  100  -,  200  acc  50  100  V 0  0 20 40 Crank Angle (deg ATDC)  I’  cc 0 -20  20 0 40 Crank Angle (deg ATDC)  -20  .‘  0 20 40 Crank Angle (deg ATDC)  29  Figures F. 1.27 to F. 1.29 : Diesel flowrate: 22.0 mg/inj IHR ratio : 0.80 Knock Ratio: 1.48 Ignition Offset: 1.55 deg  233  VII-A-3 28- 16-20-47a  Ce  €  a) 0  Knock:1.0 bar IHR: 1604 kJ/m 3 300  C.)  E  100  200 .  50  V  1\  -D a) C) 0  0-20  0  20  40  Crank Angle (deg ATDC)  20  40  Crank Angle (deg ATDC)  Vu-A- 1 31-16-10-47a  Ce  I  1100 0” -20  30  Knock:1.2 bar IHR: 1614 Id/rn 3  c3) a)  V 300  E  100  200 .  -\ I  50  >s  C) V C) V C  1100  92  0  20  920  40  Vll-A-1 32-16-10-47b  Ce  31  Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  Knock:3.6 bar IHR: 275 kJ/m 3  0) a)  V 300 C)  E  100  200 .  50  100  I)  I  fl  C) V C  0  20  40  0 -20  g  0  20  Vu-A- 1 33-16-1O-47c  200  -v a) (a 0 V C  50  V  0-20  0  20 Crank Angle (deg ATDC)  32  40  33  Knock:1.2 bar 11IR: 178 kJ/m 3  n  100  .  40  Crank Angle (deg ATDC)  Crank Angle (deg ATDC) Ce  -  0 -20  I’  I” IJ 20  Crank Angle (deg ATDC)  Figures P.1.31 to F. 1.33 : Diesel flowrate: 13.3 mg/mi IHRratio : 0.65 Knock Ratio: 0.33 Ignition Offset: 3.67 deg  234  Vu-A- 1 34-16-1O-47-145GRITa  e.  Knock: 1.6 bar THR: 1573 kJ/m 3 300 E  100  -  200  0  I K  I  a)  5.  50/  4  a)  K K  100  o  a  K  a) )  K  0 -20  0-20  20 0 40 Crank Angle (deg ATDC)  VII-A- 1 35-16-1O-47-145GRITb  8  20 40 u Crank Angle (deg ATDC)  Knock:2.6 bar IHR:291kJ/m 3 300 2  100  a-  $1  -,  ‘7  50/  -r\  200 a)  a)  100  a) o  0 -20  Vu-A-i 36-16-1O-47-145GRITc  8  0 -20  20 40 Crank Angle (deg ATDC) 0  .  rw.UA  .  .  .  20 40 0 Crank Angle (deg ATDC)  Knock:1.3 ba ilIR: 65 kJ/m  11) V.jjj  C.,  2  100  -,  200  0  It  i5  K  HK 1K  50/ 0)  a) o  0 -20  20 40 0 Crank Angle (deg ATDC)  .  -20  20 40 Crank Angle (deg ATDC)  36  Figures F.1.34 to F.1.36 : Diesel flowrate: 13.6 mg/inj IHRratio : 0.22 Knock Ratio: 0.51 Ignition Offset: 1.44 deg  235  Vu-A- 1 37-16- 1O-47-345GRITa (0 C,,  Knock:2.O bar IHR: 1570 kJ/m 3 K  100  200  0 a)  50  7”  I’i\.  100  V a) C> -D C  J  0-20  0 20 40 Crank Angle (deg ATDC)  20  40  37  Crank Angle (deg ATDC)  Vu-A- 1 38-16-1O-70-345GRITb  Knock:2.3 bar  ll{R: 258 kJ/m 3 0  E  100  200 .  I  50  V a)  oO -20  0  20  100 l  *0  40  Crank Angle (deg ATDC)  Vu-A- 1 39-16-1O-47-345GRITc  a)  V  50 /_‘  a)  >.  0  ci)  V  a)  a)  cci C  -20  38  Knock: 1.3 baç IHR: 33 kJ/m 300  A  -,  a)  V  40  E  100  .  20  Crank Angle (deg ATDC)  0 20 40 Crank Angle (deg ATDC)  a)  z  1  200  100  I’ n  —,-.—---  .—-  -20  20 40 Crank Angle (deg ATDC)  —  39  Figures F.1.37 to F.1.39 : Diesel flowrate: 11.8 mg/inj ratio: 0.13 Knock Ratio: 0.59 Ignition Offset: 4.90 deg  236  V11A4 40-16-20-70-345GRITa  Knock:1.6 bar IHR: 1544 kJ/m 3 300 S  10o  200  0  100 CD  o  o-. -20  20 40 0 Crank Angle (deg ATDC)  -20  V11A4 41-16-20-70-345GRITb  ‘a  g  20 0 40 Crank Angle (deg ATDC)  40  Knock:2.1 bar IIIR: 327 kJ/m 3 300  S  100  -,  200  0  50  / 100 <a  C) 0  -  -20  20 0 40 Crank Angle (deg ATDC)  -20  V11A4 42-16-20-70-345GRITc  ‘a  20 40 Crank Angle (deg ATDC)  41  Knock: 1.5 bar IHR: 224 kJ/m 3 S  ioo  200 I  a) 50/  100  a)  I  -  C)  0 -20  20 0 40 Crank Angle (deg ATDC)  <  0 -20  \\%_  \_ ii20 Crank Angle (deg ATDC)  40  42  Figures F.1.40 to F.1.42 : Diesel flowrate: 19.6 mg/inj IFIRratio : 0.68 Knock Ratio: 0.74 Ignition Offset: 3.70 deg  237  V11A4 43-16-20-70-545GRITc  Knock:2.1 bar IHR: 290 kJ/m 3 30O E  10o  -  200  0  .5’  50 a)  (_) CI)  o  0 -20  0 -20  20 40 0 Crank Angle (deg ATDC)  V11A4 44-16-20-70--155GRITc  e.  0 20 40 Crank Angle (deg ATDC)  Knock:2.8 bar IHR: 289 kJ/m 3 300 E  100  -  200  0  50  Z ci  a) o  100  a)  0 -20  .  20 40 0 Crank Angle (deg ATDC)  100  0 -20  A 0 20 40 Crank Angle (deg ATDC)  238  Appendix F.2- Test Series Vu-A -1200 RPM Pressure and HRR Curves  239  VII-A29 20-1O-47-045a 300  1120  a’  a  ioo  o C) :5  60 40  E 200  I  80 C  Knock: 1.7 bar IHR: 1536 id/rn 3  a)  /  150  /  a) 0 %  1  50  20 0 -20  100  20 40 0 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-A-29 21-10-47-045b  Knock:3.1 bar IHR: 264 id/rn 3  120 a)  ioo 0  0  250  ,/Ir\.\  200  80  60J 7 .9 40 20 0 -20  \J  150 a)  1o0  x  0 20 40 Crank Angle (deg ATDC)  50 -  0-20  V1PA29 22-10-47-045c a’  a  -  200  80 0 C  II  250  100  2  Knock: 1.2 bar IHR: -12 kJ/m 3  300  1120  a’-’-’ 0 20 40 Crank Angle (deg ATDC)  ____  a)  150  60  I’ I ‘  a) 0  40  100  20  50  V  .E  -20  0 20 40 Crank Angle (deg ATDC)  0 -20  --  20 40 Crank Angle (deg ATDC)  3  Figures F.2. 1 to F.2.3 : Diesel flowrate: 14.2 mg/inj IHR ratio : -0.05 Knock Ratio: 0.38 Ignition Offset: 7.39 deg  240  VII-A-29 20-10-47- 145a ‘  3OU  120  a 250  100  E 200  80 •  C >‘  o  60 40  Knock:0.8 bar IHR: 1265 kJ/m 3  a  /  150  N  100 50  20 20  0a -20  0 20 40 Crank Angle (deg ATDC)  VII-A-29 20- 10-47-045b 120  0  100  nn 250  E 200  80  a a(1)  C >,  040  150 100  a 50  20 0 -20  z 0 20 40 Crank Angle (deg ATDC)  n  -20  VII-A-29 20-1 0-47-045c 120  300  .  C  5  Knock:0.5 bar IHR: -129 kJIm 3  250 E  80  a Cu  60  a  40 0 •0 C  “0 20 40 Crank Angle (deg ATDC)  a  100 U)  -  0)  —  0  0  Knock:3.8 bar IHR: 333 kJ/m 3  a  •60  0 •0  4  0) •0  D U) 0  0 20 40 Crank Angle (deg ATDC)  a  100  20 a  C— -20  I  0 20 40 Crank Angle (deg ATDC)  0-20  40  6  Crank Angle (deg ATDC)  Figures P.2.4 to F.2.6 : Diesel flowrate: 11.0 mg/inj : -0.39 Knock Ratio: 0.14 Ignition Offset: 6.00 deg 0 j 1 IHR  241  VII-A-29 26- 1O-47-095a 300  120  Knock:1.6 bar 1HR: 1783 kJ/m 3  a,  .0  80  250 E 200  .60  150  100  .  0  a,  C >‘  100  040 a,  a,50  20 C  0— -20  V.—..  20 40 0 Crank Angle (deg ATDC)  -20  VII-A-29 27- 1O-47-95b 300  120  •,  0 20 40 Crank Angle (deg ATDC)  7  Knock:4.0 bar IHR: 319 kJ/m 3  a, a, 0 0  0 C  E -  80  .  200  a,  150  60 a, 0  0  40  a, a,  20  C  250  100  20 0 40 Crank Angle (deg ATDC)  Vll-A-29 28-1 0-47-095c a, .0  k 4  0 -20  U  -20  100  120 100  0  300  0 20 40 Crank Angle (deg ATDC)  8  Knock:O.6 bar IHR: -105 kJ/m 3  0) CD  250  2 200 a,  a,  150  C  100 50  C.)  rs ‘  C  92o00 Crank Angle (deg ATDC)  40  9  Crank Angle (deg ATDC)  Figures F.2.7 to F.2.9 : Diesel flowrate: 14.3 mg/inj IHRratio : -0.33 Knock Ratio: 0.14 Ignition Offset: 6.50 deg  242  V11A29 29-10-47-095a2 1120  0) ci)  e.  250  100 U) U)  0  Knock: 1.3 bar IHR: 1601 3 id/rn  E 200  80 ci)  150  60  (I) C.)  100  40  0  a) i C) :5  20 0 -20  10  0 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  V11A29 30- 10-47-095a2 300  120  Knock: 1.4 bar IHR: 1607 id/rn 3  0)  250  100  E  Cl)  80  .  200  0  a)  .60  150  40  100  a) 20  50  1’  ci) U)  ‘  J  0  \  : .  C -20  0--20  0 20 40 Crank Angle (deg ATDC)  V11A29 31-10-47-125a 300  1120  Knock:0.7 bar 3 1073 kJ/m  IIIR:  250 E 200  100  .  11  a)  U)  0  0 20 40 Crank Angle (deg ATDC)  80  150  60 a)  100  40 i  20  cci  50  x -20  0 20 40 Crank Angle (deg ATDC)  -20  20 0 40 Crank Angle (deg ATDC)  12  243  VII-A-29 32-1O-47-125b  Knock:4.0 bar IHR: 316 kJ/m 3 U1j  120 1100 0 0  80  0)  250  //\\\\  200 150  60 C  c’4o o :5  20 0 -20  100  N.. ..,—.-.  .  50 0,. -20  20 40 0 Crank Angle (deg ATDC)  VII-A-29 33-10-47-125c 300  120  a  •0  100  0 20 40 Crank Angle (deg ATDC)  13  Knock:0.8 bar IHR: -129 kJ/m 3  250  80  200  •60  150  40  100  0  •0  50  20 •  0 -20  I  20 0 40 Crank Angle (deg ATDC)  ,—% —-  (I  -20  20 40 Crank Angle (deg ATDC)  14  Figures F.2.10 to F.2.14 : Diesel flowrate: 11.5 mg/inj IHRratio : -0.41 Knock Ratio: 0.19 Ignition Offset: 6.87 deg  244  VII-A-29 34-10-47-1 lOa ‘  120  Knock:0.9 bar IHR: 1220 kJ/m 3 0) •0  100  200  80  150  •60  a a 100 a 50  40 20 U  -20  -6  0 20 40 Crank Angie (deg ATDC)  120  15  nn  Knock:4.2 bar IHR: 311 kJ/m 3  0)  a  •0  100  250  E -)  80  40  o Crank Angle (deg ATDC)  VII-A-29 35-10-47-1 lOb  60  250  E  /\  20 0— -20  200  a 150 a 0 a 100 a a 50  z 0 20 40 Crank Angle (deg ATDC)  C -20  VII-A-29 36-10-47-1 lOc 120  0 20 40 Crank Angle (deg ATDC)  16  Knock:0.7 bar IHR: -123 kJ/m 3 0,  a  250  100  E  j80  a 150 a 0 a 100 a a 50  60 40 20 0— -20  ,  )C  I  0 20 40 Crank Angie (deg ATDC)  -20  20 40 Crank Angle (deg ATDC)  17  Figures F.2.15 to F.2.17 : Diesel flowrate: 10.6 mg/mi IFIRrauo: -0.40 Knock Ratio: 0.17 Ignition Offset: 6.88 deg  245  V11A29 10-1O-47a 300  120  0) a,  100  /  0  80 60 C  Knock:1.9 bar IHR: 1690 kJ/m 3  / -  /  250 E 200 150  is  a,  4  040  100  20  50  I’ I  a,  0 -20  20 40 0 Crank Angle (deg ATDC)  -20  V11A29 111047b 300  ‘120  ioo  250  80  200  ‘  20 0 40 Crank Angle (deg ATDC)  18  Knock:3.4 bar IHR: 366 kJ/m 3  cci 0  60 C  40  /\  1  ci,  150 II  a,  100 a,  a,  50  20 a,  0 -20  Z  0 20 40 Crank Angle (deg ATDC)  -20  VII-A-29 14- 1O-47b 300  120  a  a,  100 80 .  C  60  //fN\\  0 20 40 Crank Angle (deg ATDC)  19  Knock: 1.3 bar IHR: 229 kJ/m 3  250 200  a,  150 ci,  ‘40  a,  100  ci)  io .E  50  20 ci)  0 -20  0 20 40 Crank Angle (deg ATDC)  o-20  0 20 40 Crank Angle (deg ATDC)  20  Figures F.2.18 to F.2.20 : Diesel flowrate: 3.4 mg/inj  246  Vll-A-30  Knock:2.1  14-1O-70-045a  bar  3 IHR: 1706 kJ/m )I1fl  120  0)  a c,250 E  100 a  \\ 40  100  J  20  50  k  920  0 20 40 Crank Angle (deg ATDC)  920  20 40 Crank Angle (deg ATDC)  V11A30  Knock:5.8  IHR: 530  15-1O-70-045b  300  120  80  200  ‘N  4O  [50 100  “S  50  20  0  -20  bar  3 kJ/m  250 E  \  i  21  0  20  .iA  0  40  -20  20  0  40  22  40  23  Crank Angle (deg ATDC)  Crank Angle (deg ATDC) VII-A-3()  Knock:1.1  IHR:  16-10-70-045c  300  120  285  bar  3 kJ/m  0)  100  —250  80 60  200 a 150  40  100  20  50  /  \  —  -20  0  20  40  -20  20  Crank Angle (deg ATDC)  F.2.21 to F.2.23 : Diesel 0.54 Knock Ratio: 0.19 “ratio  Figures  Crank Angle (deg ATDC) fLowrate:  12.3  mg/inj  Ignition Offset:  5.81  deg  247  V11A30 17-10-70-345a  Knock:1.0 bar 11IR: 1613 kJ/m 3  3CC  120 a 100  a)  250  a> a>  0  so  200 a>  150  50 a>  100  40 a)  20 .5 0 -20  50 a, I  0 20 40 Crank Angle (dog ATDC)  0  VII-A-30 181070345b 120 a>  ioo a>  2  0  24  Knock:4.9 bar IHR: 605 kJ/m 3  30G  —  a  0 20 40 Crank Angle (deg ATDC)  80 60/)  250 8 200 a, 150  \  ,  a)  11 H  100  40  \  a)  20  a> a>  0 20 40 Crank Angle (deg ATDC)  50 -20  0 20 40 Crank Angle (dog ATDC)  25  Figures F.2.24 to F.2.25 : Diesel flowrate: 10.7 mg/inj V11A30 06-10-70a  Knock:1.8 bar ll{R: 2224 kJ/m 3  30G  120 a  a> ,  100  250  U,  U)  200  280  a)  150 .9  a>  100  40  80  20 a)  0  0 20 40 Crank Angle (deg ATDC)  0  VII-A-30 074070b 30C  1120 a 100 0  0  2 a  80  ,,/J  250 E ,200  .60  &150  40  a)  C  20  100  20 40 Crank Angle (deg ATDC)  26  Knock:33 bar IHR: 634 kJIm 3  A  Ii  II  H  50 a)  -20  20 40 Crank Angle (deg ATDC)  -20  0 20 40 Crank Angle (dog ATDO)  27  Figures F.2.26 to F.2.27 : Diesel flowrate: 17.0 mg/inj  248  V11A30 08-1O-70a 1120  a>  a  100 0 0  0  Knock:2.O bar IHR: 2135 id/rn 3  8o,7/  250 E p200  60  150  40  ioo  .  I’  C  I  -20  /  50  20 20 0 40 Crank Angle (deg ATDC)  0—-• -20  VII-A-30 09-1O-70b 300  1120 a  20 40 Crank Angle (deg ATDC)  28  Knock:3.7 bar 1HR: 632 kJ/m 3  a>  250  100 Co  80  p200  60  150  (‘40  100 a>  20  50  a-  a>  0 -20  x 20 0 40 Crank Angle (deg ATDC)  0—’ -20  0 20 40 Crank Angle (deg ATDC)  29  Figures P.2.28 to F.2.29 : Diesel flowrate: 14.5 mg/inj  249  V1LA31 05-20-47-345a 300  120  Knock:1.3 bar 3 IHR: 1670 kJ/m  C) a)  250  100 0,  200  80 a)  60  150  /  a) 0) a)  a) C)  100 50  20 0 -20  0— 0  0 20 40 Crank Angle (deg ATDC)  V11A3 1 06-20-47-345b 300  1120  .  /  0 -20  200  150  60/ .9 7 40 a) 20 C)  II 50 a)  0 -20  40 0 20 Crank Angle (deg ATDC)  VII-A-31 07-20-47-345c 300  1120  20 40 Crank Angle (deg ATDC)  31  Knock:1.3 bar IHR: 141 kJ/m 3  -  li  200  80  40  *  250  100  60  Knock:2.5 bar 3 IHR: 353 kJ/m  E  80  a-  30  250  100 C0  .E  20 0 40 Crank Angle (deg ATDC)  / \,  a)  150 a) 100 50  20 a)  .E -20  0 40 20 Crank Angle (deg ATDC)  0-20  40  32  Crank Angle (deg ATDC)  Figures F.2.30 to F.2.32 : Diesel flowrate: 23.2 mg/inj IHR ratio : 0.40 Knock Ratio: 0.51 Ignition Offset: 3.93 deg  250  V11A31  Knock:5.9  08-20-47-345a  1HR: 1907  300  120  bar 3 kJim  250  100  }  80  200  I’  1QdL  50  20 0 -20  0 -20  0 20 40 Crank Angle (deg ATDC)  40 0 20 Crank Angle (deg ATDC)  VIPA3 1  33  bar kJ/m 3  Knock:4.7  092047345b  LHR: 673  ‘2’,’,  ‘120 250  jioo  I  80  200  4  60/”  N  ‘40  ioo  :  50  20 -20  20 40 0 Crank Angle (deg ATDC)  -20  —“  V11A3 1 10-20-47-345c  40 0 20 Crank Angle (deg ATDC)  34  Knock:4.6 bar IHR: 525 kJ/m 3  300  120  *  (U  jwo  c250  80  200  C.)  /  20  — CU  A  50  (U  o -20  F.2.33 to F.2.35 : Diesel IHRratio : 0.78 Knock Ratio: 0.98 Figures  0 -20  *  0 20 40 Crank Angle (deg ATDC) flowrate:  25.8  *0  40 20 Crank Angle (deg ATDC)  35  mg/inj  Ignition Offset:  3.53  deg  251  VII-A-3 1 01 -20-47a  Knock:1.7 bar IHR: 275 kJ/m 3 sf11  120  0) •0  100  250  E -)  200  a)  150 1:0  40 0  20  Cu  a)  •0  0-20  z  36  20 0 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-A-3 1 02-20-47a 120 D  Knock:1.6 bar  250  100  E  Co  200  80 0 .  a)  150  60 a) a) Cu  () •0 ci) Cci C.) •0 C  IHR: 1827 kJ/m 3  ffl( 0) a)  a) a)  100  f  50  20 a)  0— -20  n  20 40 0 Crank Angle (deg ATDC)  -20  VII-A-3 1 03-20-47b 300  120 D Co Co  80  •60  37  Knock:2.7 bar IHR: 405 3 kJ/m -  0) a)  250  100  0  ‘0 20 40 Crank Angle (deg ATDC)  E  /)  ci) Cu  C  100  40  C— -20  A  I  20 0 20 40 Crank Angle (deg ATDC)  0 20 40 Crank Angle (deg ATDC)  38  Figures F.2.36 to F.2.38 : Diesel flowrate: 19.4 mg/inj IHRratio . : 0.68 Knock Ratio: 0.60 Ignition Offset: -0.48 deg  252  VII-A-31 O1-20-47-045a 300 120 (U  (U  250  ioo C:,  //Z\  a-  Knock:1.2 bar 1HR: 1669 kJ/m 3  200 U)  150  •60 t  U) (U  U 100 (U 0  50  20 0 -20  0 40 20 Crank Angle (deg ATDC)  -20  V1PA32 02-20-70-045b 300  (U120 e. 0)  / /  .  39  Knock:3.1 bar IHR: 364 kJ/m 3  250 E 200 a) iso  ioo 80  20 0 40 Crank Angle (deg ATDC)  (U (0  100  40 V (U  50  20 :5 •  U)  0 -20  0.-20  40 20 0 Crank Angle (deg ATDC)  V11A32 03-20-70-045c 300  1120 100 80 60  .  40  z /  40  Knock:1.5 bar IHR: 159 3 kJ/m  250 E 200 150  1’  100  20  50 (U  C -20  A  (U  -5  :5  .  20 40 Crank Angle (deg ATDC)  (U  (U 0  *  (U  (I)  .  4  0 20 40 Crank Angie (deg ATDC)  C -20  0 g  40  41  Crank Angle (deg ATDC)  Figures F.2.39 to F.2.41 : Diesel flowrate: 21.1 mg/inj il-IRratio : 0.44 Knock Ratio: 0.47 Ignition Offset: 2.56 deg  253  VII-A-32 1 1-20-70-045a ‘  300  120  0) a, •0  .0  100  200 a, Co  60  &‘40 0 a,  20 -V n ‘.  -20  250  E  80 .  0 20 40 Crank Angle (deg ATDC)  I  150 100 50 0-20  VII-A-32 1 2-20-70-045b  20 40 Crank Angle (deg ATDC)  42  Knock:5.6 bar IHR: 628 kJ/m 3 3Ui,  120  0)  a)  •0  100  250  E -)  80 C  Knock:7.O bar IFIR: 614 kJ/m 3  200  a) Co  60  150  a)  >‘  Ji  a) 100  040  a,  •0  20 I  :20 Crank Angle (deg ATDC)  VII-A-32 1 3-20-70-045c Co .0  120  43  Knock:6.4 bar ilIR: 634 kJ/m 3 0) a,  250  100  E  Co 0 a,  .  0  a,  a,  a  C  a, Cl)  200  100  (0 C, C  :20  2O 40 Crank Angle (deg ATDC)  0 20 40 Crank Angle (deg ATDC)  -20  IL  I  L’ 1 L  0 20 40 Crank Angle (deg ATDC)  44  Figures F.2.42 to F.2.44 : Diesel flowrate: 22.1 mg/inj Iratio: 1.01 Knock Ratio: 1.15 Ignition Offset: 1.46 deg  254  VI1-A-32 04-20-70a  Knock:3.3 bar  IHR: 646 kJ/m 3  300  f120 9,  /  100 0 0  ?  80  250  —  a 200 9)  —,  /  60  150  ‘  40  100  20  50  -20  V  C.. -20  0 20 40 Crank Angle (deg ATDC)  0 20 40 Crank Angle (dog ATDC)  VII-A-32 05-20-70b 120 o € 100 0 0  a•  a>  45  Knock:3.4 bar IHR: 673 lcJ/m 3  300  9)  250  80  / /  60 40  lj  9)  a 20 .20 -20  I  a200 a 150  )  0 20 40 Crank Angle (dog ATDC)  50 a x C—-20  20 40 Crank Angle (deg ATDC)  46  Figures F.2.45 to F.2.46 : Diesel flowrate: 21.2 mg/inj 1HRratio : 0.96 Knock Ratio: 0.95 Ignition Offset: 1.00 deg .  V11A41 12-10-47a-LB 300 a> a) >.250  120 €  ioo 0 0  ?  0.  80 60  C) a  Knock:1.6 bar IHR: 1804 kJ/rn 3  ,  —  200  A  .2  7/  150 a 100  40 20  50 a  I  -20  0 20 40 Crank Angle (dog ATDC)  -20  20 40 Crank Angle (dog ATDC)  V11A41  Knock:1.0 bar  13-10-47b-LB 120  a>  a. .2  250  80  .  60  /  20  /  150 a 100 ...  0 -20  200  9)  —  40  .2  ll{R: 234 Id/rn 3  300  a)  100 0 0  47  0 20 40 Crank Angle (dog ATDC)  50 a) 10 -20  -  0 20 40 Crank Angle (dog ATDC) -.  48  Figures F.2.47 to F.2.48 : Diesel flowrate: 11.0 mg/inj  255  Appendix F.3- Test Series VII-B-800 RPM Pressure and HRR Curves  256  V11B1 18-1O-47a 300  120  Knock:l.7 bar IHR: 833 kJ/m 3  ci)  250 E 200  100 Co  Co 2  0  80  ci)  150 C  0)  100 0)  ‘N  I’  a)  20 .2 -o .E 0 -20  ‘‘%  50  20 40 0 Crank Angle (deg ATDC)  20  V11B1 18-1O-47b 300  120  20 40 0 Crank Angle (deg ATDC)  Knock:O.2 bar IHR: 27 kJ/m 3  0) (I)  250  100  -  CO  200  80  ci,  150  60/ -5 C)  40  /  /  0) ci)  100  0)  50  20 ci,  •  0 -20  I  0 20 40 Crank Angle (deg ATDC)  OH -.)  VIPB1 18-1O-47c 300  1120  0 20 40 Crank Angle (deg ATDC)  2  Knock:O.2 bar IHR: 28 kJ/m 3  ci)  100  -  250  CO  .200  80  a,  .  C  150  60 0)  40 a,  20  N  ioo  \N  50 I  0  -20  0 20 40 Crank Angle (deg ATDC)  0 20 40 Crank Angle (deg ATDC)  3  Figures F.3.1 to P.3.3 : Diesel flowrate: 19.9 mg/inj IHRratio : 1.04 Knock Ratio: 0.93 Ignition Offset: 0.00 deg  257  VII-B-2 18-1O-70a ‘120  Knock:1.2 bar 3 IHR: 970 kJ/m 0) a>  .0  250 E . 200  100  -,  j80  ci, Cu  •60 C  C  4  U  -20  0 40 20 Crank Angle (deg ATDC)  -20  VII-B-2 18-1O-70b 300  120  Knock:2.3 bar IHR: 279 kJ/m 3  E -)  80  .  ZIN  •60  a) <0  /  200  a> Cu  100  a>  ci,  5  20 0 -20  40 0 20 Crank Angle (deg ATDC)  -20  VI1-B-2 18-10-70c 120  300  II J i  20 0 40 Crank Angle (deg ATDC)  Knock:l.0 bar IHR: 129 kJ/m 3  a)  0  100  250  E 200  80 a> Cu  60  1\  150  a>  U>  40  100  a>  20  Cci ci)  •0  z :20  5  0>  —  .0  C  4  250  100  040  0 20 40 Crank Angie (deg ATDC)  0) a>  .0  C >  I  100  40 a> 20  0 40 20 Crank Angle (deg ATDC)  50 0’-20  A.,]’  .  20 0 40 Crank Angle (deg ATDC)  6  Figures F.3.4 to F.3.6 : Diesel flowrate: 8.3 mglinj IHRratjo: 0.46 Knock Ratio: 0.42 Ignition Offset: 2.09 deg  258  0 r 2  VII-B-3 1 8-20-47a  Knock:2.2 bar 3 IHR: 837 kJ/m C, a, V  [00  250  E -)  80  200  a,  150  •60 a, >.  o  40  100  20  50  V a, V  c-20  Crank Angle (deg ATDC)  VII-B-3 18-20-47b  1  V  100  7  Knock:O.2 bar 3 IHR: 23 kJ/m  nn  120  g  0 20 40 Crank Angle (deg ATDC)  250  E -  a, (a  a:  •60 >  200 150 100  040 V a,  50  20 92  d -L)  0 40 20 Crank Angle (deg ATDC)  VII-B-3 1 8-20-47c C) a,  250  100 CO CO  ?  8  Knock:O.2 bar 3 IHR: 25 kJ/m  300  1120  0 20 40 Crank Angle (deg ATDC)  E 200  80  a, Ca  a:  150  a, Co  100  a,  a:  20 C)  Ca  z  .E -20  50  a,  V  0 40 20 Crank Angle (deg ATDC)  ri  L.. 20 0 40 Crank Angle (deg ATDC)  9  Figures F.3.7 to F.3.9 : Diesel flowrate: 24.0 mg/inj 1.08 Knock Ratio: 1.01 Ignition Offset: 0.00 deg ‘ratio  259  V11B4 18-20-70a  Knock:2.3 bar IHR: 1041 kJIm 3 .3VU  ‘120 Co  €  250  100 C)  200  80  0)  150  60 Cl) ‘  40”  cu100 Cl)  Co i  20  50  C)  .  C -20  0—. -20  0 20 40 Crank Angle (deg ATDC)  V11B4 18-20-70b 120  300  100  250  20 40 Crank Angle (deg ATDC)  10  Knock:2.1 bar IHR: 344 kJ/m 3  U)  0 .  ,200  80 60  /  40” i  C.)  20 C -20  .  ‘N  W  150 100 .  50 0— -20  0 20 40 Crank Angle (deg ATDC)  V11B4 1 8-20-70c 300  120  20 40 0 Crank Angle (deg ATOC)  11  Knock:2. 1 bar 1}IR: 209 kJ/m 3  0)  250  100 U) Co  -,  80 Cl,  60  /\  150 0  ioo  ‘  -  20 o .  C -20  ‘  50 — CO Co  0 20 40 Crank Angle (deg ATDC)  C— -20  iJ 20 40 Crank Angle (deg ATDC)  12  Figures F.3.10 to F.3.12 : Diesel flowrate: 20.1 mg/inj IHR ratio : 0.61 Knock Ratio: 1.00 Ignition Offset: 1.00 deg  260  VII-B-5 24-1O-47a 300  120  Knock:1.0 bar IHR: 88 kJ/m 3  0) a) V  .250 E 200  ioo U)  80  a .  C)  60  150  / /  a)  40  100  20  50  -20  0— -20  0 20 40 Crank Angle (deg ATDC)  VIIB-5 24- 10-47b 300  1120  20 40 Crank Angle (deg ATDC)  13  Knock:1.5 bar IHR: 201 kJ/m 3  0) a, •0  -.250  ioo U)  a.  C  200  80  a)  60  /  150 a, U)  /  40  (I)  100  V a)  t C.) :5  20  50  z  C — ‘-  -20  0 20 40 Crank Angle (deg ATDC)  0— -20  VII-B-5 24-10-47c 300  1120  20 40 Crank Angle (deg ATDC) -  -  14  Knock:1.3 bar IHR: 95 kJ/m 3  0) ci) V  —250  ioo U)  80  200  .60  150  ci) C  t  20  a  -20  i  100  40  a, ‘a)  50  I 0 20 40 Crank Angle (deg ATDC)  1 j 20 40 Crank Angle (deg ATDC) •%___  -20  15  Figures F.3.13 to F.3.15 : Diesel flowrate: 10.2 mg/inj IHRratio : 0.47 Knock Ratio: 0.86 Ignition Offset: 1.83 deg  261  VII-B-6 24- 1O-70a  Knock:3.1 bar 3 IHR: 1427 kJ/m  300  120 a,  250  100  E -) a, a,  •60 C  >‘  0  a, Cl)  a, a, a,  40 20  a, ID  200 150  .i.  100  Iip  50  z  92o Crank Angle (deg ATDC)  VII-B-6 24- 1O-70b 250  100 0  E 200  80  •60 C  040  a,  /  150  t  a, Cl)  100  a,  50  20 C  16  Knock:4.1 bar IHR: 617 kJ/m 3  30C  120  20 40 Crank Angle (deg ATDC)  0 -20  t  0 40 20 Crank Angle (deg ATDC)  o  ., i —  -20  VII-B-6 24-1O-70c 120  20 40 0 Crank Angle (deg ATDC)  17  Knock:2.9 bar 1HR: 508 kJ/m 3 a,  250  100  E .  80  200  a, .  60  C >‘  0  40  /  20  N  150 a, 0 ioo a,  50  i\ I  a, C  C— -20  0 40 20 Crank Angle (deg ATDC)  0— -20  20 V 40 Crank Angle (deg ATDC)  18  Figures F.3.16 to F.3.18 : Diesel flowrate: 11.8 mg/inj IHRratio : 0.82 Knock Ratio: 0.71 Ignition Offset: 0.42 deg  262  V11B6 24-10-70a-15.10.20  Knock:1.3 bar 3 IHR: 1244 kJ/m  ‘120  € 100  ci,  250  Co  80  200  •60  150  0  ci,  /‘  t  ci)  100 I’  i  20  ci)  0 -20  20 40 0 Crank Angle (deg ATDC)  50 -20  VII-B-6 24-10-70b-15.10.20  A  /  20 40 Crank Angle (deg ATDC)  19  Knock:0.8 bar 11IR: 448 kJ/m 3  120  a  ci)  250 E 200  _100 Co  80  150  /  .9  ioo  40” ci)  50  20 •  0 -20  0— -20  0 20 40 Crank Angle (deg ATDC)  V1FB6 24-10-70c-15. 10.20 300  120  a  ..4 N 1  -  20 40 Crank Angle (deg ATDC)  20  Knock:0.4 bar I}IR: -30 kJ/m 3  250 E 200  100 (I)  8O  0  •60 .9  /  -  ci)  150  Co  100  40 J)  .2  .E  50  20 0 -20  x 0 20 40 Crank Angle (deg ATDC)  0— -20  0 0 40 Crank Angle (deg ATDC)  21  Figures F.3.19 to F.3.21 : Diesel flowrate: 11.3 mg/inj IHR ratio : -0.07 Knock Ratio: 0.42 Ignition Offset: 16.71 deg  263  V11B7 24-20-47a 300  120  Knock:1.4 bar 3 IHR: 103 kJ/m  a)  250  100 Cl) Cl)  7\  80  __!  •60 / 5’ 040  200 a)  150  / /  a) Cl)  100  a)  rI  50  20 0 -20  0— -20  20 40 0 Crank Angle (deg ATDC)  VII-B-7 24-20-47b 300  A  40 0 20 Crank Angle (deg ATDC)  22  Knock:1.5 bar IHR: 290 kJ/m 3  250  100 0  .200  80 0  60 5’ 040  7  150  /  Cl)  100 50  20 &  C -20  0 -20  20 0 40 Crank Angle (deg ATDC)  V11B7 24-20-47c 300 120 a)  23  Knock:1.8 bar 3 IHR: 124 kJ/m  250 E 200  100  —  80 0  14  a)  60  / /  40  150  io 0 a -  .2  20 40 0 Crank Angle (deg ATDC)  o)  Cl)  .  11  50  20 a)  -20  0 20 40 Crank Angle (deg ATDC)  C  -20  j 4 r 20 40 Crank Angle (deg ATDC)  24  Figures F.3.22 to F.3.24 : Diesel flowrate: 28.1 mg/inj IHR ratio : 0.43 Knock Ratio: 1.24 Ignition Offset: 1.00 deg  264  VII-B-7 24-20-47a 120  Knock:1.1 bar IHR: 934 kJ/m 3 0)  a) a)  250  100  E  0  .  0  200  a)  a)  150  0  100 50  a) C) V C  a)  92o Crank Angle (deg ATDC)  Jt -20  VII-B-7 24-.20-47b  20 40 Crank Angle (deg ATDC)  25  Knock:1.2 bar IHR: 185 kJ/m 3  120 250  100  E -)  80  200  a)  150  •60 C  a) 0  >‘  040  100  I’  20 V C  20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-B-7 24-20-47c 300  120  26  Knock:1.6 bar 3 IHR: 74 kJ/m  )  100  250  (1  80  200  •60  150  i  40  100  C  a)  20  50  V C  20  0 40 20 Crank Angle (deg ATDC)  20  & I  20 40 Crank Angle (deg ATDC)  27  Figures F.3.25 to F.3.27 : Diesel flowrate: 19.5 mglmj IHRratio : 0.40 Knock Ratio: 1.27 Ignition Offset: 1.68 deg  265  VII-B-8 24-20-70a Co .0  300  120  0) Cl)  250  2 100 Co Co  0)  0 0)  Knock:4.1 bar IHR: 1847 kJ/m 3  E 200  80 a,  150  60  C >‘  0  4/  100  ,  CO  C) C  20 C -20  I  i[!  0)  ‘\  —.  o  0 20 40 Crank Angle (deg ATDC)  VII-B-8 24-20-70b  28  Knock :4.1 bar IHR: 659 kJ/m 3  300  120  20 40 Crank Angle (deg ATDC)  -  .0  250  100 Co Co  2  E  80 0) CU  0) >  ()  40  0) Co C.)  20  C  11  11  100  I’  0 -20  0 20 40 Crank Angle (deg ATDC)  20 40 Crank Angle (deg ATDC)  VII-B-8 24-20-70c  Knock:3.9 bar IHR: 521 kJ/m 3  300  120 .0  0)  250 E  2 100 Co Co  280  0.  0)  150  •60 C  4o 0) o  0  29  100 -  20  Jkj\  C --  -20  0 20 40 Crank Angle (deg ATDC)  -20  0 20 40 Crank Angle (deg ATDC)  30  Figures F.3.28 to F.3.30 : Diesel flowrate: 13.5 mg/inj 0.79 Knock Ratio: 0.94 Ignition Offset: 0.47 deg ratio 11  266  VII-B-8 24-20-70b 120  300  ioo  —250  —  Knock:4.2 bar 3 IHR: 647 kJ/m  80  50  2O c -20  0 -20  0 40 20 Crank Angle (deg ATDC)  VII-B-8 24-20-70c 300  —  120  20 40 Crank Angle (deg ATDC)  31  Knock:4.2 bar 3 IHR: 512 kJ/m  250  jloo  200  80  I:  iL?. I• ‘  i”’  ‘N\  50  20 a  -20  0 40 20 Crank Angle (deg ATDC)  -20  1  i  1i  20 40 Crank Angle (deg ATDC)  32  Figures F.3.31 to F.3.32 : Diesel flowrate: 20.1 mg/inj IHRratio : 0.79 Knock Ratio: 1.00 Ignition Offset: 0.50 deg  267  V11B8 24-20-70a- 15.10.00 300  120 a 100  250 E 200  a)  80  .  \  •60 .9 40”  /  a  150 0) C’,  100 50  20 .9  C -20  0 -20  0 20 40 Crank Angle (deg ATDC)  VII-B-8 24-20-70b-15.10.00 300  120 a 100 80 .9 ‘  60  /  / 40”  I\  \_  c -20  o— -20  0 20 40 Crank Angle (deg ATDC)  300  120 a 100 0 .  20 40 Crank Angle (deg ATDC)  34  Knock:3.2 bar IHR: 593 3 kJ/m  250 E p200  80  ii  0)  150  60  40  a 100 a  20  50  C -20  Knock:3.1 bar 1HR: 736 kJim 3  a  C  .9  33  250 E .200 a 150 a C’) 100  V11B8 24-20-70c-15. 10.00  Co Co  20 40 Crank Angle (deg ATDC)  50  20 .9  j  a  0)  .  Knock:3.4 bar IHR: 1547 kJ/m 3  x 0 20 40 Crank Angle (deg ATDC)  0— -20  1’ h  :  I  0 20 40 Crank Angle (deg ATDC)  35  Figures F.3.33 to F.3.35 : Diesel flowrate: 18.1 mg/inj IHRratio : 0.81 Knock Ratio: 1.03 Ignition Offset: 0.88 deg  268  V11B9  Knock:2.2  29-10-47-al  bar  IHR: 1460 kJ/m  3  300 ° 12 Ca  e. 100 Co  a-  250 E 2o0  I”-” \\  80  150  60/ .9 — 040  ID  100 I  •6  20 :5 .9  50  -,  c  -20  0-20  40 20 0 Crank Angle (deg ATDC)  VII-B-9  20 40 Crank Angle (deg ATDC)  36  Knock:4.2 bar  29-10-47-b 1  3 IHR: 545 kJ/m  300  120  ioo  250  80  200  0  0  .60  a)  //  150 0  100 a)  040  Ii  a)  50  20 5  C -20  0 -20  40 20 0 Crank Angle (deg ATDC) V11B9  40 0 Crank Angle (deg ATDC)  30C  120  IHR:  373  bar kJ/m 3  p200  9  150  H H  250  100 CO  a  80  60 40 °  / /  a)  20  —  ‘.  -20  I!  100  C  —  37  Knock:1.9  29-10-47-cl 19.  *  0 40 20 Crank Angle (deg ATDC)  50 C-20  J  20 40 Crank Angle (deg ATDC)  38  Figures F.3.36 to F.3.38 : Diesel flowrate: 13.1 mg/inj IHRratio : 0.69 Knock Ratio: 0.46 Ignition Offset: 1.30 deg .  269  VII-B-9 29-1O-47a  Knock:2.2 bar IHR: 1069 kJ/m 3  120  C> ID V  00  200  80 0  ID  150  •60 C  ID C’,  >  100  040  a, 50  20 V C  250  E  z  I.,  -20  20 40 0 Crank Angle (deg ATDC)  VII-B-9 29-1O-47b ‘  300  120 e  100  0 2a 40 Crank Angle (deg ATDC)  39  Knock:3.7 bar IHR: 459 id/rn 3  250 E -,  80  i  0— -20  .  200  a,  •60  C’)  E C  100  20 I,  -20  20 40 0 Crank Angle (deg ATDC)  20 40 Crank Angle (deg ATDC)  VII-B-9 29- 1O-47c ‘  AI  ci,  C  40  300-  120  40  Knock:2.4 bar IHR: 246 kJ/m 3  C)  ci)  250 E  100  II  -,  80 U) .  C  (a  60  40  100  20  5  C  20 40 0 Crank Angle (deg ATDC)  -20  -  0 20 40 Crank Angle (deg ATDC)  41  Figures F.3.39 to F.3.41 : Diesel flowrate: 16.7 mg/inj IHRratio : 0.54 Knock Ratio: 0.65 Ignition Offset: 0.96 deg  270  V11B10 29-10-70-a 300  1120  I\\\\  j100 Cl, Cl,  ?  80  0  Knock:1.8 bar 3 IHR: 1761 kJ/m  250 E 200 15O  60  0)  40  \  a)  20 •  C -20  100 50  x 0 20 40 Crank Angle (deg ATDC)  0--20  V11B10 294070b1 300  120  /  \  20 40 0 Crank Angle (deg ATOC)  42  Knock:1.1 bar IHR: 1046 kJ/m 3  0)  250  1oo In  80  ,  200  a)  60 .  40  150  /  Cl) Co  100 11)  I  ci)  50  20  \ .  C -20  0--20  0 20 40 Crank Angle (deg ATDC)  Vu-B-b 29-10-70-cl 1120  ,  300  20 40 Crank Angle (deg ATDC)  43  Knock:1.2 bar IHR: 421 U/rn 3  Cl)  250  100  II  80 150  60 40 a)  i o  20  PN  Cl)  100 50 I  -20  0 20 40 Crank Angle (deg ATDC)  -20  i-J 0 g  20 40 Crank Angle (deg ATDC)  44  Figures F.3.42 to F.3.44 : Diesel flowrate: 12.8 mg/inj IHR ratio : 0.40 Knock Ratio: 1.10 Ignition Offset: 1.31 deg  271  Vu-B-b 29-1 0-70a 300  ‘120  a  Knock:2.O bar IHR: 1385 kJ/m 3 -  a)  250  100  2  0  -  0  80 0  •60 C  100  40 20  .9  c  -20  0 20 40 Crank Angle (deg ATDC)  0’ 20 40 Crank Angle (deg ATDC)  Vu-B- 10  Knock:1.6 bar  29- 10-70b  IHR: 791 kJ/m 3  300  120  250  100  2  0  -  80  .  a •60  200  e  ii  150  C  a)  40  100  P  20  50  i’  0  -20  0 20 40 Crank Angle (deg ATDC)  92o  Vu-B- 10 29-1 0-70c 30C  120  :A  2 -  80  •60 C  /P\  e 150  40  100  20  50  -20  20 0 40 Crank Angle (deg ATDC)  46  Knock:1.7 bar IHR: 415 3 kJ/m  250  100  a  0 20 40 Crank Angle (deg ATDC)  a)  a 0 Cl)  45  -20  I 1’:  I%  V 0’ 20 40 Crank Angle (deg ATDC)  47  Figures F.3.45 to F.3.47 : Diesel flowrate: 14.2 mg/inj IHRratio: 0.52 Knock Ratio: 1.05 Ignition Offset: 0.69 deg  272  Vu-B- 10 29-10-70a-15. 10.59 ‘  120  Knock:1.5 bar  3 1HR: 1700 kJ/m C) ci, V  100  250  E -,  80  1  200  I’  11)  150  •60 C  1  ci, 0)  >‘  040  I  100  I  ci,  20  Cu Cu  C  50 0— -20  Crank Angle (deg ATDC)  VII-B-10 29-10-70b-15. 10.59 Cu  300  120  a Cu C >..  0  C’)  40 0 20 Crank Angle (deg ATDC)  48  Knock:0.4 bar IHR: 418 kJ/m 3  250  E -)  80 60  I’  L  C) ci)  100 U) Cl)  I  //\  200  Cu  150 Cu 0  40  100  ci) Cu C) V C  20  Cu ci,  I  I, ‘.4  -20  20 0 40 Crank Angle (deg ATDC)  50 C -20  Vu-B- 10 29-10-70c-15. 10.59 1120  .%  /%%  20 0 40 Crank Angle (deg ATDC)  49  Knock:0.5 bar IHR: 53 kJ/m 3 C) ci)  250  100  E  80  .  200  ci)  •  150  60 40  /  0— -20  I\  100  I  I  50  20 V C  ‘  40 0 20 Crank Angle (deg ATDC)  -20  4  \_ 20 40 Crank Angle (deg ATDC)  50  Figures F.3.48 to F.3.50 : Diesel flowrate: 14.0 mg/inj IHRratio: 0.13 Knock Ratio: 1.26 Ignition Offset: -3.37 deg  273  V11B11 29-20-47-al  ‘nfl  120  a, 250 E 200 a 150  1O0 0  8O  0  •6O  :  .9  4o  0  ca 100 —  20  50  C  0  —________  •0  .  Knock:1.8 bar IHR: 1624 kJ/m 3  -20  20  0  40  -20  Crank Angle (deg ATDC)  0  20  40  51  Crank Angle (deg ATDC)  Vu-B-i 1 29-2047b1  Knock:4.1 bar IHR: 540 kJ/m 3 300  120  e. 250  100 0  80 .  .200  a  j’N  60  C  150  a  ioo a  40/’  1  1’ 50  20  a —  0 -20  I.  -20  20 40 0 Crank Angle (deg ATDC)  V11B11 29-20-47-c 300 a 250  120 100  ——  20 40 Crank Angle (deg ATDC)  52  Knock:2.0 bar JHR: 409 kJ/m 3 .  .  0  C  80  .200  60 40  150 a 100  20  50 a  Z  C  —  -20  20 40 0 Crank Angle (deg ATDC)  0 -20  A  14  I  •I  20 40 Crank Angle (dag ATDC)  53  Figures F.3.5 ito F.3.53 : Diesel flowrate: 18.0 mg/inj IHRratio : 0.76 Knock Ratio: 0.50 Ignition Offset: 1.18 deg .  274  V11B1 1 29-20-47a  Knock:4.5 bar 111Th 1355 id/rn 3  300  —  0) a)  120 a)  250  jioo  40  100  20  50  -20  0 -20  20 40 0 Crank Angle (deg ATDC)  V11B 11 29-2047b a)120  54  Knock:4.4 bar IHR: 571 kJ/m 3  300  —  A  ‘ 20 40 0 Crank Angle (deg ATDC)  a)  jioo  c250  80  200  60 40  /  150  /  a)  100 a)  •0 a)  50  20  0  -20  20 40 0 Crank Angle (deg ATDC)  V11B4 1 29-20-47c 300  —  120  ——  -20  20 40 Crank Angle (deg ATDC)  55  Knock:4.1 bar IHR: 461 id/rn 3  0)  a)  0  ioo 200  80  /  0  60/ ‘  %t’’\  a)  150  I  100  40’ 2: 20 40 Crank Angle (deg ATDC)  d 2040 Crank Angle (deg ATDC)  56  Figures F.3.54 to F.3.56 : Diesel flowrate: 21.3 mg/inj IHR ratio : 0.81 Knock Ratio: 0.93 Ignition Offset: 0.50 deg  275  V11B42 29-20-70a 120  300  ioo  250  80  .200  —  \  60/_  150  040  Knock:4.6 bar IHR: 1866 kJ/m 3  ‘1  (1) a,  50  20  x 20  4o  57  Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-B-12 29-2070b 120  300  ioo  250  a  60  200  \  80  Knock:4.5 bar IHR: 971 kJ/m 3  150  /  h  a,  100 50  20 -20  0 -20  0 20 40 Crank Angle (deg ATDC)  V11B12 29-20-70c 30C Cu  ioo 80  20  a, •0  IL 20 40 Crank Angle (deg ATDC)  58  Knock:4.2 bar IHR: 793 kJ/m 3 I  250  Is  J  \  20 40 Crank Angle (deg ATDC)  200  920  2O4O Crank Angle (deg ATDC)  59  Figures F.3.57 to F.3.59 : Diesel flowrate: 19.7 mg/inj IHRratio : 0.82 Knock Ratio: 0.93 Ignition Offset: 0.00 deg  276  V11B13 18-1O-52LBa 300  a  Knock:1.3 bar 3 IHR: 1097 kJ/m  250  100 0  200 Il Co Co 100 Co  :5  20 40 0 20 Crank Angle (deg ATDC)  -20  -20  VII-B-13 18-1O-52LBb 300  120  a  0* 20 40 Crank Angle (deg ATDC)  60  Knock:0.8 bar 3 LHR: 69 kJ/m  250  100 Co  ,200  80  a, 150 ‘1, 0  40  50  20 •  100  .5  0 -20  0—S-20  0 20 40 Crank Angle (deg ATDC)  V11B13 18-10-52LBc 300  120  a  ki  250 E 200  0  —  60 40  61  Knock:1.5 bar 3 IFIR: -14 kJ/m  a,  100  150  /  50  20 C., .  . 20 40 0 Crank Angle (deg ATDC)  z  C -20  20 0 40 Crank Angle (deg ATDC)  0-20  20 40 Crank Angle (deg ATDC)  —  62  Figures F.3.60 to F.3.62 : Diesel flowrate: 12.7 mg/inj IFIRratio : -0.20 Knock Ratio: 1.89 Ignition Offset: 0.00 deg .  277  VII-B-15 1 8-20-48LBa ‘  30C  120  Knock:1.7 bar IHR: 1211 kJ/m 3  DC  cC)  250  100  E 200  80 ci, .  C  60  /7  150 ci)  c’40  Cl) a)  20  Ca a)  100  I  50  “  I  C V  -20  0 20 40 Crank Angle (deg ATDC)  20  Vu-B- 15 18- 10-7OLBb 30G  120  0 40 Crank Angle (deg ATDC)  63  Knock:O.5 bar IHR: -21 kJ/m 3  V  100  2 -,  80  .  200  a)  150 100 a)  5  20 .20 -20  A.  :20  0 20 40 Crank Angle (deg ATDC)  Vll-B- 15 18-1O-7OLBc 120  15 20 40 Crank Angle (deg ATDC)  64  Knock:0.7 bar IHR: -84 kJ/m 3 DC ci)  250 2 . 200  100  -,  80  ci)  60  150  40  100  C  I  20 V C  -20  0 20 40 Crank Angle (deg ATDC)  J  20 40 Crank Angle (deg ATDC)  65  Figures F.3.63 to F.3.65 : Diesel flowrate: 15.6 mg/mi IIIR.ratio : 3.99 Knock Ratio: 1.44 Ignition Offset: 0.00 deg  278  Vu-B- 15 18- 1O-7OLBa  Knock:1.3 bar IHR: 1335 kJ/m 3  120 U)  80  250 E 200  .60  150  ioo 0  —  .9  ‘40  ci,  /  ci) U)  *  4  50  20  4  ci)  :6  .9  4  134’  100  1,  -20  0 -20  41  —.__________  20 40 0 Crank Angle (deg ATDC)  Vu-B- 15 18- 1O-7OLBb  20 40 0 Crank Angle (deg ATDC)  66  Knock:O.2 bar 300  e.  3 IHR: 25 kJ/m  250  100 U)  80  200  .60  150  .9  ci,  U)  100  40  50  20 ci)  :5  z  CC  —  -20  20 40 0 Crank Angle (deg ATDC)  20 40 0 Crank Angle (deg ATDC)  Vu-B- 15 18- 1O-7OLBc  67  Knock:1.2 bar kJ/m IHR: 220 3 QLIU  120 <I)  250  100  E I4  -‘  80  , ci)  .  11  60  I  C  ci)  20 .9,1  :20  I’ 1’  100  40  0  20  Crank Angle (deg ATDC)  40  1  -20  Ai  20 40 0 Crank Angle (deg ATDC)  68  Figures F.3.66 to F.3.68 : Diesel flowrate: 11.8 mg/inj IHRrauo: 8.91 Knock Ratio: 7.69 Ignition Offset: 22.49 deg  279  VII-B-16 1 8-20-7OLBa 1120  Knock:3.2 bar IHR: 1489 kJ/m 3 a) 250 E . 200  100  -,  80  1)  4  150 a) 0  100  a)  50 a)  I’  0 -20  —-  20 40 0 Crank Angle (deg ATDC)  Vu-B- 16 1 8-20-7OLBb 300  1120  *  I!  20 0 40 Crank Angle (deg ATDC)  69  Knock:2.3 bar IHR: 329 kJ/m 3  ci)  [00  C.)  250  E -,  80  0  200  ci) Ca  •60  /  C  >‘  040  150  a) a)  20  (a  100 50  ci) n  -20  0— -20  0 20 40 Crank Angle (deg ATDC)  Vu-B- 16 1 8-20-7OLBc  ft  L  20 40 Crank Angle (deg ATDC)  70  Knock:3.4 bar kJ/m IHR: 355 3  120 ci)  250  100  E -,  80  200  ci)  •60  Icu  C  40  /  20 20  0 20 40 Crank Angle (deg ATDC)  I  n.  1  100  I  50  ,  20 Crank Angle (deg ATDC)  71  Figures F.3.69 to F.3.71 : Diesel flowrate: 15.8 mg/inj : 1.08 Knock Ratio: 1.48 Ignition Offset: 0.59 deg 0 j 1 IHR  280  V11B17 24-1O-47LBa  Knock:1.5 bar 3 IHR: 1383 kJ/m “vu  120 :2. 100  250 .  ci,  80  .200  •60  150  0 C  0  100  40 a,  20  50  •  za,  0 -20  20 40 0 Crank Angle (deg ATDC)  0— -20  V11B17 24-1047LBb 300  40 20 0 Crank Angle (deg ATDC)  72  Knock:0.3 bar IHR: 27 kJ/m 3  1120 250  100 0  80 .60 C .  040  /  /  150 a, 0  100 .5 50  20 .9  I  C -20  20 40 0 Crank Angle (deg ATDC)  0 20 40 Crank Angle (deg ATDC)  VII-B-17 24-10-47LBc  73  Knock:1.8 bar 3 77 kJ/m  IHR: 300  120 a,  250  100 0  0 .  200  80  a,  60  150  a,  100 a) .5 50  40 20  I’\\%  a,  .9  I  C -20  20 0 40 Crank Angle (deg ATDC)  0 -20  J  40 0 20 Crank Angle (deg ATDC)  74  Figures F.3.72 to F.3.74 : Diesel flowrate: 12.6 mg/inj IHR ratio : 2.90 Knock Ratio: 6.25 Ignition Offset: 22.40 deg .  281  V11B48 24-10-7OLBa-15.10.40 30C  1120 a 100 a 80 •  Knock:1.3 bar IHR: 1360 kJ/m3  250  60  a 10°  40  I\  •0  a  50  20  /1  a  C -20  40 0 20 Crank Angle (deg ATDC)  I  -20  V11B 18 24-10-7OLBb-15.10.40 30C  1120  a  a  100 a 80  200  60  150  40 a 20 0  5°  40 20 Crank Angle (deg ATDC)  75  Knock:0.2 bar 3 ll{R: 21 kJ/m  250 a a6, 100  c  -20  40 20 0 Crank Angle (deg ATDC)  40 0 20 Crank Angie (deg ATDC)  V11B18 24-10-7OLBc1-15.10.40 30C  1120  a  76  Knock:0.2 bar 3 IHR: 2OkJ/m  a 250  100  /\ 6,  80  60 9 4Ø V a 20 .  \  200 a 150 a  ioo 50 a  ç  I  40 0 20 Crank Angle (clog ATDC)  40 0 20 Crank Angle (deg ATDC)  V11B18 24-10-7OLBc1-15. 10.40 30C  1120 a 100 C-  Knock:1.3 bar IHR: 1901 U/rn3  a  1  250 E  a  80  200  60//J\N 40  a 6,  a 150 100  77  I  50  20 a  c  -20  40 0 20 Crank Angle (deg ATDC)  o— -20  40 0 20 Crank Angle (deg ATDC)  78  Figures F.3.76 to F.3.78 : Diesel flowrate: 19.2 mg/mi ll{Rratio : 93.02 Knock Ratio: 7.34 Ignition Offset: 27.94 deg  282  Vll-B49 24-20-47LBa1 300  120  Knock:3.2 bar IHR: 1551 kJ/m 3  —250  ioo  U  200 iso  60  ;.c  ‘  100  40”  50  20 0 -20  40 0 20 Crank Angle (deg ATDC)  0 -20  V11B49 24-20-47LBb 300  120  jioo  v.250  80  200 $3 150  \  20 40 Crank Angle (deg ATDC)  79  Knock:1.6 bar IHR: 161 kJ/m 3  100 —50  20 40 20 Crank Angle (deg ATDC)  920  VIIB-19 24-20-47LBc 300  40 20 Crank Angle (deg ATDC)  Knock:3.1 bar IHR: 194 kJ/m 3  4  250 200  80  80  I :/‘ [: E  50  20  c  -20  40 20 0 Crank Angle (deg ATDC)  -20  ii i[$ 20 40 Crank Angle (deg ATOC)  81  Figures F. 3.79 to F. 3.81: Diesel flowrate: 13.1 mg/inj 1.20 Knock Ratio: 1.93 Ignition Offset: 1.09 deg T1 “ ratio  283  VII-B19  Knock:3.1  24-20-47LBa  bar  3 IHR: 1522 kJ/rn  30G  120  —250 200  50  20 0 -20  0—. -20  20 0 40 Crank Angle (deg ATDC) VII-B-19  20 40 Crank Angle (deg ATDC) Knock:0.5  24-20-47LBb  82  bar  3 THR: 31 kJ/m  300  —  120  —250  ioo  200  80  150 C)  40  100  20  5°  0 -20  0-•— -20  0 40 20 Crank Angle (deg ATDC)  V11B19 24-20-47LBc 300  —  120  •  20 40 0 Crank Angle (deg ATDC)  83  Knock:3.3 bar 3 IHR: 209 Id/rn  250 E 200  1O0 80  IZ” 50  20 Z  -20  0 20 40 Crank Angle (deg ATDC)  0 -20  1  (t  i Jj  20 40 Crank Angle (deg ATDC)  84  F.3.82 to F.3.84 : Diesel flowrate: 13.3 mg/mi 6.67 Knock Ratio: 6.24 Ignition Offset: 0.56 deg ‘ratio  Figures  284  VII-B21 29-1O-47LBa 300  1120 € 100  Knock:L2 bar IHR: 1573 kJ/m 3  p250  0  0  200 .5 150  80 60  I  a,  50  20 •0 .5  a  C -20  40 0 20 Crank Angle (deg ATDC)  85  -20 Crank Angle (deg ATDC)  VII-B-22 24-1O-7OLBa 30C  g12o  Knock:L9 bar 3 IHR: 1849 kJ/m  it  1100 U) 6Q  0  80 60  —  C  40  /  /  a 150 a ioo a  I  0  0 20 40 Crank Angle (deg ATDC)  c— -20  V11B22 241O7OLBb 300  1120 € 100 80 60 .  / (  0 20 40 Crank Angle (deg ATDC)  86  Knock:O.2 bar IHR: 25 U/rn 3  250 E 200 a 150 a 100  U)  0  it  j  50  20  c  \  I I’  a  —  /  /  40  50  20  a  •6  I  c  -20  0 20 40 Crank Angle (deg ATDC)  0 20 40 Crank Angle (deg ATDC)  V1FB22 24-1O-7OLBc 300 a 25O  1120 e.  ioo  87  Knock:1.4 bar IHR: 513 kJ/m 3  0  0.  2200 a, 150  80 / /  60 .  I’  a 100  40 -U  50  20 a  -  c  I \L  -20  20 0 40 Crank Angle (deg ATDC)  -20  1$ 20 40 Crank Angle (deg ATDC)  88  Figures F.3.86 to F.3.88 : Diesel flowrate: 10.8 mg/inj IHRratio : 20.29 Knock Ratio: 7.58 Ignition Offset: 22.01 deg .  285  V1PB22 29-1O-7OLBa-15. 11.25  Knock:1.2 bar 3 IHR: 1300 kJ/m  300  120 a)  V 100 0 0 )  80  .  250  200  •60  150  40  caloo a)  1  V a)  .2  20  CU a)  V .  C -20  50  0— -20  20 40 0 Crank Angle (deg ATDC)  40 0 20 Crank Angle (deg ATDC)  VII-B-22 29-1O-7OLBb-15.11.25  Knock:0.3 bar 111Th 242 kJ/m 3  30C  120 V  100  89  250  0  20C a, 150  8O .60 C  a) U)  ci)  a) 20 C.) :5 0  ioo  .50 Ca (1)  C— -20  20 0 40 Crank Angle (deg ATDC)  V11B22 29-1O-7OLBc- 15.11.25 30C  120 & 100  40 0 20 Crank Angle (deg ATDC)  90  Knock:0.7 bar IHR: 382 kJ/m 3  C’,  250  0)  60  ,200 a) 150  40  100  20  50  80 .  a)  E  i .2 V .E  0 -20  0 20 40 Crank Angle (deg ATDC)  -20  /  \  40 0 20 Crank Angle (deg ATDC)  91  Figures F.3.89 to F.3.91 : Diesel flowrate: 14.3 mg/inj IHR ratio : 1.58 Knock Ratio: 2.60 Ignition Offset: 2.53 deg .  286  V11-B-23 29-20-47LBa1 300  1120 0) Co Co  0 0) >  Dl  a)  250  100  E 200  80 0) Cu  60  01 (0  0  40  CO  20  Knock:1.2 bar IHR: 1482 kJ/m 3  150  I’  100  ci)  C  C -20  40 20 0 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-B-23 29-20-47LBb1 CO  0) CO  0)  250  100  E -  80 60  0  40  CO  20  0 V C  200  Li)  C  >  Knock:0.8 bar IHR: 240 id/rn 3  300  120  Co  a  //\\  (a 0) Co Cu 0) 01 CO  100  s 4  50  a)  z -20  40 20 0 Crank Angle (deg ATDC)  0  V11-B-23 30C  120  a)  ci) V C  250 E  80  .  200  a) 150  60 0)  40 C) V C  93  Dl 0)  100 0  20 4 Crank Angle (deg ATDC)  Knock:2.2 bar kJ/m I[IR: 417 3  29-20-47LBc 1  Co Co  92  0  II)  100  a)  C]  -20  /%  50  20 40 20 0 Crank Angle (deg ATDC)  X  —-  -20  I  20 0 40 Crank Angle (deg ATDC)  94  Figures F.3.92 to F.3.94 : Diesel flowrate: 7.2 mg/inj JHR ratio : 1.74 Knock Ratio: 2.74 Ignition Offset: 0.08 deg  287  VII-B-24 24-20-7OLBa  Knock:4.2 bar IHR: 2080 kJ/m 3  120 a) V  100  250  E 200  80 a)  150  •60 C >‘  040  100  20  50  C  C -20  0 -2 )  20 0 40 Crank Angle (deg ATDC)  Vll-B-24 24-20-7OLBb Cu a) 0 0  120  30C  100  250 E -) . 200  80  0 a) C >.  0  V a) Cu C) V C  60 40  U  20 40 Crank Angle (deg ATDC)  95  Knock:1.7 bar IHR: 549 kJ/m 3  a, Cu  /  a) 0  100  /  a)  50  20 U  -20  0 20 40 Crank Angle (deg ATDC)  20  VII-B-24 24-20-7OLBc  20 40 Crank Angle (deg ATDC)  96  Knock:4.5 bar 3 IHR: 651 kJ/m 5VU  120 a)  250  100  E -  80 60 40  /  N  a) a) a)  200  150  i (\  100  j  a)  50  20 0— -20  a) I  0 20 40 Crank Angle (deg ATDC)  n  i\  —  -20  20 40 Crank Angle (deg ATDC)  97  Figures F.3.95 to F.3.97 : Diesel flowrate: 23.3 mg/inj IHR ratio : 1.19 Knock Ratio: 2.68 Ignition Offset: 0.30 deg  288  Appendix F.4- Test Series VII-B-1200 RPM Pressure and HRR Curves  289  VII-B-26 18-1O-70a-14. 13.25  (U  Knock:0.6 bar 1494 kJ/m 3  )3QQ E  Co  U)  -  a)  a) C  200  It  50  a)  100  a)  a)  cc  a) (U C)  -  C  -20  0 20 40 Crank Angle (deg ATDC)  -20  d’  20  40  Crank Angle (deg ATDC)  VII-B-26 18-1O-70b-14. 13.25  CU  Knock:1.4 bar .30C  3 IHRPCE. 540 kJ/m  S  100  I  I  -  200  a) (U  50  7  cc  N  a)  /1  100  a) a)  cc  CU  ‘I  -20  0 20 40 Crank Angle (deg ATDC)  a)  I  , -—  -20  VII-B-26 18-1O-70c-14. 13.25  (U  30C—  0 20 40 Crank Angle (deg ATDC)  2  Knock:0.5 bar 367 kJ/m 3 PCE 1111  a) U) Cl)  S  100  -  -  0  a)  a) C  cc  50  (1)  >,  0  a) cc  1) C) C  0  -20  200  0  20  40  Crank Angle (deg ATDC)  a)  I  100  A I’  0  -20  0”” 20 40 Crank Angle (deg ATDC)  3  Figures E.4.1 to E.4.3 : Diesel flowrate: 10.0 mg/inj IHRratio: 0.68 Knock Ratio: 0.32 Ignition Offset: 2.78 deg  290  VII-B-29 24-1O-47a  Knock:1.1 bar 3 IITRPCE. 120 kJ/m  8  .300  ioo  E -200  50/  C)  /  cc C)  100 ci) C)  ci)  /  cc  0  20 40 Crank Angle (cleg ATDC)  VIFB29 24-1O-47b  (U  8  20 40 Crank Angle (deg ATDC)  4  Knock:2.6 bar 285 kJ/m 3 .300  ioo  E  c)  -, -  0 .9  920  C)  200  cc  50  a)  100  ci)  ‘5 20 40 Crank Angle (deg ATDC)  I  0 -20  VIIB29 24-1O-47c  -  (U  300  8  20 40 Crank Angle (cieg ATDC) Knock:1.0 bar 3 111 kJ/m ‘ 93 PCE:  ioo  (I)  -,  -200 C)  I.  a)  5’  cc  °  100  0  a)  I’  cc  C)  (U  20 40 Crank Angle (deg ATDC)  O  0 20 40 Crank Angle (deg ATDC)  6  Figures E.4.4 to E.4.6 : Diesel flowrate: 8.1 mg/inj IHR : 0.33 Knock Ratio: 0.40 Ignition Offset: 4.59 deg ratio  291  VII-B-30 24-1O-70a 8  300  Dioo  E  Knock:1.9 bar 3 PCE 1997 kJ/m 1 ‘  200 .9  /  \  50/  4.  •  E  IVy  Crank Angle (cieg ATDC)  ‘‘ 0.—•*” -20 0 20 40 Crank Angle (deg ATDC)  VII-B-30 24-1O-70b  Knock:5.1 bar 509 kJ/m 3  0  -20  0  20  40  X  7  300  100  Co  I  I  \  / 50/ c_)  -)  -200 ci)  a4  100  (8  1It  —  0 -20  0 20 40 Crank Angle (deg ATDC)  Z  0 -20  VII-B-30 24-1O-70c 300  100  0  20  40  8  Crank Angle (deg ATDC)  Knock:1.3 bar 3 ‘PCE 486 kJ/m  -  00  .2  E  0  -20  0 20 40 Crank Angle (deg ATDC)  o-. -20  A 20  40  Crank Angle (deg ATDC)  Figures E.4.7 to E.4.9 : Diesel flowrate: 11.8 mg/inj IHRru: 0.95 Knock Ratio: 0.25 Ignition Offset: 3.42 deg  292  VII-B-30 24-1O-70a-14.13.OO  a 100  0)  300  Knock:1.7 bar 3 IHRE: 1579 kJ/m  -  0 -  200  a)  50  /—r-,\  a)  ci)  A  100  a) —  0  cci  -20  0 20 40 Crank Angle (deg ATDC)  920  VII-B-30 24-1O-70a-15.9.30  a  0)  300  K  *—  0 20 40 Crank Angle (deg ATDC)  10  Knock:0.7 bar 3 IHRE: 1443 kJ/m  E  100 0  -200 a)  cci  50/  100  ‘N  . E  20  20 0 40 Crank Angle (deg ATDC)  920  VII-B-30  Knock:1.0 bar 3 INRE: 1491 kJ/m  24-1O-7Ob-15.9.3O  a  •*—  0 20 40 Crank Angle (deg ATDC)  “3—  -  100 Cl>  -,  200  /  0  I’  cci  50/  a)  ioo  I’  a) ‘  E  a)  -20  20 0 40 Crank Angle (deg ATDC)  12  Crank Angle (deg ATDC)  VII-B--30 24-1O-70b-15.9.31  -  a ioo  Knock:0.5 bar 406 kJ/m 3 .3OC  “3— E  /\  200 a) cci  50/  100 ci) ‘—,.  (I)  00 -20  —  I  0 20 40 -ko 20 40 Crank Angle (deg ATDC) Crank Angle (deg ATDC) Figures F. 10 to F. 14: Diesel flowrate: 9.0 mg/inj  13  293  Vll-B-31  Knock:2.5 bar 3 IHRPCE: 314 kJ/m  24-20-47a1  300  8  ioo Cn  E -,  r\\\ /  cj  1  50/  200  100 a)  V a) C)  —  9  <a  20 40 Crank Angle (deg ATDC)  (1)  I  0 20* 0 20 40 Crank Angle (deg ATDC)  VII-B-31  Knock:1.6 bar 3 111 219 kJ/m ‘ PCE  24-20-47b1  -  14  a E  ioo Cl)  200  \  a)  V  50l.  \  100 a)  V ci)  4.  I  .  —  <a  C)  20  40  9d  Crank Angle (deg ATDC) VII-B-31 24-20-47c1  -  20  a  .300  15  100  E  Knock:1.3 bar 195 kJ/m 3 PCE 1 ’ 11  Cl)  0  40  Crank Angle (deg ATDC)  It  <a  / 50/  a)  100  V a)  ci)  o  cci ci) I  20 40 Crank Angle (deg ATDC)  (  0 -20  , L 20  40  16  Crank Angle (deg ATDC)  Figures E.4.14 to E.4.16 : Diesel flowrate: 37.9 mg/inj IHR ratio : 0.89 Knock Ratio: 0.81 Ignition Offset: 1.18 deg  294  VIFB31 24-20-47  Knock:1.3 bar  122 kJ/m 3 300 (.3—  100  50  E  0 -20  0 -20  —  Z  20 40 0 Crank Angle (deg ATDC)  VII-B-31 24-20-47b  17  Knock:1.9 bar 300  100  20 40 0 Crank Angle (cieg ATDC)  “PCE  328 kJ/m 3  -  -200  i .2  50/\  0 -20  20 40 0 Crank Angle (deg ATDC)  [00  Z  o-*-•• -20 0 20 40 Crank Angle (deg ATDC)  V11B31 24-20-47c  18  Knock:2.1 bar 153 kJ/m 3 w300 e  100  I’  -  CO  -200 /  1  /  .950/  ..  \  a)  -  1  100  -  .2  E  0 -20  0 20 40 Crank Angle (deg ATDC)  I  0  -20  *.MA1  W  0 20 40 Crank Angle (deg ATDC)  19  Figures E.4.17 to E.4.19 : Diesel flowrate: 20.4 mg/inj ‘ratio 0.47 Knock Ratio: 1.07 Ignition Offset: 3.31 deg  295  V11B32 24-20-70a1  Knock:1.1 bar 1624 kJ/m 3 300  ioo  C’)  0  c.3  E  7//\  a)  200  a)  50  a) 100 ci)  °  E  C -20  0 20 40 Crank Angle (deg ATDC)  I  0 -20  VII-B-32 24-20-70b1 .300  ioo  0  20  40  20  Crank Angle (cieg ATDC)  Knock:4.5 bar 671 kJ/m 3 PCE:  E  Co  —)  j  200  0  50  100  °  a)  V  a)  a) °  -20  0  20  40  I  -20  Crank Angle (deg ATDC)  VII-B-32 24-20-70c1 e.  0  20  40  21  Crank Angle (deg ATDC)  300  Knock:1.0 bar 532 kJ/m 3 PCE 1111  E  100 C,)  -) a)  /\ / 50/  200  I  ioo  C)  a)  V a)  I\ _J  j\  C)  cci  0  0  20  40  Crank Angle (deg ATDC)  ,.  20  0  20  40  Crank Angle (deg ATDC)  Figures E.4.20 to E.4.22 : Diesel flowrate: 8.1 mg/inj IHR ratio : 0.79 Knock Ratio: 0.23 Ignition Offset: 2.07 deg  296  V11B32 24-20-70a 1  Knock:2.6 bar PCE T ‘ 1 1819  300  3 kJ/m  ci CI) Ci)  E  IJV  -  I  \  V .  .2 •E  50  _)  200 a,  ,  0 -20  I\ 0 20 40 Crank Angie (deg ATDC)  0 -20  --  Z  0 20 40 Crank Angle (deg ATDC)  VII—B-32 24-20-70b1  -  23  Knock:4.3 bar  IHR: 710 kJ/m 3 300  a,  k  -200 50/ 100 .2  —  0 -20  0 -20  --  0 20 40 Crank Angle (deg ATDC)  VII-B-32  e. D (0  24  Knock:2.9 bar  24-20-70c1  -  0 20 40 Crank Angle (deg ATDC) II-IR.  649 kJ/m 3  300 E  --  IVV  -  Cl)  -  -200 50 100  920  20 40 Crank Angle (deg ATDC)  Z  -20  A 020 40 Crank Angle (deg ATDC)  25  Figures E.4.23 to E.4.25 : Diesel flowrate: 16.8 mg/inj THRratio: 0.91 Knock Ratio: 0.66 Ignition Offset: 0.74 deg  297  VIJB-32 24-20-70a1  Knock:1.1 bar 3 111 1599 kJ/m ‘ PCE  .300 c.5 E  8  ioo U)  -  a)  0 2  200  a)  a)  50 C.)  a,  -  0 40 ‘S  Z  20 0 40 Crank Angle (cleg ATDC)  0— -20  V11B32 24-20-70b1 300 c.62  8  ioo  I -  20 40 Crank Angle (deg ATDC)  a  Knock:2.7 bar 322 kJ/m 3 PCE 1 ‘  -)  U)  -  200  a)  .2  >  50 -  100 a)  o 0  a)  o  0 -20  20 0 40 Crank Angle (deg ATDC)  Z  -20  VIFB-32 24-20-70c1 300  ioo U)  2  .2  50  /  a)  \f%%%  a) 100 2 I  a)  0  20 40 Crank Angle (deg ATDC)  27  it  -200  ,\\  a)  0 20 40 Crank Angle (deg ATDC)  Knock:1.2 bar 197 kJ/m 3  8  /  A  ‘20  A0  j  20 40 Crank Angle (deg ATDC)  ae  Figures E.4.26 to E.4.28 : Diesel flowrate: 8.9 mg/inj IHR ratio : 0.61 Knock Ratio: 0.43 Ignition Offset: 2.34 deg  298  V11B32 24-20-70a  Knock:55 bar 3 IHR: 613 kJ/m  30O a) Cl)  200 .  50/  100 .%-  V  I  2?  0  -20  0 20 40 Crank Angle (deg ATDC)  -—-i ---  0  I  -20  V11B32 24-20-70b  .  .a  Knock:4.4 bar 3 IHRE: 582 kJ/m  300 Cl) Ci)  L..  0 20 40 Crank Angle (deg ATDC)  E  inn “-“-‘  I  i  50/  .1 /  _••)  ‘..  -200 11)  4.  (1)  ioo  a) 0)  V  2?  —,.——.  -20  0 20 40 Crank Angle (deg ATDC)  VThB-32 24-20-70c 100  I  0— -20  0  20  40  Crank Angle (deg ATDC)  Knock:53 bar 638 kJ/m 3  Ø3J -,  0  200 V50/ 00 0)  0 -20  0 -20  ._  0 20 40 Crank Angle (deg ATDC)  I  N-  0 20 40 Crank Angle (deg ATDC)  :i  Figures E.4.29 to E.4.3 1: Diesel flowrate: 19.8 mg/inj Wratio 1.10 Knock Ratio: 1.21 Ignition Offset: 2.36 deg  299  VII-B-33 28-1O-47a 300  Knock:1.7 bar 1486 kJ/m 3 “PCE  CD  100 200 a,  50/ !:100 V CD  G)  .2 •E  —  0 -20  0 20 40 Crank Angle (deg ATDC)  I  WTt  0---— 40 -20 0 20 Crank Angle (deg ATDC)  V11B33 28-1O-47b  -  r a  Knock:5.8 bar 464 kJ/m 3 300  0) inn 0  E -,  0  -200  i  I  a,  1 V .  c  I  50/  v -20  40 0 20 Crank Angle (deg ATDC)  I  0 ——X 20 40 -20 0 Crank Angle (deg ATDC)  V11B33 28-1O-47c  -  s  Knock:1.1 bar 291 kJ/m 3  a,  100 0  _)  -200 a,  I  cc  50 0 V  R  100 ‘  a,  cc kj  0 -20  0 20 40 Crank Angle (deg ATDC)  Z  -20  4k,.  40 0 20 Crank Angle (cieg ATDC)  Figures E.4.32 to E.4.34 : Diesel flowrate: 9.5 mg/inj IHRratio : 0.63 Knock Ratio: 0.20 Ignition Offset: 3.54 deg  300  VII-B-33 28-1O-47a  Ca  U) U)  -5  ci,  ci,  a)  0  -20  20 40 0 Crank Angle (deg ATDC)  VII-B-33 28-1O-47b  ci, I  0) ci 0  92 Crank Angle (deg ATDC)  35  Knock:5.8 bar 512 kJ/m 3 PCE 1 ‘  300  2  100  -5  200  0  a)  //  I.  cc  50  a)  ci)  100  I  ccci)  ci) Ca C) •0 C  ui  cc (a  0  a) •0 C  100  a) a)  •0 a)  U) U)  200  cc  50  >‘  C  CCC  E  100  a•0 C  0,  Knock:1.8 bar 3 1506 kJ/m PCE  -20  20 40 0 Crank Angle (deg ATDC)  a)  I  I  i  —_4 ki  -20  VII-B-33 28-1O-47c 300  20 40 Crank Angle (deg ATDC)  36  Knock:1.2 bar 3 309 kJ/m 111 ‘ PCE  ci U) U)  E  100  -3 -  a a) •0 C  cc  50  ci)  >‘  0  100  a) a)  •0 a) 0 0 C  200  a)  <a  I  cc -20  0 20 40 Crank Angle (deg ATDC)  A  ci)  I  0  b 1  It t:  I  4 j fi  20 40 Crank Angle (deg ATDC)  37  Figures E.4.35 to E.4.37 : Diesel flowrate: 11.4 mg/inj IHRratio : 0.60 Knock Ratio: 0.21 Ignition Offset: 4.77 deg  301  V11B33 28-1O-47a  Knock:3.0 bar 952 kJ/m 3 300  cs— 100  -,  (I)  -200  I  50/’  100  C.)  1i’tjti  .2  0 -20  0 -20  —  20 40 0 Crank Angle (deg ATDC)  V11B33 28-1O-47b 300 100  ii  I1JIlIII  40 20 0 Crank Angle (deg ATDC)  a  Knock:4.8 bar “PCE 475 kJ/m  -  U) ci)  -,  -200  ioo  2  0 -20  0 20 40 Crank Angle (deg ATDC)  I  0 -20  VII-B-33 28-1O-47c  20 40 0 Crank Angle (deg ATDC) Knock:1.2 bar IHR: 262 kJ/m 3  .300 100  -  U)  -200 50/\  I  a: -  E  0 -20  0 20 40 Crank Angle (deg ATDC)  I  0 -20  0  20  40  40  Crank Angle (cieg ATDC)  Figures E.4.38 to E.4.40 : Diesel flowrate: 21.1 mg/inj ‘ratio 0.55 Knock Ratio: 0.25 Ignition Offset: 2.29 deg  302  VII-B34 28-1O-70a  Knock:1.4 bar 3 IHRPCE: 1888 kJ/m 300 E  100  -)  0  -  200  a)  A  a) 50,  a)  100 a)  a) —  o  20  0 20 40 Crank Angle (deg ATDC)  VIFB34 28-1O-70b  —  d”  0 20 40 Crank Angle (deg ATDC)  41  Knock:3.7 bar 872 kJ/m 3 .300 E  //1  0  -,  \\\  50  a) 0)  200 A  100  0)  a)  o  —  0 -20  0 20 40 Crank Angle (deg ATDC)  Z  -20  V11B34 28-1O-70c  -  0  20  40  42  Crank Angle (deg ATDC)  Knock:O.9 bar 3 IHRPcE. 705 kJ/m 300  ioo U)  -,  0  a)  E 200  I-  0)  50/  a)  1  100  a) V 0)  o  (0  20 40 Crank Angle (deg ATDC)  920  f) \  20 40 Crank Angle (deg ATDC)  Figures E.4.41 to E.4.43 : Diesel flowrate: 12.3 mg/inj IFIR ratio : 0.81 Knock Ratio: 0.24 Ignition Offset: 2.24 deg  303  VIIB-34  Knock:1.4 bar  28-1O-70a  3 IHRE: 1972 kJ/m  a a?  .300  E  100  (I)  -  a?  -  0  ci  7/  200  a:  50,  100 a,  -D  a:  a) C)  20 40 0 Crank Angle (deg ATDC)  a,  Z  U__  -20  VIFB.34 28-1O-70b  Cu  .300  40 20 Crank Angle (deg ATDC)  Knock:3.9 bar 3 IFWE: 899 kJ/m  E  100 Cl)  -,  /1  -  a,  200  cu  a:  It  4  100  o  a,  -V >‘  a,  a: -20  a a? ioo  40 0 20 Crank Angle (deg ATDC)  Z  jii  o  -20  V11B34 28-1O-70c  *0  20 40 Crank Angle (deg ATDC)  Knock:0.8 bar 3 779 kJ/m 300 E  Cl)  a?  200  0  a,  a:  50  a)  100 a,  %-N  a:  a,  Cu  C)  20 40 Crank Angle (deg ATDC)  920  20 40 0 Crank Angle (deg ATDC)  Figures E.4.44 to E.4.46 : Diesel flowrate: 13.3 mg/inj IHRratio : 0.87 Knock Ratio: 0.21 Ignition Offset: 2.83 deg  304  V11B34 28-1O-70a  Knock:1.8 bar 3 IIIlPCE. 1584 kJ/m .300  8  E  ioo c)  -,  -  I.  a)  / 50/  /\  E  cc €100 a)  cc  a) °  200  0)  0  -20  0 20 40 Crank Angle (deg ATDC)  Z  0--  -20  VII-B-34 28-1O-70b  -  47  Knock:2.9 bar 924 kJ/m 3 .300  8  2  ioo 0  -  2OO  0  a)  cc  50,/’  a)  \  V a)  20 40 Crank Angle (deg ATDC)  fij  a)  cc I  -20  V11B34 28-1O-70c  -  A  0 20 40 Crank Angle (deg ATDC)  a  Knock:1.3 bar 3 IHRPCE. 758 kJ/m 300  8  ioo  2  Cl>  -,  -200  50/ 5’  \  fr f  o  V  cc  00  E  A  0 20 40 Crank Angle (deg ATDC) -  -20  0 20 40 Crank Angle (deg ATDC)  96  hd  0 20 40 Crank Angle (deg ATDC)  Figures E.4.47 to E.4.49 : Diesel flowrate: 19.9 mg/mi IHRratio : 0.82 Knock Ratio: 0.44 Ignition Offset: 2.17 deg  305  VII-B-35 28-20-47a  Knock:1.9 bar 1533 kJ/m 3 300  100 -  C  E  50  \  0 -20  0 20 40 Crank Angle (deg ATDC)  200  F X  -20  V11B35 28-20-47b  a  /  so  Knock:5.3 bar 3 IHl{PCE. 469 kJ/m  300  ci 100  20 40 Crank Angle (deg ATDC)  200  50 100  c  -20  0 -20  ——-—  0 20 40 Crank Angle (deg ATDC)  VII-B-35 28-20-47c  0 20 40 Crank Angle (deg ATDC)  51  Knock:1.4 bar 347 kJ/m 3 30C  0)  100 0)  -,  -200 /F\\% 50  20 40 Crank Angle (deg ATDC)  92 Crank Angle (deg ATDC)  Figures E.4.50 to E.4.52 : Diesel flowrate: 15.7 mg/inj : 0.74 Knock Ratio: 0.27 Ignition Offset: 3.58 deg 0 IHRj  306  V11B35 28-20-47a  Knock:4.7 bar 472 kJ/m 3 300  a a)  100 Cl)  0  /  \  -  a)  50 /‘  c  0 -20  200  100  0 20 40 Crank Angle (deg ATDC)  Z  0 -20  k’—• 0  20 40 Crank Angle (deg ATDC)  V11B35 28-20-47b a Cl Cl)  s  Knock:5.7 bar 3 IHRE: 509 kJ/m  300 E  IVV  -200 50/  •  \\ 100  0  -20  0 •-.-*.“.‘  20 0 40 Crank Angle (deg ATDC)  -20  V11B35 28-20-47c a  0 20 40 Crank Angle (deg ATDC)  Knock:4.6 bar 420 kJ/m 3 300  ci)  100  -  Cl)  1  -200 50/  ioo  I 0 -20  0 20 40 Crank Angle (deg ATDC)  Z  0 -20  &jf  •—  0 20 40 Crank Angle (deg ATDC)  Figures E.4.53 to E.4.55 : Diesel flowrate: 18.4 mg/inj IHR : 0.82 Knock Ratio: 0.81 Ignition Offset: 2.65 deg ratio  307  VII-B-35  Knock:4.9 bar 3 “PCE 1097 kJ/m  28-20-47a 300 c’5  100  -  -200  4.  -s  G100  C.)  .2  0 -20  0 -20  -----  0 20 40 Crank Angle (deg ATDC)  VII-B35 28-20-47b 300 100  I  I’  ii  —2  .  0 20 40 Crank Angle (deg ATDC)  ‘c  Knock:4.1 bar 3 kJ/m 487 “PCE:  -  :r’  t  20 4057 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VIIB-35 28-20-47c  -  300  Knock:4.7 bar 468 kJ/m 3 “PCE•  ci)  100 Cl)  -,  -200  50/  ioo ..  c  u -20  0 20 40 Crank Angle (deg ATDC)  ci)  Z  0 -20  4.\.  0 20 40 Crank Angle (deg ATDC)  Figures E.4.56 to E.4.58 : Diesel flowrate: 15.9 mg/inj IHRrauo: 0.96 Knock Ratio: 1.13 Ignition Offset: 2.50 deg  308  VII-B-36 28-20-70a .0  Knock:2.2 bar 0)  ci)  E  100 C,)  a) 0  a)  a) 0 C  cc a) CO Cu 0) a)  -o a)  cc  C) a)  z  C  59  Crank Angle (deg ATDC)  VII-B-36 28-20-70b  cci .0  Crank Angle (deg ATDC)  0) a)  a) CO  -)  ci)  - 200  0  a)  50  ioo  >.  C) •0 a) C) C  300  2  100  Cl)  a) •0 C  Knock:4.7 bar 3 IHRPCE. 944 kJfm  0  -20  cci .0  I t.  !  0 20 40 Crank Angle (deg ATDC)  VII-B.-36 28-20-70c  ii  ‘I  20  40  60  Crank Angle (cleg ATDC)  0) a)  Knock:1.9 bar 3 IFWPCE 922 kJ/m  a)  E  100 (0 a)  0  a)  I.  0) 0 C  Cu  cc  50  a) CO Cu ci) a)  >..  0  cc  cci C)  0 20 40 Crank Angle (deg ATDC)  Cu a)  z  20 0 40 Crank Angle (deg ATDC)  61  Figures E.4.59 to E.4.61 : Diesel flowrate: 15.2 mg/inj 0.98 Knock Ratio: 0.40 Ignition Offset: 1.43 deg ‘ratio  309  V11B36 28-20-70a  Knock:3.8 bar 3 IIIRPCE. 951 kJ/m  .300 ci)  ioo  7/1  200  1  .50 100  (3’  I 0 -20  0 20 40 Crank Angle (deg ATDC)  Z  0 -20  —  VII-B-36 28-20-70b  20 40 Crank Angle (deg ATDC)  Knock:6.0 bar 961 kJ/m 3  IHR. 11)300  100 200 50/  ;  920  0 20 40 Crank Angle (deg ATDC)  VIFB36 28-20-70c  2040 Crank Angle (deg ATDC)  Knock:4.4 bar 913 kJ/m 3  )  300 100  I  I  \  -200 cii  .9507  • E  o  -20  I  —  0 20 40 Crank Angle (deg ATDC)  0 -20  % L 1 IIj  VJ  ---  Z  0 20 40 Crank Angle (deg ATDC)  64  Figures E.4.62 to E.4.64 : Diesel flowrate: 15.9 mg/inj : 0.95 Knock Ratio: 0.73 Ignition Offset: 1.27 deg ratio  IHR  310  V11B36 28-20-70a  Knock:4.6 bar 1649 kJ/m 3  IHR: .300  ci)  100 U)  0..  -200  4  i5  S  ci)  /_%é  920  20 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VIIB36 28-20-70b  Knock:4.5 bar 960 kJ/m 3 300  100  -200  /“\  50  :  Cu •  100  —  I: —.  0  0  -20  0  20  40  Z  -20  h 20  40  Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-B-36 28-20-70c  Knock:4.8 bar 805 kJ/m 3  Crank  ci  .300  ci)  Dioo  E 200  50  .S  \  —  0 -20  0  20  : 40  Crank Angle (deg ATDC)  Z  0 -20  Ill L’i X XIi1!  0 20 40 Crank Angle (deg ATDC)  Figures E.4.65 to E.4.67 : Diesel flowrate: 19.8 mg/inj IHRratio : 0.84 Knock Ratio: 1.07 Ignition Offset: 1.37 deg  311  VII-B-36 29-20-70a-14.12.27  Knock:4.5 bar  IHR: 1082 kJ/m 3 300  200  /7 100  •  0 -20  0 20 40 Crank Angle (deg ATDC)  Z  kw  0-—--20 20 40 Crank Angle (deg ATDC)  VII-B-36 29-20-70b-14.12.27  -  Knock:4.7 bar IHR: 1105 kJ/m 3 300  a)  100 0  -,  I  -200  —  :1  50/  920  -  a)  Il I  N  100  0 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-B-36 29-20-70c-14.12.27  Knock:4.3 bar 3 HIRE: 1056 kJ/m 300  [00  50  //  A  200  1  100  i-.1Iii•  E  0 -20  0 20 40 Crank Angle (cieg ATDC)  I  0 -20  M  20 40 Crank Angle (deg ATDC)  70  Figures E.4.68 to E.4.70 : Diesel flowrate: 17.1 mg/inj 0.96 Knock Ratio: 0.90 Ignition Offset: 0.02 deg ’ratio 1 “  312  VII-B-36  Knock:4.9 bar IHR: 1690 kJ/m 3  29-20-47a-14.12.40  300 100  j 1  5O/  920  N  -200  9  0 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  VII-B-36 29-20-47b-14.12.40  300 Cl) Cd)  inr W  E  -200  0 0  H  50_  >‘  •  Knock:4.9 bar 3 IIflE: 702 kJ/m  ioo  0 -20  —  0 20 40 Crank Angle (deg ATDC)  I  0 -20  VII-B-36 29-20-47c-14. 12.40  -  I  I  0 20 40 Crank Angle (deg ATDC) Knock:4.0 bar I1): 604 kJ/m 3  300  a, 100  I  C-  U)  -200  0  50  I  / 20  40  Crank Angle (deg ATDC)  92d O20 40 Crank Angle (deg ATDC)  Figures E.4.71 to E.4.73 : Diesel flowrate: 15.0 mg/inj IHRrauo: 0.86 Knock Ratio: 0.83 Ignition Offset: 2.50 deg  313  VII-B-37 18-1O-47LBa-18-12.47  Knock:2.2 bar .300  a  2 ioo U) 2 a  3 lHRE: 1506 kJ/m  E  -,200  II  1  11  H  50/ 100  o  ci) ci)  I  o  0  20 40 Crank Angle (deg ATDC)  Z  -20  VII-B-37 18-1O-47LBb-18  a  .3O0  2  ioo  ‘  •* 0  20 40 Crank Angle (deg ATDC)  74  Knock:0.2 bar 3 IHRE: 26 kJ/m  E  (0  _)  2  -200 ci,  ci)  100 ci)  V ci)  -  .2  r  c  -0  —  0 20 40 Crank Angle (cleg ATDC)  VII-B-37 18-1O-47LBc  a 2  0)  A  -)  D)  a-  - 200 CI)  50/  40  75  Knock:3.0 bar : 364 kJ/m 3 PCE  E  .  20  IHR  ioo V  0  Crank Angle (deg ATDC)  A  ci)  100 ci)  00  •E  -20  2O  Id  %. 20 40 Crank Angle (deg ATDC) ,  0 20 40 Crank Angle (deg ATDC)  76  Figures E.4.74 to E.4.76 : Diesel flowrate: 11.6 mglmj 11Wratio : 14.18 Knock Ratio: 12.38 Ignition Offset: 16.15 deg  314  V11B37 18-1O-47LB-zaftera  Knock:1.4 bar  e. ioo U)  30O  a  -  •  3 IHRE: 1592 kJ/m  E -,  50/’  200  N\ :ioo 0)  Ct  0)  o  (0  920  e. ioo  0 20 40 Crank Angle (deg ATDC)  a, I  0 -20  V11B37 18-1O-47LB-zafterb  11  40 0 20 Crank Angle (deg ATDC)  n  Knock:O.3 bar 26 kJ/m 3  IHRPCE:  .300 E -,  -200 50  a)  100 a)  0 0)  cc  o  a)  .  20 40 Crank Angle (deg ATDC)  I  -  -.J  VII-B-37 18-1O-47LB-zafterc  e. ioo  .300  40 0 20 Crank Angle (deg ATDC)  78  Knock:1.0 bar 3 IHRE: 1743 kJ/m  E  Ci)  _)  -200  R  cc 50/  100  I  Ct (0  C) -  Crank Angle (deg ATDC)  9  0 20 40 Crank Angle (deg ATDC)  79  Figures E.4.77 to E.4.79 : Diesel flowrate: 13.5 mg/inj IHRratio : 66.81 Knock Ratio: 4.08 Ignition Offset: 30.50 deg  315  VII-B-37  Knock:1.2 bar IFT1: 1622 kJ/m 3  18-1O-47LBa-18-13.40.53  8 U) Co  V  E  100  -3  a)  200  (U  V  cc  50/  ci)  I’  100 a)  I’  a)  V a)  cc  C)  -20  0 20 40 Crank Angle (deg ATDC)  a,  0-20  VII-B-37 18-1 O-47LBb  a, 300  8  0 20 40 Crank Angle (deg ATDC)  80  Knock:0.2 bar 3 IITRpCE: 26 kJ/m  V  E  ioo C)  -,  200 /S\  a)  cc  50 —  a) Co 80  V ci)  CD  C)  c  100  cc  0 -20  0 20 40 Crank Angle (deg ATDC)  80 CD Z  d-0 20 40 Crank Angle (deg ATDC)  VII-B-37 18-1O-47LBc  300  8 U) C,,  200 50  a)  100  V ci) C)  Knock:0.2 bar 169 kJ/m 3 ’PCE 1  100  0 V  81  a)  cc  0 -20  (U  0 20 40 Crank Angle (deg ATDC)  a)  Z  -20  0 20” 4Ô Crank Angle (deg ATDC)  Figures E.4.80 to E.4.82 : Diesel flowrate: 7.8 mg/inj IHRratio : 6.39 Knock Ratio: 1.18 Ignition Offset: 45.15 deg  316  VII-B-38 18-1O-7OLBa  -  .300  8  Knock:1.2 bar 3 I[TRPCE. 1807 kJ/m  c.  E  ioo  -200 50,  7—j%’\  \  0  G100  o  ci)  cc  ci) ‘3  (  -0  0 20 40 Crank Angle (deg ATDC)  20 40 Crank Angle (deg ATDC)  V11B38 18-1O-7OLBb  Knock:0.2 bar .300  8  ioo  3 IHRPCE. 27 kJ/m  E  0  -,  -200 0)  ‘  50/’’’\  cc 0)  ioo  ci)  cc  ci,  o  0 -20  0 20 40 Crank Angle (deg ATDC)  Z  -i..)  VIFB-38 18-1O-7OLBc  —  .300  8  0 20 40 Crank Angle (deg ATDC)  Knock:1.0 bar 265 kJ/m 3 141 ‘ PCE  E  ioo ci)  _)  N\_,_  A  -200  ‘1  50/” —  o  00 ci)  cc  0  c  0 20 40 Crank Angle (deg ATDC)  920  20 40 Crank Angle (deg ATDC)  Figures E.4.83 to E.4.85 : Diesel flowrate: 8.0 mg/inj IHR ratio : 9.85 Knock Ratio: 6.21 Ignition Offset: 18.58 deg  317  VII-B-38  Knock:1.3 bar  IHRE:  18-1O-7OLBa  e.  300  ioo  2  1864 kJ/m 3  0 -  200  a)  i5 50  I’  H’  a) a) —  o o 9  20 40 Crank Angle (deg ATDC)  a)  I  0 -20  VII-B-38 18-1O-7OLBb .300  e.  40 20 Crank Angle (deg ATDC)  86  Knock:O.1 bar 3 IIPCE: 27 kJ/m  03  2  ioo 0  -,  -200  0..  a) cx  50  100  C-)  a)  V a)  ir  o  0 -20  .  0 20 40 Crank Angle (deg ATDC)  I  fl—. -i.)  0  40  20  87  Crank Angle (deg ATDC)  VII-B-38 18-1O-7OLBc .300  Knock:O.8 bar 3 IHRE: 215 kJ/m  0  E  ioo (0  0 a)  5’ 0  A  -  -200 / 50/  100  V a)  ° -20  0 20 40 Crank Angle (deg ATDC)  Z  o—-W  -20  1’ / \%  ‘  0 20 40 Crank Angle (cieg ATDC)  Figures E.4.86 to E.4.88 : Diesel flowrate: 8.7 mg/inj IHR ratio : 7.92 Knock Ratio: 5.62 Ignition Offset: 16.45 deg  318  VIFB38  Knock:0.7 bar 3 IHRPCE. 847 kJ/m  18-1O-7OLBa-14.13.43  E  ioo (1)  -,  -200 a)  50 ci)  100  -  -  I.,.  C)  9o  20 40 Crank Angle (deg ATDC)  92d  20 40 Crank Angle (deg ATDC)  V11B38  Knock:1.1 bar 347 kJ/m 3  18-1O-7OLBb-14.13.43  300 2  ioo Id)  -  -  0  200  a)  50/  a)  100  Cl)  a)  0 -20  0 20 40 Crank Angle (deg ATDC)  I  0— -20 0 20 40 Crank Angle (deg ATDC)  VII-B-38  18-1O-7OLBc-14.13.43 e.  0)  300  Knock:0.4 bar 3 IHRpCE. 159 kJ/m  2  ioo (I)  -  ci,  50  a) 100  A,  •0 a)  o  20 40 Crank Angle (deg ATDC)  20 40 Crank Angle (deg ATDC)  91  Figures E.4.89 to E.4.91 : Diesel flowrate: 9.9 mg/inj JHRratio : 0.46 Knock Ratio: 0.34 Ignition Offset: 1.32 deg  319  VIFB.-39 18-20-47LBa .300 2 ioo 0 2 aa, 50  E -200  100  0 -20  -  Z  0 20 40 Crank Angle (cieg ATDC)  0.—.. -20 20 40 Crank Angle (deg ATDC)  VII-B-39 18-20-47LBb  Knock:0.2 bar 29 kJ/m 3 .300  a 2  ioo  2  Co  -,  2  -200  /.—\\  0  ci)  1  ft  50/  100  0  a,  •0  a,  cc o 9  a,  20 40 Crank Angle (deg ATDC)  Z  -J  V11B39 18-20-47LBc  0 20 40 Crank Angle (deg ATDC)  Knock:1.7 bar 2499 kJ/m 3  a 2  .30C c’3  ioo  E  U)  -,  2  -200  a,  50  /-r  0)  E  1 I  -  ci)  o  :1  ci)  0  °  Knock:2.4 bar 3 IHRR: 2233 kJ/m  ft ci,  -  100 a,  ft  0 -20  0 20 40 Crank Angle (deg ATDC)  0  0 20 40 Crank Angle (deg ATDC)  Figures E.4.92 to E.4.94 : Diesel flowrate: 13.8 mg/inj IHR : 85.50 Knock Ratio: 8.68 Ignition Offset: 28.16 deg ratio  320  V11B39 18-20-47LBa  Knock:2.4 bar 3 IHRPCE: 1689 kJ/m  -  -200  0  /  cc  50/  a) 100  a’ I  a) a) ‘  cc —  0 -20  0 20 40 Crank Angle (deg ATDC)  I  0 -20  0 20 40 Crank Angle (deg ATDC)  VII-B-39 1 8-20-47LBb  Knock:0.2 bar 3 IHRPCE: 26 kJ/m  .3O0  ioo  E  U)  200  0  a)  cc  50  a)  100  a)  cc 00  -20  0 20 40 Crank Angle (deg ATDC)  a)  Z  -  -.)  V11B39 18-20-47LBc  0 20 40 Crank Angle (cieg ATDC)  96  Knock:0.2 bar IHR PCE : 3 kJ/m 3 JJ  ioo  E  0..  200 Cl)  50  /7  cc Cl)  o  a)  a)  cc  .2 0  0  ,  -10  0 20 40 Crank Angle (deg ATDC)  100  ti-I  )C  0 20 40 Crank Angle (deg ATDC)  97  Figures E.4.95 to E.4.97 : Diesel flowrate: 15.9 mg/inj IHRratio : 0.12 Knock Ratio: 1.07 Ignition Offset: 0.00 deg  321  VII-B40 18-20-7OLBa  Knock:5.O bar 3 PCE 1782 kJ/m 11  300 100 CO  .9  j.L.I..’’..._  17 P I 50,’  \  -200  a)  cc  /  100  >  c  .2  I  cc  c  -20  0 -20  VII-B-40 18-20-7OLBb  .  --  --  ..  .  20 40 Crank Angle (deg ATDC)  Knock:2.1 bar 300  100  F  •.._  0 20 40 Crank Angle (deg ATDC)  3 IHRPCE: 382 kJ/m  -  0  5N /1 0 \  200  9 100 :  0 -20  0 20 40 Crank Angle (deg ATDC)  I  -20  VII-B-40 18-20-7OLBc  20 0 40 Crank Angle (deg ATDC)  Knock:5.1 bar 300  3 ‘ P 1 CE 362 kJ/m  ci)  10o  -  U)  -,  -20O  50/  0  -20  0 20 40 Crank Angle (deg ATDC)  J\  100  I  0 -20  0  20  40  Crank Angle (deg ATDC)  Figures E.4.98 to E.4.100 : Diesel flowrate: 13.4 mg/inj IHR ratio : 0.95 Knock Ratio: 2.44 Ignition Offset: 0.98 deg  322  VIFB40 18-20-7OLBa  a.  Knock:5.0 bar 3 IHRPCE: 1812 kJ/m  300  a)  100  200 50  100  C)  (1)  V  Ii  .2  llI  0  —  -20  0 20 40 Crank Angle (deg ATDC)  Z  0 -20  VII-B-40 18-20-7OLBb a.  I  ;  —  300  ,  0 20 40 Crank Angle (deg ATDC)  101  Knock:3.7 bar 399 kJ/m 3 “PCE  a) U)  E  ‘-‘  200  >‘  ° .2  0  -20  0  20  40  Z  0  -20  Crank Angle (deg ATDC)  20  40  io  Crank Angle (deg ATDC)  VII-B-40 18-20-7OLBc  a.  jhii Knock:4.9 bar 3 IHRPCE: 338 kJ/m  300  a)  10o U)  -,  0% /  0 -  V  50  /  -200 ci)  -  100 A  .2 -20  0 20 40 Crank Angle (deg ATDC)  Z  0  -20  0  20  40  ioa  Crank Angle (deg ATDC)  Figures E.4.lO1 to E.4.103 : Diesel flowrate: 16.2 mg/inj 0.85 Knock Raflo: 1.32 Ignition Offset: 0.45 deg ratio  323  VII-B-4 1 24-1 O-47LBa 120  a) V  a) U)  LAJ  E 250  100  U)  1  80 a)  Knock:1.1 bar 3 IHR.E: 1956 kJ/m  -)  200 a) (U  cc  60/  C  150  I 1  a)  U)  >.  (U a)  0 V a)  a) cc  C) V C  (U a)  0  -20  0 20 40 Crank Angle (deg ATDC)  z  100 50 0 -20  “  0 20 40 Crank Angle (deg ATDC)  VII-B-41 24-1 O-47LBb U)  Knock:0.2 bar 3 IHRpCE: 29 kJ/m  C)  a) 300  120 D  104  V  E 250  100  U)  0  a)  a) V C  (U  cc  200 150  a) a)  cca)  V  a)  0 V C  0 20 40 Crank Angle (deg ATDC)  VII-B-4 1 24-1 O-47LBc (U  za)  100 50  1h Crank Angle (deg ATDC)  a) nn C)  9. 120  Knock:1.4 bar 3 IHRPCE. 162 kJ/m  a  c  E 250  10O  -)  200 a)  cc 150 a) 100 a) a) cc 50  ‘40  E20  (U 0 V C  CU a)  92  0 20 40 Crank Angle (deg ATDC)  0-  -20  I”  .J_.  I  I  ‘ 4  20 40 Crank Angle (deg ATDC)  Figures F.4.104 to P.4.106 : Diesel flowrate: 4.7 mg/inj IHRratio: 5.56 Knock Ratio: 6.53 Ignition Offset: 16.22 deg  324  VII-B-42 24-1O-7OLBa  Knock:1.4 bar 3 THRE: 2386 kJ/m 30O  120  250  100 -)  (I)  -200 a, 150  80 /7  60  a,  ,40  ioo a,  20  50  C.,  • .  0 -20  0 20 40 Crank Angle (deg ATDC)  —  -20  VII-B42 24-1O-7OLBb  20 40 Crank Angle (deg ATDC)  107  Knock:0.2 bar 3 THRPCE. 29 kJ/m ‘3O0  120  V  ci)  ioo  E 2 50  0)  0  -200  80  a,  150 a) 100 20 o .  50  0 -20  0 20 40 Crank Angle (deg ATDC)  a,  0 20 40 Crank Angle (deg ATDC)  VII-B-42 24-1O-7OLBb-14.13.OO  108  Knock:0.2 bar IHRPCE:  26 kJ/m 3  300 -o  120 a,  ioo -,  Cl)  -200 a) 150 ci) 100 a,  80 -  -  60 40  /  /\  20 0  .E  50 .---  0 -20  0 20 40 Crank Angle (deg ATDC)  . a,  0 20 40 Crank Angle (deg ATDC)  109  325  V11B42 24-10-7OLBc-14.13.00  300  120 a)  Knock:0.2 bar 3 IHRPCE: 5 kJ/m  c.3  ioo 0  -200  80  .$  20  50 —  .  -20  0 20 40 Crank Angle (deg ATDC)  -2ô  0 20 4Q Crank Angle (deg ATDC)  110  Figures F.4.108 to F.4.1 10 : Diesel flowrate: -7.4 mg/inj Sauo 0.19 Knock Ratio: 0.95 Ignition Offset: 0.00 deg T VIF.B-42  Knock:1.0 bar  24-1 0-7OLBc 1-14.13.00  3OO  12O a,  ioo  ° 2 E 5  80  -200  3 IHRPCE: 414 kJ/m  s .E  0 -20  0 20 40 Crank Angle (deg ATDC)  -20  0 20 40 Crank Angle (deg ATDC)  111  Figures F.4.109 to F.4.1 11: Diesel flowrate: 39.6 mg/inj ‘ratio 84.17 Knock Ratio: 5.60 Ignition Offset: 20.98 deg V11B43 24-20-47LBa  300  120  Knock:5.1 bar IHR PCE : 311 kJ/m 3  c 250  U)  2.200  80  V .S  .E  20 0 -20  50 0 20 40 Crank Angle (deg ATDC)  0 — -20  ‘.!iI  IL )CW  !i;4i  •!I.Lthi 0 20 40 Crank Angle (deg ATDC)  112  326  VIIB-43 24-20-47LBa2 300  €120  Knock:2.2 bar 3 IHRPCE: 160 kJ/m  40 a)  20 0 -20  -  0 20 40 Crank Angle (deg ATDC)  50 -20  0 20 40 Crank Angle (deg ATDC)  113  Knock:0.2 bar IHR: 29 kJ/m 3 •‘300  120 a)  1oo  2250  80  -200  0  60/ / 40 .?  150 \  a)  ioo  \  cc  20  50  ,  .E  -20  0 20 40 Crank Angle (cieg ATDC)  -.)  VII-B-43 24-20-47LBc  0 20 40 Crank Angle (deg ATDC)  114  Knock:3.2 bar 208 kJ/m 3 300  €120 a)  ioo  I .2  2250  IE____  20 0 -20  cc 0 20 40 Crank Angle (deg ATDC)  50  0 -20  gUIA1 0 20 40 Crank Angle (deg ATDC)  115  Figures F.4.113 to F.4.115 : Diesel flowrate: 17.6 mg/inj IHRratio: 7.13 Knock Ratio: 13.93 Ignition Offset: 14.32 deg  327  VII-B-44 24-20-7OLBa  Knock:2.8 bar 3 IHRPCE: 2438 kJ/m  300  120  c5  ioo  E 2 5  I  (3’  40  100  20  1  50 —  .  0 -20  I  0 -20  0 20 40 Crank Angle (deg ATDC)  ‘—-  VIFB44 24-20-7OLBb  0 20 40 Crank Angle (deg ATDC)  116  Knock:0.2 bar IHR: 29 kJ/m 3 •300  12O a)  1oo  ° 2 E 5  80  e20O  60/  15O  40  10O  -  20 .E  50__ 0  0  -20  0 20 40 Crank Angle (deg ATDC)  -)  VIFB-44 24-20-7OLBc 300  12O  ‘/  /  -2O0 150  40  100  20  50  -20  Knock:4.O bar 3 IIIRPCE. 738 kJ/m  E250  80  C  117  c)  ioo 60  0 20 40 Crank Angle (deg ATDC)  20 40 Crank Angle (deg ATDC)  -20  1  .  1. w  0 20 40 Crank Angle (deg ATDC)  118  328  Figures F.4.1 16 to F.4.1 18 : Diesel flowrate: 14.5 mg/inj IHRratio : 25.11 Knock Ratio: 25.17 Ignition Offset: 14.51 deg  VII-B-43 24-20-47LBb  Knock:0.2 bar  26 kJ/m 3 300  12O  V Co  ioo Cl  -,  80  -20O  6O/  150  40  1O0 a) 50  20 C)  •.5  .E  0 -20  —..  0 20 40 Crank Angle (deg ATDC)  ci)  ii--  0 20 40 Crank Angle (deg ATDC) -  119  329  Appendix F.5- Test Series VIII-A Pressure and HRR Curves  330  VIII-A-4 121-3-6 cci ci) D I.  U)  ) ci)  300  Knock:1.2 bar IHR: 1431 kJ/m 3 -  25  100 -  80  -  200  150 (‘40  100  jJ  20 -20  0 20 40 Crank Angle (deg ATDC)  -20  0 20 40 Crank Angle (deg ATDC)  1  Figure F.5.1 : Diesel flowrate: 13.8 mg/inj  VIII-A-4 121-3-11  Knock:1.4 bar  cci  -o D 0  ci  300  IBR: 1431 kJ/m 3  E 250  100  200 ci)  150 ci) Cl)  cci Cl) ci)  20  Cci ci)  -20  0 20 40 Crank Angle (deg ATDC)  I  100 50 0 -20  I  !  •1jiL  0 20 40 Crank Angle (deg ATDC)  2  Figure F.5.2 : Diesel flowrate: 19.9 mg/mi  VIII-A-5 121-3-5 300  Cci .0 ci 1... D U)  Knock:2. 1 bar IHR: 1439 kJ/m 3  250  100 -  200  ci) cci  c.40  N  ci) 0 cci ci)  100  0 -20  50 0 20 40 Crank Angle (deg ATDC)  jL  !\j ,  20  9264’2O4O Crank Angle (deg ATDC)  3  Figure F.5.3 : Diesel flowrate: 14.5 mg/inj  331  VIII-A-5 121-3-8 (U .0  a> C,)  0)  Knock:2.3 bar THR: 1430 kJ/m 3  300  -  E 250  100  200 a)  150 a, a, a)  2O  100  R  50  4r a) ‘-I  -20  0 20 40 Crank Angle (deg ATDC)  0 -20 Crank Angle (deg ATDC)  4  Figure F.5.4 Diesel flowrate: 14.9 mg/inj VIII-A-5 121-3-9 300  CU .0  a)  I  C,)  Knock:3.9 bar IFIR: 1444 kJ/m 3  250  100  200 a) a)  (I)  ca 100 a)  2O -20  50 (U a) z 0 -20  ___1  0 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  5  Figure F.5.5 : Diesel flowrate: 23.1 mg/inj VIII-A-5 121-3-12 30C  .0  Knock:3.7 bar IIJR: 1445 kJ/m 3  0  a)  250  D 100 U)  80  -  a,  200  (U  60  a)  40  a)  a,  20 0-20  50  a)  0 20 40 Crank Angle (deg ATDC)  z  h  100  .20  fi 0 20 40 Crank Angle (deg ATDC)  6  Figure F.5.6 : Diesel flowrate: 22.1 mglinj  332  VIII-A-5 121-3-17 a) a) I D  Cl)  Knock:1.4 bar fiR: 1446 kJ/m 3  ;uu  -  2 250  100  -)  80  a)  a)  200 150  C  a)  (40  a) a)  20  a)  20  z  0 20 40 Crank Angle (deg ATDC)  100 50  t  0 -20  0 20 40 Crank Angle (deg ATDC)  7  Figure F.5.7 : Diesel flowrate: 16.3 mg/inj VIII-A-5  Knock: 1.1 bar [FIR: 1455 kJ/m 3  121-3-18 300 Cl) 0)  C’  100  250  -  200  a)  ; ‘  a)  Cl)  (40  c5 100  1)  a)  20  a)  -20  0 20 40 Crank Angle (deg ATDC)  I  50  -(  0 -20  Crank Angle (deg ATDC)  8  Figure F.5.8 : Diesel flowrate: 14.2 mgfinj VIII-A-5 121-3-19 .  Cl)  Knock:1.3 bar 30O  120  IHR: 1415 kJ/m 3  -o  250  100  80  -  200  a)  60 40  .S  N  f;A  a)  K1  0)  Cu  Cl)  a)  20  a)  -20  100  50  0 —  (  0 20 40 Crank Angle (deg ATDC)  I’  /\  0 —‘-_———.-wJ -20 20 40 Crank Angle (deg ATDC)  9  Figure F.5.9 : Diesel flowrate: 18.4 mg/mi  333  VIII-A-5 121-3-20 300  .0 a) U,  250  100  IE  -  200  a)  cc CI)  100  20  20  50 0 20 40 Crank Angle (deg ATDC)  92o Crank Angle (deg ATDC)  Figure F.5.1O : Diesel flowrate: 18.9 mg/inj VIII-A-6 121-3-7  U,  Knock:1.0 bar IHR: 1449 kJ/m 3  120  300  100  250  80  -  10  Knock:2.4 bar 11IR: 1466 kJ/m 3  200  ci)  I  cc a) a) ci,  cc  20  100  I‘  50  a)  —  -20  I  0 20 40 Crank Angle (deg ATDC)  20 40 Crank Angle (deg ATDC)  Figure F.5.11 : Diesel flowrate: 11.9 mg/inj VIII-A-6 121-3-10 >30C  .0  a)  I  z  C’)  11  Knock:4.O bar IHR: 1454 kJ/m 3  V C’  100  E  -  200  ci) Cu  cc  a)  1M3  i  a)  20  0 -20  250  cc 0 20 40 Crank Angle (deg ATDC)  50  ‘Ii I  p  20 40 Crank Angle (deg ATDC)  12  Figure F.5.12 : Diesel flowrate: 21.2 mg/inj  334  VIIFA7 121-3-14 2120  300 a)  100 D (0  250 E  80  200  CO  Knock:3.0 bar IHR: 1417 kJ/m 3  I’  I  60  150  40  10o  I’  a)  20  50  C) —  (,  a)  -20  I  0 20 40 Crank Angle (deg ATDC)  /  0 —-—X t -20 0 20 40 Crank Angle (deg ATDC)  13  Figure F.5.13 : Diesel flowrate: 16.8 mg/inj VIIFA8 121-3-13 2120  c300 a)  100  250 E  80  200  CO CO  a)  150  60 C  ci)  (‘40  100 a)  20  50  C) : —  Knock:0.9 bar IHR: 1412 kJ/m 3  a) -20  0 20 40 Crank Angle (deg ATDC)  0 -20  1  / \ /  ‘  0 20 40 Crank Angle (deg ATDC)  14  Figure F.5.14 : Diesel flowrate: 15.3 mglmj VIIFA8 121-3-16 300  2120 a) 100  Knock:0.9 bar [FIR: 1408 kJ/m 3  (1)  250 E  U)  c)  -  200  80  ci)  150  60 C  a)  40  100  20  50  0  920  -  Crank Angle (deg ATDC)  0 20 40 Crank Angle (deg ATDC)  15  Figure F.5.15 : Diesel flowrate: 13.9 mgfinj  335  VIII-A-9 121-3-15 (U  -o  Knock: 1.0 bar IHR: 1445 kJ/m 3 a,  -  a)  5UU  E 250  D 100  -  200 ci  150 a, 40  1)  a)  20 (U 0  —  -20  0 20 40 Crank Angle (cleg ATDC)  100  5__,1  (U ci :i:  0 20 40 Crank Angle (deg ATDC)  16  Figure F.5. 16 : Diesel flowrate: 11.3 mg/inj VIII-A-28 121-3-3 a, 300 0)  CU  .0  a, 0)  Knock:1.1 bar IHR: 696 kJ/m 3  •0  E 250  100  a)  200 150  a) a,  a,  100  50  a,  —  -20  0 20 40 Crank Angle (deg ATDC)  z  920i7 Crank Angle (deg ATDC)  Figure F.5.17 : Diesel flowrate: 12.4 mg/inj  VIII-A-29 121-3-1 0) a, 0  CU  .0  (I)  30G  Knock:0.9 bar IHR: 713 kJ/m 3  S 250  100  a, (U  200 150  a,  c’40  a, a,  20 0-20  0 20 40 Crank Angle (deg ATDC)  100  50  a, I  40 Crank Angle (deg ATDC)  18  Figure F.5.18 : Diesel flowrate: 13.9 mg/mi  336  VIII-A-29 121-3-4 a) 300  ( CD U)  Knock:0.9 bar ll{R: 704 kJ/m 3  E 250  100  200 a)  cc 150 a)  100  cc  50  a, a,  20 0 -20  U) Cu  1)  0 20 40 Crank Angle (deg ATDC)  z  0 -20  1’  I’  M 2 _J —  0 20 40 Crank Angle (deg ATDC)  19  Figure F.5.19 : Diesel flowrate: 16.6 mg/inj V111A30 121-3-2 .  a) U) U)  a) 300  Knock:0.6 bar ll{R: 687 kJ/m 3  0)  120  0  E 250  100  200  80 CD  cc 150  60  a)  40  a) a)  cc  20 ••  0 -20  CD  0 20 40 Crank Angle (cieg ATDC)  :i:  100 50 0-20  0 20 40 Crank Angle (deg ATDC)  20  Figure F.5.20 : Diesel flowrate: 15.2 mg/inj  337  Appendix F.6- Test Series VIII-B Pressure and HRR Curves  338  VIII-B-10  Knock:2.7  bar  E-1B  iliR: 1018  3 kJ/m  ci)  a?z  U) U,  a?  -  0  ci) V C  ci)  /\  200  50  4  j’  V a) 0 V C  0  0  -20  0  20  —-.  -20  40  Crank Angle (deg ATDC) Figure  F.6.1  : Diesel flowrate:  0  20  40  Crank Angle (deg ATDC)  1  13.4 mg/inj  VIII-B- 10 E-1D ‘ V  -  a? U) U)  d.I  100  Knock:2.4  bar  IITR: 1055  3 kJ/m  30C  -  a?  -200  I  0  a)  V C  50’ 100  C) V C  0 -20  0  20  Figure  F.6.2  : Diesel flowrate:  Crank Angle (deg  12.7  2  ATDC)  mg/inj  VIII-B-1 1 E-1A  Knock:1.6 bar IHR: 1046 kJ/m 3 ‘ V  a?  30C  —  E  U) U)  -,  a?  -  0  ci) ix ci)  50/  >‘  0  ci)  V ci) 0 -D C  j  0_  40  Crank Angle (deg ATDC)  ci) V C  k4  ,  t’  js  V ci)  ‘__%__  .5  ix  o  -20  0  20  Crank Angle (deg Figure  F.6.3  40  14.1  100  1  j\iç. 20  ATDC)  : Diesel flowrate:  200  Crank Angle (deg  40 ATDC)  3  mg/inj  339  VIII-B-1 1 E-1E .0  Knock:2.O bar IHR: 1065 kJ/m 3 a)  a)  z  U)  (I)  E  100  -,  -  0 I-  a) V C  50  z  a)  0  -20  .0  C0 a)  A  I 92o  20 40 Crank Angle (deg ATDC)  4  Knock:1.2 bar LFIR: 948 kJ/m 3 a) V  orr  -  a)  50  a)  —  >  C) V a) C  ‘  2  C’)  C) V  100  a)  0 20 40 Crank Angle (deg ATDC)  Figure F.6.4 : Diesel flowrate: 13.0 mg/inj VIII-B- 12 E-1C  a) V C  200  a)  V a) C) V C  a)  a) a)  0 -20  .0  100  -20  /  A I  U—  0 20 40 Crank Angle (deg ATDC)  Figure F.6.5 : Diesel flowrate: 12.6 mg/inj VIII-B- 12 E-1F  200  \  j  0 20 40 Crank Angle (deg ATDC)  5  Knock:1.5 bar IHR: 959 kJ/m 3 a) V  2  Co  (0 -  0 ci V C  a)  50  /  200  100  V a) C) V C  1  -20  0 20 40 Crank Angle (deg ATDC)  n -20  /i’  I .4  kf  0’ 20 40 Crank Angle (deg ATDC)  6  Figure F.6.6 : Diesel flowrate: 14.9 mg/inj  340  V111B4 D- 1 D  e.  Knock:3.4 bar IHR: 2204 kJ/m 3 30O  /  100  I  / .  -,  -200  \  50  100  o  ‘%%  0 -20 —  e.  Iii.. ‘1U  C -20  20 40 0 Crank Angle (deg ATDC)  Figure F.6.7 Diesel flowrate: 14.4 mg/inj VIII-B-5 D-1E  ‘  40 0 20 Crank Angle (deg ATDC)  Knock:4.8 bar IHR: 2239 kJ/m 3 -30O  100  50  /  200  N a)  —  -20 —  20 40 0 Crank Angle (deg ATDC)  I  ‘1 -20  Figure F.6.8 : Diesel flowrate: 15.5 mg/inj V111B6 D-1F  20 40 0 Crank Angle (deg ATDC)  8  Knock:5.7 bar IHR: 2302 kJ/m 3 30O  100  50”  .s —  C -20  N  IJ  I  —  ji  t144  c’  0 20 40 Crank Angle (deg ATDC)  H  200  -20  (I  fi  20 40 Crank Angle (deg ATDC)  Figure F.6.9 : Diesel flowrate: 15.0 mg/inj  341  VIIFB43  Knock:3.1  bar  3 IHR:1466kJ/m  13-15-05-55-13  €300 c.3  E  /  100  200 .  50  -  /  I  100 ,  .2  0 -20  20 40 0 Crank Angle (deg ATDC)  —  z  0.— -20  20 0 40 Crank Angle (deg ATDC)  10  Figure F.6.10 : Diesel flowrate: 17.0 mg/inj VIII-B-13  Knock:2.8  bar  1}IR: 2143 kJ/m 3  D- lB a)  0  a? co Co  -,  /  iuv  E  /  a?  -,  -200 a)  100 a)  C.)  E .2  C -20  —  Figure  x  F.6. 11: Diesel flowrate: 11.4 mg/inj VIII-B-14 11-15-10-55-13  20 40 0 Crank Angle (deg ATDC)  Knock:3.4 bar IITR:1461k3/m 3  €300  a?  E  ci)  200  / I-  .  0 -20  —-  20 0 40 Crank Angle (deg ATDC)  I  /  50 /  a)  C.)  100  a)  1920 —  Figure  20 40 0 Crank Angle (deg ATDC)  F.6.12  : Diesel flowrate:  z  40 20020 Crank Angle (deg ATDC)  12  17.2 mg/inj  342  VIII-B- 14 D-1A  Knock:2.9 bar ll{R: 2235 kJ/m 3  .0 D Cl)  (I,  300 E  100  -,  - 200 a)  0 ci>  50  100  ,  ‘i  i•I  fl  I f  a) 0 V  0 -20  20 40 0 Crank Angle (deg ATDC)  0  Figure F.6.13 : Diesel flowrate: 18.4 mg/mi VIII-B- 15 12-15-05-55-13 300 E  100  -  0 a) V C  a)  50  /  a)  a)  a)  20 40 0 Crank Angle (deg ATDC)  -20  I  Figure F.6.14 : Diese’ flowrate: 15.5 mg/inj VIII-B- 15 D-1C .0  Cl) Cl)  100  !4  I Iy  j /_  I-  a)  50  (I  .———-  -20  Th 20 40 Crank Angle (deg ATDC)  14  Knock:2.5 bar 3 1HR: 2341 kJ/m ci)  /  __‘\  -  - 200 a)  N  i’% -  V a) cci 0 V C  I  (  E  100  0 V C  200  a) a)  -v 0 V C  13  Knock:2.5 bar IHR: 1467 kJ/m 3  .0  Cl) Cl)  0 20 40 Crank Angle (deg ATDC)  -20  0 20 40 Crank Angle (deg ATDC)  0 -20  /\J  \\  ..  20 40 Crank Angle (deg ATDC)  15  Figure F.6.15 : Diesel flowrate: 11.8 mg/inj  343  VIIIB-16 06-08-05-55-13-1500  \\\%:-  Cl,  100  a,  D  7/ a,  a,  •  50’  /  Knock:4.5 bar IEIR: 2132 3 kJ/m 300 200  /  Co  100  a,  %_  Ci)  Ci)  /  —  00  -20  a,  0 20 40 Crank Angle (deg ATDC)  Figure F.6.16 : Diesel flowrate: 16.3 mg/inj V111B46 10-15-05-55-13-1500  9204O Crank Angle (deg ATDC)  oo  16  Knock:5.4 bar IHR: 2143 3 kJ/m  V  (0  100  /7/  E —  200  0  a,  50  10O  0 a,  e.  1 j  a,  -  .2  11 I  1)  I  0 -2  0 20 40 Crank Angle (deg ATDC)  z  0 — -20  Figure F.6.17 : Diesel flowrate: 14.3 mg/mi VIIFB46 B-lB ,  N  100 /V}  a-  17  Knock:5.0 bar 3 IHR:2166kJ/ m  300 a, V  h 11  -)  a,  200  I’  cr  50  20 40 Crank Angle (deg ATDC)  I  Co •  V  a,  a,  .20 -20  0 20 40 Crank Angle (deg ATDC)  a,  z  -20  20 40 Crank Angle (deg ATDC)  18  Figure F.6.18 : Diesel flowrate: 15.2 mg/inj  344  VIIFB16 C-lB  Knock:4.5 bar fiR: 2091 kJ/m 3  e. [oo//\  300  -200  /  ci)  50/  0 -20  z  0 20 40 Crank Angle (deg ATDC)  —  0  -20  0  20  40  Crank Angle (deg ATDC)  Figure F.6.19 : Diesel flowrate: 18.2 mg/inj V111B16 B-iD  19  Knock:6.7 bar 3 IHR:2193kJ/ m -8300  ,  2  -200  5O/ 100  0 -20  0  Crank  —  20 40 Angle (deg ATDC)  z  0 -20  Figure F.6.20 : Diesel flowrate: 15.3 mg/inj V111B16 C-iD  e. a, ioo  1 I I  /  20 40 Crank Angle (deg ATDC)  Knock:6.1 bar 3 IIIR:2137kJ/ m -8300  \  (I  c’—  I -  I •00  -20 —  0 20 40 Crank Angle (deg ATDC)  20  x  c—i -20  I L  0 20 40 Crank Angle (deg ATDC)  21  Figure F.6.21 : Diesel flowrate: 17.8 mg/inj  345  V111B17 08-15-10-55-13-1500 300 a)  Knock:1.9 bar fiR: 2171 3 kJ/m  E  ioo  Co  -,  / / /  200  r\  50 o  a)  Co100  •0 4  .2  v  -20  a)  0 20 40 Crank Angle (deg ATDC)  92’ 20 40 Crank Angle (deg ATDC)  Figure F.6.22 : Diesel flowrate: 14.0 mg/inj V111B47  Knock:2.5  B-i A  22  bar  fIR: 2220 kJ/m 3 300  Ci)  E  100  -  -200 a)  I’  50 .5’ 0  a)  100  .20  -20  ca a)  0 20 40 Crank Angle (deg ATDC)  Figure F.6.23 : Diesel flowrate: 14.8 V111B17  I  I  z  Crank Angle (deg ATDC)  mg/inj Knock:1.9 IHR: 2203  C-lA  °30o Cl)  bar 3 Id/rn  E  100  200 a)  / .5’ 0  1  100 a)  a)  .2  23  0 -20  I  \  —  ca  0  20  Crank Angle (deg ATDC)  a)  20  0 20 40 Crank Angle (deg ATDC)  24  Figure F.6.24 : Diesel flowrate: 17.9 mg/inj  346  VIII-B- 17 B-1E  I U) Co  100  ___/‘  ‘c  Knock:2. 1 bar IHR: 2202 kJ/m 3  a) V  2 200  I’  a, C  50  io  (3’ V a, () C  0 -20  0 20 40 Crank Angle (deg ATDC)  92’’’ 2040 Crank Angle (deg ATDC)  Figure F.6.25 : Diesel flowrate: 16.5 mg/inj VIII-B- 17 C-1D2 C,) (1)  a I-.  a) V C  >.  I \ 25  Knock:3.8 bar JHR: 2177 kJ/m 3  a,  -D  \ ::,‘  2 -200 I,  100  0  /  ‘\  V a, C.)  0 20 40 Crank Angle (deg ATDC)  —  Crank Angle (deg ATDC)  Figure F.6.26 : Diesel flowrate: 16.3 mglinj VIll-B- 17 C-1E 0) V D (0 (I,  I-  50/”  N  -  1  0  S.—  V a)  a)  cc 0 -20  300  200  100  >.  C) V C  Knock:2.0 bar IHR: 2201 kJ/m 3  E  100  a a) V C  26  0 20 40 Crank Angle (deg ATDC)  a)  / 20 40 Crank Angle (deg ATDC)  27  Figure F.6.27 Diesel flowrate: 16.5 mg/inj  347  VIIIB-18 09-15-15-55-13-1500  Knock:2.7 bar 3 1TIR:2250kJ /m 30C  1)  100  50/  / ioo  o  I vti.  /A  / -20  0 20 40 Crank Angle (deg ATDC)  —  x  -20  Figure F.6.28 Diesel flowrate: 14.0 mglinj V111B18 B-iC  ç  0 20 40 Crank Angle (deg ATDC)  28  Knock:l.4 bar IHR:23l4kJ/m 3 300 E  100  50  100  I V  -20  0 20 40 Crank Angle (deg ATDC)  —  z  0 -20  Figure F.6.29 Diesel flowrate: 15.8 mg/inj VIIFB48 C-iC  0 20 40 Crank Angle (deg ATDC)  29  Knock:1.8 bar [FIR: 2288 kJ/m 3 300  100  /  7 /  -200  ‘N.  5..  o  100  I ,  Cl)  ( -20 —  0 20 40 Crank Angle (deg ATDC)  —)(  0 -20  0 20 40 Crank Angle (deg ATDC)  30  Figure F.6.30 : Diesel flowrate: 17.5 mglinj  348  VIII-B-18 B-iF  e. ioo  300 a)  Knock:1.2 bar 3 1HR:2256kJ /m  2  :  -,  a)  f  a)  ‘N  200  A  100 a)  %...  It  a) /  C)  -20  e.  0 20 40 Crank Angle (deg ATDC)  9200 20 40 Crank Angle (deg ATDC)  Figure F.6.31 Diesel flowrate: 14.2 mg/inj V111B18 C-iF 300 a)  31  Knock:1.4 bar 3 IIIR:2181kJ/ m  c)  2  ioo U)  - 200  A  a)  a) 50’  a)  ioo  /  I  a)  92 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  Figure F.6.32 : Diesel flowrate: 15.5 mg/inj V111B22 02-08-05-55-13  32  Knock:2.2 bar 11IR: 2160 kJ/m 3 V  /7  100 U)  \\  E _,  0  200  4.’  a)  50  /  /  a)  100  C.)  —  V  a)  a)  0 —  0 20 40 Crank Angle (deg ATDC)  0 -20  0 20 40 Crank Angle (deg ATDC)  33  Figure F.6.33 : Diesel flowrate: 6.3 mglinj  349  V111B22 A- lB  €  Knock:3.7 bar IHR: 2215 kJ/m 3 300  a)  I  i  E  I:____ 0 -20 —  0 20 40 Crank Angle (deg ATDC)  .  z  0 -20  Figure F.6.34 : Diesel flowrate: 6.9 mg/inj V111B22 A-iD 300 a)  --  20 40 Crank Angle (deg ATDC)  Knock:2.9 bar 3 IIIR:2172kJ/ m  200  100  -20 —  0 20 40 Crank Angle (deg ATDC)  z  c  -20  Figure F.6.35 Diesel flowrate: 8.9 mg/inj V111B23  Q)  Knock:2.9 bar IHR: 2267 kJ/m 3  01-08-10-55-13  e.  20 40 Crank Angle (deg ATDC)  300 2  IUV  200 .9  50  CD —  0 -20 —  0 20 40 Crank Angle (deg ATDC)  Figure F.6.36 Diesel  100  .  z  C -20  w 1 ‘ 0 20 40 Crank Angle (deg ATDC)  36  flowrate: 8.6 mglirij  350  V111B23  IHR:  300  I  J  ioo  C’,  bar 2242 kJ/m 3  Knock:2.3  A-i A  A  11  -,  -200  —  5O/ 100  ‘S.  o  ci)  ci)  .2  I  -  0 -20  —  .-‘  —  0 20 40 Crank Angle (deg ATDC)  z  0 —— -20  0 20 40 Crank Angle (deg ATDC)  Figure F.6.37 : Diesel flowrate: 7.4 mg/inj V111B23  Knock:1.7  LHR:  A-1E  bar  2243 kJ/m 3  CL) -D  —  200  i\ I  50’ 100  0 -20 —  Figure  0 20 40 Crank Angle (deg ATDC)  F.6.38  : Diesel flowrate:  •  z  0— -20  \  0 20 40 Crank Angle (deg ATDC)  38  9.6 mglinj  VIIIB24 03-08-15-55-13  Knock:2.4 bar IHR: 2442 kJ/m 3 300  0  100  —  -  -S  -,  -200 “ 50 100 a:  0 -20 —  Figure  0 20 40 Crank Angle (deg ATDC)  F.6.39  : Diesel flowrate:  0—•---20  *0  ‘  20 40 Crank Angle (deg ATDC)  7.0 mg/mi  351  V111B24  Knock:1.2 bar IHR: 2401 3 kJ/m  A-i C  .0  .  (I)  300  2  100  S.-  a-  -  a)  N  50  200  \  /  ci) —  a)  I  00  -20  0  20  0  ci)  40  -—-  -20  z  Crank Angle (deg ATDC)  20  40  Crank Angle (deg ATDC)  40  Figure F.6.40 : Diesel flowrate: 11.0 mg/inj  V111B24 A-iF  e.  Knock:1.5 bar IHR: 2359 3 kJ/m c3  (I,  2  100  —S  a-  -200  50  a)  ci)  I’  I’ I’  V ci) —  .2  Cu  -2” —  Cl)  0 20 40 Crank Angle (deg ATDC)  z  92o 40 Crank Angle (deg ATDC)  41  Figure F.6.41 : Diesel flowrate: 10.7 mg/mi  VIWB-25  Knock:1.6 bar IHR: 2232 3 Id/rn  05-08-10-55-13-1500 300 0)  E  100  -,  -200  a-  a)  i5  Cu  \\N  100  V  .2 V c  0 -2” ‘  —  (  /  \  —  Cl)  0 20 40 Crank Angle (cleg ATDC)  Figure F.6.42 : Diesel flowrate:  13.1  z  c -  20 40 Crank Angle (deg ATDC)  42  mg/inj  352  VIIIB25 B-1A2  (U  8 Co  Knock: 1.9 bar ll-IR:221OkJ/m 3  , 0  30C E  iOO I\  r.  —200  /  50  100  >•  (1)  /  a)  o  -20  z  0 20 40 Crank Angle (deg ATDC)  Figure F.6.43 : Diesel flowrate: 13.7 mglrnj VIII-B26 O7-O8-15-5513-15OO D 0 Co  //%%\  iu0  E 0)  a 50’  20 40 Crank Angle (deg ATDC)  Knock:1.4 bar ifiR: 2325 kJ/m 3  0300  8  —II  C i—-.— -20  \  200  0  100  /  a) —  -20  0 20 40 Crank Angle (deg ATDC)  0 -20  \.  IC’”  0 20 40 Crank Angle (cieg ATDC)  Figure F.6.44 : Diesel flowrate: 11.1 mg/inj LOW CO SLOW TEST  8  300  (0  -,  I—  0)  -o .  Knock:2.9 bar IHR: 2394 kJ/m 3  ,  50  -  \  a)  200  ioo O  0 a)  2  0 -20  0 20 40 Crank Angle (deg ATDC)  0 -20  1W  ----  20 40 Crank Angle (deg ATDC)  Figure F.6.45 : Diesel flowrate: 11.4 mglinj  353  VIII-B-35  Knock:1.6 bar IHR: 1487 3 kJ/m  REP-08-02-26  e.  30O E  coO 0  200 ci)  0  a)  7 50’/  ci  I’ 00  ci) °  0 -20  —  0 20 40 Crank Angle (deg ATDC)  -20  0 20 40 Crank Angle (deg ATDC)  46  Figure F.6.46 : Diesel flowrate: 12.4 mg/inj V111B36  Knock:1.7 bar IHR: 1527 3 kJ/m  REP-08-02-28 ci) D30C  I  E  coO 10 0  _,  a)  ci) 50/  u  a) 0100  \N  -  o  200  a)  ci)  0 -20 Figure  —  ca  0 20 40 Crank Angle (deg ATDC)  F.6.47  : Diesel flowrate:  13.3  0 -20  0 20 40 Crank Angle (deg ATDC)  mg/inj  354  Appendix F. 7- Test Series VIII-B2 Pressure and HRR Curves  355  VIII-B2-1 Vill-1A  1’ 8  Knock:1.8 bar IHR: 1090 kJ/m 3 30C  100 -  200  5O7  .2  C -20  —  8 0  [\ —  0 20 40 Crank Angie (deg ATDC)  z  0 -20  Figure F.7.1 : Diesel flowrate: 14.2 mg/mi VIII-B2-1 VIII-1  0 20 40 Crank Angle (deg ATDC)  Knock:1.9 bar 11IR: 1084 kJ/m 3 300 (.5— E  IViJ  0  :  -,  -200  50  C  -20 —  8  0 20 40 Crank Angle (deg ATDC)  z  100  0  -—--  -20  Figure F.7.2 : Diesel flowrate: 14.4 mg/inj VIII-B2-1 Vill-1B  0 20 40 Crank Angle (deg ATDC)  2  Knock:1.6 bar IHR: 1113 kJ/m 3 300 E 2OO  U.  -20 —  0 20 40 Crank Angle (deg ATDC)  z  0 -20  *..  0 20 40 Crank Angle (cieg ATDC)  Figure F.7.3 : Diesel flowrate: 15.3 mg/inj  356  V111B21 Vill-1D  e. (0 (I)  Knock:1.8 bar 3 IIIR:1O83kJ /m 300 E  IVi)  -20O .50 •5.  / /•_  o  100  -  .  .s  0 -20  —  i.  —  0 20 40 Crank Angle (deg ATDC)  z  -20  Figure F.7.4 : Diesel flowrate: 14.2 mg/inj V111B2 1 VIII-1D2 € (1)  0 20 40 Crank Angle (cieg ATDC)  Knock: 1.9 bar IHR:1237kJ/m 3 300 E  100  -  /‘  \  -200 a,  50  0 -20 —  0 20 40 Crank Angle (deg ATDC)  x  Figure F.7.5 : Diesel flowrate: 14.6 mg/inj V111B22 VIII-2A €  0 -20 0 20 40 Crank Angle (deg ATDC)  Knock:2.O bar 1}IR: 1115 kJ/m 3  ,,  300  100  (0  -200 50  -  a  -20 —  0 20 40 Crank Angle (deg ATDC)  z  0 -20  4C-  20 40 Crank Angle (deg ATDC)  6  Figure F.7.6 : Diesel flowrate: 14.1 mg/inj  357  V111B22 VIII-2  Knock: 1.7 bar IIIR: 1040 kJ/m 3 QVV  c3 U) U)  E  IUJ  -  -200 a)  50  .2  100  0 -20  0 20 40 Crank Angle (deg ATDC)  —  e..  i  0 -20  Figure F.7.7 : Diesel flowrate: 14.3 mglinj V111B22 VIII-2B  0 20 40 Crank Angle (deg ATDC)  Knock: 1.7 bar IHR: 1096 kJ/m 3 3OO  0) U) Cl)  E  IUV  -,  -200  1  50 ,•  0 -20  e. C) U)  .  0 20 40 Crank Angle (deg ATDC)  z  -20  Figure F.7.8 : Diesel flowrate: 14.6 mg/mi V111B23 VIII-3A  20 40 Crank Angle (deg ATDC)  8  Knock:2.5 bar IHR:1036kJ/m 3 300 E  IJJ  -,  -200 .E > C) —  50  CI)  t  100  j  CI)  o  -20 —  7  Ij 1 N 1  111  0 20 40 Crank Angle (deg ATDC)  0 -20  —————-  z  .  1  0 20 40 Crank Angle (deg ATDC)  Figure F.7.9 : Diesel flowrate: 14.5 mg/inj  358  VIII-B2-3 VIII-3  Ct  a, V  2  Co 100  -,  Co  -  a,  0 1) V  Knock:1 .4 bar IHR: 1036 kJ/m 3  200  50  1  V  i  a,  C) V  Ct  Lt  -20  20 40 0 Crank Angle (deg ATDC)  ——--  Crank Angle (deg ATDC)  Figure F.7.10 : Diesel flowrate: 14.0 mg/inj VIII-B2-3 VIII-3B 0) a)  0 Cl) (0  a,  nn 200  Ct  a,  C  Knock: 1.7 bar IHR: 1072 kJ/m 3  2  100  0 V  50  4  [00  >.  C) V  a,  .  C)  V  C  -20  20 0 40 Crank Angle (deg ATDC)  -20  Figure F.7.1 1: Diesel flowrate: 15.3 mglinj VIII-B2-3 VIII—3D 0) a,  V (0 (I)  a,  I-  Knock:2.4 bar IHR: 1166 kJ/m 3 ,,,‘,  200  50  /7  100  1i øL TIfr A 1  a, Ct C)  0 C  11  Ct  a,  C  a  ‘4•Hi ‘IM AALA 0 40 20 Crank Angle (deg ATDC)  E  100  0  -o  10  Cl  0  20 0 40 Crank Angle (deg ATDC)  .___.__  -20  tti1Ji Th  20 40 Crank Angle (deg ATDC)  12  Figure F.7.12 : Diesel flowrate: 14.8 mg/inj  359  V111B24 VIII-4A  Knock:1.8 bar IHR: 2192 kJ/m 3 300  \  ioo C’)  -  -)  -200  j / 50  100  C,  0  -20  0 20 40 Crank Angle (deg ATDC)  •  z  0 20 40 Crank Angle (deg ATDC)  Figure F.7.13 Diesel flowrate: 15.1 mglrnj V111B24 VIII-4  a ioo  I  -  -200  “N.  50/  Knock:2.0 bar IHR: 2193 kJ/m 3 .300  /\  C)  13  % .  a) .  —  0 -20  0 20 40 Crank Angle (deg ATDC)  z  0 -20  Figure F.7.14 : Diesel flowrate: 16.0 mglinj V111B24 VIll-4B  0 20 40 Crank Angle (deg ATDC)  14  Knock:2. 1 bar fiR: 2265 kJ/m 3  10O -200 a)  / 50/  100  >. C)  ,-,  V  U)  -  0 -20  0 20 40 Crank Angle (deg ATDC)  o-20  0 20 40 Crank Angle (deg ATDC)  15  Figure F.7.15 : Diesel flowrate: 15.8 mg/inj  360  VIILB24 VIII-4D S S 10o  /  Knock:1.9 bar IHR: 2148 kJ/m 3 300  \  S  -200 50  100  / A  .  0 -20  r  0 20 40 Crank Angle (deg ATDC)  —  0 -20  Figure F.7.16 : Diesel flowrate: 15.2 mg/inj VIll-B2-5 VIII-5A S  16  Knock:2.8 bar IHR: 2135 kJ/m 3  Co  ioo  .n—,/  \ 200  1’  I  100  0 -20  0 —....iIiHThJ4 -20 20 40 Crank Angle (deg ATDC)  0 40 20 Crank Angle (deg ATDC)  —  I  0 20 40 Crank Angle (deg ATDC)  300  S  .  A  Figure F.7.17 : Diesel flowrate: 14.9 mg/inj VIIhB25 VIII-5  € S 0100  17  Knock:2.4 bar iliR: 2118 kJ/m 3 300 E  N N ;  200  50/  we  o  .2 —  &U  U  -20  0 20 40 Crank Angle (deg ATDC)  r  0 -20  ‘W  20 0 40 Crank Angle (deg ATDC)  18  Figure F.7.18 : Diesel flowrate: 15.7 mglinj  361  VIILB25  Knock:2.3 bar IHR: 2155 3 kJ/m  VIII-5B  a  300 100 200 50  100  0 -20  ‘  0 20 40 Crank Angle (deg ATDC)  —  Figure F.7.19  Diesel flowrate: 16.1  x  0 -20  19  mg/inj  V1WB25  I  20 40 Crank Angle (deg ATDC)  Knock:2.4 bar 3 IHR:2119kJ/ m  VIII-5D  300  ci>  100 a.  S  -200  /  I-  z  ci> c  cc  -  50/  ‘N.  .ro0 0 -20 —  0 20 40 Crank Angle (deg ATDC)  Figure F.7.20  I a  0 ‘-j -20  I  Diesel flowrate: 15.2 mg/mi VIILB26 VIII-6A  tk.Q 0 20 40 Crank Angle (deg ATDC)  20  Knock:4.5 bar IHR: 1975 kJ/m 3 300  200  :  100 C)  frI .2 ‘  —  0 -20  0 20 40 Crank Angle (deg ATDC)  IJf  iI.JH ill  flli’:%i&J 0 ‘*.._t!h1 -20 20 40 Crank Angle (deg ATDC)  21  Figure F.7.21 Diesel flowrate: 15.2 mg/inj  362  VIII-B26 VIII-6  Knock:2.5 bar IHR: 2038 kJ/m 3 300  a,  E ‘  a,  N  50  °  >  200  t  100  1 I  a)  a:  0  -20  0 20 40 Crank Angle (deg ATDC)  —  z  i  0 —........g  -20  Figure F.7.22 : Diesel flowrate: 15.7 mg/inj V111B26 VIII-6B  v’i  0 20 40 Crank Angle (deg ATDC)  22  Knock:3.1 bar IHR: 2048 kJ/m 3 300  (I)  U)  E  IVU  -,  -200 50  //  u -20 —  I  100  0 -20  0 20 40 Crank Angle (deg ATDC)  Figure F.7.23 : Diesel flowrate: 15.5 mglrnj VIIIB2-6 VllI-6D  0 20 40 Crank Angle (deg ATDC)  23  Knock:2.7 bar IHR: 2038 kJ/m 3 300 2  -  0  200  i5 50/  r100  C.)  ‘%.  .  a: 0 -20 —  0 20 40 Crank Angle (deg ATDC)  0 -20  20 40 Crank Angle (deg ATDC)  24  Figure F.7.24 : Diesel flowrate: 14.9 mg/inj  363  VIII-B2-7 VIII-7A  Cu  e.  300  D..  (0  Knock:2.3 bar IHR: 2237 kJ/m 3  100  E  •1  /d  0  A  -)  200  0  50/  /  cr a)  0100 cu  0  I I  -  a, —  9  02 -  20 40 Crank Angle (deg ATDC)  Figure F.7.25 : Diesel flowrate: 14.5 mglrnj VIllB27 , VIII-7 (I)  100  0  a) 50/  /  Knock:2.5 bar lHR: 2210 kJ/m 3  N  E -  a,  ‘N  200  100 a,  11 I’  I  I I  a)  0 -20  20 40 Crank Angle (deg ATDC)  —  Cl) Cl)  25  300 a)  %%  Cu  0 20 40 Crank Angle (deg ATDC)  Figure F.7.26 : Diesel flowrate: 14.5 mg/inj VIII-B27 VIII-7B 00  0  /1  20 40 Crank Angle (deg ATDC)  26  Knock:2.2 bar IHR: 2227 kJ/m 3 300  \\  2 -200  I’  a,  50 %%  a, 0100 Cu a,  i  \  a) -  —  0  20 40 Angle Crank (deg ATDC)  0 -20  40 20 Crank Angle (deg ATDC)  27  Figure F.7.27 : Diesel flowrate: 14.5 mg/inj  364  V111B28 VIII-8A  8 U) CI)  300 a)  Knock:1.3 bar IHR:2211kJ/m 3  E  IL)iJ  -,  -  a  1)  200  50/ 100 o  a)  ‘_‘%_  .5  .  a)  I  9o  0 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC) Figure F.7.28 : Diesel flowrate: 14.7 V111B28  8 0  28  mglinj  VIII-8  o,  \  (1) 300  Knock: 1.4 bar JHR: 2177 3 kJ/m  iOO -,  a 50  /  (1)  N  200  a)  °)100 I  a)  20 40 Crank Angle (deg ATDC)  —  Figure  F.7.29  : Diesel flowrate:  14.7  -20  29  mg/inj  V111B28 VIII-8B  (  20 0 40 Crank Angle (deg ATDC)  Knock:1.3 bar IHR: 2243 3 kJ/m  8  300 c5  0  100  a-  /  50  ,f\  -,  -200 a)  7/ a)  j  100 I  2  0 -20  0 20 40 Crank Angle (deg ATDC)  0 X -20  “  0 20 40 Crank Angle (deg ATDC)  30  Figure F.7.30 : Diesel flowrate: 14.5 mg/inj  365  VIII B29 VIII-9A  Knock:1.7 bar 11IR: 1987 kJ/m 3  a)  200 50/ %%  A  (Hl  I k  11  -  C -20  •  0 20 40 Crank Angle (deg ATDC)  —  j  0 ‘* -20 0 20 40 Crank Angle (deg ATDC)  Figure F.7.3 1: Diesel flowrate: 14.0 mg/inj V111B29  Knock:2.2 bar fiR: 1879 3 kJ/m  VIII-9  a  31  300  100  -  - 200 a) 50/ ,  kH  -20  0 20 40 Crank Angle (deg ATDC)  —  Figure F.7.32 : Diesel flowrate: V111B29  x  0 -20  !!l  20 40 Crank Angle (deg ATDC)  32  15.2 mg/mi  VIII-9B  100  Knock:1.9 bar 1946 kJ/m 3  IHR:  a) -o -  200 -  .  /  50/  (1) -  -20 —  0 20 40 Crank Angle (deg ATDC)  Figure F.7.33 : Diesel  z  100  0 -20  1 20 40 Crank Angle (deg ATDC)  flowrate: 14.8 mg/inj  366  VIII-B2- 10  Knock:1.6  VIII-1OA  U) U)  E  100  -  0  a)  200  I-  k , )\  11) •0 C  a)  >  C) a)  C  100  a) a)  •0  C) -D  -20  ID  40 20 0 Crank Angle (deg ATDC)  Figure F.7.34 : Diesel flowrate:  I  -  X’-  20 0 40 Crank Angle (deg ATDC)  34  15.4 mg/mi  VIII-B2- 10  Knock: 1.4  VIlI-lo 0) 0)  bar  IHR: 1100 kJ/m 3  0) a) •0  IHR: ci)  bar  1082 kJ/m 3  300  100 200  0  a)  •0 C  100  >‘  C)  ‘  0 ci) 0 •0 C  :200 Crank Angle (deg ATDC)  Figure F.7.35  : Diesel flowrate:  -  16.0  35  mg/inj  VIII-B2- 10 Vill-1OB U) (1)  20 40 0 Crank Angle (deg ATDC)  Knock: 1.0 bar IHR: 1039 3 kJ/m  0)  a) 0  1: 200  0 I-  a)  •0 C  10o ci) a) 0 0  C  ‘J  -20  Figure  .  20 0 40 Crank Angle (deg ATDC)  z  0-20  20 40 Crank Angle (deg ATDC)  36  F.7.36 : Diesel flowrate: 14.8 mg/inj  367  VIIFB211 Vill-11A  e. ioo C’)  Knock:1.5 bar IHR:1068kJ/m 3  ,  “300 E _,  200  a  a)  0  a)  a)  “0 -20  gi  100  i\  0 20 40 Crank Angle (deg ATDC)  20 40 Crank Angle (deg ATDC)  Figure F.7.37 : Diesel flowrate: 16.0 mg/inj VIII-B24 1 VIll-1 1A ,  37  Knock:1.3 bar IHR: 1075 kJ/m 3  0)300  V  Cs.  ioo  E  09  200  0-  0)  I  a,  •—‘‘  —  0 -20  0 -20  0 20 40 Crank Angle (deg ATDC)  Figure F.7.38 Diesel flowrate: 16.0 mglinj Vffl-B241 “300  V  2 -,  -200  0  50  a>  -  0100  0 a) V o  38  Knock:1.3 bar IHR: 1065 3 kJ/m  VIII-1 lB  e. ioo 0  20 40 Crank Angle (deg ATDC)  0 -20  0 20 40 Crank Angle (deg ATDC)  0 -20  fri  1  0 20 40 Crank Angle (deg ATDC)  Figure F.7.39 : Diesel flowrate: 15.0 mg/inj  368  V111B242 VIII-12A  (U  Knock:1.5 bar IHR: 1069 kJ/m 3  300 CD  e.  E  gi,iOO Cl)  -,  -200  0 I  50  o  /__  0 -20  cc  N  0  ° 10 (U ci)  CU  0 20 40 Crank Angle (deg ATDC)  A 0 -20  Figure F.7.40 Diesel flowrate: 16.1 mglinj VIIIB242 VIII- 12B  ioo  0 20 40 Crank Angle (deg ATDC)  40  Knock:1.2 bar IHR: 1002 kJ/m 3  300 CD  e.  \  —-  E  U)  _,  a_  -200  CD  cc 0100 (U  ci>  50  N%  ci)  V  o  Q -20  CU  0 20 40 Crank Angle (deg ATDC)  -20  Figure F.7.41 : Diesel flowrate: 14.8 mg/inj V111B212B VIII-12  20 40 Crank Angle (deg ATDC)  Knock:1.1 bar IHR: 1059 kJ/m 3 300  e.  V  E  ioo U)  -,  -200 ci)  ci>  cc  50  ci>  100  V  .  0 -20  0 20 40 Crank Angle (deg ATDC)  0 -20  --  u 20 40 Crank Angle (deg ATDC)  42  Figure F.7.42 : Diesel flowrate: 15.6 mglinj  369  VIIFB2 13 VIII-13A  Knock:2.5 bar LHR: 2191 kJ/m 3 E  100  200  0  50  a)  ioo  920  C -20  40 20 Crank Angle (deg ATDC)  Figure F.7.43 Diesel flowrate: 16.6 mg/inj V111B213 VIII-13  N  ioo 0) a50/  300 a) A  -  -200 a)  100  100  a50/  0 -20  0 20 40 Crank Angle (deg ATDC)  Figure F.7.44 : Diesel flowrate: 16.2 mg/inj V111B213 VIII-13B C’)  Knock:3.1 bar IHR: 2199 kJ/m 3  a)  0 -20  .1  /J  0 20 40 Crank Angle (deg ATDC)  E  a) °  I 4 I  a)  a)  —  A  -)  \  I 0 20 40 Crank Angle (deg ATDC)  Knock:3.1 bar THR: 2227 kJ/m 3 a) .30C  2 200 I  a)  I  “S_  a) .  0 -20  0 20 40 Crank Angle (deg ATDC)  a)  z  0 -20  20 40 Crank Angle (deg ATDC)  Figure F.7.45 Diesel flowrate: 16.3 mg/inj  370  V111B214 VIII-14A  8  Knock:1.8 bar IHR: 2127 kJ/m 3 300 a,  ioo  -  -200  0  a,  50  a,  0100  o  cc  a,  ‘20  (U  (I,  1’  cc  0 -20  20 40 0 Crank Angle (deg ATDC)  Figure F.7.46 : Diesel flowrate: 16.7 mg/inj V111B214 VIII-14  0 40 20 Crank Angle (deg ATDC)  46  Knock:1.8 bar IHR: 2238 kJ/m 3 300 E  100  -  0  200  cc  50  a, 0  (U  100  a,  cc 0  (U  8  20 0 40 Crank Angle (deg ATDC)  -20  Figure F.7.47 : Diesel flowrate: 15.8 mg/mi V111B214 VIII-14B  40 20 0 Crank Angle (deg ATDC)  47  Knock:1.7 bar IHR: 2236 kJ/m 3  V c5  oiOO 0  a,  \N V  200  cc a,  cc 0  20 40 0 Crank Angle (deg ATDC)  0 -20  \ 20 40 0 Crank Angle (deg ATDC)  48  Figure F.7.48 : Diesel flowrate: 16.5 mglinj  371  VIILB215 VIII-15A  Knock:1.5 bar IHR: 1970 kJ/m 3 300 a) 0  a. a? oiOO  S  0  -,  - 200  0  a)  cc  a)  1% 1  a)  100  N  ‘  cc  a) 0  Q  20 40 Crank Angle (deg ATDC)  —  -20  Figure F.7.49 Diesel flowrate: 16.9 mg/inj VIILB245 VIII-15 a. S D 0  S  iOu  -  o  200  cc  /7  a)  0100 Ct a)  V  cc  C1  0 -20  4I  0 20 40 Crank Angle (deg ATDC)  /;  a__Si -20  Figure F.7.50 Diesel flowrate: 16.4 mglinj VIILB245 VIII-15B  50  -D  S -,  200  fr\  cc  N  a)  100 ‘¼  0 a)  a)  cc —  -20  N  Knock:1.5 bar IITR:2017kJ/m 3  a)  U  I  I  0 20 40 Crank Angle (deg ATDC)  S S oiOO 0 S  .2  49  °300 V  50 5  C.  x  20 40 Crank Angle (deg ATDC)  Knock:1.5 bar IIHER: 2009 kJ/m 3  a)  .5  ‘4.  C.;  S  .2  ¶  Ct  20 0 40 Crank Angle (deg ATDC)  0 -20  I  I  M 4 —4 Yt  X  20 40 Crank Angle (deg ATDC)  51  Figure F.7.51 Diesel flowrate: 16.8 mglinj  372  V111B216 VIII-16  Knock:4.5 bar IHR: 2218 kJ/m 3 .  300  c)  E  100 0)  —,  -200 o  0  1  50  / ‘  I  C)  o  0 -20  20 40 0 Crank Angle (deg ATDC)  /  0 20 40 Crank Angle (deg ATDC)  Figure F.7.52 : Diesel flowrate: 17.9 mg/inj V111B216 VIII-16B  52  Knock:4.8 bar IHR: 2215 kJ/m 3 300  co iOO 0)  -200 C)  11  / 50/  -  ‘I, %__  o 9  /\  C)  C) V a)  .  20 40 Crank Angle (deg ATDC)  /1  92 d 1 20 40 Crank Angle (deg ATDC)  Figure F.7.53 : Diesel flowrate: 17.9 mg/inj VIll-B2-17 VIII-17A  53  Knock:2.3 bar IHR: 2278 kJ/m 3 E  C’)  -)  0  -200 a)  1  100  I’  5o ‘%  C)  V C)  00  -20  20 0 40 Crank Angle (deg ATDC)  0 —-/ -20 20 40 Crank Angle (cleg ATOC)  Figure F.7.54 : Diesel flowrate: 16.0 mg/inj  373  VIII-B2- 17 VIII- 17  Knock:2.0 bar IHR: 2194 kJ/m 3 .  U) 0)  300  c)  a?  E  -,  100  a)  200 a)  0  if  I’ I (  a,  50  100  >‘  C.)  a)  a, C) -D C  -20  0  20  40  a)  I  92o’  Crank Angle (deg ATDC)  Figure F.7.55 : Diesel flowrate: 17.5 mg/inj VIII-B2- 17 VIII- 17B a?  D 0  U,  a,  -)  a? 0 C  Knock:2.3 bar 11IR: 2202 kJ/m 3 E  100  -  a,  50  >‘  /  1  *—II’  0  20  40  Crank Angle (deg ATDC)  Cu  p  100  a) .20  200  I  C.) Cu 0 0 C  55  Crank Angle (deg ATDC)  Figure F.7.56 : Diesel flowrate: 17.0 mg/mi VIII-B2- 18 VIII-18A  -D  56  Crank Angle (deg ATDC)  Knock:1.8 bar IHR: 1956 kJ/m 3  a)  a? U, U,  100  a? 0  -  N  a, C  200  50/  /  1)  0 0 C  920  0  20  Crank Angle (deg  0 ATDC)  \  AVAA  20  Crank Angle (deg ATDC)  40  57  Figure F.7.57 : Diesel flowrate: 18.3 mg/inj  374  V111B218 VIII-18  a w 0)  Knock:1.4 bar IHR: 2206 kJ/rn 3 300 a) t  S  iOu  a  S  a)  0  i5 5% o  50  / %%%,  0 a)  iN /  Q)  a:  92o  a  200  20 40 Crank Angle (cieg ATDC)  Figure F.7.58 : Diesel flowrate: 17.1 rng/inj VIII-B2- 18 VIII-18B  20 Crank Angle (deg ATDC)  Knock: 1 1 bar IHR: 2263 kJ/rn 3 *  ‘3OO  V  S wiOO 0 S  S -,  - 200  0  a)  a:  so/  a)  5%  00  o  a)  V a)  CC  %o  o  4o  Crank Angle (deg ATDC)  a  Figure F.7.59 : Diesel flowrate: 16.5 rng/inj VIII-B2-19 VIII-19  59  20 40 Crank Angle (deg ATDC)  60  Knock:1.7 bar IITR: 1087 kJ/rn 3 300 a)  S -,  -  0  a)  I-  200  a:  50  a)  13  100  /_  It  a)  -D  a)  a:  a)  o  40 Crank Angle (deg ATDC)  V  S oiOO 0 S  ‘  58  0 -20  0 20 40 Crank Angle (deg ATDC)  a)  Z  4 w 92(10  Figure F.7.60 : Diesel flowrate: 13.9 mg/inj  375  VIIFB219  Knock:1.3 bar IHR:1133kJ/m 3  VIII-19A 300 a)  e. co  Co  E  iOO  -,  ‘  200  i5 a)  50  5  C-)  -  0  0  20  40  0 -20  Crank Angle (deg ATDC)  u  20  40  Crank Angle (deg ATDC)  Figure F.7.61 : Diesel flowrate: 12.2 V111B219  61  mglinj Knock:1.4 bar 1}IR: 1118 3 kJ/m  VIII- 1 9C  e.  300 V  S  D  ci,  -  a)  200  I  ci)  50  a)  ioo 4  20  0 c  40  -  -  Crank Angle (deg ATDC)  20  40  Crank Angle (deg ATDC)  Figure F.7 .62 : Diesel flowrate: 12.1 VI11B220  62  mglinj Knock: 1.4 bar LHR: 1041 3 kJ/m  VIII-20  300  e.  V  ioo  S  a-  -200 a)  Cl)  50  a) 100  o  a)  V Cu  00  -20  0  20  40  Crank Angle (deg ATDC)  Figure  F.7.63  : Diesel flowrate:  0  z  •—  -20  I *-  -  20  Crank Angle (deg ATDC)  40  63  12.3 mg/inj  376  V111B220 VIII-20A a  Knock:1.2 bar 11IR: 1067 kJ/m 3 300 a,  E  ioo  -)  (I)  -200  0  a, a,  100  I  a,  0  20 40 Crank Angle (deg ATDC)  —  40  64  20 40 -00 Crank Angle (deg ATDC)  65  Crank Angle (deg ATDC)  Figure F.7.64 : Diesel flowrate: 11.6 mg/inj V111B220 VIII-20B a  Knock:1.2 bar IHR: 1062 kJ/m 3 300 a,  E  ioo 0)  200  ——f.\  I...  G) ci  1oo  -  °  C -20  A  —  40 0 20 Crank Angle (deg ATDC)  Figure F.7.65 : Diesel flowrate: 12.8 mg/inj V111B221 VIII-21 a  Knock:l.0 bar 1}IR: 990 kJ/m 3 300 a, V  E  ioo C)  -200 a,  50  a, a,  7  icx  a) .  V  0 -2’V  —  20 0 40 Crank Angle (deg ATDC)  C -20  f 20 40 Crank Angle (deg ATDC)  66  Figure F.7.66 : Diesel flowrate: 12.0 mg/inj  377  V111B221A VIII-21A  Knock:1.1 bar IHR: 1052 kJ/m 3 3O0 t5  0 0  E  iOO  -,  -200 a) I’  50  Il  0100  1  ci) V 0)  ‘20  e.  0 -20  20 40 0 Crank Angle (deg ATDC)  Figure F.7.67 : Diesel flowrate: 11.8 mg/inj V111B221B VIII-21B  )--‘-  ‘‘  -  0 20 40 Crank Angle (deg ATDC)  67  Knock:1.3 bar IIIR:1039kJ/m 3 300 V E  øiOO (I)  -,  -200  0 I-  a)  50  It  ci)  100  V  a)  0-20 __dAtM 0  20 40 0 Crank Angle (deg ATDC)  —  20 40 Crank Angle (deg ATDC)  Figure F.7.68 : Diesel flowrate: 12.2 mg/inj V111B222 VIII-22  Knock:2.9 bar IHR: 2204 kJ/m 3 V  0  ‘11  2  100  68  200  a  ii  50  0)  100  (ss V a)  —  00  -20  0 20 40 Crank Angle (deg ATDC)  I  I  1  0)  “---—-——  -20  0 20 40 Crank Angle (deg ATDC)  69  Figure F.7.69 : Diesel flowrate: 12.4 mg/inj  378  VIII-B2-22 VIII-22B C’, Cl)  100  0 ci) 0 C  50/  /zj  \  Knock:3.3 bar IHR: 2213 kJ/m 3  0)  a)  0  I  E -  a,  il I  200  0 0  ,\  4’  -20  0 -Ic’ -20 40 20 0 Crank Angle (deg ATDC)  20 0 40 Crank Angle (deg ATDC)  Figure F.7.70 : Diesel flowrate: 11.7 mg/inj VIII-B2-22 VIII-22B 0)  -  a)  a) 0 C  cc  >  0  C)  CU  0 a)  cc  (U  a, a,  -20  20 40 0 Crank Angle (deg ATDC)  I  11  200  ,  1  I  ,  :  100  I  a,  ‘I  20 40 Crank Angle (deg ATDC)  3:  Figure F.7.71 : Diesel flowrate: 12.0 mg/inj VIII-B2-23 VIII-23 0)  Knock:1.8 bar IHR: 2245 kJ/m 3  ft  E  100  -  a,  0  200  I’  L.  a)  •0 C  cc  50  a) a) a)  0  a)  0 0 C  71  30C  a) 0 0  I  E  100  70  Knock:2.7 bar IHR: 2198 kJ/m 3 300  0  C) •0 C  I  I  100  a)  -D  I  cc  •0  C) •0 C  i  100  cc 20 40 0 Crank Angle (deg ATDC)  a)  3:  n  -20  /\  I  \%4 —.—---—  20 40 0 Crank Angle (deg ATDC)  .- -  72  Figure F.7.72 Diesel flowrate: 11.9 mg/inj  379  V111B223 VIII-23B  cci  0  300  100  /\\  50  -200 a, I  a, 100  o 0  a  0 -20  20 40 0 Crank Angle (deg ATDC)  I I  a  a)  a)  -  (ci  Knock: 1.8 bar 3 IHR: 2224 kJ/m  Figure F.7.73 : Diesel flowrate: 12.0 mg/inj V11LB223 VIII-23B  0 20 40 Crank Angle (deg ATDC)  Knock:1.6 bar IHR: 2243 kJ/m 3  1  300  S 0  a) 50  /\\  -  200 I’  100 ci)  I  -  a,  0  (ci  20 40 0 Crank Angle (deg ATDC)  -20  Figure F.7.74 : Diesel flowrate: 11.6 mg/inj V111B224 VIII-24  0 20 40 Crank Angle (deg ATDC)  74  Knock:1 .3 bar IHR: 2293 kJ/m 3 300 a) V  S  ioo C’)  -200 50  o  (ci  /  :\  a)  0100  V  a)  —-_-  0  0 20 40 Crank Angle (deg ATDC)  0 -20  20 40 Crank Angle (deg ATDC)  Figure F.7.75 : Diesel flowrate: 12.4 mg/inj  380  V111B224 VIII-24B  Knock: 1.3 bar IHR: 2242 kJ/m 3 €300  Cl) C’)  E  IVU  -  200 50/Z  0 -20  40 20 0 Crank Angle (deg ATDC)  —  i  0 -20  Figure F.7.76 : Diesel flowrate: 11.9 mg/inj V111B224 VIII-24B  40 0 20 Crank Angle (deg ATDC)  76  Knock:1.4 bar IHR: 2242 kJ/m 3 €300  Ci)  E  IUJ  1::  50 V  20 40 0 Crank Angle (deg ATDC)  -20 —  e.  z  C  -20  Figure F.7.77 : Diesel flowrate: 11.6 mg/mi VIII-B2-25 VIII-25 I  colOO 0  Knock:3.4 bar IHR: 2223 kJ/m 3 S —,  1  -..  -200 50/  0  -20 —  20 40 0 Crank Angle (deg ATDC)  €300  D  .  (4  U)  / N. 20 0 40 Crank Angle (deg ATDC)  ioo x  -20  I’ it 1  /  20 40 Crank Angle (deg ATDC)  78  Figure F.7.78 : Diesel flowrate: 13.3 mg/in  381  V111B225 VIII-25B  a Co  Knock:3.8 bar IHR: 2173 kJ/m 3 300  \‘  S  200 50/  / a)  a)100  V  a)  a)  20 40 Crank Angle (deg ATDC)  920’0 40 20 Crank Angle (deg ATDC)  iOO  S  I  S  2  1  -  -20O a)  50  a) Co a)  o  aa)  V a)  100 ‘  6  I  )4ILfl’L  o  -20  _.%‘, -20 40 0 20 Crank Angle (deg ATDC)  20 0 40 Crank Angle (deg ATDC)  Figure F.7.80 : Diesel flowrate: 12.2 mg/inj VIII-B2-26 VIII-26  °3oc  80  Knock: 1.5 bar IFIR: 2221 kJ/m 3  V  S Co  300  C.)  a  a  79  Knock:4.1 bar 11IR: 2213 kJ/m 3 V  \\  s  I  Figure F.7.79 : Diesel flowrate: 12.6 mg/inj VIIFB225 VllI-25B  S co  ‘I  -  0  a  1  2  100  2  100  -)  S  - 200  a)  so/  \  a)  100  aa)  V a)  20 40 Crank Angle (deg ATDC)  A  a)  I I I  o  -20  20 40 0 Crank Angle (deg ATDC)  81  Figure F.7.81 : Diesel flowrate: 13.4 mglinj  382  V111B226 VIII-26B  c  0  Knock:1.6 bar IHR: 2207 kJ/m 3 300  100  0  ci)  50  ci)  100 .22 ci)  o ci)  —  .2  200  u  -20  20 40 0 Crank Angle (deg ATDC)  i  0 -20  Figure F.7.82 : Diesel flowrate: 12.4 mglinj V111B226 VIII-26B  0  300  E  iOO  -,  22  -  0  ci)  i5 50  1)  200  100  ‘  / /  ,.  ci)  92 20 40 Crank Angle (deg ATDC)  Crank Angle (deg ATDC)  e.  82  Knock:1.6 bar LHR: 2193 kJ/m 3 •  Cl)  0 20 40 Crank Angle (deg ATDC)  Figure F.7.83 : Diesel flowrate: 12.6 mg/inj V111B227 VIII-27  83  Knock:1.1 bar IITR: 1992 kJ/m 3 300  22  E  oiOO  22  -  0  / 50  %\  ci)  200  100 ci)  0  K  —  .o  0 -20  —  0 20 40 Crank Angle (deg ATDC)  -20  ‘  X  0 20 40 Crank Angle (deg ATDC)  84  Figure F.7.84 : Diesel flowrate: 13.3 mg/inj  383  V111B227 VIII-27B  Knock:1.1 bar IHR: 1957 kJ/m 3  €  30O E -  200  50 1100 C -20 —  e.  .  0 20 40 Crank Angle (cleg ATDC)  z  C _ick/%fri -20 20 40 0 Crank Angle (deg ATDC)  Figure F.7.85 : Diesel flowrate: 12.3 mg/inj V111B227 VIII-27B  85  Knock:1.4 bar IHR: 1920 kJ/m 3 30C E 200  “N  (100  J  ci)  ?  —  V  -20 —  0 20 40 Crank Angle (deg ATDC)  z  -20  Figure F.7.86 : Diesel flowrate: 12.1 mg/inj V111B228 VIII-28 €  0 20 40 Crank Angle (deg ATDC)  86  Knock:1.6 bar IHR: 1101 kJ/m 3 300 E 200  ‘N .S —  0 -20  100  I 4 “  0 20 40 Crank Angle (deg ATDC)  z  -20  0 20 40 Crank Angle (deg ATDC)  87  Figure F.7.87 : Diesel flowrate: 15.8 mg/inj  384  V111B228 VIII-28B  a D 0 U)  Knock:2.1 bar IHR: 1080 kJ/m 3 300 E  -,  IVy  -200  so7’ 0 -20 —  a  \ 0 -20 40 20 Crank Angle (deg ATDC)  40 20 0 Crank Angle (deg ATDC)  Figure F.7.88 : Diesel flowrate: 14.8 mg/inj V111B228 VIII-28B  88  Knock: 1.7 bar IHR: 1108 kJ/m 3 300 E 200  ‘4 ,  .2  0 -20  —  a  20 0 40 Crank Angle (deg ATDC)  z  0 —-* -20 20 0 40 Crank Angle (deg ATDC)  Figure F.7.89 : Diesel flowrate: 14.6 mg/inj VJII-B2-29 VIII-29  89  Knock:2.8 bar IHR: 1072 kJ/m 3 300  ci  0 0  E  IVy  200  j —  100  a  -20 —  IIk  20 40 0 Crank Angle (deg ATDC)  0 -20  ‘—‘——  40 20 Crank Angle (deg ATDC)  90  Figure F.7.90 : Diesel flowrate: 15.5 mg/inj  385  VIII-B2-29  1038  3 kJ/m  a) V  € S (0  bar  Knock: 1.7 1HR:  VIII-29B  a) 2  100  -,  S  -  0  a)  200 II  50  a)  7  C)  a) a)  V  a)  .2  0 -20  a)  40 0 20 Crank Angle (deg ATDC)  I  1\  100  A LV jvUw t  (‘a  -20  0 20 40 Crank Angle (deg ATDC)  91  Figure F.7.91 : Diesel flowrate: 13.7 mg/inj  VIII-B2-29 VIII-29B  C. (‘S .0  Knock:2.2 bar JHR: 1103 kJ/m 3 a) V  e3  S (0 (0  300  2  100  S  0  a)  200  a) C  a)  a’  a) a)  a)  rr  Cu  C) V C  920  0 20 40 Crank Angle (deg ATDC)  ft,  100  Au EM 2 9  ØJqiA  fIIIft9  Crank Angle (deg ATDC)  92  Figure F.7.92 : Diesel flowrate: 13.5 mg/inj  VIII-B2-30 VIII-30 SD (0 (0  Knock:2.0 bar IHR: 1094 kJ/m 3 a) V  2  100  S  S  0  a) V C  a)  50  a)  a’  200 tL  100  a) a)  V a) Cu  C)  V C  300  920  Figure  0 -20  0 20 40 Crank Angle (deg ATDC)  F.7.93  : Diesel flowrate:  15.5  A  Jrtti 20 40 Crank Angle (deg ATDC)  93  mg/inj  386  VIII-B2-30 VIII-30B  a  Knock:1.6 bar IHR: 1028 kJ/m 3  a) 0  Co Co  9  100  -  a)  0  i5 .9  200  )\  50  ,  100  a, C.)  .9  (ci  a Co Co  0 -20  20 40 0 Crank Angle (deg ATDC)  92 Crank Angle (deg ATDC)  Figure F.7.94 : Diesel flowrate: 13.6 mg/inj VIII-B2-30 VIII-30B  Knock:1.8 bar IHR: 1031 kJ/m 3 0  300  E  1OO  -  0  i3 .9  ci)  50  (“ CU  C.)  200  .9  (U  a  0 -20  I,  100  a,  lip  a, C.)  94  /, A  20 40 0 Crank Angle (deg ATDC)  92’ Crank Angle (deg ATDC)  Figure F.7.95 : Diesel flowrate: 14.4 mg/inj REP-08-06-09 REP-08-06-09  A  95  Knock:2.0 bar fiR: 1499 kJfm 3 300  c) O Co  E  100  -  -200 a) a,  50  7  ioo cci  a) ci)  a,  .9  0 -20  Cci  20 0 40 Crank Angle (deg ATDC)  ci)  i  0 -20  20 40 Crank Angle (deg ATDC)  96  Figure F.7.96 : Diesel flowrate: 13.8 mg/inj  387  REP-08-06-1O REP-08-06-1O  Knock:2.4 bar IHR: 1551 kJ/m 3 300  Co  E  IvJ  1:  50  o  ci)  -o  5  —  -20 —  0 20 40 Crank Angle (deg ATDC)  i  -20  Figure F.7.97 : Diesel flowrate: 14.2 mg/inj REP-08-06-1 1 REP-08-06-11  IP  40 20 Crank Angle (deg ATDC)  Knock:1.9 bar IHR:1545kJ/m 3 300  ci) Co Co  E  I’Jv  -,  7.  0  -200  \  N  .‘  100  C)  -  ci)  •0  .s  a  -20  —  —  0 20 40 Crank Angle (deg ATDC)  0 -20  Figure F.7.98 : Diesel flowrate: -278527.0 mg/inj REP-08-06-12 REP-08-06-12  20 40 0 Crank Angle (deg ATDC)  98  Knock:1.7 bar 3 IHR:1514kJ/m  300 c5 C) Co  E  IVU  -,  1:::  50  -  0 -20 —  0 20 40 Crank Angle (deg ATDC)  z  A 4  0 -20  Ui  20 40 Crank Angle (deg ATDC)  Figure F.7.99 : Diesel flowrate: 13.6 mg/inj  388  REP-08-06- 13 REP-08-06- 13  Cu  D  C’) C’)  C)  -) C) Cu  C)  50  C)  C)  -20  20 40 0 Crank Angle (deg ATDC)  C)  L  100  92&4o Crank Angle (deg ATDC)  Figure F.7.100 : Diesel flowrate: 13.6 mglinj VIII-B2- 1 VIII- 1D 0) C)  •0 Cl)  200  C) C)  0  C0  ann 0))  E  100  0  Cu  Knock: 1.6 bar IHR: 1541 kJ/m 3  Knock:1.7 bar IHR: 1078 kJ/m 3 300  E  100  0  C)  200  n  I.  ci)  50  C)  ‘,  -20  Cu  100  C) C)  C) 0  100  20 40 0 Crank Angle (deg ATDC)  C)  I  921(b  20 Crank Angle (deg ATDC)  Figure F.7.101 : Diesel flowrate: 14.5 mg/inj VIII-B2-1 1 VIII-1 1 0)  101  Knock:1.2 bar 11IR: 1061 kJ/m 3  ci  0  C) D  C’) C,)  E  100  0  C)  cx  C)  50  C)  /%  C)  C  I’ I’  100  C) C)  C)  -o  200  -20  0 20 40 Crank Angle (deg ATDC)  C)  I  0 -20  20 40 Crank Angle (deg ATDC)  102  Figure F.7.102 : Diesel flowrate: 14.6 mglinj  389  V111B219 VIII-19B  Ca  300 a)  Knock:1.1 bar IFIR: 1037 kJ/m 3  -o  E  ioo Co  -  -200  0  50 5’ C.)  zN  a) a) C’, Ca  100  a)  0 i  20 40 0 Crank Angle (deg ATDC)  -  920 20 Crank Angle (deg ATDC)  Figure F.7. 103 : Diesel flowrate: 11.7 mg/mi VIII-B227 VIII-27  300 a)  103  Knock:0.8 bar IHR: 922 kJ/m 3  -o  E  coiOO  -  09  -  -200 a)  a) 50  (1)  cn 100 Ca  0 a)  -4  a)  0 -20  20 40 0 Crank Angle (deg ATDC)  0 -20  20 40 Crank Angle (deg ATDC)  104  Figure F.7.104 : Diesel flowrate: 13.3 mg/mi  390  

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