Residual Gas Mixing in Engines

by

Andrew G. Bright

A thesis submitted in partial fulfillment

of the requirements for a degree of

Master of Science

(Mechanical Engineering)

at the

University of Wisconsin – Madison

2004

ii

i

Abstract

The mixing of fresh charge with residual gases was studied in a spark-ignition engine

using planar laser-induced fluorescence (PLIF) of a homogenous air/fuel/tracer mixture. An

adjustable, dual-overhead cam cylinder head and throttled operation provided a range of

elevated residual gas fractions. The bulk residual fraction was measured with a sampling

valve and exhaust emissions were recorded for 15 experimental conditions covering two

engine speeds and five valve overlap strategies.

Residual gas fractions ranged from 24% to 40% at 600 RPM and 21% to 45% at 1200

RPM. Indicated mean effective pressure ranged from 146 kPa to 271 kPa across all

conditions, with variability levels consistently below 6%. Calculated heat release confirmed

the high dilution levels with universally slow burning rates.

A non-intensified CCD camera was used to capture the PLIF signal and operated with

a peak signal-to-noise ratio of 21:1. The negative-PLIF imaging technique was verified with

a quantitative measure of intake charge homogeneity, and a fuel-cutoff experiment that

isolated unwanted fluorescence signal from residuals. Data images were analyzed with first

and second statistical moments of pixel intensity, as well as an ensemble PDF curve.

All fired conditions showed a clear increase in spatial variation from the

homogeneous condition, a trend that was qualitatively verified visually in the corrected data

images. Inhomogeneity in the compressed charge increased rapidly above 35% residual gas

fraction, independent of engine speed or overlap strategy. The intake cam advance valve

overlap strategy was found to provide reduced spatial variation over equivalent symmetric

valve overlaps and exhaust cam retard overlaps.

ii

Acknowledgements

First to thank for the completion of this project is my advisor, Professor Jaal B.

Ghandhi for giving me the opportunity to pursue my graduate work at the Engine Research

Center. Prof. Ghandhi has been an exceptional point of reference for the myriad challenges

that have presented themselves over the past two years.

The support staff at the ERC also have to be thanked, particularly Sally Radecke and

Susan Strzelec in the office, for tolerating my approach to procedure and paperwork. Also,

Ralph Braun has provided the supplies and access to shop facilities essential to completing

this project.

Very little would have been accomplished without the help of fellow students here,

past and present. Matt Wiles got me started in the engine lab and familiarized me with all

aspects of the laser imaging procedure. Randy Herold has been an invaluable aid throughout

the project with the optical system and emissions analyzers. Lonny Peet provided his time in

completing the accumulator fuel system, which has been a major improvement in the lab.

Brian Albert, Dennis Ward, Bob Iverson, Tongwoo Kim, Soochan Park, Jared Cromas, Nate

Haugle, Karen Bevan, Daniel Rodriguez and Anton Kozlovsky have all given substantial

help along the way. Cheers to all.

The Wisconsin Small Engine Consortium generously assumed funding support mid-

way through this project. The representatives of Briggs & Stratton, Fleetguard/Nelson,

Harley-Davidson, Kohler, Mercury Marine, MotoTron and the Wisconsin Department of

Commerce are to be thanked. Preliminary funding came through a grant from the National

Science Foundation, to which I am equally grateful.

iii

Table of Contents

ABSTRACT..............................................................................................................................I

ACKNOWLEDGEMENTS .................................................................................................. II

TABLE OF CONTENTS .....................................................................................................III

LIST OF FIGURES ..............................................................................................................VI

LIST OF TABLES ................................................................................................................. X

1. INTRODUCTION........................................................................................................... 1 1.1. MOTIVATIONS FOR RESIDUAL GAS STUDY ................................................................ 1

1.1.1. Small Engines Issues............................................................................................. 2 1.1.2. High-Dilution Automotive Engines....................................................................... 3 1.1.3. Homogeneous-Charge Compression-Ignition ...................................................... 4

1.2. PROJECT OBJECTIVES................................................................................................. 6 1.3. OUTLINE .................................................................................................................... 6

2. BACKGROUND ............................................................................................................. 8 2.1. RESIDUAL GAS EFFECTS ON COMBUSTION ................................................................ 8

2.1.1. Combustion Thermodynamics............................................................................... 8 2.1.2. Flame Speed Effects.............................................................................................. 9 2.1.3. Oxides of Nitrogen Formation............................................................................ 11 2.1.4. Cycle-to-Cycle Variations................................................................................... 12

2.2. BULK RESIDUAL GAS FRACTION MEASUREMENT.................................................... 13 2.2.1. Measurement Principle....................................................................................... 13 2.2.2. Sampling Valves.................................................................................................. 14 2.2.3. Sampling Valve Operation.................................................................................. 15

2.3. ONE-DIMENSIONAL STUDIES OF RESIDUAL GAS...................................................... 17 2.3.1. Early Work .......................................................................................................... 17 2.3.2. Recent Work ........................................................................................................ 19

2.4. PLANAR LASER-INDUCED FLUORESCENCE .............................................................. 22 2.4.1. Laser Source ....................................................................................................... 23 2.4.2. Tracer Chemical Selection.................................................................................. 23 2.4.3. Camera................................................................................................................ 25

2.5. PLIF MEASUREMENTS IN ENGINES.......................................................................... 26 2.5.1. 2-d Quantification of SI Engine Flow Inhomogeneity ........................................ 27 2.5.2. Direct Visualization of Residual Gas.................................................................. 30 2.5.3. Negative Visualization of Residual Gas.............................................................. 33

3. EXPERIMENTAL SETUP.......................................................................................... 36 3.1. SINGLE-CYLINDER RESEARCH ENGINE.................................................................... 36

iv3.1.1. Base Engine ........................................................................................................ 36 3.1.2. Optical Access..................................................................................................... 37 3.1.3. Cylinder Head and Combustion Chamber.......................................................... 38 3.1.4. Valvetrain Timing System ................................................................................... 40 3.1.5. Dynamometer...................................................................................................... 43 3.1.6. Engine Fluid Systems.......................................................................................... 43 3.1.7. Engine Aspiration Systems.................................................................................. 44 3.1.8. Fuel Delivery System .......................................................................................... 45 3.1.9. Engine Control System........................................................................................ 48

3.2. COMBUSTION DATA ACQUISITION ........................................................................... 49 3.2.1. Cylinder Pressure Measurement......................................................................... 49 3.2.2. Sampling Valve ................................................................................................... 51 3.2.3. Emissions Bench ................................................................................................. 53

3.3. OPTICAL MEASUREMENT SYSTEM ........................................................................... 55 3.3.1. Laser Source ....................................................................................................... 55 3.3.2. Laser Optics ........................................................................................................ 56 3.3.3. Camera................................................................................................................ 58 3.3.4. Optical Triggering .............................................................................................. 60

4. ENGINE OPERATING CONDITIONS..................................................................... 63 4.1. SELECTION CRITERIA............................................................................................... 63

4.1.1. Optical Engine Considerations........................................................................... 63 4.1.2. Establishing Engine Conditions.......................................................................... 64

4.2. COMBUSTION ANALYSIS .......................................................................................... 67 4.2.1. Cylinder Pressure Data ...................................................................................... 68 4.2.2. Heat Release Analysis......................................................................................... 69

4.3. EXHAUST GAS EMISSIONS MEASUREMENT.............................................................. 74 4.3.1. Emissions Measurement Procedure.................................................................... 75 4.3.2. Emissions Analysis.............................................................................................. 75 4.3.3. Emissions Measurements .................................................................................... 77

4.4. BULK RESIDUAL GAS FRACTION MEASUREMENT.................................................... 79 4.4.1. Sampling Valve Measurement Technique........................................................... 79 4.4.2. Residual Gas Fraction Calculations................................................................... 83 4.4.3. Residual Gas Fraction Measurements................................................................ 84

5. IMAGING SYSTEM DEVELOPMENT AND ANALYSIS ..................................... 86

5.1. PLIF IMAGE PROCESSING ........................................................................................ 86 5.1.1. Image Acquisition Procedure ............................................................................. 86 5.1.2. Image Correction Procedure .............................................................................. 89 5.1.3. Median Filtering ................................................................................................. 91 5.1.4. Image Statistics ................................................................................................... 92 5.1.5. Probability Distribution Function ...................................................................... 94 5.1.6. Image Presentation ............................................................................................. 95

5.2. IMAGING SYSTEM PERFORMANCE............................................................................ 96 5.2.1. Camera Selection................................................................................................ 96

v5.2.2. Region of Interest and Spatial Resolution .......................................................... 97 5.2.3. Signal-to-Noise Ratio.......................................................................................... 99 5.2.4. MicroMax Comparison with Intensified CCD.................................................. 101

5.3. ASSESSMENT OF INTAKE CHARGE HOMOGENEITY................................................. 102 5.3.1. First and Second Moments of Homogeneous Data........................................... 102 5.3.2. Homogeneous Image PDF................................................................................ 104

5.4. DIRECT-INJECTION TEST OF IMAGING TECHNIQUE ................................................ 105 5.4.1. Skip-Direct Injection Experiment ..................................................................... 106 5.4.2. Skip-DI Imaging and Results ............................................................................ 109

6. RESIDUAL GAS MIXING........................................................................................ 111 6.1. SAMPLE IMAGING DATA ........................................................................................ 111 6.2. CORRELATION OF SPATIAL-MEAN PIXEL INTENSITY WITH MEASURED RESIDUAL GAS FRACTION .......................................................................................................................... 113 6.3. CORRELATION OF RESIDUAL GAS FRACTION TO IMAGE INTENSITY VARIATION.... 115

6.3.1. Cycle-Averaged Image Intensity COV Correlation .......................................... 115 6.3.2. Lower Residual Fraction Case-to-Case Comparison....................................... 118 6.3.3. Higher Residual Fraction Case-to-Case Comparison...................................... 121

6.4. PRIOR-CYCLE EFFECT ON IMAGE INTENSITY VARIATION ...................................... 123 6.5. ENGINE OPERATING CONDITIONS EFFECT ON DATA IMAGE INTENSITY VARIATION 126

6.5.1. Symmetric Overlap Increase............................................................................. 128 6.5.2. Intake Cam Advance ......................................................................................... 129 6.5.3. Exhaust Cam Retard ......................................................................................... 133

7. SUMMARY AND CONCLUSIONS ......................................................................... 134 7.1. PROJECT SUMMARY ............................................................................................... 134 7.2. RESULTS SUMMARY............................................................................................... 135 7.3. CONCLUSIONS........................................................................................................ 138 7.4. RECOMMENDATIONS FOR FUTURE WORK .............................................................. 140

REFERENCES.................................................................................................................... 141

APPENDIX A – ENGINE OPERATING CONDITIONS .............................................. 144

APPENDIX B – IMAGE STATISTICS............................................................................ 149

vi

List of Figures FIGURE 1.1. STRATEGIES PURSUED FOR HCCI CONTROL IN CURRENT RESEARCH. REPRINTED

FROM [9]. ........................................................................................................................... 5 FIGURE 2.1. EXPERIMENTAL MEASUREMENTS OF GASOLINE LAMINAR FLAME SPEED IN

EXHAUST GAS-DILUTED MIXTURES RELATIVE TO UNDILUTED MIXTURES, SU(0), FOR A RANGE OF DILUENT FRACTIONS, EQUIVALENCE RATIOS AND INITIAL BOMB PRESSURES. REPRINTED FROM [3]. ...................................................................................................... 11

FIGURE 2.2. SAMPLE CYLINDER PRESSURE DATA FOR IN-CYLINDER SAMPLING IN A SMALL 2-STROKE ENGINE, WITH VALVE LIFT DURATION MEASURED BY AN INDUCTIVE PROXIMITY SENSOR SHOWN. REPRINTED FROM [12]. ......................................................................... 16

FIGURE 2.3. CORRELATION OF MEASURED [CO2] TO LOCAL N2 TEMPERATURE USING CARS. THE PLOT ON THE LEFT IS FOR DATA ACQUIRED AT 30° BTDC WITH A CORRELATION COEFFICIENT OF 0.486. THE PLOT ON THE RIGHT IS AT 5° BTDC WITH A CORRELATION OF 0.420. REPRINTED FROM [20].......................................................................................... 18

FIGURE 2.4. EXPERIMENTAL SETUP FOR RAMAN SCATTERING MEASUREMENTS IN A MODERN 4-VALVE PENT-ROOF COMBUSTION CHAMBER. REPRINTED FROM [8]. ................................ 19

FIGURE 2.5. RESIDUAL GAS MOLE FRACTION VS. CRANK ANGLE, BASED ON ENSEMBLE-AVERAGED CONCENTRATION MEASUREMENTS OF VARIOUS SPECIES. REPRINTED FROM [8].......................................................................................................................................... 21

FIGURE 2.6. LEVELS OF VARIANCE IN DATA FOR ENSEMBLE-AVERAGED MEAN RESIDUAL GAS MOLE FRACTION GIVEN IN FIGURE 2.5. REPRINTED FROM [8]. ......................................... 22

FIGURE 2.7. ABSORPTION AND EMISSION PROPERTIES OF 3-PENTANONE IN LIF APPLICATIONS [17]. ................................................................................................................................. 24

FIGURE 2.8. MEASURED TEMPERATURE DEPENDENCY OF LIF SIGNAL OF ACETONE AT ATMOSPHERIC PRESSURE, NORMALIZED TO ROOM TEMPERATURE CONDITION. REPRINTED FROM [18]. ....................................................................................................................... 25

FIGURE 2.9. MEAN H2O PLIF SIGNAL TREND WITH INTAKE MAP. REPRINTED FROM [22]..... 31 FIGURE 2.10. CYCLIC VARIATION IN H2O PLIF SIGNAL FOR INCREASING LOAD. REPRINTED

FROM [22]. ....................................................................................................................... 31 FIGURE 2.11. CORRELATION OF LOAD-NORMALIZED RESIDUAL GAS FLUCTUATION TO CCV OF

0-0.5% HEAT RELEASE DURATION USING H2O PLIF. REPRINTED FROM [22].................. 32 FIGURE 2.12. COMPARISON OF FLOWFIELD EFFECT ON RESIDUAL GAS DISTRIBUTION AS

MEASURED BY NEGATIVE-PLIF. BOTH CONDITIONS ARE 1200 RPM, ΗVOL = 0.6. REPRINTED FROM [23]. .................................................................................................... 34

FIGURE 2.13. MEAN RESIDUAL GAS DISTRIBUTION ACROSS COMBUSTION CHAMBER (DIRECTION ALONG PENT-ROOF AXIS) FOR TWO BULK FLOWFIELD CONDITIONS. IMAGE DATA TAKEN WITH NEGATIVE-PLIF AT SPARK TIMING (27° BTDC). 1200 RPM, ΗVOL = 0.6. REPRINTED FROM [23]. .................................................................................................... 35

FIGURE 3.1. VALVETRAIN TIMING LAYOUT FOR DOHC CYLINDER HEAD................................ 41

viiFIGURE 3.2. COMPARISON OF MEASURED CYLINDER PRESSURE TRACES AT WALL-MOUNT

LOCATION TO CONVENTIONAL ROOF-MOUNT. MOTORING ENGINE CONDITION WITH OHV HEAD, 1200 RPM............................................................................................................. 51

FIGURE 3.3. IN-CYLINDER SOLENOID-ACTUATED SAMPLING VALVE MOUNTED TO BLOCK-HEAD SPACER RING. TEFLON SAMPLED GAS LINE TRAVELS TO AN ADJACENT ICE BATH AND THEN TO THE ANALYZER. .......................................................................................................... 52

FIGURE 3.4. 266 NM LASER PULSE SEPARATION AND DELIVERY OPTICS (PLAN VIEW).............. 57 FIGURE 3.5. LASER SHEET-FORMING OPTICS SETUP FOR 266 NM PLIF IMAGING. .................... 57 FIGURE 3.6 MICROMAX CAMERA MANUAL SUMMARY OF DIF-MODE TIMING. IMAGE

EXPOSURE TIMES ARE SHOWN IN THE SECOND LINE. READY AND SCAN ARE OUTPUT SIGNALS FROM THE CAMERA CONTROLLER, EXT. SYNC IS THE INPUT TRIGGER TTL, LASER OUTPUT SHOWN IS FOR A DOUBLE-PULSE LASER, THIS EXPERIMENT ONLY USES THE FIRST PULSE. REPRINTED FROM [24]. ......................................................................................... 59

FIGURE 3.7 SCHEMATIC FOR TTL TIMING OF LASER PULSE AND CAMERA, SYNCHRONIZED WITH MOTOTRON SKIP-FIRING IGNITION BY A “ONE-AND-ONLY-ONE” CIRCUIT. ....................... 62

FIGURE 4.1 SUMMARY OF FOUR VALVE OVERLAP STRATEGIES. BASELINE CAM TIMING IS INDICATED BY THE DASHED LINE IN ALL PLOTS. ARROWS INDICATE CAM SHIFT FROM BASELINE. THE BASELINE OVERLAP DURATION IS 20°, THE 600 RPM EXTENDED OVERLAPS ARE 30° DURATION, AND THE 1200 RPM CONDITIONS ARE 60° OVERLAP DURATION. ....................................................................................................................... 66

FIGURE 4.2 HEAT RELEASE RATE AND CUMULATIVE HEAT RELEASE FOR ALL CAM STRATEGIES AT 600 RPM LOW LOAD.................................................................................................. 70

FIGURE 4.3 HEAT RELEASE RATE AND CUMULATIVE HEAT RELEASE FOR ALL CAM STRATEGIES AT 600 RPM MID LOAD. ................................................................................................. 71

FIGURE 4.4 HEAT RELEASE RATE AND CUMULATIVE HEAT RELEASE FOR ALL CAM STRATEGIES AT 1200 RPM MID LOAD. ............................................................................................... 72

FIGURE 4.5 SKIP-FIRING SEQUENCE EXAMPLE (1200 RPM BASELINE OVERLAP SHOWN). SAMPLING VALVE IS ACTUATED ON COMPRESSION STROKE OF SKIP-FIRED CYCLE (SEE TABLE 4.5)....................................................................................................................... 80

FIGURE 4.6 SAMPLE PRESSURE DATA FOR SKIP-FIRED CYCLE WITH SAMPLING VALVE ACTUATION. THE AVERAGE FIRED CYCLE PRESSURE TRACE AND THE SAMPLING VALVE LIFT TRANSDUCER SIGNAL FOR THAT SKIP-FIRED CYCLE (NO PHYSICAL UNITS) ARE OVERLAYED. 1200 RPM EXHAUST CAM RETARD CONDITION SHOWN.............................. 81

FIGURE 4.7 FREQUENCY HISTOGRAM OF PRIOR-CYCLE IMEP FOR SKIP-FIRING OPERATION AT 600 RPM LOW LOAD SYMMETRIC OVERLAP INCREASE CONDITION. DATA COMPILED FROM 100 CONSECUTIVE SAMPLED CYCLES. .............................................................................. 82

FIGURE 4.8 FREQUENCY HISTOGRAM OF PRIOR-CYCLE IMEP FOR SKIP-FIRING OPERATION AT 1200 RPM EXHAUST RETARD CONDITION. DATA COMPILED FROM 100 CONSECUTIVE SAMPLED CYCLES............................................................................................................. 83

FIGURE 5.1 SAMPLE 100-IMAGE MEAN BACKGROUND IMAGE. PIXEL INTENSITY SCALE IS ON RIGHT............................................................................................................................... 87

FIGURE 5.2 100-IMAGE MEAN FLATFIELD IMAGE, 30° BTDC 600 RPM MID LOAD EXHAUST RETARD CONDITION. FLATFIELD IMAGES HAVE BEEN BACKGROUND-SUBTRACTED. ....... 88

FIGURE 5.3 SAMPLE RAW DATA IMAGE (NO CORRECTIONS), 30° BTDC 1200 RPM EXHAUST RETARD CONDITION. ........................................................................................................ 89

viiiFIGURE 5.4 SAMPLE HOMOGENEOUS IMAGES ACQUIRED AT 30° BTDC FOR THE 1200 RPM,

ZERO OVERLAP CONDITION DEMONSTRATING VERTICAL BANDING IN THE CORRECTED IMAGES. SEE SECTION 5.1.6 FOR IMAGE PRESENTATION CONVENTION. ........................... 93

FIGURE 5.5 LOCATION OF ROI WITHIN COMBUSTION CHAMBER, DOHC CYLINDER HEAD. DISTANCE H IS BETWEEN LASER SHEET PLANE AND PISTON FACE, AND IS TABULATED FOR IMAGE TIMINGS IN TABLE 5.1........................................................................................... 97

FIGURE 5.6 CAMERA NOISE CHARACTERIZATION, AS A FUNCTION OF SIGNAL INTENSITY - MICROMAX FRAME-STRADDLING CCD. REPRINTED FROM [14]. .................................... 99

FIGURE 5.7 COMPARISON OF THEORETICAL SHOT NOISE INTENSITY VARIATIONTO MEASURED

HOMOGENOUS PIXEL INTENSITY VARIATION ( )y yσ µ

................................................ 103 FIGURE 5.8 PROBABILITY DISTRIBUTION FUNCTION FOR PIXEL INTENSITY IN HOMOGENEOUS

IMAGE SETS AT FOUR IMAGE TIMINGS FOR ALL THREE ENGINE SPEED/LOAD POINTS. BASELINE VALVE OVERLAP. EACH PDF CURVE CONTAINS INFORMATION ABOUT 100 CORRECTED HOMOGENOUS IMAGES. .............................................................................. 105

FIGURE 5.9 DIRECT-INJECTION EXPERIMENT CYLINDER PRESSURE TRACE COMPARISON WITH DOHC BASELINE VALVE OVERLAP. 600 RPM. ............................................................. 108

FIGURE 5.10 DIRECT-INJECTION EXPERIMENT CYLINDER PRESSURE TRACE COMPARISON WITH DOHC BASELINE VALVE OVERLAP. 1200 RPM. ........................................................... 108

FIGURE 6.1 SAMPLE HOMOGENEOUS IMAGE SEQUENCE, 60° BTDC. ..................................... 111 FIGURE 6.2 SAMPLE DATA IMAGE SEQUENCE, HIGH RESIDUAL FRACTION CONDITION, 60°

BTDC. ........................................................................................................................... 111 FIGURE 6.3 SAMPLE DATA IMAGE SEQUENCE, MID-RANGE RESIDUAL FRACTION, 60° BTDC. 112 FIGURE 6.4 SAMPLE DATA IMAGE SEQUENCE, LOW RESIDUAL FRACTION CONDITION, 60°

BTDC. ........................................................................................................................... 112 FIGURE 6.5 CORRELATION OF MEAN IMAGE INTENSITY RATIO TO MEASURED RESIDUAL

FRACTION FOR ALL 15 EXPERIMENT CONDITIONS. .......................................................... 114 FIGURE 6.6 PIXEL INTENSITY COV VS. RESIDUAL GAS FRACTION FOR ALL ENGINE CONDITIONS

AT 30° BTDC. SHOT NOISE-LIMITED MAXIMUM SNR WAS ~22:1 FOR THIS IMAGE TIMING........................................................................................................................................ 116

FIGURE 6.7 PIXEL INTENSITY COV VS. RESIDUAL GAS FRACTION FOR ALL ENGINE CONDITIONS AT 45° BTDC. SHOT NOISE-LIMITED MAXIMUM SNR WAS ~20:1 FOR THIS IMAGE TIMING........................................................................................................................................ 116

FIGURE 6.8 PIXEL INTENSITY COV VS. RESIDUAL GAS FRACTION FOR ALL ENGINE CONDITIONS AT 60° BTDC. SHOT NOISE-LIMITED MAXIMUM SNR WAS ~18:1 FOR THIS IMAGE TIMING........................................................................................................................................ 117

FIGURE 6.9 PIXEL INTENSITY COV VS. RESIDUAL GAS FRACTION FOR ALL ENGINE CONDITIONS AT 99° BTDC. SHOT NOISE-LIMITED MAXIMUM SNR WAS ~15:1 FOR THIS IMAGE TIMING........................................................................................................................................ 117

FIGURE 6.10 SAMPLE DATA IMAGES FOR 600 RPM, LOW-RESIDUAL CONDITION. ................. 119 FIGURE 6.11 SAMPLE DATA IMAGES FOR 1200 RPM, LOW-RESIDUAL CONDITION. ............... 119 FIGURE 6.12 100-IMAGE PIXEL INTENSITY PDF FOR 600 RPM LOW-RESIDUAL CONDITION. . 120 FIGURE 6.13 100-IMAGE PIXEL INTENSITY PDF FOR 1200 RPM LOW-RESIDUAL CONDITION.121 FIGURE 6.14 SAMPLE DATA IMAGES FOR 600 RPM, HIGH-RESIDUAL CONDITION. 45° BTDC.

....................................................................................................................................... 123

ixFIGURE 6.15 SAMPLE DATA IMAGES FOR 1200 RPM, LOW-RESIDUAL CONDITION. ............... 123 FIGURE 6.16 PRIOR-CYCLE IMEP VS. IMAGE INTENSITY COV. 600 RPM LOW LOAD, SYM.

INCREASE 60° BTDC. YR = 40.4%, IMEP=152 KPA, COVIMEP = 6.0%, ( )y yσ µ

=5.2%............................................................................................................................ 124 FIGURE 6.17 PRIOR-CYCLE IMEP VS. IMAGE INTENSITY COV. 1200 RPM, SYM. INCREASE 60°

BTDC. YR = 43.7%, IMEP=253 KPA, COVIMEP = 1.2%, ( )y y nσ µ

=7.3%............... 125 FIGURE 6.18 MEAN IMAGE INTENSITY VARIATION VS. CA AT 600 RPM LOW LOAD, ALL

OVERLAPS. ..................................................................................................................... 126 FIGURE 6.19 MEAN IMAGE INTENSITY VARIATION VS. CA AT 600 RPM MID LOAD, ALL

OVERLAPS. ..................................................................................................................... 127 FIGURE 6.20 MEAN IMAGE INTENSITY VARIATION VS. CA AT 1200 RPM, ALL OVERLAPS. ... 127 FIGURE 6.21 INTAKE ADVANCE DATA IMAGES AT 600 RPM MID LOAD. 45° BTDC. ........... 130 FIGURE 6.22 EXHAUST RETARD DATA IMAGES AT 600 RPM MID LOAD. 45° BTDC. ........... 130 FIGURE 6.23 INTAKE ADVANCE DATA IMAGES AT 1200 RPM. 45° BTDC. ........................... 130 FIGURE 6.24 EXHAUST RETARD DATA IMAGES AT 1200 RPM. 45° BTDC............................ 130 FIGURE 6.25 INTAKE ADVANCE 100-IMAGE PIXEL INTENSITY PDF AT 600 RPM MID LOAD, 45°

BTDC. ........................................................................................................................... 131 FIGURE 6.26 EXHAUST RETARD 100-IMAGE PIXEL INTENSITY PDF AT 600 RPM MID LOAD, 45°

BTDC. ........................................................................................................................... 131 FIGURE 6.27 INTAKE ADVANCE 100-IMAGE PIXEL INTENSITY PDF AT 1200 RPM 45° BTDC.

....................................................................................................................................... 132 FIGURE 6.28 EXHAUST RETARD 100-IMAGE PIXEL INTENSITY PDF AT 1200 RPM 45° BTDC.

....................................................................................................................................... 132

x

List of Tables TABLE 1.1. SAMPLE RESULTS FROM A HIGH-DILUTION STOICHIOMETRIC DISI ENGINE. CASE 1

REPRESENTS THE BASELINE ENGINE RUNNING THROTTLED WITH PORT FUEL INJECTION. CASE 2 IS A 70-CAD WIDENED VALVE OVERLAP WITH DIRECT INJECTION, SUPPLEMENTED WITH A SECONDARY AIR INJECTION AND A HIGH-ENERGY VARIABLE-GAP IGNITION SYSTEM. BOTH CONDITIONS ARE AT 1500 RPM AND 400 KPA BMEP. [5] ....................... 4

TABLE 3.1. FIXED INTERNAL DIMENSIONS OF GM-TRIPTANE ENGINE. VALVE TIMINGS ARE FOR INTERNAL SINGLE CAMSHAFT USED FOR OHV ENGINE OPERATION.................................. 37

TABLE 3.2. MAJOR COMBUSTION CHAMBER DIMENSIONS FOR GM-TRIPTANE ENGINE WITH DOHC ADJUSTABLE-CAM CYLINDER HEAD. .................................................................... 40

TABLE 3.3 FUEL PROPERTIES FOR PURE ISO-OCTANE AND THE 20% 3-PENTANONE TRACER BLEND USED FOR THIS EXPERIMENT. ................................................................................ 46

TABLE 3.4. HORIBA EXHAUST EMISSIONS ANALYZER BENCH SUMMARY. ............................... 54 TABLE 3.5 TRIGGER TIMING DELAYS FOR OPTICAL MEASUREMENT SYSTEM. DELAYS ARE

RELATIVE TO THE LEADING EDGE OF THE TRIGGER SIGNAL FROM THE CRANKSHAFT ENCODER.......................................................................................................................... 61

TABLE 4.1. AIR/FUEL ENGINE OPERATION PARAMETERS FOR THE THREE EXPERIMENTAL SPEED/LOAD POINTS. THESE VALUES WERE HELD CONSTANT FOR EACH CAM STRATEGY. 67

TABLE 4.2 MEAN EFFECTIVE PRESSURE DATA FOR 100-CYCLE AVERAGE PRESSURE DATA AT ALL EXPERIMENTAL CONDITIONS. PERCENTAGES SHOWN ARE CHANGES RELATIVE TO THE BASELINE OVERLAP CONDITION FOR THE INDIVIDUAL SPEED/LOAD POINTS AT EACH CAM STRATEGY. ....................................................................................................................... 68

TABLE 4.3 FLAME DEVELOPMENT ANGLES AND OVERALL BURNING ANGLES FOR DIFFERENT OVERLAP STRATEGIES, DETERMINED BY A SINGLE-ZONE HEAT RELEASE CODE. PERCENTAGES INDICATED ARE CHANGES RELATIVE TO THE BASELINE OVERLAP CONDITION AT EACH SPEED/LOAD POINT. ........................................................................................... 73

TABLE 4.4 SUMMARY OF EXHAUST EMISSIONS SPECIES MEASUREMENTS, CONCENTRATIONS SHOWN ARE CORRECTED TO A WET BASIS FROM THE RAW READINGS. AIR/FUEL RATIO AND COMBUSTION EFFICIENCY COEFFICIENT HAVE BEEN CALCULATED FROM THE CONCENTRATION DATA. ................................................................................................... 77

TABLE 4.5 SAMPLING VALVE OPERATION FOR ALL EXPERIMENTAL CONDITIONS. SAMPLING FREQUENCY IS LISTED AS THE NUMBER OF FIRED CYCLES BETWEEN SAMPLED CYCLES (SEE FIGURE 4.5). .................................................................................................................... 80

TABLE 4.6 SUMMARY OF BULK RESIDUAL GAS FRACTION MEASUREMENTS AT ALL EXPERIMENTAL CONDITIONS. PERCENTAGES SHOWN ARE CHANGES RELATIVE TO THE BASELINE OVERLAP CONDITION AT EACH INDIVIDUAL SPEED/LOAD POINT. ...................... 85

TABLE 5.1 DISTANCE FROM PISTON FACE TO LASER SHEET ROI FOR EXPERIMENT IMAGE TIMINGS. .......................................................................................................................... 98

xiTABLE 5.2 VALUES OF SPATIAL-MEAN DATA IMAGE INTENSITY AND RESULTING SHOT NOISE-

LIMITED MAXIMUM SNR FOR THREE SPEED/LOAD POINTS. EACH SET IS THE MEAN VALUE FOR THE FIVE VALVE OVERLAP STRATEGIES................................................................... 100

TABLE 5.3 DIRECT INJECTION EXPERIMENT ENGINE CONDITIONS AND UNBURNED HYDROCARBON EMISSIONS MEASUREMENTS. * INDICATES THE APPROXIMATE IGNITION TIMING. .......................................................................................................................... 106

TABLE 5.4 DIRECT INJECTION EXPERIMENT IMAGING RESULTS. 100-IMAGE MEAN SIGNAL LEVEL FOR FLATFIELD, SKIP-FIRED, AND MOTORED SKIP-DI PLIF DATA........................ 109

TABLE 6.1 COMPARISON OF LOWER-RESIDUAL CONDITIONS AT 600 AND 1200 RPM.

DEVELOPMENT OF IMAGE ( )y yσ µ

[%] WITH CRANK ANGLE..................................... 119 TABLE 6.2 COMPARISON OF HIGHER-RESIDUAL CONDITIONS AT 600 AND 1200 RPM.

DEVELOPMENT OF ( )y yσ µ

[%] WITH CRANK ANGLE. ............................................... 122

1

1. Introduction

1.1. Motivations for Residual Gas Study

Residual gas plays an important role in the combustion development process in four-

stroke cycle spark-ignition (SI) engines. This type of internal combustion has to this day

been the dominant prime-mover in automobiles and utility engine applications. Residual gas

is present in all engines and has important implications to the designer in terms of engine

stability and pollutant emissions.

Residual gas is especially significant in its role as a diluent species during

combustion. This property provides the major benefit to increased residual gas fractions –

reduction in NOx generation during combustion. NOx is a major pollutant species in internal

combustion engine exhaust.

The advent of variable valvetrain actuation (VVA) systems in recent years has

provided much more freedom to the spark ignition engine designer to utilize the exhaust

residual for pollutant reduction and load control, in addition to improvements in volumetric

efficiency across the engine speed and load range. VVA, commonly performed by

mechanical or electro-hydraulic phase-shifting of the camshaft, is becoming increasingly

common on new automotive engine designs.

More information about the participation of residual gas in engine flows preceding

combustion reactions will be critical to achieving the maximum potential (in terms of SI

engine emissions and efficiency) of this and other dilution-controlling technologies.

21.1.1. Small Engines Issues

Small engines can be defined as the category of internal combustion engines below

500 bhp used for non-automotive applications, principally in power equipment, motorcycles

and marine transportation. Despite sharing similar if not identical operation fundamentals,

small engines have unique engineering considerations to automotive SI engines. When faced

with new challenges related to emissions regulations, small engine manufacturers do not

have the luxury of simply adopting mature technologies from the automotive industry.

Of particular concern is NOx emissions, which have only been reduced to

environmentally acceptable levels in cars by universal use of three-way exhaust catalysts

(TWC). For many small engines, the unit cost of the automotive TWC exceeds that of the

entire engine, and as such this technology is not deemed practical in the category. Instead of

aftertreatment, focus is being placed on charge dilution strategies for NOx reduction, and the

simplest delivery mechanism is through internal recirculation via residual gas.

Since VVA systems also fall outside the cost-acceptable realm of most small engine

designs, elevated residual gas fractions will likely be provided by fixed camshaft profiles.

This presents a strong challenge to the combustion chamber designer, with the need to

accommodate high-dilution mixtures throughout the engine speed and load range without

negatively impacting performance felt by the user. More must be learned about charge

composition development at high dilution levels in small engines for this worthy goal to be

achieved.

31.1.2. High-Dilution Automotive Engines

New applications of high residual gas dilution occur in novel engine designs.

Olafsson et al. in [5] describe a high-dilution spark ignition engine designed at Saab to

reduce fuel consumption and NOx emissions. The engine has a similar objective as seen with

direct injection spark ignition (DISI) engines which typically operate without intake

throttling and thus enjoy large improvements in part-load fuel efficiency. The critical

drawback to DISI engines is that by using excess fresh air, the highly effective and durable

three-way catalyst cannot be used to control NOx, CO and HC emission. By utilizing the

exhaust gas residual instead of excess air, Olafsson et al. were able to operate at overall

stoichiometric conditions with a 10% reduction in part-load fuel consumption from the

conventional SI engine. This engine design requires complicated engine systems such as

continuously variable camshaft phasers to control residual dilution, air-assisted in-cylinder

fuel injection, and most notably, a variable spark plug gap to consistently ignite dilute

mixtures. Sample results from this project are presented in Table 1.1.

4 Case 1 Case 2 Change MAP [kPa] 50 93 --- BMEP [kPa] 400 400 --- PMEP [kPa] 54 11 --- COV of IMEP [%] 1.0 1.5 --- BSFC [g/kWh] 265 228 - 14 % BSNOx [g/kWh] 16 0.6 - 96 % BSHC [g/kWh] 6 9 + 50 % BSCO [g/kWh] 19 9 - 50 % Exhaust Temp [C] 560 450 --- 0-10% HR [CAD] 24 35 --- 10-90% HR [CAD] 20 22 --- IGN timing [bTDC] 25 41 ---

Table 1.1. Sample results from a high-dilution stoichiometric DISI engine. Case 1 represents the baseline engine running throttled with port fuel injection. Case 2 is a 70-CAD widened valve overlap with direct injection, supplemented with a secondary air injection and a high-energy variable-gap ignition system. Both conditions are at 1500 RPM and 400 kPa BMEP. [5]

1.1.3. Homogeneous-Charge Compression-Ignition

Homogeneous Charge Compression Ignition (HCCI) is a rapidly developing new

engine combustion strategy that could combine some of the best operating characteristics of

SI and diesel engines. In particular, HCCI can achieve the part-load fuel efficiency of diesel

engines with substantially reduced in-cylinder soot and NOx emissions on the level of SI

engines. Like knock in homogeneous charge SI engines, HCCI involves a controlled

autoignition that can be obtained with a variety of petroleum-based fuels. Controlling the

autoignition of a mixture is separated into 2 strategies: altering the fuel mixture reactivity

kinetics and altering the time-temperature history of the mixture. Cooled external EGR is

often explored for the former, given the usual need to delay the onset of compression

5ignition. The latter strategy commonly involves significant heating of the fuel/air charge

which can encourage the onset of autoignition in engines with lower compression ratios.

Figure 1.1. Strategies pursued for HCCI control in current research. Reprinted from [9].

This lower-compression ratio configuration would enable dual-mode operation with

part-load HCCI combustion transitioning to full-load spark ignition combustion. Intake air

heating, while convenient in a laboratory, is not deemed practical for mobile applications.

Instead, the focus is being placed on the use of VVA to deliver high residual fractions for

heating of the charge. High-dilution operation may be a likely application of HCCI for

improving the efficiency of gasoline automotive engines [9, 10]. For this and a variety of

other reasons, the mixing and chemical kinetics of the exhaust gas residual is a growing topic

of research.

6

1.2. Project Objectives

Four broad objectives have been identified for this research:

1. To provide high-quality, spatially and temporally resolved, two-dimensional

quantification of residual gas mixing with fresh homogenous air/fuel charge through a

range of positions in the SI engine cycle.

2. To supplement and correlate the mixing data with engine-out operating information

such as cylinder pressure data and exhaust emissions analysis for a range of residual

gas dilution levels.

3. To extract conclusions from the residual gas mixing measurements and engine

performance data that will be helpful to the field in designing high-dilution engines.

4. To aid in the development of Planar Laser-Induced Fluorescence as an invaluable

combustion diagnostic in SI engines.

1.3. Outline

This thesis will be divided into six subsequent chapters. Chapter 2 presents the

project background in the form of a literature review of residual-effected SI combustion,

sampling valve measurements, prior optical studies of residual gas and the use of PLIF in

engines. Chapter 3 contains a detailed, design-oriented discussion of the experimental

facility including the research engine, combustion diagnostic instrumentation, and the optical

system. Chapter 4 will present the engine operating conditions covered in the project,

including the basis for their selection and the measurements of bulk residual gas fraction at

7each condition. Chapter 5 will discuss the development of the imaging technique,

particularly the selection criteria for the hardware and processing steps and subsequent

performance of the data images. Chapter 6 will contain the residual gas mixing data derived

from the PLIF images, with discussion. Finally, chapter 7 contains project summary,

conclusions and recommendations.

8

2. Background

2.1. Residual Gas Effects on Combustion

Recycled exhaust gas has a substantial effect on combustion processes by acting as a

diluent, meaning that it does not participate in the oxidation of the fuel but is present and

absorbing the released energy in a quantity significant enough to reduce flame speed and gas

temperature [2]. Decreasing flame front speed inherently lengthens the time to reach 10, 50,

and 90% mass-fraction burned levels, extending combustion reactions further into the

expansion stroke. If the engine control system is not able to adjust other parameters

properly, residual gas dilution can slow the burning rate to a point where partial-burn and

misfire cycles emerge with severe penalties on emissions and performance. The

temperature-mitigating effect of residual gas is well-known as a strategy for reducing oxides

of nitrogen (NOx) production in internal combustion engines.

2.1.1. Combustion Thermodynamics

Residual gas in a spark-ignition engine running at a stoichiometric air/fuel ratio is

composed predominantly of N2, CO2, H2O and O2. Engines that operate fuel-rich of

stoichiometry, such as small air-cooled utility engines, will see significant CO and H2 and

very little remaining O2 in the residual gas. In most SI engines, pollutant species such as

NOx and unburned hydrocarbon compounds (HC) normally sum to 1% or less by volume [1].

9Based on this composition, it can be seen that when added to a mixture of vaporized fuel and

air, residual gas will lower the mass-specific heating value of the mixture. For constant-

volume combustion, the first law of thermodynamics can be expressed as

reactants products ad f( , ) = ( , )i iU T p U T p (2.1)

where Tad is called the adiabatic flame temperature and is easily calculated from a balanced

reaction equation by assuming adiabatic conditions, ideal gas behavior, and no dissociation

of reactants or products into minor species [4]. These assumptions make exact calculations

difficult but the trend of in-cylinder flame temperature vs. initial reactant composition

becomes clear. Residual gas species reduce the total enthalpy (formation plus sensible) of

the reactants, which is related to the initial internal energy by the universal gas constant, and

thus reduce the flame temperature from that of undiluted air/fuel mixtures.

2.1.2. Flame Speed Effects

The effect of reducing adiabatic flame temperature is observed in reduced burning

velocity. Combustion in an SI engine occurs via a turbulent, thin-sheet wrinkled flame

structure, which, despite being inherently complex is locally modeled closely by laminar

flame propagation rates. The laminar flame speed, SL has been measured [24], and for

conventional hydrocarbon fuels has been found to obey the power law equation:

10

,00 0

uL L

T pS ST p

α β

=

(2.2)

where the reference values are standard temperature and pressure and SL,0, α and β are

tabulated constants for particular combinations of fuel and equivalence ratio. The term Tu

represents the unburned gas temperature just ahead of the reaction zone in the flame front.

Rhodes and Keck [3] studied gasoline combustion with controlled residual concentration in a

constant-volume bomb experiment and quantified a laminar flame speed correction factor for

Equation (2.2) given the inclusion of a residual gas fraction in the reaction, based on the data

of figure 2.1:

0.77( ) ( 0)(1 2.06 )L r L r rS x S x x= = − (2.3)

Decreasing the flame temperature and velocity represents a significant challenge to

maintaining appropriate engine performance. If, for whatever reason, reactant preheating

temperatures fall below 1900 K, flame velocity will be at or near the partial-burn and misfire

lower limit [5]. This situation might typically arise if the exhaust valve opens prior to

completion of flame propagation, or if the flame is prematurely extinguished [1]. Partial

burn and misfire are extreme symptoms of cycle-to-cycle variation (CCV) in engine power

output. Besides contributing to unwanted engine roughness characteristics, the incomplete

combustion of the fuel charge represents a very significant emission of HC pollutants.

11

Figure 2.1. Experimental measurements of gasoline laminar flame speed in exhaust gas-diluted mixtures relative to undiluted mixtures, Su(0), for a range of diluent fractions, equivalence ratios and initial bomb pressures. Reprinted from [3].

2.1.3. Oxides of Nitrogen Formation

Another major consequence of the dilution effect of residual gas is reduced NOx

formation. NOx is a primary ingredient in photochemical smog found in the lower

atmosphere mainly above major cities. It also is known to contribute to acid rain. NOx is

also regrettably known for being somewhat inextricably linked with engine performance and

efficiency. Rate equations for the formation of NOx are non-linear functions of time,

elevated temperature and availability of nitrogen and oxygen molecules. Peak NOx

formation at optimal combustion phasing occurs close to stoichiometric air/fuel ratio, which

also represents the operating point for peak engine stability, power output and efficiency [4].

122.1.4. Cycle-to-Cycle Variations

Increased residual fractions are expected to locally affect small-scale mixture

homogeneity, which describes imperfect distribution of fuel vapor within the air and residual

charge. It is assumed that low to moderate spatial inhomogeneity will affect combustion

only during the earliest stages near the discharge of the spark plug and the formation of a

flame kernel. The scales of non-uniformity are larger or of the same order of the enflamed

volume during these critical early instants. As the flame front area grows much larger, the

effect of inhomogeneity is averaged out in a global sense [7, 8].

The variation of air/fuel ratio and residual dilution in the vicinity of the spark gap has

an important effect on cycle-to-cycle variations (CCV) in SI engines. Local mixtures outside

the ignition limit or too dilute to rapidly transition into a fully developed turbulent flame are

common causes of misfire and high CCV [1]. In their literature review of cyclic variation,

Ozdor et al. [6] summarized several studies of mixture inhomogeneity on flame development.

They point to a general uncertainty in applicable length scales of non-uniformities, but to a

demonstrated effect of controlled in-cylinder turbulence (particularly swirling motion) at

time of spark on reducing CCV. At the time of writing (1994), they point out that none of

the dozens of papers reviewed were able to quantify the impact of spatial inhomogeneity of

residual gas on CCV.

13

2.2. Bulk Residual Gas Fraction Measurement

In this project, residual gas mixing quantifications will be performed for varying

levels of residual gas fraction. This quantity, denoted yr, is defined as the mass of burned

exhaust gases carried over from the previous cycle’s combustion process relative to the total

cylinder mass. Like most other in-cylinder quantities, yr is subject to cycle-by-cycle

variation in magnitude. However, cycle-averaged values can be measured using in-cylinder

gas sampling as will be discussed in this section.

2.2.1. Measurement Principle

The exhaust gas emissions analyzer bench has become a standard engine test cell

instrument and typically provides concentration measurements of CO2, CO, O2, NO and HC

present in a stream of exhaust gas. Given this measurement capability, the most direct way

of quantifying total cylinder residual gas fraction is by the relation:

%( )%( )

CO2

CO2

compr

exh

xx

x= (2.5)

which defines a ratio of mole fractions of CO2 in the cylinder during the compression stroke

(after IVC) and the exhaust system downstream of the engine, typically after passing through

a mixing volume. It is important that this calculation be made on a “wet basis,” where the

absence of water vapor in NDIR CO2 analyzers is accounted for. Water is always condensed

14out of the exhaust sample lines since it is damaging to instruments. There are a few

techniques for correction and they typically involve knowledge of fuel chemistry, CO2 and

CO “dry basis” readings and intake air relative humidity [1].

2.2.2. Sampling Valves

Extracting an emissions analyzer sample during the compression stroke from the

closed cylinder is most directly performed with a category of hardware known as the fast-

acting sampling valve. Sampling valves have been employed as early as 1927 to aid the

study of chemical and physical processes in engine combustion.

Zhao and Ladommatos [14] document a more comprehensive summary of valve

designs employed in the engine literature. Most sampling valves covered were either of the

outward-opening poppet type or inward-opening needle type. Needle valves hold advantages

of smaller tip diameters, which can be advantageous in space-confined combustion chamber

surfaces, and also a lack of physical intrusion into the combustion chamber volume. Poppet

valves benefit from better sealing performance, aided by combustion pressures and potential

for smaller crevice volumes via flush-mount machining. It is proposed by the authors that

needle valve sampling volumes will be slightly larger in reach across the combustion

chamber.

Although mechanical and electro-hydraulic sampling valves have been used for

engine studies in the past, the most popular actuation mechanism is electromagnetic force.

Typically driven by a linear solenoid, this design must feature a high traction force to

counteract a strong return spring used for valve sealing and high armature acceleration for

15minimum lift duration [11]. Utilization of programmable research/calibration-type digital

engine control systems has greatly improved control of valve response. Additionally,

monitoring the valve stem lift with an inductive proximity sensor in the back side of the

valve body can provide necessary feedback for exact location of the valve window [12].

2.2.3. Sampling Valve Operation

For sampling of residual gas mixtures, the ignition system should be synchronized to

shut off during the cycle of valve actuation to prevent alteration of the residual concentration.

Monitoring the effect of skip-firing the engine is important in controlling the quality of the

analyzed residual gas mixture. It is expected that after the misfire of the sampled cycle, the

following cycle will be strong due to the residual gas being composed of additional unburned

fuel/air. It is necessary to ensure that the next sampled cycle follows a cycle that is

representative of the steady-state engine performance. One example from the literature is

that Hinze & Miles, in [7], found that the third cycle following the skip-fired cycle had an

average IMEP equal to the steady 100-cycle average for a 32 kPa MAP, 800 RPM condition.

For residual fraction measurement, sampling valve opening frequency must be optimized for

maximum sample gas flow rate and minimum deviation of sampled cycle characteristics

from steady-state conditions.

16

Figure 2.2. Sample cylinder pressure data for in-cylinder sampling in a small 2-stroke engine, with valve lift duration measured by an inductive proximity sensor shown. Reprinted from [12].

One other concern with global residual fraction measurements with fast-acting

sampling valves is that the volume of sampled gas must be representative of the total cylinder

charge. In designing the UW/ERC poppet-type sampling valve in [15], Foudray referenced

sources that indicated that a minimum of 10% to 25% of cylinder volume is adequate to

characterize cylinder composition, depending on degree of stratification. Although that

research was focused on 2-stroke cycle engine exhaust scavenging, the same criteria are

believed to hold for the 4-stroke cycle engine. Using a bellows flow meter, Foudray

estimated a sampling mass flow to be within a range of 33% to 66% of per-cycle cylinder

mass. Leakage was measured to be approximately 3% of the sample flow rate and neglected

in calculations.

17

2.3. One-Dimensional Studies of Residual Gas

Raman scattering has been used for many years to provide in-cylinder temporally-

resolved measurements in IC engines. Three papers are reviewed here where this one-

dimensional optical technique has been used to characterize residual gas participation in SI

engine flows.

Line spectroscopy studies hold advantages over two-dimensional imaging in the

reduced impact of optical access and the ability in many cases to track individual chemical

species without the use of tracers. They are inherently limited by their one-dimensional

nature and within that, a limited spatial resolution.

2.3.1. Early Work

Lebel and Cottereau in [20] performed an early study of residual gas effects on SI

combustion. They measured simultaneous CO2 concentration and N2 temperature using a

Coherent Anti-Stokes Raman Scattering (CARS) setup, with a fixed measurement region 1

cm long and 100 µm in diameter. CO2 was chosen to track residual gas, while charge

temperature was monitored to ensure that same-cycle burned gases in the firing engine were

not present in the measurement region. Laser beam intensity referencing was used to allow

comparison of single-shot measurements. Correlations were reported, at a single operating

condition, between [CO2] and temperature, cycle peak cylinder pressure (PP) and location of

peak pressure (LPP) at instants before and after ignition and two locations near and far from

the spark plug.

18Very poor correlation was found between [CO2] and PP/LPP in measurements taken 1

mm from the spark plug and 5° bTDC (considered end of ignition delay). Since this is

counter-intuitive, the authors conclude that, given their limited measurement region, it

indicates that the residual gas is not perfectly mixed at the end of the compression stroke.

The only meaningful correlation reported in this paper is between increasing [CO2] and

increasing T (figure 2.3), which is somewhat obvious given the charge heating property of

residual gas. As local temperature readings did not correlate with pressure data, this would

reinforce the statement that residual gases (and thus local charge temperatures) are stratified

late in the compression stroke. Direct correlations of [CO2] with PP/LPP yielded coefficients

from -0.2 to 0.2, limiting the authors to very basic conclusions for effects of local residual

gas concentrations on engine performance with this technique.

Figure 2.3. Correlation of measured [CO2] to local N2 temperature using CARS. The plot on the left is for data acquired at 30° bTDC with a correlation coefficient of 0.486. The plot on the right is at 5° bTDC with a correlation of 0.420. Reprinted from [20].

192.3.2. Recent Work

Hinze and Miles at Sandia National Laboratories performed two subsequent line-

imaging studies of residual gas mixing [7, 8], developing a detailed statistical quantification

for mean and fluctuating inhomogeneity components. Both studies utilized a laser

measurement volume in an axially centered position, in which CO2, H2O, N2, O2 and C3H8

concentrations were recorded. Binning on the CCD array divided the volume into individual

adjacent measurement points which established the spatial resolution. Data was presented in

15 CAD increments from start of intake to TDC compression. Homogenous propane/air

mixtures were supplied at stoichiometric conditions. Neither paper presents engine

performance data.

Figure 2.4. Experimental setup for Raman scattering measurements in a modern 4-valve pent-roof combustion chamber. Reprinted from [8].

20Ensemble-averaged measurements were taken to describe mean stratification of fresh

charge and residual gas, while 500-cycle single-shot images were analyzed to establish a

cycle-to-cycle fluctuating component. These data were used to generate spatial covariance

functions of species mole fractions (based on the adjacent measurement points), which were

broken down into fluctuation components coming from system noise, turbulence, and bulk

composition. These covariance functions, once developed, could be used to extract integral

length scales of local residual gas fraction fluctuation (the scale over which turbulent

fluctuations remain correlated.)

In their first paper [7], Miles and Hinze utilized a side-valve, side-spark optical

engine to test this technique at the same engine operating conditions in two bulk flowfields –

a semi-quiescent condition and a high-swirl condition. The measurement volume was 11 mm

long and 0.49 mm in diameter, divided into 12 measurement points. The quiescent flow was

shown to homogenize rapidly, with fluctuations in residual gas concentration nearly

eliminated by 150° bTDC. For the swirling flow, the measurement volume was radially

traversed away from the centerline to two additional measurement regions. Gradients were

observed throughout the cycle in the mean concentration data between these volumes which

suggested a bulk charge stratification which persisted throughout the compression stroke.

Rms fluctuations in the mixture composition at spark time were 5 times higher in the swirling

condition (5% vs. 1% for quiescent at -15 CAD.) Mixing length scales for both conditions

were found to vary from 2 to 5 mm.

In the second paper [8], Hinze and Miles moved to a more conventional pent-roof, 4-

valve cylinder head for their measurements and chose to focus on a single engine condition

representative of idle. Figure 2.5 shows the reported development of the ensemble-averaged

21residual gas fraction during the engine cycle. In this experiment, the measurement volume

was 14.5 mm long and 0.27 mm in diameter divided into 16 sub-regions, improving the

spatial resolution by nearly a factor of two. During the intake stroke, the authors were able to

track residual gas backflow into the intake and a later period where all the residual gas has

been re-inducted away from the measurement volume. The largest gradients in the

measurement volume occurred at BDC, as shown in Figure 2.6, with significant gradient

breakdown during compression similar to the first project. Length scales encountered at -180

CAD were on the order of 1 cm. Rms fluctuation (1%) and mixing length scale range (2-4

mm) at spark time were comparable to the previous experimental computations.

Figure 2.5. Residual gas mole fraction vs. crank angle, based on ensemble-averaged concentration measurements of various species. Reprinted from [8].

22

Figure 2.6. Levels of variance in data for ensemble-averaged mean residual gas mole fraction given in figure 2.5. Reprinted from [8].

2.4. Planar Laser-Induced Fluorescence

Planar laser-induced fluorescence (PLIF) is an increasingly popular advanced

combustion diagnostic. PLIF has the ability to provide quantitative two-dimensional

measurements in single-phase or multi-phase flows with exceptional spatial and temporal

resolution. A general summary of a PLIF measurement system is a high-energy, pulsed laser

sheet propagating through a flowfield containing a suitable fluorescent tracer species

resulting in absorption and subsequent emission of photons at a characteristic wavelength of

the tracer molecules. With a process time response on the order of nanoseconds, individual

laser shots can be captured by a CCD camera for correction and analysis.

23Detailed discussion of PLIF theory has been presented in the literature [13, 15] and

will not be repeated here. Instead, a summary of the important characteristics of the system

components used in this project are covered, including laser source, camera, and tracer

chemical.

2.4.1. Laser Source

The traditional laser source for PLIF work in engines is the Nd:YAG laser, which

offers high-power laser pulses at four harmonic wavelengths, 1064 nm, 532 nm, 354 nm and

266 nm. Laser pulses are delivered at an optimal repetition rate, most commonly 10 Hz.

Individual pulses are on the order of 8 ns duration with maximum energies exceeding 100

mJ. Nd:YAG lasers can operate with external triggering and can thus be synchronized with

engine events, although the low repetition rate typically precludes sequential measurements

in the engine cycle. Pulsed laser operation requires attention to shot-to-shot variation in laser

beam intensity and profile when making quantitative measurements.

2.4.2. Tracer Chemical Selection

Since neither air nor iso-octane fluoresce under the range of wavelengths supplied by

the Nd:YAG laser, a tracer chemical is doped into the intake charge at a controlled

concentration. Tracer addition can occur by either on-the-fly seeding of the intake air or by

pre-mixing in solution with the fuel, depending on the targeted measurement. Maximum

tracer concentration must yield maximum fluorescence signal without significant laser power

24attenuation or influence on combustion performance. The most popular class of tracers for

combustion PLIF is the di-ketone group, and the preferred match for iso-octane research is 3-

pentanone, based on its closely-related distillation curve. Tracer-matching is far more

important in multi-phase PLIF where evaporation rates must be matched than in pre-

vaporized homogenous charge studies.

1.0

0.8

0.6

0.4

0.2

0.0Rel

ativ

e A

bsor

ptio

n, F

luor

esce

nce

500450400350300250 λ (nm)

Absorption Fluorescence

Optical Properties of 3-Pentanone

Figure 2.7. Absorption and emission properties of 3-pentanone in LIF applications [17].

The excitation wavelengths for di-ketones fall in the ultraviolet, with an absorption

range of 225-320 nm [17]. Thurber et al. performed important studies on the temperature

[18] and pressure [19] dependence of acetone fluorescence at various excitation wavelengths.

It was shown that temperature dependence is practically eliminated on the range of 300-700

K using 289 nm. Likewise, an optimal wavelength for neglecting pressure effects is shown

25to be 308 nm. Making the extension of the acetone behavior to 3-pentanone, tuning the laser

wavelength to a value near 289 nm is highly beneficial in quantifying engine flows which are

at all temperature-stratified.

Figure 2.8. Measured temperature dependency of LIF signal of acetone at atmospheric pressure, normalized to room temperature condition. Reprinted from [18].

2.4.3. Camera

The di-ketone tracer group emits photons in a broadband range of 350-550 nm [17].

This visible light is best collected by a high-resolution scientific-grade CCD camera.

Charge-coupled devices contain a photo-sensitive pixel array, which when impacted by

photons, convert the photon energy to electron charge potentials with a quantum efficiency

26that is a property of the device. The individual pixel charges are read out sequentially into a

registry where they are amplified and digitized for computer processing [14].

There are four sources of noise important in making quantitative measurements with

CCD images: dark, read, pattern and shot noise. Dark noise arises from thermal generation

of electrons in the array and is limited with cooled (thermo-electric or cryogenic) CCD chips.

Read noise is a property of the array readout circuit and the programmed readout rate. Fixed

pattern noise can be traced from sources on either the CCD chip or the imaging subject, and

is unique in this discussion in that it can be eliminated with standard background and flatfield

image correction. Shot noise is typically the limiting noise element in high-fidelity CCD

imaging such as found in PLIF studies. Shot noise is completely independent of the CCD

type and arises from the probabilistic nature of photon impingement on the pixels. The shot-

noise limited signal-to-noise ratio is equal to the square root of the number of photons

incident per CCD pixel, based on Poisson statistics [13].

2.5. PLIF Measurements in Engines

As mentioned in the previous section, planar laser-induced fluorescence is a powerful

IC engine diagnostic tool due to its two-dimensional nature and superior spatial and temporal

resolution. Previous studies at the UW/ERC have achieved sufficient spatial resolution to

calculate scalar dissipation and used it to quantify the degree of mixedness in stratified DISI

flows [15, 16]. Additionally, using two high-shuttering speed intensified CCD cameras,

Rothamer [13] was able to simultaneously image unburned and burned mixtures to quantify

27flame-front equivalence ratio in a stratified-charge DISI engine. For the current study of

residual gas mixing in engines, it is important to first present basic techniques for quantifying

spatial charge inhomogeneity from PLIF intensity data and then introduce the limited

literature on residual gas studies using this technique.

2.5.1. 2-d Quantification of SI Engine Flow Inhomogeneity

Baritaud and Heinze conducted an early application of PLIF in an SI engine at the

Institut Français du Pétrole (IFP) in 1992 [21]. The subject of their experiment was

quantification of the development of fuel/air stratification in a PFI engine. A major portion

of this paper discusses the statistical means for describing charge inhomogeneity in PLIF

images.

The authors define a total standard deviation for a set of N single-shot images, based

on the idea that a single image’s inhomogeneity can be quantified by its standard deviation

about the spatial mean (σn). By ensemble-averaging this value after normalizing each by the

mean image intensity ( nI ), the influence of the pulse-to-pulse variation in laser intensity is

removed:

1

1 Nn

totnnN I

σσ=

= ∑ (2.6)

28The total standard deviation σtot is presented as a relative value, since absolute measures of

charge inhomogeneity cannot be correlated with individual engine cycles without bias error

from the pulse energy variations.

To extract the maximum potential information from the data images, the simple

standard deviation was broken down into fine-scale and large-scale contributions by

employing a basic spatial Fourier transform. First, a 3x3 smoothing procedure was twice

performed on the I x J pixel data image, with the resulting smooth field termed Φ(In(i,j)).

The large scale contribution to the inhomogeneity, arising from gradients in large-scale

structures in each data image n is:

( )( )( )2

n,lf,

1 , nni j

I i j IIJ

σ = Φ −∑ (2.7)

After ensemble averaging, the relative large scale variation is:

n,lf

1

1 N

LFniN I

σσ

=

= ∑ (2.8)

Likewise, small-scale fluctuations in each image can be tracked by examining the fluctuation

in the raw image intensities relative to the smoothed image:

( )( ) ( )( )2

n,hf,

1 , ,n ni j

I i j I i jIJ

σ = Φ −∑ (2.9)

29

This value is again ensemble averaged on a normalized basis:

n,hf

1

1 N

HFniN I

σσ

=

= ∑ (2.10)

If the ensemble-averaged pixel intensity field ( ),nI i j is used in place of the single-

image data in equation (2.9), a “hybrid” fluctuation arises which can describe the variation of

the large-scale inhomogeneities from cycle-to-cycle:

( )( ) ( )( )2

n,cyc,

1 , ,n ni j

I i j I i jIJ

σ = Φ −∑ (2.11)

Importantly, ( ),nI i j is biased by laser pulse variations, which limited its usefulness in this

initial study. Finally, this value can also be ensemble-averaged to a relative basis.

n,CCV

1

1 N

cycniN I

σσ

=

= ∑ (2.12)

The authors indicate that it is difficult using metrics such as σtot, σLF, σHF, and σcyc to

separate single-cycle inhomogeneity effects from cycle-to-cycle variations captured in the

data images.

302.5.2. Direct Visualization of Residual Gas

Direct visualization of combustion residual species such as H2O and NO2 is possible,

although challenging, with PLIF. In [22], Johansson et al. used water as a residual tracer,

which required use of strategy known as “2-photon” LIF, which is unique in its requirement

for an interaction of two photons at 248 nm to detect the water molecule. This approach

yields inherently lower signal levels than a single-photon LIF study like those done on fuel

tracers. Additionally, the authors were unable to provide a homogeneous distribution of

water molecules at a known concentration, which prevented signal calibration and therefore

quantification of the H2O intensity data.

The objective of this study was to observe the influence of residual gases on cycle-by-

cycle variations in engine power output. The optical access system required a vertical laser

sheet only 6 mm in height. The laser sheet centerline was passed 4.5 mm below the spark

plug and water concentration images were obtained for a range of engine loads (based on

intake MAP.) Cylinder pressure-derived heat release data were compiled to correlate

residual gas levels with initiation and propagation of SI combustion. The engine was

operated on homogeneous natural gas at 700 rpm, and the images were acquired 1° before

spark time. Imaging was performed with an intensified CCD gated to 100 ns exposure.

Resulting noise levels due to low signal strength and maximum intensifier gain were roughly

20%.

The conclusions made on ensemble-averaged water intensity data were fairly basic,

essentially confirming predicted trends in increasing residual gas concentration near the

spark plug with decreasing load. When normalized by the equivalence ratio of the data set,

31the duration of 0-0.5% heat release was shown to correlate well with the CCV of the water

concentration normalized by load point. This is thought to strengthen the argument that

fluctuation in residual gas near the spark plug is a major contributor to CCV in SI engines.

Unfortunately, quantitative values of the observed fluctuations were not available.

Figure 2.9. Mean H2O PLIF signal trend with intake MAP. Reprinted from [22].

Figure 2.10. Cyclic variation in H2O PLIF signal for increasing load. Reprinted from [22].

32Johansson et al. also attempted correlations with pressure and heat release data for the

single-cycle measurements. Although laser power intensity fluctuations were corrected in

this experiment by shot-resolved power meter readings, the poor SNR and small imaging

region created a large amount of scatter in these correlations. The correlation between

duration of 0-0.5% HR and [H2O] was optimized for radius of ROI within the image. At a

low-load condition, a peak 60% correlation was shown at a radius of 2.9 mm. This

correlation degraded with decreasing residual fraction, which was satisfactory since the

magnitude of the fluctuations relative to the image noise was expected to also decrease.

Figure 2.11. Correlation of load-normalized residual gas fluctuation to CCV of 0-0.5% heat release duration using H2O PLIF. Reprinted from [22].

332.5.3. Negative Visualization of Residual Gas

Residual gas can also be tracked with PLIF images by examining the negative of the

intensity field provided by a homogeneous air/fuel/tracer charge. Following up on the early

work described in Section 2.5.1, Deschamps and Baritaud at IFP [23] performed a negative-

PLIF visualization of burned gas distribution in an SI engine. Because this project sought to

observe separately the distributions provided by external EGR as well as internal residual

gas, the upstream intake air was chosen to be seeded with biacetyl. Air seeding via a

carburetor imparted more uncertainties and challenges than premixed fuel solutions. A 25-

mm wide horizontal laser sheet was passed 4 mm below the spark plug parallel to the ridge

of the cylinder head’s pent roof.

For the internal residual gas study, five engine effects were examined: fuel type, fuel

distribution, tumble level, spark plug location and volumetric efficiency. Mean image

intensity profiles in the direction of the sheet across the pent roof were examined, but only in

a qualitative manner.

The enhanced tumble experiment was conducted with propane to remove fuel

stratification effects. With enhanced tumble, mixing along the roof ridge direction was

observed to be more difficult during the intake stroke than during compression, where it is

assumed that the tumble motion normal to the laser sheet is broken down by turbulence.

However, by the end of compression, the enhanced tumble condition shows both a higher

concentration and flatter linear distribution than the standard case. The increased

concentration suggested that lower tumble levels leave a portion of the residual gas trapped

in the bottom of the combustion chamber. Increased charge motion then not only helps

34distribute the residual gas vertically in the combustion chamber, but laterally to create a more

homogenous mixture. Another property of enhanced tumble operation proposed by the

authors is improved SI combustion efficiency which often correlates with increased intake

MAP, reducing bulk residual fraction.

Figure 2.12. Comparison of flowfield effect on residual gas distribution as measured by negative-PLIF. Both conditions are 1200 RPM, ηvol = 0.6. Reprinted from [23].

35

Figure 2.13. Mean residual gas distribution across combustion chamber (direction along pent-roof axis) for two bulk flowfield conditions. Image data taken with negative-PLIF at spark timing (27° bTDC). 1200 RPM, ηvol = 0.6. Reprinted from [23].

With varying volumetric efficiencies, changes in the distribution of residual gas in the

data images taken at -30 CAD are explained primarily through assumed changes and

asymmetries in the intake port flows, imparting different bulk flowfields. The residual gas

concentration in the image ROI decreases with increasing volumetric efficiency as expected.

Deschamps and Baritaud conclude in this section of the paper that the interacting

parameters they studied were too complex for control of residual gas distribution in an

engine, and suggest choosing external EGR as a delivery mechanism instead. The remainder

of the paper discusses EGR effects in a similar manner, only with the addition of emissions

work.

36

3. Experimental Setup

3.1. Single-Cylinder Research Engine

This project was performed on a single-cylinder, optically-accessible research engine

mated to a regenerative AC dynamometer. For improved control of residual gas dilution, a

dual overhead cam cylinder head was integrated. Calibrated air flow was delivered from a

critical flow orifice rack and control of air-assisted fuel injection and spark timing was

provided by a commercial engine control and calibration system.

3.1.1. Base Engine

The base engine block for this project is the GM Research “Triptane Base 4”,

originally designed for alternative fuels research in the late 1950’s. It is of two-part

construction, with cast iron crankcase and cylinder barrel. The crankcase contains a

balancing shaft and a single fixed two-lobe camshaft for pushrod actuation of an overhead-

valve system. The cylinder barrel has been re-lined recently and contains a liquid coolant

jacket. The firedeck surface includes a groove for an o-ring seal with the cylinder head

spacer ring. The major fixed dimensions of the Triptane engine are provided in table 3.1.

37Bore [mm] 92.4

Stroke [mm] 76.2

Displacement [cc] 511

Connecting Rod Length [mm] 144.8

Exhaust Valve Open [CAD] 115

Exhaust Valve Close [CAD] 365

Intake Valve Open [CAD] 349

Intake Valve Close [CAD] -180

Table 3.1. Fixed internal dimensions of GM-Triptane engine. Valve timings are for internal single camshaft used for OHV engine operation.

3.1.2. Optical Access

The major feature of the Triptane engine is the Bowditch-type optical-access

piston/cylinder geometry. The extended-height cylinder barrel accommodates the aluminum

Bowditch piston and allows for mounting of the 45° mirror, which passes through the

cylinder barrel and allows for a periscope view of the combustion chamber via a transparent

piston cap.

The piston cap is fabricated of aluminum and is fastened to the Bowditch piston with

an internally threaded steel retaining ring. The cap contains an axially-centered 47 mm-

diameter 10 mm-thick sapphire window. The fit of the cap into the retaining ring is indexed

and the assembly locks with a small screw-fastened key. Cylinder sealing for the window-

cap and cap-retainer surfaces is performed with Viton O-rings.

38The piston rings used for this experiment are common to optical engine studies and

unique in that they operate without a lubricating oil film on the cylinder wall. Custom

manufactured by the C. Lee Cook Company based on dry gas compression technology, they

are composed of a spring-loaded oil control ring and a bronze-impregnated Nylon rider ring

below the mirror and an additional rider ring above the mirror. The single compression ring

is of a butt-cut design and is made of Vespel. The compression ring groove is located in the

steel retaining ring at a maximum height that does not cross the firedeck surface gap.

The final component of the optical access system is the steel spacer ring fastened

between the block and head, with an inside diameter matching the engine bore. The 25 mm-

tall ring contains four equally-sized window ports. Two ports contain 16.5 mm-thick quartz

windows for laser sheet propagation through the combustion chamber. The other two ports

are utilized for combustion diagnostics described in Section 3.2. Although the piston cap

crosses the plane of the windows near TDC, the compression ring stays below the spacer ring

throughout the cycle.

3.1.3. Cylinder Head and Combustion Chamber

In the interest of generating a range of residual gas fractions for this experiment, a

means of independent cam phasing was required. Since the base engine’s OHV camshaft is

of fixed geometry and difficult to access within the crankcase, a dual overhead camshaft

(DOHC) single-cylinder research cylinder head of near-identical bore was obtained from GM

Research Labs. Originally designed and used for gasoline direct- injection (GDI) studies, the

39cylinder head contains intake and exhaust cams that are independently phase adjustable via

taper-split drive pulleys.

As an additional lab improvement, the DOHC cylinder head provided a combustion

chamber geometry that is consistent with modern multi-valve SI engines. The pent-roof

cylinder head contains two intake valves and two exhaust valves with an axially-centered

M14 spark plug. The GDI injector bore (tangential, wall-guided orientation) was plugged in

this project. One intake port is cast in a helical approach for swirl generation, which can be

varied with a butterfly throttle on this port alone. This throttle was left full-open for this

project to generate maximum flowfield turbulence.

The major consideration in the integration of this cylinder head was its effect on

compression ratio. The DOHC head’s pent-roof occupies a 49.5 cc volume, whereas the

Triptane engine’s traditional OHV head is designed with a flat “pancake” roof. The

requirements of the optical access system prevented modification to either the spacer ring or

the Bowditch piston, so the highest-compression flat-top piston crown was used in this

project. With an OHV setup, this piston yields a CR of over 12:1 for compression ignition

studies, but with the DOHC head, we are able to obtain only 5.95:1. The upside of this

arrangement is that the engine free-spins without any valve-piston interference, allowing

infinite valve timing flexibility and reduced risk of catastrophic engine damage. During

initial testing, the low compression engine was demonstrated to operate stably at elevated

dilution conditions. The major combustion chamber dimensions for this project are

summarized in Table 3.2.

40Compression Ratio 5.95:1

Top-Ring Crevice Volume [cc] 4.13

Exhaust Valve Inner Seat Diameter [mm] 29.5

Intake Valve Inner Seat Diameter [mm] 34.4

Table 3.2. Major combustion chamber dimensions for GM-Triptane engine with DOHC adjustable-cam cylinder head.

Mating the 116 mm square bolt pattern of the DOHC head to the 117x86 mm

rectangular pattern of the Triptane firedeck required fabrication of four hardened steel mating

blocks for an offset 2-screw fastening method. The deck surface of the DOHC head was

lowered by 0.030” to accommodate a replaceable graphite head gasket to seal against the

spacer ring. As mentioned, the spacer ring sealed to the firedeck with a Viton O-ring.

Additional lab modifications for the DOHC setup are described in subsequent sections.

3.1.4. Valvetrain Timing System

The major feature of the DOHC cylinder head is its independent cam phasing

adjustment. This is accomplished with indexed taper-split pulleys on each camshaft. The

inner flange half is permanently fastened and keyed to the camshaft along with a backing

graduated degree wheel. The outer pulley half is fastened through radial slots against the

internal taper. The slots were machined so as to allow access to any practical cam phasing

arrangement.

41Prior to this project, an external, belt-driven half-speed shaft was added to the

laboratory to provide a timing signal for engine control software. For the DOHC

arrangement, this half-speed shaft was linked to the camshaft pulleys at a 1:1 ratio via a

Gates 1”-width 52”-length 3/8”-pitch trapezoidal tooth timing belt. An automotive OEM 1”

torsional belt tensioner was added to the half-speed shaft assembly for final belt tensioning.

Despite its apparent complexity, this dual-belt timing system eliminated the need for a very

long single timing belt and camshaft extensions, as well as providing required modularity

with the OHV engine setup.

Figure 3.1. Valvetrain timing layout for DOHC cylinder head.

42The degree wheels for each camshaft were first calibrated to valve open/close events.

The graduations on the degree wheels are to be read against markers bolted to the valve cover

with the engine rolled to TDC compression. This baseline point was used because it is

assumed that both cams will always be on the base circle at this time. Using a 0.006” valve

lift threshold of open/close timing and proper oil pressurization of the hydraulic valve lifters,

the engine is manually rolled over with a crankshaft degree wheel, watching a dial indicator

on the appropriate valve stem. This procedure does suffer from significant degree value

uncertainty of valve timing, due to the effect of engine rotational speed on hydraulic lifter

response. However, engine operation data have shown the timing system to be highly

repeatable in terms of engine-out performance.

A related procedure was developed for selecting valve timings during the experiment.

The engine was rolled to TDC compression, and the cam taper was broken with the timing

belt still taught. The pulley half is backed off slightly (not completely off) to allow relative

rotation between the two halves. The flange half contains a hex nut that can be used to turn

the camshaft to a different degree wheel position relative to the timing belt, which remains

locked to the crankshaft. The pulley was then re-fastened on the taper. It is important in this

procedure that the engine always is manually rotated in its proper counter-clockwise

direction to avoid the multi-degree backlash in the belt tensioner and that the timing belt

tension is preserved throughout the process.

433.1.5. Dynamometer

The crankshaft of the Triptane engine is connected, via a flywheel, to a three-phase

440 VAC General Electric dynamometer. The control system is a Reliance Electric Max Pak

Plus VS Drive box. The dyno system can provide motoring or generating operation up to

1500 RPM and 30 kW load. Manual selection of engine speed is performed with a rheostat

and feedback control loop, which is periodically optimized.

3.1.6. Engine Fluid Systems

The Triptane engine and DOHC cylinder head are liquid cooled in a conventional

block-thru-head loop, linked by external hoses. A 50/50 water-antifreeze mixture was used

for corrosion resistance. The external, motor-driven circulating pump is pressure-fed by a

standing column reservoir. An electric water heater is operated continuously during

experiments to bring the engine to operating temperature and a feedback-controlled solenoid

valve is used to meter cold building water through a copper counter-flow heat exchanger to

maintain a system set point during operation. Previous optical engine projects have

determined the optimum coolant temperature of 68° C.

Oil pressure and flow rate was also provided by an external pump, with a commercial

in-line filter. For this project, a Triptane internal post-main bearing oil gallery was selected

to provide an external high-pressure feed to the DOHC head cam journals and valve lifters.

A low-pressure drain line was also installed from the head to the reservoir in the crankcase.

Engine oil selected for optical studies is SAE 40, for its higher viscosity and resistance to

44infiltrating the combustion chamber. With low speeds, loads and temperatures, engine oil

grade selection is not considered critical to Triptane performance.

To provide further defense against oil fouling of the optical access system, vacuum

pumps are applied to both the crankcase and the cylinder head valve cover. Crankcase

vacuum cuts down on blow-by past the oil control ring, which can quickly foul the turning

mirror, camera lens and back surface of the piston window. Valve cover vacuum was

necessary in this project to reduce oil migration past the valve stem seals, which fouls all

combustion chamber windows.

3.1.7. Engine Aspiration Systems

Intake air is metered through a critical flow orifice rack, where the upstream pressure

was varied to obtain set mass flow rates. Three orifice diameters are used to provide an

adequate range of air flow at the common engine operating speeds of 600 and 1200 RPM:

0.125”, 0.100” and 0.050”. All three orifices are calibrated with a bellows flow meter for the

range of upstream pressures providing choked flow. Mass air flow rate is then determined

from a density correction. During the experiment, the supply air (separated, filtered and

humidity-controlled central compressed air) is monitored for consistent upstream properties

with the calibration condition.

For this project, only the smallest 0.050” orifice is used, since all experiment

conditions were “throttled” or sub-atmospheric intake manifold absolute pressure (MAP).

Also for this consideration, a new pressure-tested rigid copper intake runner was fabricated to

link the DOHC head with the existing 14 gal intake air surge tank. Intake MAP is monitored

45with a Wallace & Tiernan 0.1-psi resolution absolute pressure gauge at the surge tank. An

atmospheric intake vent is opened for all transient dyno speed selection periods.

For the exhaust, a new steel runner was fabricated. Approximately 10 cm

downstream of the cylinder, the runner contains an axially-located K-type thermocouple used

to confirm thermal steady-state firing operation. Engine exhaust emissions are sampled from

the near-exit centerline of a 12 gal mixing tank located 2.25 m downstream of the ports. This

surge tank is positioned to allow for modularity with the OHV head setup and contains a

perforated tube diffuser entrance for gas mixing. Exhaust back pressure is manually

controlled with a gate valve and monitored on an absolute pressure gauge. All conditions in

this project were set to 1.0 bar absolute back pressure.

3.1.8. Fuel Delivery System

The fuel used in all measurements of this experiment is 80% iso-octane, 20% 3-

pentanone, by volume. The tracer concentration was set as being the maximum level that did

not attenuate the laser sheet intensity across the combustion chamber. Since this experiment

involved a homogeneous, pre-vaporized mixture, the tracer was assumed to faithfully track

the fuel. More thorough discussions of the co-evaporation properties of iso-octane and 3-

penatnone are found in [14] and [15], where direct-injection spray studies mandated the extra

consideration. Relevant properties of the fuel mixture are given in Table 3.3.

46 Iso-octane 20% 3-pentanone / 80% iso-octane Molecular Weight 114.23 106.16 H:C Ratio 2.25 2.20 Stoich. Air/Fuel Ratio 15.1 14.25

Table 3.3 Fuel properties for pure iso-octane and the 20% 3-pentanone tracer blend used for this experiment.

This optical project required that the intake air and vaporized fuel/tracer solution be

thoroughly pre-mixed before entering the combustion chamber. To accomplish this, the fuel

injector is mounted far (1.3 m) upstream of the intake ports. Based on previous optical

engine studies, an Orbital air-assisted fuel injection system is used to provide the highest

level of initial fuel/tracer atomization. Degree of homogeneity is analyzed in Section 4.3.

The Orbital air-assist injector is primarily used in North America in the Mercury

Marine Optimax 2-stroke outboard engine line, where it operates in a GDI arrangement. In

this system, a fuel injector draws fuel from a pressurized fuel rail and fires into a mixing

volume on the entrance to an adjacent air injector, which is within a separate 80 psi

compressed air rail. A fixed delay of 4 ms occurs for initial air/fuel mixing before the air

injector fires the mixture into the intake runner for a fixed duration of 3 ms. The critical

operation parameter with this system is that the fuel supply rail be held at 10 psid above the

compressed air rail (= 90 psi). Fuel delivery rate is varied with the initial fuel injection pulse

width.

Due to the use of pure iso-octane fuel and 3-pentanone tracer, the OEM high pressure

pump and recirculating pressure regulator system are not used for the Orbital fuel rail.

Pressurized fuel/tracer and air are instead supplied by a clean accumulator system. Mixed

47fuel/tracer solution is drawn into the Tobul piston-type accumulator (Model 4.5A20-8-5763)

by a vacuum pump. The accumulator, which contains Teflon-encapsulated seals for

chemical resistance, is then pressurized with nitrogen. After passing through a safety shut-

off valve and a 0.5 micron filter, the fuel is regulated with a Go single-stage regulator (Model

00-HO2073) to the necessary 10 psi differential pressure. An Orange Research differential

pressure gauge (Model 1516D6) is used to directly monitor this value during the experiment.

Regulated compressed air for the injector is supplied by a medical-grade cylinder.

Fuel delivery rate was calibrated before the experiment using a gravimetric technique

for varying durations of fuel injector pulsewidth. The engine control system was operated

using its internal timing generator mode set at 600 RPM. Permanent injector characterization

settings along with fuel-air delay and air injector duration were fixed in software, with only

the fuel injector duration varied. The air-assist injector rail was mounted on an atmospheric-

pressure mixing volume and drained through a ¼” tube into a capped glass beaker on an

Ohaus Scout II digital scale. Overall mass flow rate (over a six minute duration after initial

flow equilibrium) was then converted to a mass per injection value. As expected, the mass

flow rate was linear across the delivery rate range of this project (10-18 mg/inj). Gain and

offset were then entered into the engine control software.

There is substantial uncertainty in the overall air/fuel ratio with this experimental

setup. Particular sources are from the just-described injector calibration, which is not

performed in a negative-pressure environment as found in the intake system, the well-known

chamber sealing inefficiencies of the optical engine piston ring pack, uncertainties in the

intake orifice calibration and small air leaks in the intake system.

483.1.9. Engine Control System

Control of fuel injection, ignition, sampling valve and camera triggering for this

project was provided by the MotoTron commercial engine control and calibration package.

Using triggers from an interpolated crankshaft encoder and a camshaft Hall-effect encoder,

the ECU software is able to generate output signals with 1/16-CAD resolution. A major

advantage of this particular package is its built-in support for the Orbital air-assist fuel

injection system, including characterization and calibration software inputs and integrated

electronic driver circuits. Furthermore, it provides up to six spark ignition TTL signals,

which can be given independent timings and pulse widths for triggering of external systems.

For this project, Mototron engineers generously provided the laboratory with a new

programmed feature for “skip-fired” operation. When enabled, this mode provides a user-

defined number of firing cycles to occur before cutting the ignition on a single cycle. During

that skip-fired cycle, a TTL signal of user-defined timing and duration is activated for

triggering of a sampling valve or camera.

The ignition coil used for this experiment is a Mercury Marine “DFI” model, with

maximum spark energy of 150 mJ. Due to the high residual dilution levels and low

compression ratio of this project, the maximum coil dwell was used at all times. Likewise,

the AC Delco spark plug (Model 41-954) was gapped to 2.1mm, a very large amount, but

one that was proven to consistently sustain spark propagation.

49

3.2. Combustion Data Acquisition

To add relevance to the optical studies of this project, multiple combustion

diagnostics were needed. Most important is cylinder pressure data, which in addition to

allowing for optimization of performance at each running condition, allows for analysis of

cylinder heat release rate and cycle-by-cycle power variations. To quantify bulk residual gas

fraction at each running conditions, a solenoid-actuated in-cylinder sampling valve is used in

conjunction with an exhaust gas emissions analyzer. This emissions bench is also used to

describe general running trends in pollutant formation.

3.2.1. Cylinder Pressure Measurement

Cylinder pressure was measured with an AVL model QC42D-E C109 water-cooled

piezoelectric transducer. The charge output was sent to a Kistler model 5010 amplifier,

operated with a medium time constant. The amplified signal was logged on a Hi-Techniques

A/D conversion PC running REVelation software. Pressure readings were recorded in sets of

100 cycles at 0.25 CAD resolution, based on a simultaneous signal from a high-resolution

BEI optical crankshaft encoder. The pressure transducer/amplifier were calibrated using a

hydraulic dead-weight tester with an excellent resulting linearity (R2 > 99.9%).

The relative pressure signal was pegged to the intake MAP at -180 CAD in software.

The REVelation program provided instant display of an averaged trace, IMEP and COV of

the recorded data in the laboratory, but additional statistics and cycle-resolved data were only

accessible using a binary data extraction program. Extracted pressure traces for each final

50running condition were entered into a single-zone heat release code which operates in the

Engineering Equation Solver (EES) environment, for analysis of cumulative and

instantaneous heat release rate.

An important consideration in this project was the location of the AVL pressure

transducer (M14 thread) within the combustion chamber. Space constraints in the 4-valve

DOHC head prevented traditional roof access. An unused window in the optical spacer ring

was selected and a special tapped window was machined. There were initial concerns about

dynamic effects on the pressure trace at this location from the piston, which passes the

window location near TDC, placing the transducer in the top ring crevice volume.

A test was performed with the OHV head, using two pressure transducers (Figure

3.2). One was mounted in that head’s roof location and the other in the window port. Both

were logged simultaneously and compared. The window transducer followed the shape of

the roof transducer exactly, except for a small (< 5%) deviation of peak pressure near TDC.

The behavior of the cylinder wall location was deemed acceptable for this project.

51

Pressure Trace Comparison

0

500

1000

1500

2000

2500

-90 -75 -60 -45 -30 -15 0 15 30 45 60 75 90

CAD

p [k

Pa]

Roof MountWall Mount

Figure 3.2. Comparison of measured cylinder pressure traces at wall-mount location to conventional roof-mount. Motoring engine condition with OHV head, 1200 RPM.

3.2.2. Sampling Valve

Residual gas fraction levels are measured by comparing the concentrations of CO2

trapped in-cylinder before ignition to that in the combustion products in the exhaust. To

obtain the in-cylinder CO2 concentration, a solenoid-actuated sampling valve is installed in

the head/block spacer ring in a modified steel window opposite the pressure transducer. The

design details of this sampling valve are covered in [12].

The sampling valve outlets into a 1 m long ¼” Teflon line connected to an ice bath.

The condensing ¼” stainless steel tube coil removes water from the stream and two external

52coalescing filters remove impurities before passing the sampled gas stream through a 10 m

long ¼” Teflon line to the emissions bench.

The sampling valve 48 VDC driver circuit is triggered by the skip-fire MotoTron

TTL signal. Skip-firing mode is used for cylinder sampling to ensure faithful measurement

of pre-ignition trapped charge composition. Sampling valve timing and duration are

optimized for the engine running condition to provide maximum flow rate to the emissions

bench. The target flow rate is 2.5 lpm, measured by a rotameter at the bench entrance.

Figure 3.3. In-cylinder solenoid-actuated sampling valve mounted to block-head spacer ring. Teflon sampled gas line travels to an adjacent ice bath and then to the analyzer.

Before the experiment, the leakage rate past the sampling valve seat was measured

during three continuous-firing baseline engine operating conditions without opening the

53valve. A Hewlett-Packard 1-10-100 ml soap bubble flow meter was used for this experiment.

At the 600 RPM 64 kPa MAP load condition, the leakage was 1.11 ml/sec. At 600 RPM 46

kPa MAP, the leakage was reduced to 0.57 ml/sec. Finally, at 1200 RPM 50 kPa MAP, the

leakage rate was 1.03 ml/sec. Given these values and a worst-case sample flow rate of 1.5

lpm, the highest possible leakage gas concentration was 4.4%. Since most conditions were

assumed to be below this value, the sealing performance of the sampling poppet valve was

deemed acceptable.

3.2.3. Emissions Bench

A five-gas Horiba emissions analyzer was used to measure steady-state exhaust

species concentration sampled from the engine exhaust mixing tank. After exiting the

mixing tank, the sample was transported to the emissions bench by an electrically heated line.

The line was temperature-controlled to 190° C to avoid hydrocarbon and water condensation.

Adequate flow is provided by a vacuum pump in the bench and a regulated manifold tree to

the individual analyzers. Before entering the infrared analyzers, water was condensed from

the stream in a 0° C refrigerant bath.

The five analyzers were CO, CO2, O2, HC and NOx, each paired to a signal

amplifier. The CO and CO2 analyzers were calibrated through a Stec gas divider to a

second-order polynomial fit of voltage vs. concentration. The remaining three analyzers

were linear in operation and required only two-point calibrations. Table 3.4 summarizes the

Horiba emissions bench hardware.

54Gas Span Level Horiba

Analyzer # Analyzer Type Horiba Amplifier

# CO2 10.1 % AIA-23 ND-Infrared OPE-135 CO 2.56 % AIA-23 ND-Infrared OPE-115 O2 1.01 % MPA-21A Paramagnetic OPE-335 NOx 101 ppm CLA-22A Chemiluminescent CLA-22A HC 6286 ppm

(C3H8) FIA-23A Flame Ionization Detector

(FID) FIA-23A

Table 3.4. Horiba exhaust emissions analyzer bench summary.

All amplifier output signals were passed through an A/D converter card and logged

on a PC using LabView 6.0. LabView was used to automatically perform the voltage

calibration and average multiple samples for final data reporting.

NOx measurement during this experiment was precluded by the inability of the

analyzer to achieve a steady-state during the short firing duration of the optical engine. The

engine could not be fired continuously for more than five minutes, where NOx readings were

still increasing for all conditions. Therefore, NOx will not be reported in the results section.

To rapidly switch from exhaust emissions measurement to the sampling valve stream

during the experiment, the front side calibration port of the CO2 analyzer was used to receive

the line from the ice bath. In this arrangement, the low-flow sampling valve stream was

supplied exclusively to the CO2 analyzer, which was the only measurement needed. Before

and after sampling valve measurement runs, the valve plumbing system was purged with

nitrogen.

55

3.3. Optical Measurement System

The mixing of the residual gas with the fresh homogeneous fuel/air charge in the

combustion chamber of the engine was performed using planar laser-induced fluorescence

(PLIF). The measurement system consisted of a laser source, beam-transport and sheet-

forming optics and a camera. The optical system was synchronized with the engine

crankshaft to capture PLIF images at several crank angles during the compression stroke.

3.3.1. Laser Source

The laser used in this project was the Spectra-Physics GCR-170 Nd:YAG. The

fundamental 1064 nm output was frequency quadrupled to 266 nm, providing a pulse

duration of 4-5 ns and a peak energy of 90 mJ/pulse. Pulse-to-pulse energy stability is listed

as <10%. The laser was externally triggered at its design repetition rate of 10 Hz, with an

optimized delay between flash lamp and q-switch trigger signals of 186 µs.

Laser power, which could only be measured safely before the last two optical

elements ~35 cm upstream of the engine entrance, was monitored with a ScienTech Mentor

MD-10 power meter with a UV-sensitive power head. Laser energy at this point was

adjusted to 30 mJ/pulse +/- 3 mJ, although this value is an integrated time average and can

not be resolved to a pulse-by-pulse basis. The Nd:YAG laser was able to supply more

power, but damage to the engine quartz windows near the focused laser sheet prevented

higher powers.

563.3.2. Laser Optics

The 266 nm output of the Nd:YAG laser was separated from the higher harmonics by

a Pellin-Broca prism located at the exit of the laser. The visible and infrared beams were

captured by beam dumps. The 266nm beam was again turned 90° by a dichroic mirror to

traverse the length of the laser table to a second Pellin-Broca Prism for final wavelength

separation.

A dichroic mirror directed the 266 nm beam into a 1 m focal length spherical lens,

designed to focus the laser at the center of the engine bore. After the spherical lens, the beam

was vertically traversed by two right-angle prisms from laser output height to engine window

height. The traverse distance is ~10 cm and was adjusted by a micrometer translation stage.

A 100 mm positive focal length cylindrical lens was located 60 cm downstream of the

spherical lens to develop the laser sheet. Finally, a 2 in diameter dichroic mirror was used to

direct the laser sheet into the engine’s through the quartz windows in the space ring.

The optical system is presented in Figures 3.4 and 3.5 for clarity.

57

Figure 3.4. 266 nm laser pulse separation and delivery optics (plan view).

Figure 3.5. Laser sheet-forming optics setup for 266 nm PLIF imaging.

583.3.3. Camera

The primary camera used for PLIF imaging in the Triptane engine was the Roper

Scientific MicroMax. This camera lies in a category of scientific CCD cameras known as the

“frame-straddling” type and is specified with a nominal quantum efficiency of 45% at the

peak fluorescence wavelength range from Figure 2.7. The MicroMax CCD is front-side

illuminated and cooled thermoelectrically to -20°C. A twin-blade fast mechanical shutter is

used to protect the device from combustion luminosity. The CCD array measures 1300x1030

pixels with a pixel size of 6.7 µm. In this experiment, the camera was binned on-chip 6-by-6

to increase PLIF signal and shorten read-out time. The binning selection is discussed in

Section 4.2.2. The readout rate for the device is 5 MHz, with 12-bit digitization. The

MicroMax camera is not intensified.

The principal design feature of the MicroMax is a “Double-Image Feature” (DIF)

mode designed for particle-image velocimetry (PIV), which allows two separate exposures to

be captured on the CCD array in rapid succession without mechanical shuttering. This is

accomplished by “interline transfer” on-chip, which is the reason why this camera was used

in the project. In DIF mode, the mechanical shutter is pre-opened once the camera is finished

reading the previous image and the chip is actively drained of charge before the exposure

trigger arrives. The CCD is divided into alternating columns of masked and unmasked pixels

in this mode and the unmasked pixels are charged during the first exposure for a time

programmed in software as short as 1 µs. The camera then performs the interline transfer of

the charged pixels over to the masked pixels, which requires 200 ns, or some longer

programmed duration. Then, a second exposure is taken on the unmasked pixel columns.

59The second exposure has no place to be shifted to, so it must be exposed until the mechanical

shutter is closed for read-out. A diagram of DIF-mode operation is shown in Figure 3.6. The

basis for selecting the MicroMax over other high-quality CCD cameras is discussed in

Section 5.2.

In this project, the exposure time was set to 10 µs, with the second image from each

engine cycle deleted from the analysis. When pixels were binned 6-by-6, the resulting read-

out time was 0.3 sec, easily fast enough to keep up with our engine skip-firing frequency.

MicroMax images were saved on a Pentium III Windows PC operating Roper Scientific

WinView/32 v. 2.4.8 as multiple-frame 12-bit grayscale TIFF files.

The lens selected for the primary camera was an 85 mm Nikkor f/1.4 model mounted

on a 20 mm Kenko extension tube in addition to a C-mount to F-mount adapter. The

selection criteria for this lens are detailed in [15].

Figure 3.6 MicroMax camera manual summary of DIF-mode timing. Image exposure times are shown in the second line. Ready and Scan are output signals from the camera controller, Ext. Sync is the input trigger TTL, Laser Output shown is for a double-pulse laser, this experiment only uses the first pulse. Reprinted from [24].

603.3.4. Optical Triggering

The requirements of the optical triggering system were threefold: 1) To supply the

Nd:YAG laser with an uninterrupted 2-pulse trigger sequence at 10 Hz; 2) to ensure that the

camera triggers were delivered on skip-fired cycles only; and 3) to gate both the primary and

reference camera with sub-microsecond resolution to capture the 5 ns laser pulse. To

perform this, the BEI crankshaft encoder and the MotoTron engine control were each utilized

with an interface at a special TTL logic circuit.

The “clock” signal for the laser and both cameras was provided by the shaft encoder,

via a TTL counter box supplied by Mercury Marine. The A-pulse (divided by 4 to one-per-

CAD frequency) and the Z-pulse (TDC) were input into the counter box. An advance or

delay value of 0-99 CAD before or after TDC could be selected with surface-mounted

control switches. One TTL pulse per engine revolution was output from the counter box at

the input advance/delay timing.

A Berkeley Nucleonics model 555 pulse/delay generator was used to supply the high-

precision trigger signals to the optical system. This device was operated in external gate

mode with an input clock signal from the counter box. At 600 RPM, the counter box

frequency (one per revolution) was already at 10 Hz, while at 1200 RPM, the pulse/delay

generator had to be operated in divide-by-2 mode, where synchronization with the

compression stroke had to be verified. The outputs of the pulse/delay generator were used to

provide the signals summarized in Table 3.5, with delays relative to the leading edge of the

counter box pulse:

61Channel Device Width Delay

A Flash Lamp 5.000 ms 0.000 µs

B Q-Switch 5.000 ms 186.0 µs

C MicroMax (primary camera) 5.000 ms 180.8 µs

D PI-Max (alternate camera) 5.000 ms 180.0 µs

Table 3.5 Trigger timing delays for optical measurement system. Delays are relative to the leading edge of the trigger signal from the crankshaft encoder.

The laser trigger pulses were delivered to the Nd:YAG directly to provide the

uninterrupted 10 Hz operation. The camera triggers were sent to two “one and only one”

TTL logic circuits (chip #4013), where they provided the “clock” input to the circuit diagram

shown in Figure 3.7. The MotoTron “skip-fire” TTL signal, previously used for activating

the sampling valve, was supplied as the “enable” input, at a timing of -180 CAD. On receipt

of the enabling signal, the circuit outputs the next clock signal and only that one pulse. This

allows the camera to capture the laser sheet that is fired on the compression stroke of the

skip-fired engine cycle and prevents the MicroMax camera from being triggered during a

combustion cycle. With the camera exposure set to 10 µs, the circuit’s insertion loss

(predicted to be ~10-20 ns) does not affect the capturing of the laser pulse.

62

Figure 3.7 Schematic for TTL timing of laser pulse and camera, synchronized with MotoTron skip-firing ignition by a “one-and-only-one” circuit.

63

4. Engine Operating Conditions

4.1. Selection Criteria

Utilizing the adjustable-camshaft feature of the cylinder head, conditions for varying

levels of residual gas dilution were established at our two operating speeds. For this project,

five categories of cam-phasing strategies were covered: a “baseline” valve overlap, an

enlarged valve overlap symmetric about TDC exhaust, an enlarged valve overlap with the

intake cam advanced from the baseline, an enlarged valve overlap with the exhaust cam

retarded from the baseline, and finally a zero-overlap (IVC=EVO) setting. At each of these

five categories, three engine conditions were established: a “low” load at 600 RPM, a “mid”

load at 600 RPM, and a “low” load at 1200 RPM. The result was a test matrix of 15 distinct

engine operating conditions.

4.1.1. Optical Engine Considerations

The objective of this project was to study conditions of high residual gas fraction.

Severe limitations in operating a fired optically accessible engine made this objective

challenging. The low compression ratio (5.95:1), despite being conducive to increased

trapped residual gas levels, impacts the ignition and combustion stability of the engine. To

address this, fuel was delivered at consistently rich conditions to aid the spark ignition (with

the added benefit of increased tracer density for optical measurement). Additionally, the

64spark plug was gapped to 2.1mm and the ignition coil was permanently set on maximum

dwell. The second major limitation was from thermal loading of the oil-less ring pack for the

Bowditch piston. Despite the relatively cold block temperature (§ 3.1.6), the Triptane engine

could not be fired continuously for more than four minutes at the 1200 RPM and 600 RPM

“mid” load conditions or six minutes for the 600 RPM “low” load conditions.

Mass loss through the non-metallic ring pack and the temporally evolving nature of

the mass loss under firing operation introduce uncertainty into cylinder pressure analyses (§

4.2). Secondly, time-limited firing operation effects steady-state exhaust gas emissions

measurement uncertainty (§ 4.3). A final mention before proceeding must be given to the

consideration of engine speed, bulk flowfield and manifold wave dynamics when extending

optical engine combustion data to conventional SI engines.

4.1.2. Establishing Engine Conditions

A baseline valve overlap duration of 20° was selected, based on a presumed-typical

value for a 4-valve engine of 510 cc displacement operating at 600-1200 RPM. For

simplicity, this overlap duration was positioned symmetrically about TDC exhaust. Intake

air mass flow rates were established at the baseline valve overlap for the three engine

speed/load combinations, and were held consistent for the other overlap cases. In

establishing the baseline air flow rates, fuel delivery and spark timing were both freely

adjusted to optimize IMEP and COV of IMEP.

The 600 RPM “low” load condition was set by varying the air mass flow rate to find a

comfortable minimum for stable combustion operation. Based on the measurement

65requirements of both the optical and sampling valve techniques, a combustion stability

criteria of <10% COV of IMEP was established. Since the objective was to increase dilution

with increased valve overlap from this baseline condition, an absolute minimum air delivery

rate was not chosen. The air mass flow rate chosen for the five 600 RPM “low” load

conditions, 144 mg/cycle, resulted in an intake MAP of approximately 50 kPa at the baseline

overlap. The 1200 RPM “low” load condition was established by adjusting the intake mass

flow rate to provide the same MAP as 600 RPM “low” load at the baseline overlap. The

result was 181 mg/cycle, although IMEP and exhaust temperatures were significantly higher

at the increased speed.

The 600 RPM “mid” load point was established by increasing the intake mass flow

rate to a comfortable upper limit for safe engine operation. At 208 mg/cycle, peak cylinder

pressure was at 15 bar and steady exhaust temperature was at 400 C, both acceptably close to

the upper limit. Intake MAP increased from the low load condition to 61 kPa. A “mid” load

point at 1200 RPM could not be established due the safety limits in combustion pressures and

temperatures. Likewise, full-load atmospheric-MAP firing operation could not be performed

at either speed, so no quantitative pressure-based reference for load points could be

established. Therefore, only intake air mass flow rate is used as a basis for engine load.

With air mass flow rates fixed, the advanced valve overlap conditions could be

established. Both fuel mass and spark timing were varied and 10% COV was used an upper

limit for combustion stability. Since the engine was less tolerant of increased overlap at 600

RPM compared to 1200 RPM, different overlap levels were tested for the two speeds. At all

600 RPM increased-overlap conditions (both loads), a total valve overlap duration of 30° was

set. This 10° increase from baseline was established from the maximum amount tolerable at

66the low-load condition at either intake-advance, exhaust-retard or symmetric-increase. At

1200 RPM, a 40° increase in overlap from baseline (60° total) was achieved for all three cam

strategies.

-50 0 500

2

4

6

8

10

12

valv

e lif

t [m

m]

Symmetric Overlap Increase

-50 0 500

2

4

6

8

10

12Intake Cam Advance

-50 0 500

2

4

6

8

10

12

valv

e lif

t [m

m]

CAD aTDC

Exhaust Cam Retard

-50 0 500

2

4

6

8

10

12

CAD aTDC

Zero Valve Overlap

Baseline 600 RPM 1200 RPM

Exh.

Exh.

Exh.

Exh.

Int.

Int. Int.

Int.

Figure 4.1 Summary of four valve overlap strategies. Baseline cam timing is indicated by the dashed line in all plots. Arrows indicate cam shift from baseline. The baseline overlap duration is 20°, the 600 RPM extended overlaps are 30° duration, and the 1200 RPM conditions are 60° overlap duration.

Fuel delivery rate was set by the highest injection mass required in the increased

overlap conditions for each of the three speed/load combinations. Fuel mass flow was then

fixed for all overlap conditions at the particular speed/load, to provide consistency in the

67optical measurements. A summary of the targeted air/fuel delivery rates for each of the

speed/load combinations is shown in Table 4.1. Spark timing was not fixed in this

experiment, but was optimized at each of the 15 test conditions to provide consistent

combustion phasing.

Mass Air Flow Fuel Injection Targeted AFR 600 RPM, Low Load 144 mg/cycle 14.5 mg/cycle 9.93:1 600 RPM, Mid Load 208 mg/cycle 18.0 mg/cycle 11.56:1

1200 RPM, Low Load 181 mg/cycle 18.0 mg/cycle 10.06:1

Table 4.1. Air/fuel engine operation parameters for the three experimental speed/load points. These values were held constant for each cam strategy.

A complete summary of engine operating conditions recorded in the laboratory can be

found in the master summary in Appendix A.1.

4.2. Combustion Analysis

Cycle-resolved cylinder pressure data were recorded throughout this experiment

using the system discussed previously. Ensemble-averaged indicated mean effective

pressure (IMEP) and pumping mean effective pressure (PMEP) were directly calculated by

the post-processing program along with COV of IMEP, which was used as a basis for

combustion stability. Averaged pressure traces were also used to compute heat release

information for each of the 15 test conditions.

684.2.1. Cylinder Pressure Data

Table 4.2 contains a summary of the IMEP and PMEP data for the test conditions.

The data are organized by valve overlap strategy and include a percent change relative to the

baseline overlap condition for the particular speed/load point. As previously discussed,

combustion phasing was kept consistent (location of peak pressure ~ 15° aTDC) for all

conditions.

Cam Strategy Speed [RPM] /Load

IMEP [kPa] (gross)

COV of IMEP [%] PMEP [kPa]

600 low 147.3 () 4.1 () 46.2 () 600 mid 260.9 () 2.05 () 35.3 () Baseline

Overlap 1200 low 228.2 () 1.32 () 51.1 () 600 low 151.7 (+ 3%) 6.03 (+ 47%) 45.0 (- 3%) 600 mid 270.7 (+ 4%) 2.07 (+ 1%) 30.3 (- 14%) Symmetric

Increase 1200 low 253.4 (+ 11%) 1.2 (- 10%) 35.3 (- 31%) 600 low 156.4 (+ 6%) 3.89 (- 5%) 45.4 (- 2%) 600 mid 269.8 (+ 3%) 2.46 (+ 17%) 31.1 (- 12%) Intake Advance 1200 low 260.6 (+ 14%) 1.84 (+ 40%) 32.2 (- 37%) 600 low 157.8 (+ 7%) 3.77 (- 8%) 43.5 (- 6%) 600 mid 271.5 (+ 4%) 2.64 (+ 20%) 29.1 (- 18%) Exhaust Retard 1200 low 240.7 (+ 6%) 3.64 (+ 176%) 35.1 (- 31%) 600 low 141.6 (- 4%) 2.99 (- 27%) 51.5 (+ 12%) 600 mid 255.1 (- 2%) 1.2 (- 41%) 39.2 (+ 11%) Zero Overlap 1200 low 229.3 (+ 0%) 0.87 (- 50%) 53.3 (+ 4%)

Table 4.2 Mean effective pressure data for 100-cycle average pressure data at all experimental conditions. Percentages shown are changes relative to the baseline overlap condition for the individual speed/load points at each cam strategy.

From the trends shown in Table 4.2, it can be seen that the baseline overlap selected

was not an ideal setting for any of the three speed/load points. The 600 RPM low load

condition was least sensitive to the cam-phasing changes, but nevertheless showed a

69consistent increase in IMEP with extended overlaps. The 600 RPM mid load points enjoyed

larger volumetric efficiency improvements (12 to 18 %) with the increased overlap durations

with minor increases in IMEP, indicating that the pumping improvements were nearly

overshadowed by losses due to charge dilution and reduced compression/expansion times.

The 1200 RPM points show the most significant reductions (> 30%) in pumping work, and

also the largest gains in IMEP, with the notable exception of the exhaust cam retard case.

The zero-overlap condition predictably reversed the trend due to penalties in pumping work.

4.2.2. Heat Release Analysis

A single-zone heat release code was used to process ensemble-averaged cylinder

pressure data. The code, which runs in the EES equation solver, uses an iterative

optimization scheme within the numerical integration to obtain wall heat transfer

characteristics. Combustion efficiency was calculated at each condition, using exhaust gas

measurements (§ 4.3.2), with a resulting range from 72% to 90%. The limits of integration

were optimized for each of the 15 engine conditions to obtain appropriate cumulative heat

release traces. The graphical results of the analysis are presented in Figures 4.2-4.4,

organized by speed/load point.

70

-30 0 30 60 90 120 1500

0.002

0.004

0.006

0.008

0.01

0.012

0.014

0.016

CAD aTDC

Hea

t Rel

ease

Rat

e [k

J/de

g]

BaselineSymmetric Incr.Zero OverlapIntake AdvanceExhaust Retard

600 RPMLow Load

-30 0 30 60 90 120 150 1800

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

CAD aTDC

Cum

ulat

ive

Hea

t Rel

ease

BaselineSymmetric Incr.Zero OverlapIntake AdvanceExhaust Retard

Figure 4.2 Heat release rate and cumulative heat release for all cam strategies at 600 RPM Low Load.

71

-30 0 30 60 90 120 1500

0.005

0.01

0.015

0.02

0.025

0.03

CAD aTDC

Hea

t Rel

ease

Rat

e [k

J/de

g]

BaselineSymmetric Incr.Zero OverlapIntake AdvanceExhaust Retard

600 RPMMid Load

-30 0 30 60 90 120 150 1800

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

CAD aTDC

Cum

ulat

ive

Hea

t Rel

ease

BaselineSymmetric Incr.Zero OverlapIntake AdvanceExhaust Retard

Figure 4.3 Heat release rate and cumulative heat release for all cam strategies at 600 RPM Mid Load.

72

-30 0 30 60 90 120 150 1800

0.005

0.01

0.015

0.02

0.025

CAD aTDC

Hea

t Rel

ease

Rat

e [k

J/de

g]

BaselineSymmetric Incr.Zero OverlapIntake AdvanceExhaust Retard

-30 0 30 60 90 120 150 1800

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

CAD aTDC

Cum

ulat

ive

Hea

t Rel

ease

BaselineSymmetric Incr.Zero OverlapIntake AdvanceExhaust Retard

Figure 4.4 Heat release rate and cumulative heat release for all cam strategies at 1200 RPM Mid Load.

73The most immediate trend from the plot sequence is the long tail of the heat release

curves for this engine, which extend far into the expansion stroke and approaches the EVO

timing. Even the baseline and zero overlap conditions demonstrate protracted burn durations,

demonstrating the influence of the low compression ratio and inherently high trapped

residual mass of the engine.

Table 4.3 summarizes the flame development duration and overall burning duration,

defined as crank angle 0-10% cumulative heat release and 10-90% heat release [1],

respectively. Once again, trends relative to the baseline 20° overlap are included at each

condition.

Cam Strategy Speed [RPM] /Load

Flame Development Angle

(0-10% HR)

Overall Burning Angle (10-90% HR)

600 low 53° () 59° () 600 mid 36° () 60° () Baseline Overlap 1200 low 36° () 68° () 600 low 61° (+ 15%) 79° (+ 34%) 600 mid 41° (+ 14%) 67° (+ 12%) Symmetric

Increase 1200 low 57° (+ 59%) 87° (+ 28%) 600 low 61° (+ 14%) 75° (+ 27%) 600 mid 43° (+ 20%) 67° (+ 12%) Intake Advance 1200 low 57° (+ 59%) 84° (+ 24%) 600 low 61° (+ 14%) 78° (+ 32%) 600 mid 42° (+ 17%) 70° (+ 17%) Exhaust Retard 1200 low 60° (+ 68%) 93° (+ 37%) 600 low 40° (- 25%) 60° (+ 2%) 600 mid 27° (- 27%) 59° (- 2%) Zero Overlap 1200 low 33° (- 7%) 72° (+ 6%)

Table 4.3 Flame development angles and overall burning angles for different overlap strategies, determined by a single-zone heat release code. Percentages indicated are changes relative to the baseline overlap condition at each speed/load point.

74The extended valve overlaps have a more pronounced effect on the heat release data

than the IMEP data of Table 4.2. At 600 RPM low load, the 10° increases in overlap

duration show up clearly in the 10-90% burn angle, with a consistent 25% increase, and a

smaller influence on the flame development angle. The overall increase in the burn duration

was lowest for the elevated load (600 mid) conditions. At 1200 RPM, where the largest

valve overlap extensions occurred and the highest residual fractions were anticipated, the

effect on early flame development was most severe. Since the baseline overlap at 1200 RPM

was predicted to be one of the lowest-residual conditions, this result was not surprising.

Higher turbulence levels at 1200 RPM help explain the lessened impact on 10-90% burn

duration at the high-overlap conditions. Once again, the 1200 RPM exhaust cam retard

condition is an outlier, showing a larger impact than either symmetric increase or intake

advance.

4.3. Exhaust Gas Emissions Measurement

Downstream exhaust gas emissions were recorded for the 15 established test

conditions as part of the residual fraction measurement described in Section 4.4. CO2

readings were critical for those measurements, but the additional CO, HC and O2 readings

are used here to better-quantify air/fuel ratio and combustion efficiency.

754.3.1. Emissions Measurement Procedure

Downstream emissions were recorded during continuously-fired operation with the

sample line feeding all five gas analyzers. Readings were acquired at 30 second intervals

during the time allowed for engine firing – four minutes at 600 RPM mid load and 1200

RPM, six minutes at 600 RPM low load. At that time, the engine had to be stopped (1200

RPM) or motored (600 RPM) for several minutes to avoid damage to the piston rings. The

readings were then graphically reviewed to determine the region of steady-state behavior,

and the corresponding measurements in that region were averaged to yield the reported

measurement value.

The NOx analyzer was not able to achieve steady-state in any of the measured

conditions. Rather than report suspect values, NOx measurements will not be included in the

project. With the highly fuel-rich operating conditions and low exhaust temperatures,

readings were predicted to be low (<< 1000 ppm) and should be negligible in combustion

calculations.

4.3.2. Emissions Analysis

Since all analyzers were recorded as “dry” measurements from the exhaust sample

line, a dry-to-wet correction factor had to be applied to raw the emissions bench data. This

value, Kexh, is defined as:

76

( )2 ,

exhexh

exh H O exh

nKn n

=+

(4.1)

where exhn and 2 ,H O exhn are derived from a carbon balance and the fuel’s combustion

stoichiometry (see §3.1.8). The concentration of H2 gas in the rich combustion products was

determined by the equation:

[ ]( )

[ ]2

:

4fuel

dry dry

H CH CO= × (4.2)

Knowledge of the incomplete combustion species, the water concentration and the

dry-to-wet correction factor allows for computation of the air/fuel ratio by way of a chemical

balance. Combustion efficiency was determined by the wet-basis mole-fraction equation

[25]:

( )[ ] [ ] [ ]{ }

[ ] [ ] [ ]

, 298 , 298 , 29822

, 298

2

% 100 (4.3)

100 (4.4)

CO T K H T K HC T K

cfuel T K

N CO h H h HC h

h

NHC CO CO

η= = =

=

+ += −

≡+ +

The molar enthalpy of combustion for the unburned HC is assumed to be equal to that of the

fuel. All concentration measurements are to be used on a “wet” basis.

774.3.3. Emissions Measurements

Steady-state downstream exhaust emissions measurements are summarized in Table

4.4. Analyzer readings shown are on a “wet” basis, after correction.

Cam Strategy

Speed [RPM] /Load [CO2] [CO] [O2] HC

[ppm C1] AFR ηc

600 low 8.1 6.5 0.4 10,629 10.7 82 % 600 mid 10.6 4.8 0.6 5769 12.1 90 % Baseline

Overlap 1200 low 10.1 5.9 0.6 6159 11.5 89 % 600 low 8.1 6.3 1.1 18,822 10.5 72 % 600 mid 10.7 4.6 0.7 5838 12.2 90 % Symmetric

Increase 1200 low 8.9 6.0 0.8 9015 11.3 85 % 600 low 8.4 6.3 0.9 14,463 10.8 77 % 600 mid 11.3 3.9 0.6 5697 12.5 90 % Intake

Advance 1200 low 10.3 5.0 0.7 6879 12.0 89 % 600 low 8.2 6.4 0.9 15,483 10.7 76 % 600 mid 11.1 4.3 0.5 5766 12.3 90 % Exhaust

Retard 1200 low 7.9 6.7 0.8 12,666 10.7 79 % 600 low 7.9 6.6 0.9 10,551 10.9 82 % 600 mid 10.5 4.8 0.7 5253 12.2 91 % Zero

Overlap 1200 low 8.4 6.4 0.8 6858 11.2 88 %

Table 4.4 Summary of exhaust emissions species measurements, concentrations shown are corrected to a wet basis from the raw readings. Air/fuel ratio and combustion efficiency coefficient have been calculated from the concentration data.

The combination of fuel-rich stoichiometry and the abnormally large crevice volume

of the Bowditch piston contributed to the very high hydrocarbon readings shown in Table

4.4. Pressure data summarized in Section 4.2.1 were taken simultaneous with these emission

measurements and did not indicate a single misfire (less than 1 bar IMEP) for any of the 15

78conditions. When the heat release data of section 4.2.2 are considered, it seems likely that

incomplete combustion at EVO was a major culprit in HC concentration, particularly at the

two low-load conditions.

It is notable that the intake cam advance strategy demonstrated the lowest CO

readings in the experiment and the lowest HC readings for extended overlap at all three

speed/load points. The 600 RPM mid load condition was least sensitive to overlap strategy

in terms of pollutant emissions. Combustion at the 600 RPM low load condition fared the

poorest at the symmetric overlap increase condition, with comparable results found at the two

asymmetric overlap extensions. At 1200 RPM, the worst-performing strategy was once

again the exhaust retard case, consistent with the pressure and heat release data.

The O2 readings were higher than would be anticipated at the equivalence ratio

indicated by our recorded CO concentrations. Since the highest O2 measurements were

consistently found at 600 RPM low load (~ 1%), where IMEP was lowest, it is proposed that

the high readings are a consequence of thermally-dependent sealing inefficiencies in the ring

pack during cylinder scavenging. The exhaust system was kept under positive back pressure

(above ambient) throughout the experiment and was newly constructed and leak-tested.

Furthermore, the emissions vacuum sampling line was tested with pure N2 gas with a

successful zero reading on the O2 analyzer.

Air/fuel ratio calculations were consistent across both of the 600 RPM load points. A

significant variation was encountered at the 1200 RPM condition, which is believed to be a

consequence of difficulty reading the time-limited measurement. 1200 RPM were the

hardest-running conditions in the experiment, with the high piston speeds and gas

temperatures. All analyzer-derived AFR readings were higher than the targeted delivery

79rates, indicating some error in the fuel injector and/or intake air orifice calibrations

(discussed in § 3.1.7-8).

4.4. Bulk Residual Gas Fraction Measurement

For this project, the primary basis for comparing optically measured residual gas

mixing phenomena is the bulk in-cylinder residual gas mass fraction (yr). To measure this

value, the fast-acting sampling valve was installed in the Triptane engine to measure in-

cylinder CO2 as described in Section 3.2.2.

4.4.1. Sampling Valve Measurement Technique

The sampling valve experiment was performed simultaneously with the exhaust

emissions measurements described in the previous section. In-cylinder CO2 readings were

taken under skip-fired operation at each condition immediately prior to switching to

continuous-firing for the downstream sampling and pressure data logging. The skip-fired

sequence is graphically presented in the pressure trace of Figure 4.5. The settings for the

sampling valve and the sampled gas flow rates are summarized in Table 4.5.

80

Figure 4.5 Skip-firing sequence example (1200 RPM baseline overlap shown). Sampling valve is actuated on compression stroke of skip-fired cycle (see Table 4.5).

600 RPM 1200 RPM Sampling Valve Open 56° bTDC 75° bTDC Valve Open Duration 17 ms 14 ms Sampling Valve Close 5° aTDC 25° aTDC Sampling Frequency 4 cycles 6 cycles

Table 4.5 Sampling valve operation for all experimental conditions. Sampling frequency is listed as the number of fired cycles between sampled cycles (see Figure 4.5).

Like the downstream emissions measurements, the skip-fired in-cylinder readings

were limited by the continuous-firing time limits of the engine. In a similar procedure, the

81steady-state reading of the CO2 analyzer was determined graphically. Measurements taken

at the two “low load” condition sets at the smaller baseline and zero overlaps demonstrated a

weaker flow to the bench, which required fully opening the flow control valve on the

analyzer entrance, resulting in flow pulsations to the instrument. All other conditions were

able to operate with a non-pulsating flow delivered to the bench at the analyzer’s optimal

flow rate of 2.5 lpm.

-100 0 100 200 300 400 5000

1

2

3

4

5

6

7

8

9

10

CAD aTDC

P [k

Pa]

Fired CycleSampled CycleValve Lift Transducer Signal

Figure 4.6 Sample pressure data for skip-fired cycle with sampling valve actuation. The average fired cycle pressure trace and the sampling valve lift transducer signal for that skip-fired cycle (no physical units) are overlayed. 1200 RPM exhaust cam retard condition shown.

82A concern with measuring in-cylinder composition with this technique is that the

mixture trapped in-cylinder on the skip-fired cycle be representative of the bulk cylinder

charge composition. Particularly concerning was the possibility of the influence of partial-

burn or misfire on the cycle prior to the sampling valve open event. Even at high dilution

conditions, this was not found to be a major problem, as can be seen for the histograms of

prior-cycle IMEP shown in Figures 4.7 and 4.8, which are the two conditions having the

slowest burning rate for the two engine speeds..

Figure 4.7 Frequency histogram of prior-cycle IMEP for skip-firing operation at 600 RPM low load symmetric overlap increase condition. Data compiled from 100 consecutive sampled cycles.

83

Figure 4.8 Frequency histogram of prior-cycle IMEP for skip-firing operation at 1200 RPM exhaust retard condition. Data compiled from 100 consecutive sampled cycles.

4.4.2. Residual Gas Fraction Calculations

The residual gas fraction was calculated on a molar basis by comparing the mole

fractions of CO2 in the compressed charge, xCO2,cc, with the downstream exhaust

measurement, xCO2,exh.

2

2

,

,

CO ccr

CO exh

xx

x= (4.5)

84The dry-to-wet factor for the exhaust, Kexh, was determined in Equation 4.1 and used to

convert the denominator term from its raw dry-basis analyzer reading. The dry-to-wet factor

is different in the compressed charge, and is defined in [26] to be:

11

exh rcc

r

K xKx

+=

+ (4.6)

Equations 4.1, 4.5, 4.6 and the raw CO2 analyzer readings can be used to iteratively solve for

xr. It is assumed that the molecular weights of the compressed charge mixture and the

exhaust gas mixture are equal, and that the mole fraction xr is then equal to the mass fraction

yr. Additionally, the small relative humidity in the dried intake air is neglected.

4.4.3. Residual Gas Fraction Measurements

The results of the bulk residual gas fraction study are presented in Table 4.6

organized by cam phasing strategy. As in previous sections, the percent change from the

baseline is provided for ry . Values for 2 ,CO ccx and ccK are in the master conditions summary

in Appendix A.1.

The intake cam advance strategy is shown to yield the smallest increases in residual

fraction. For the two 600 RPM loads, the symmetric overlap extension provided the largest

residual gas fraction increase. Since the baseline overlap condition at 600 RPM low load was

already at 37 % residual fraction, it is not surprising that the engine was not especially

tolerant of increased valve overlap durations there. The largest residual gas fraction was

85measured at 1200 RPM with the exhaust cam retarded, which is consistent with the analysis

of the pressure, heat release, and emissions measurements. In general, the range of residual

fractions covered (21.9% - 44.8%) is believed to be representative of those encountered in a

high-dilution engine design. It is not a major concern for this project that the lower end of

the measured ry range does not correlate well with typical values in conventional SI engines

(5% - 25%).

Cam Strategy Speed [RPM] /Load yr

600 low 0.377 () 600 mid 0.296 () Baseline Overlap 1200 low 0.273 () 600 low 0.404 (+ 7%) 600 mid 0.358 (+ 21%) Symmetric

Increase 1200 low 0.437 (+ 60%) 600 low 0.387 (+ 3%) 600 mid 0.325 (+ 10%) Intake Advance 1200 low 0.408 (+ 50%) 600 low 0.399 (+ 6%) 600 mid 0.327 (+ 10%) Exhaust Retard 1200 low 0.448 (+ 64%) 600 low 0.287 (- 24%) 600 mid 0.245 (- 17%) Zero Overlap 1200 low 0.219 (- 20%)

Table 4.6 Summary of bulk residual gas fraction measurements at all experimental conditions. Percentages shown are changes relative to the baseline overlap condition at each individual speed/load point.

86

5. Imaging System Development and Analysis

5.1. PLIF Image Processing

Before discussing the selection of the camera for the experiment, an overview of the

procedure for correcting planar laser images in engines will be covered. First the image

acquisition sequence at each measurement condition will be described, followed by the

numerical corrections applied to the images to extract a faithful representation of the tracer

molecules in the laser sheet. Finally, the statistical processes used to evaluate and quantify

the inhomogeneity of the fresh charge in the engine are discussed.

5.1.1. Image Acquisition Procedure

Each PLIF measurement consists of three series of TIFF-format grayscale intensity

images – a background image sequence, a flatfield image sequence, and a data image

sequence. These images are all acquired at the same crank angle timing (engine

motoring/firing), with the camera focused on the laser sheet plane and the room lights turned

off.

Fifty background images are acquired before the data measurement and 50 more are

acquired afterward. The background images are taken with the laser sheet firing through the

combustion chamber, but without tracer addition (i.e. no fuel injection). Background images

contain signal from blemishes in the turning mirror and piston window, as well as from

87scattered laser sheet light impinging on cylinder head roof surfaces. These images are

subtracted from all subsequent images to better isolate the laser sheet. A sample 100-image

mean background is shown in Figure 5.1, where all four valves, the spark plug and the

periphery of the piston window are all visible in the full-CCD image.

Figure 5.1 Sample 100-image mean background image. Pixel intensity scale is on right.

Following the first set of 50 background images, 100 flatfield images are acquired

with the fuel injector activated and the ignition disabled. With the far-upstream fuel injection

(§ 3.1.8), this mode is considered to provide a very nearly homogeneous in-cylinder mixture.

Laser-induced fluorescence images of the homogeneous tracer distribution are used to

perform a “flatfield correction” of the data images, whereby the laser sheet intensity profile is

normalized. In this project, the flatfield image sets are additionally useful, since they can be

used to define the homogeneous image condition, when the mean flatfield correction is

applied to the 100 individual flatfield images. Since the flatfield images are acquired while

motoring, the residual gas contains fuel/tracer and these images can be used for comparison

88with the fired data images which contain residual gases. Provided the intake charge is

thoroughly pre-mixed (§ 5.3), the corrected homogeneous images provide the statistical

definition for a “completely mixed” cylinder charge, limited by the detection system signal-

to-noise characteristics. Figure 5.2 contains a mean flatfield image acquired 30° bTDC, and

demonstrates the spatial variation in laser sheet intensity.

Figure 5.2 100-image mean flatfield image, 30° bTDC 600 RPM Mid Load Exhaust Retard condition. Flatfield images have been background-subtracted.

Data images were acquired with the engine in skip-fired operation. Data images

contain fresh air/fuel/tracer charge mixing with combustion residuals. Like the homogeneous

images, data images are background subtracted and flatfield corrected. A sample raw data

image is shown in Figure 5.3. The correction technique will be developed in the following

section.

89

Figure 5.3 Sample raw data image (no corrections), 30° bTDC 1200 RPM Exhaust Retard condition.

5.1.2. Image Correction Procedure

Image processing in this experiment was performed using the Matlab programming

environment. First, 100 background images are read in and converted from a 12-bit integer

value to double precision (to avoid truncation of low signal in subsequent computations). An

ensemble mean background image,,i j

B , is computed by the formula:

( )( ) ( )

100

, ,, 1

th, ,

1 (5.1)100

B intensity of pixel , of n background image for all , in CCD array

i j ni j n

i j n

B B

i j i j=

=

=

∑

The 100 flatfield images are then read in and converted to double precision. An

ensemble mean flatfield image is computed by Equation 5.2:

90

( )( ) ( )

100

, ,, ,1

th, ,

1 (5.2)100

F intensity of pixel , of n flatfield image for all , in CCD array

i j ni j i jn

i j n

F F B

i j i j=

= −

=

∑

The mean maximum pixel value from the 100 flatfield images, maxF , is stored and used as a

scaling factor for a normalized mean flatfield image used for data image correction, 0 1 ,i jF − .

This image will be divided into the individual data images. In order to preserve the full range

of the original pixel count scale after division, 0 1 ,i jF − is created as a normalized zero-to-one

double precision array.

( )

,0 1 ,

max

(5.3)

for all , in mean flatfield image

i ji j

FF

F

i j

− =

To avoid amplifying noise-level pixels in the flatfield correction, pixels in 0 1 ,i jF − below

25% of the maximum value are forced to zero. Zero-value pixels are not included in the

flatfield correction.

The homogeneous images and data images are corrected by the normalized flatfield

after background subtraction (Equations 5.4 and 5.5, respectively).

91

( )

, , ,, ,

0 1 ,

,

(5.4)

for each of n=100 flatfield images

i j n i ji j n

i j

i j

F BH

F

F

−

−=

( )

, , ,, ,

0 1 ,

,

(5.5)

for each of n=100 engine raw data images

i j n i ji j n

i j

i j

Q BD

F

Q

−

−=

In this project, the data and homogenous image corrections are performed on a region of

interest completely located within the laser sheet. The coordinates of the ROI within the

CCD array vary based on the exact alignment of the laser sheet at each condition. The ROI

properties are discussed in Section 5.3.1.

5.1.3. Median Filtering

Homogeneous and data image sets ( , , , , and i j n i j nH D ) are transformed with the 3-by-3

median spatial filter built in to Matlab’s Image Processing Toolbox. This operation was

performed after all corrections from the previous section and before any statistical

calculations or output presentation. Median filtering has shown to be a valuable technique

for noise removal in images while preserving gradients.

925.1.4. Image Statistics

The two-dimensional spatial mean pixel intensity and standard deviation were

computed on the homogenous and data image ROI using built-in array operators in the Image

Processing Toolbox. Since the fluorescence intensity of the 3-pentanone tracer was not

calibrated to an absolute scale (due to shot-to-shot laser power variation and thermal

gradients in-cylinder), a coefficient of variation had to be defined to quantify inhomogeneity

in each image n ( nCOV ), relative to the mean signal.

( )( )

( )

( )

(5.6)

where = standard deviation of pixel intensities on ROI of image n

and = spatial mean pixel intensity on ROI of image n

I nn

n

I n

n

COVI

I

σ

σ

=

During the experiment, a significant error in the flatfield correction occurred,

imposing vertical bands on the corrected images (Figure 5.4). This phenomenon has been a

common occurrence in prior PLIF studies [13, 15, 16] and can been attributed to slight

temporal variation in the laser sheet profile during the measurement. The non-physical

structures introduce error into the two-dimensional statistics, so a modified “column COV”

was used to calculate the relative standard deviation only in the direction of the vertical

banding. This quantity, although technically a coefficient of variation, will be referred to

symbolically as ( )y yσ µ indicating statistics performed in the vertical direction across an

entire image ROI and then ensemble-averaged.

93

( ) ( )1

th

1 (5.7)

where and are the mean pixel intensity and standard deviation

for each pixel column j in the n image ROI ( columns wide)

Nj

y y nj j

j j

N I

I

J

σσ µ

σ

=

= ∑

If the laser sheet is aligned with the camera’s pixel array, the column COV term better

indicates variation in pixel intensity due to engine flows only. As a convention in the

nomenclature, over-bars will used to indicate spatial-mean values within an image and angle

brackets < > will used to denote ensemble-mean values.

Yr = 21.87%, <(σy / µy)> = 1.74%

1200 RPM Homogeneous, 30° bTDC

220

240

260

280

300

320

340

360

380

0.9 × Imax

0.5 × Imax

Figure 5.4 Sample homogeneous images acquired at 30° bTDC for the 1200 RPM, zero overlap condition demonstrating vertical banding in the corrected images. See Section 5.1.6 for image presentation convention.

945.1.5. Probability Distribution Function

The term ( )y yσ µ was developed in the previous section to quantify intensity

variation in individual data images, and also the mean stratification for a particular set of

images. Since this ensemble mean encompasses a large amount of pixel values, the

probability density function (PDF) is a relevant tool for presenting graphically the measure of

central tendency for a set of 100 images.

Pixel intensities were gathered from the regions of interest in both corrected data and

homogeneous images. A two-dimensional count of these data would be biased by the

vertical intensity banding shown in Figure 5.4. Therefore, as in the ( )y yσ µ calculations, a

vertical column-based analysis was developed. For each column j in each image ROI, the

mean intensity, jI , was computed and used to generate a “pseudo-image”, ,i jD′ , of the ROI

where each pixel was normalized by it’s column mean.

( )

( )

, ,1.. 1..

1

, for corrected data image (5.8)

1 , for each column in the ROI (5.9)

i j i jj i M j N

M

ji

D i jD D

I

I D i j jM

= =

=

′ =

= ∑

The relative intensities of ,i jD′ on the (M x N) region of interest are decoupled from

the flatfield correction errors and can be processed by two-dimensional statistics. Each

image n in the 100-image data set is analyzed by the image histogram function in the Matlab

Image Processing Toolbox. Since this function only operates on unsigned integers, the

95double-precision values of ,i jD′ are first multiplied by 10,000 counts to avoid truncation

before conversion to16-bit integer class. The histogram bin widths are set to 4 counts (total =

16,384 bins) and the bin counts from each image are summed for the data set to form the full

histogram.

The full histogram is integrated and then normalized by the integral value to form the

PDF curve. The bin scale is normalized to the mean intensity from the ,i jD′ pseudo-images.

At each condition where the PDF is examined, the homogenous image set is overlaid on the

skip-fired data for comparison.

5.1.6. Image Presentation

Figure 5.4 provides an example of the PLIF data image presentation format for this

project. Regions of interest from 25 successive corrected images are compiled into a tiled

array. The images are in successive left-to-right, top-to-bottom order from the 100-frame

raw data file, taken at the same image timing at the engine cycle imaging frequency

(typically one every 6th cycle). Figure 5.4 also includes an intensity scale at right. As

labeled, the upper and lower limits of this scale are the 90% and 50% levels, respectively, of

the maximum intensity found in the ROI at that imaging condition. These cutoff levels were

used for visual representation help to better illustrate regions of high and low fresh charge

concentration in the fired data images. In all future images this same grayscale criteria has

been employed.

96

5.2. Imaging System Performance

An objective of this project was to apply the highest-fidelity image-based

measurement system to residual gas mixing processes. The evaluation criteria for the camera

are signal-to-noise ratio and spatial resolution. The MicroMax CCD camera, described in

Section 3.3.3, was selected after comparison with other alternatives. The lens selection

resulted in a small, centrally-positioned laser sheet region of interest in the combustion

chamber. A test target was used to determine the spatial resolution of the system. Finally,

image data were analyzed to determine a maximum range of SNR.

5.2.1. Camera Selection

The MicroMax camera is categorized as a frame-straddling CCD, and was analyzed

in a comparative study on PLIF detection systems by Rothamer and Ghandhi in [14]. The

highest-performing device in that study (assuming a strong PLIF signal) was from the slow-

scan category of cameras. The Apogee AP7 CCD, which had produced SNR greater than

80:1 in prior non-firing direct-injection mixing studies [15, 16], was not capable of shuttering

residual background luminosity during the skip-fired compression stroke. The 16-bit CCD

saturated at the lowest shutter duration time (20 ms), and was eliminated from consideration

in the project. The MicroMax, with the electronic interline transfer on-chip shuttering

feature described in Chapter 3, was able to easily reject this luminosity at an exposure of 10

µs.

975.2.2. Region of Interest (ROI) and Spatial Resolution

As demonstrated in the background image in Figure 5.1, the full CCD array captures

a region of the combustion chamber located below the spark plug and slightly off-center

toward the exhaust valve-side of the pent-roof axis. Figure 5.5 more exactly locates the ROI

relative to the cylinder head and bore diameter, while Table 5.1 contains the vertical distance

separating the ROI from the piston face at the four experimental image timings.

Figure 5.5 Location of ROI within combustion chamber, DOHC cylinder head. Distance h is between laser sheet plane and piston face, and is tabulated for image timings in Table 5.1.

98Image Timing [CAD aTDC] h [mm]

- 30 6.4 - 45 13.7 - 60 22.9 - 99 49.0

Table 5.1 Distance from piston face to laser sheet ROI for experiment image timings.

The small pixel size (6.7 µm) of the 1300-by-1030 pixel MicroMax CCD required

binning the pixels 6-by-6 to obtain an equivalent pixel size in the imaging plane on the order

of 185 µm. This value was measured by Wiles [15] in a bench experiment to optimize the

photonic flux on the CCD array given our use of a large-aperture lens and also the fixed

dimensions of the lab’s Bowditch piston (~45cm camera-ROI distance). Bin size in this

experiment was determined prior to test target measurement using a less-precise scale image.

The equivalent pixel size, and thus the physical size of the ROI, was computed using

an image taken of the USAF 1951 optical test target. The target was back-illuminated and

placed in the laser sheet plane above the piston with the cylinder head removed. Using the

standard 4% contrast criterion, the image’s spatial resolution was measured to be 4.0 line

pairs per mm (lpmm). The equivalent pixel size was measured on the target to be 174 µm in

the imaging plane. The binned pixel size is comparable to that of the focused laser sheet

thickness [15] and therefore does not reduce system spatial resolution.

All regions of interest in this project were 163 pixels long in the direction parallel to

the laser sheet propagation. Condition-to-condition variation in laser alignment resulted in a

range of ROI widths (across sheet profile) of 103-130 pixels. Therefore, the ROI sizes

99recorded in Chapter 6 range from 18-by-28 mm to 23-by-28 mm in real size within the

cylinder.

5.2.3. Signal-to-Noise Ratio

Rothamer and Ghandhi [14] demonstrated that, in the presence of a sufficient

threshold signal level, the frame-straddling camera will operate in the shot noise-limited

regime. This is important, as it reduces the influence of other CCD noise sources (§ 2.4.3)

from the calculation of the maximum possible SNR. This condition is verified graphically in

Figure 5.6, where a characteristic curve for the MicroMax camera is shown with a reference

line indicating the shot noise-limited slope of ½.

Figure 5.6 Camera noise characterization, as a function of signal intensity - MicroMax frame-straddling CCD. Reprinted from [14].

100Signal-to-noise evaluations of the data images were based on the two-dimensional

mean intensity level in the background-subtracted data images (not flatfield-corrected). This

value was calculated on the laser sheet ROI described in the previous section. Day-to-day

and also set-to-set variation in laser power caused a range of mean signal levels to occur.

Table 5.2 contains mean signal level and shot noise-limited SNR (Equation 5.10) for each of

the four experiment image acquisition times. Due to the increased number density from

compression, the later timings were expected to yield higher signal, and thus SNR.

Image Timing [aTDC]

Mean Signal Intensity [counts]

Shot Noise-Limited Maximum SNR

- 30 102.3 21.2 : 1 - 45 89.5 19.8 : 1 - 60 68.2 17.3 : 1

600 RPM Low Load

- 99 45.0 14.1 : 1 - 30 125.7 23.5 : 1 - 45 105.7 21.6 : 1 - 60 86.5 19.5 : 1

600 RPM Mid Load

- 99 58.4 16.0 : 1 - 30 106.4 21.6 : 1 - 45 90.1 19.9 : 1 - 60 70.8 17.6 : 1

1200 RPM Low Load

- 99 46.0 14.2 : 1

Table 5.2 Values of spatial-mean data image intensity and resulting shot noise-limited maximum SNR for three speed/load points. Each set is the mean value for the five valve overlap strategies.

The signal levels of Table 5.2, when located on the abscissa of Figure 5.6 provide a

similar range of SNR as published by Rothamer and Ghandhi. The equation used for shot

noise-limited maximum signal-to-noise ratio is:

101

max * (5.10)outSNR I G=

The output gain term outG is related to digitization and equal to 4.4 e-/ADU for the

MicroMax camera. On examination of the characteristic curve in Figure 5.6, it can be seen

that the camera is operating in the shot-noise limited region, although the lower-signal

images at 99° bTDC are very near the transition away from shot noise-limited behavior.

Flatfield images, and thus homogenous images, have higher signal levels since they do not

contain residual gases. Due to the marginal signal level at some data image conditions, there

was interest in exploring potential benefits to using an intensified slow-scan camera in place

of the MicroMax.

5.2.4. MicroMax Comparison with Intensified CCD

Intensified cameras are commonly used in combustion PLIF measurements since their

extremely short gating time capabilities allow maximum light rejection. The MicroMax

camera was compared with the Roper Scientific PI-Max intensified slow-scan camera. In a

moderate signal level environment, the two devices were found to provide very comparable

signal-to-noise ratios, and the MicroMax yielded a slightly lower measure of spatial variation

on the homogeneous image condition.

Given the moderate-signal conditions and the intensified camera’s fundamental

tendency to blur gradients, the MicroMax camera was selected to maximize spatial

performance of the detection system.

102

5.3. Assessment of Intake Charge Homogeneity

A major experimental assumption of this project is that the fresh charge contains only

air, fuel and tracer and that it is homogenously mixed before IVO. Far-upstream air-assisted

fuel injection of pure hydrocarbon fuels has previously been verified for similar laboratory

setups in [13, 15, 16]. Nevertheless, the homogenous PLIF images will be analyzed here to

quantify degree of homogeneity.

5.3.1. First and Second Moments of Homogeneous Data

The homogenous images were tested by comparing their intensity variation

( )y yσ µ with the theoretical level of shot noise for the mean flatfield signal level at the

four timings at all 15 engine conditions. The column statistics were performed on unfiltered

corrected flatfield images, since the non-linear filtering is not predicted by shot noise theory.

The mean signal level mξ was that of the 100 flatfield images after background subtraction.

From this value, the theoretical shot noise level can be defined in terms of a normalized

standard deviation as:

( ) (5.11)

1where

mshot

m

out

AD

AD G

ξσ µ

ξ=

=

103A plot of the homogeneous image spatial variation compared to the theoretical shot

noise flatfield is shown in Figure 5.7. A 1-to-1 line is shown indicating the shot noise floor

for the ( )y yσ µ metric. All homogeneous images lie correctly above the shot noise floor,

although the small offset, close grouping and similar slope of the data points indicate that the

variations measured in all homogeneous conditions are primarily influenced by shot noise on

the CCD. The small offset can be partially attributed to slight misalignment of the laser sheet

vertical banding to the pixel columns where ( )y yσ µ is calculated as well as the

contribution of read and dark noise.

0 1 2 3 4 5 6 7 8 9 100

1

2

3

4

5

6

7

8

9

10

<(σ / µ)>shot [%]

<(σ

y / µ y)>

hmg

[%

]

Figure 5.7 Comparison of theoretical shot noise intensity variation ( )shot

σ µ to measured

homogenous pixel intensity variation ( )y yσ µ.

104

5.3.2. Homogeneous Image PDF

As a second check of the homogeneous fresh charge and also a demonstration of the

pixel intensity PDF function described in Section 5.1.5, the 12 homogeneous image sets

taken at the baseline valve overlap are shown in Figure 5.8. This plot shows the strong

grouping about the mean pixel intensity in the corrected images. The 600 RPM Low Load

condition is the curve showing the lowest peak, and also the condition of highest ( )y yσ µ

and lowest SNR in the experiment. Additional PDF figures will be shown in the next

chapter, accompanying data images, which will indicate that most homogeneous image sets

fall near the taller curves in Figure 5.8 (peaks near 30) and also that the skip-fired residual

gas data PDF’s are substantially lower in profile.

105

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.50

5

10

15

20

25

30

I / Imean

PD

F

Figure 5.8 Probability distribution function for pixel intensity in homogeneous image sets at four image timings for all three engine speed/load points. Baseline valve overlap. Each PDF curve contains information about 100 corrected homogenous images.

5.4. Direct-Injection Test of Imaging Technique

Residual gas mixing was measured in this experiment using what amounts to a

“negative PLIF” approach. The presence of residual gases in the image ROI is denoted by

regions of low fluorescence intensity. The fresh charge, in which the tracer molecules are

homogenously distributed, is assumed to provide the entire PLIF signal. If molecules

trapped or re-inducted with the residual gas fluoresce in the laser sheet, that signal would

then skew the identification of the residual in the image. A test was then performed to isolate

this signal from the predominant PLIF signal from the fresh air/fuel/tracer charge. Gasoline

direct-injection (GDI) hardware and the MotoTron skip-fire sequence were used to do this.

1065.4.1. Skip-Direct Injection Experiment

In order to isolate the fluorescence signal of the residual gases, fuel injection had to

be shut off on the image cycle. Direct fuel injection is capable of achieving this, so the

Triptane engine’s pushrod OHV cylinder head, which is outfitted with access for a high-

pressure Chrysler Pressure-Swirl injector, was used. The details of this combustion chamber

and fuel injection strategy are well documented in [13, 15, 16]. A summary of the engine

conditions for the GDI experiment are covered in Table 5.3

600 RPM 1200 RPM

CR 9.8 : 1

airm•

[mg/cycle] 212 172

fuelm•

[mg/cycle] 16.80 13.37

AFR 12.62 12.86

IMEP [kPa] 180 151

COV of IMEP [%] 13.2 7.7

iMAP [kPa] 55.84 46.53

EOI [° aTDC] -270 -270

IGN [° aTDC] -20* -20*

Texh [° C] 320 450

ry [%] 12.28 13.25

HC [ppm C1] 13,044 7706

Table 5.3 Direct injection experiment engine conditions and unburned hydrocarbon emissions measurements. * indicates the approximate ignition timing.

107To isolate the residual gas fluorescence signal on the image cycle, the direct injection

was disabled on the intake stroke preceding the image timing. The only means to accomplish

this was by using the one available skip-fire TTL output from MotoTron to perform both fuel

injection and ignition timing. The MotoTron sequence was set to trigger the SOI timing, via

a Berkeley Nucleonics Model 555 pulse/delay generator. The pulse generator provided the

fuel injection driver signal and, on a time delay, the ignition coil signal. This operation

disabled both fuel injection and ignition on the image cycle, but introduced an uncertainty

into the ignition timing, due to the time-based delay from the ECU signal. This resulted in

the high COV of IMEP shown for early-injection GDI operation.

The remainder of Table 5.3 shows that the OHV cylinder head provides a

significantly higher geometric compression ratio than the DOHC head (CR = 5.95:1). This is

reflected in the much lower residual gas mass fraction measured here (compare with Table

4.6). Based on the significant change in combustion chamber geometry, the experimental

conditions covered in Chapter 5 could not be matched. Instead, worst-case conditions for HC

emissions were set at 600 and 1200 RPM (within engine load limits). Hydrocarbon

emissions were considered indicative of the presence of unburned residual 3-pentanone (or

other fluorescent compounds) in the residual gas mixing images. By advancing the ignition

timing to phase location of peak pressure (LPP) near TDC, HC levels from crevice volume

outgassing were maximized. Measured HC was found to be comparable to the imaging

conditions in Table 4.4. Figures 5.9 and 5.10 present pressure data at both 600 RPM and

1200 RPM relative to the baseline overlap condition at each speed.

108

-180 -90 0 90 180 270 360 450 5400

200

400

600

800

1000

1200

1400

1600

1800

2000

CAD (aTDC)

p cyl [k

Pa]

Direct injection, OHV headBaseline overlap, DOHC head

Figure 5.9 Direct-injection experiment cylinder pressure trace comparison with DOHC baseline valve overlap. 600 RPM.

-180 -90 0 90 180 270 360 450 5400

200

400

600

800

1000

1200

1400

1600

CAD (aTDC)

p cyl [k

Pa]

Direct injection, OHV headBaseline overlap, DOHC head

Figure 5.10 Direct-injection experiment cylinder pressure trace comparison with DOHC baseline valve overlap. 1200 RPM.

1095.4.2. Skip-DI Imaging and Results

Two engine speeds were examined (Table 5.3); images were acquired at 10° bTDC

compression. First, the flatfield condition was imaged using upstream homogenous fuel

injection. The mean signal from the 100 flatfield images was compared to two direct-

injection conditions. First, the engine was operated with both skip-direct injection and skip-

firing enabled. This operation provided the “best case” for low concentration of unburned

fluorescent compounds in the combustion chamber, since the prior cycle involved complete

combustion. Second, the skip-fire ignition signal was disabled and the engine was motored

with the skip-direct injection. This condition provided the “worst case” for hydrocarbon

concentration in the residual gas, simulating a prior cycle with partial burn or misfire quality.

The results of the imaging experiment are shown in Table 5.4, with all mean intensity values

representing that of an ROI within the laser sheet after background subtraction.

600 RPM 1200 RPM

Motored Flatfield 165.5 119.0

Skip-Fire / Skip-DI 4.2 2.7

Skip DI, motoring 48.2 30.5

Table 5.4 Direct injection experiment imaging results. 100-image mean signal level for flatfield, skip-fired, and motored skip-DI PLIF data.

The data of Table 5.4 indicate a very low signal from the fired residual gas – less than

3% of the flatfield intensity for both engine speeds. Under non-fired operation, where

110misfire residual gas is simulated, the signal level rose to nearly 30% of the flatfield. The

engine operating conditions outlined in the next chapter were established to avoid high

cyclic-variability operation. Therefore, provided that the combustion is nearly complete, the

low fluorescence intensity levels shown by the skip-DI/skip-fire test indicate that the

technique is faithful in depicting residual gas mixing with homogenous fresh air/fuel charge.

111

6. Residual Gas Mixing

6.1. Sample Imaging Data

Before entering into statistical analysis of residual gas mixing, sample data images

representative of engine conditions with high, medium, and low residual gas fractions are

presented in Figures 6.2-6.4. A sample homogeneous image set from the same engine timing

is shown in Figure 6.1. All engine images are presented in sets of 25 consecutive frames, and

the grayscale assignment follows the procedure outlined in Section 5.1.6.

Yr = 0%, <(σy / µy)> = 1.94%

600 RPM Homogeneous Images, 60° bTDC

160

180

200

220

240

260

2800.9 × Imax

0.5 × Imax

Figure 6.1 Sample homogeneous image sequence, 60° bTDC.

Yr = 44.9%, <(σy / µy)> = 8.9%

1200 RPM Exh. Retard, 60° bTDC

90

100

110

120

130

140

1500.9 × Imax

0.5 × Imax

Figure 6.2 Sample data image sequence, high residual fraction condition, 60° bTDC.

112

Yr = 35.7%, <(σy / µy)> = 4.7%

600 RPM Sym. Increase, 60° bTDC

70

75

80

85

90

95

100

105

110

115

1200.9 × Imax

0.5 × Imax

Figure 6.3 Sample data image sequence, mid-range residual fraction, 60° bTDC.

Yr = 21.9%, <(σy / µy)> = 4.0%

1200 RPM Zero Overlap, 60° bTDC

100

110

120

130

140

150

160

170

1800.9 × Imax

0.5 × Imax

Figure 6.4 Sample data image sequence, low residual fraction condition, 60° bTDC.

All data image shown have been background-subtracted, flatfield-normalized and 3 x

3 median-filtered. All pixel variance data, ( )y yσ µ , are ensemble-mean values, computed

in the vertical direction (§5.1.4). This figure sequence serves to introduce the data produced

by the measurement system outlined in Chapters 3 and 5, and also visually demonstrates the

effect of elevated residual gas fractions on the homogeneous air/fuel/tracer mixture shown in

Figure 6.1. A more thorough discussion of the correlation between residual fraction and

( )y yσ µ will be covered in Section 6.3.

113

6.2. Correlation of Spatial-Mean Pixel Intensity with Measured Residual Gas Fraction

The ratio of the spatial-mean fluorescence signal of the skip-fired data image to that

of the motored flatfield image can be assumed to correlate with the bulk residual gas fraction

if important assumptions are made. The region of interest captured by the imaging system (§

4.3.1) is small relative to the combustion chamber and inherently two-dimensional, and

therefore extensions of the ROI image properties to entire cylinder volume are uncertain.

Ensemble-averaging of the maximum number of fired data images available can improve the

characterization, but nevertheless assumptions about the ROI have to be made. Secondly,

knowledge of the different in-cylinder conditions between skip-fired and motored operation

indicates an influence of the temperature-dependent fluorescence intensity on any

calculations. The flatfield condition, without any combustion products, contains a cooler

mixture, while the skip-fired data images contain tracer molecules that have been heated by

the residual gases shown to be mixing with them. Figure 2.8 demonstrates the di-ketone

group’s decreasing intensity yield with increasing temperature at 266-nm laser excitation,

and therefore a ratio of fired data to motored data will slightly over-predict our residual gas

fraction (Equation 6.1).

For these calculations, only background subtraction was applied to the raw CCD

image data, to preserve the detection system’s absolute scaling at each image timing and

engine condition. For each of the 15 engine speed/load/overlap conditions, the intensity ratio

was calculated as the mean of the four image acquisition timings by Equation 6.1:

114

30 45 60 99

114

data data data dataratio

ff ff ff ff

I I I III I I I

− − − −

= − + + +

(6.1)

The correlation between ratioI and ry is shown in Figure 6.5. The correlation can

be considered surprisingly good, given our assumptions about the ROI. The temperature

dependence of fluorescence appears in this plot as the offset between the slope of the data

point grouping, which is not far from parallel to the 1:1 line. The influence of residual gas

charge heating on PLIF uncertainty indicated by this correlation certainly has important

implications for the results of this project. A likely source of scatter in Figure 6.5 is slight

inconsistencies in mean laser sheet power between corresponding flatfield and skip-fire

imaging measurements.

0 10 20 30 40 50 60 70 800

10

20

30

40

50

60

70

80

Yr [%]

< I ra

tio >

[%

]

Figure 6.5 Correlation of mean image intensity ratio to measured residual fraction for all 15 experiment conditions.

115

6.3. Correlation of Residual Gas Fraction to Image Intensity Variation

The relationship between the degree of mean spatial variation in the corrected data

images and the bulk residual gas fraction was investigated across all 15 experiment

conditions. Four image timings (30°, 45°, 60°, and 99° bTDC) were studied at each

condition and a consistent trend of increasing charge stratification with increasing residual

fraction was found for all timings. This trend invited further exploration of our image data

based on residual fraction, independent of engine speed or load.

6.3.1. Cycle-Averaged Image Intensity COV Correlation

Figures 6.6 through 6.9 present the correlation observed between measured residual

gas fraction and the ensemble-mean pixel intensity variation, ( )y yσ µ , captured in the

skip-fired data images. Separate plots are prepared for each of the four image timings and

each plot contains the appropriate reading for each of the 15 experiment conditions.

Additionally, the spatial variations for the corresponding homogeneous image data are shown

as a relative measure of the absolute shift in stratification when measuring the fired engine

flow.

The sequence in Figures 6.6-6.9 includes the mean SNR values calculated from all

data points at the individual image timings (Table 5.2). This value clearly decreases with the

lower-charge density images at the advanced timings. This effect is also seen in the

incremental upward shift in homogeneous image variation level at the advanced timings.

116

0 5 10 15 20 25 30 35 40 45 500

2

4

6

8

10

12

30° bTDC

Yr [%]

<(σ

y / µ y)>

[%

]

Fired DataHomogeneous

Figure 6.6 Pixel intensity COV vs. residual gas fraction for all engine conditions at 30° bTDC. Shot noise-limited maximum SNR was ~22:1 for this image timing.

0 5 10 15 20 25 30 35 40 45 500

2

4

6

8

10

12

45° bTDC

Yr [%]

<(σ

y / µ y)>

[%

]

Fired DataHomogeneous

Figure 6.7 Pixel intensity COV vs. residual gas fraction for all engine conditions at 45° bTDC. Shot noise-limited maximum SNR was ~20:1 for this image timing.

117

0 5 10 15 20 25 30 35 40 45 500

2

4

6

8

10

12

60° bTDC

Yr [%]

<(σ

y / µ y)>

[%

]

Fired DataHomogeneous

Figure 6.8 Pixel intensity COV vs. residual gas fraction for all engine conditions at 60° bTDC. Shot noise-limited maximum SNR was ~18:1 for this image timing.

0 5 10 15 20 25 30 35 40 45 500

2

4

6

8

10

12

99° bTDC

Yr [%]

<(σ

y / µ y)>

[%

]

Fired DataHomogeneous

Figure 6.9 Pixel intensity COV vs. residual gas fraction for all engine conditions at 99° bTDC. Shot noise-limited maximum SNR was ~15:1 for this image timing.

118The most notable feature of Figures 6.6-6.9 is the quasi-exponential growth in

( )y yσ µ at the highest recorded residual gas fractions. This trend, particularly the

transition range of 35% to 40% residual fraction where both 600 RPM and 1200 RPM data

points are located, indicates that there is a residual dilution level at which mixture

composition inhomogeneity begins to rapidly increase, independent of speed or load. Since

the engine could only exceed 40 % residual fraction at 1200 RPM, the maximum ( )y yσ µ

points naturally occur at 1200 RPM only.

Figures 6.6 through 6.9 also demonstrate the absolute magnitude shift in the intensity

variation metric from the motored flatfield condition. Corrected homogeneous images fall

near 1-2% ( )y yσ µ , depending primarily on the shot noise encountered at the image

timing. The difference between the fired data points and the homogeneous points is a clear

and consistent sign of the presence of residual gas unmixedness.

6.3.2. Lower Residual Fraction Case-to-Case Comparison

At the lower end of the measured scale in Figures 6.6-6.9, the nearest match between

600 RPM and 1200 RPM conditions was for the baseline overlap 1200 RPM set ( ry =27.3%)

and the zero overlap 600 RPM low load set ( ry =28.7%). The ( )y yσ µ data for these

conditions are presented in Table 6.3. With the exception of the early 99° bTDC timing, the

1200 RPM data show consistently lower variation than 600 RPM. Both data sets

demonstrate an increasing image intensity variation at 30° bTDC.

119% ry 30° bTDC 45° bTDC 60° bTDC 99° bTDC

600 RPM, Zero Overlap 28.7 5.21 4.37 4.44 5.47 1200 RPM, Baseline OV 27.3 4.18 3.99 4.09 5.90

Table 6.1 Comparison of lower-residual conditions at 600 and 1200 RPM. Development of

image ( )y yσ µ [%] with crank angle.

The data of Table 6.3 would seem to suggest an engine speed influence on residual

gas mixing at this dilution level, with the higher engine speed doing a better job of mixing

the residual gas with the homogeneous fresh charge in the imaging plane. Sample image

data acquired for the two conditions at 45° bTDC are presented in Figures 6.10 and 6.11.

Yr = 28.7%, <(σy / µy)> = 4.37%

600 RPM Low Load, Zero OV, 45° bTDC

110

120

130

140

150

160

170

180

0.9 × Imax

0.5 × Imax

Figure 6.10 Sample data images for 600 RPM, low-residual condition.

Yr = 27.3%, <(σy / µy)> = 3.99%

1200 RPM, Baseline OV, 45° bTDC

120

130

140

150

160

170

180

190

0.9 × Imax

0.5 × Imax

Figure 6.11 Sample data images for 1200 RPM, low-residual condition.

120Visually comparing the 600 RPM data images in Figure 6.10 with the 1200 RPM set

in Figure 6.11 indicates the difficulty in qualitatively distinguishing two conditions. The

( )y yσ µ values for these are in fact similar, with the brighter and more numerous flow

structures shown at 600 RPM likely accounting for the difference. The PDF is a more

quantitative comparison between the two conditions, as shown in Figures 6.12-6.13.

Figures 6.12-6.13 are an introduction to the nature of the residual gas data image

series PDF, where a substantial deviation from the homogeneous mixture data is seen. At

600 RPM, where the mean intensity variation was lower and shows a slightly lower peak in

the distribution, although the homogenous data was also less tightly grouped.

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.50

5

10

15

20

25

30

45° bTDC600 RPM Low LoadZero Overlap

I / Imean

PD

F

HomogeneousFired Data

Yr = 28.7 %<(σy / µy)> = 4.37 %

Figure 6.12 100-image pixel intensity PDF for 600 RPM low-residual condition.

121

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.50

5

10

15

20

25

30

45° bTDC1200 RPMBaseline Overlap

I / Imean

PD

F

HomogeneousFired Data

Yr = 27.3 %<(σy / µy)> = 3.99 %

Figure 6.13 100-image pixel intensity PDF for 1200 RPM low-residual condition.

6.3.3. Higher Residual Fraction Case-to-Case Comparison

Since the engine was more tolerant of elevated residual fractions at 1200 RPM,

comparison of image data between engine speeds was not possible for the maximum dilution

levels. The closest match at the high-end occurred between the 600 RPM low load,

symmetric increased overlap ( ry = 40.4%) and the 1200 RPM intake cam advance condition

( ry = 40.8%). Again, the development in ( )y yσ µ is presented in Table 6.4.

122% ry 30° bTDC 45° bTDC 60° bTDC 99° bTDC

600 RPM, Sym. Incr. OV 40.4 5.93 5.19 5.15 6.68 1200 RPM, Intake Advance 40.8 5.92 7.09 7.78 8.02

Table 6.2 Comparison of higher-residual conditions at 600 and 1200 RPM. Development of ( )y yσ µ [%] with crank angle.

At the latest image timing, the variation was identical for the two conditions.

However, the earlier development of intensity variation was much different, with the 600

RPM condition reaching an early minimum at 60° bTDC and the 1200 RPM data decreasing

steadily from a high initial level of 8%. At this dilution level, the influence of engine speed

seems to be reversed, as the longer mixing time found at 600 RPM produced lower mixture

variation in the ROI. Given the conflicting conclusions on engine speed, it would then

appear that bulk residual gas fraction is a more applicable parameter in predicting the mixture

inhomogeneity. Sample data images for the two conditions of this section are shown in

Figures 6.14-6.15.

123

Yr = 40.4%, <(σy / µy)> = 5.19%

600 RPM Low Load, Sym. Incr., 45° bTDC

110

120

130

140

150

160

170

180

0.9 × Imax

0.5 × Imax

Figure 6.14 Sample data images for 600 RPM, high-residual condition. 45° bTDC.

Yr = 40.8%, <(σy / µy)> = 7.09%

1200 RPM Intake Advance, 45° bTDC

100

110

120

130

140

150

160

1700.9 × Imax

0.5 × Imax

Figure 6.15 Sample data images for 1200 RPM, low-residual condition.

6.4. Prior-Cycle Effect on Image Intensity Variation

Cylinder pressure was recorded for all engine cycles during the image acquisition

periods and single-cycle IMEP data could be extracted in software. Since the experiment

involved skip-firing, same-cycle pressure data was not relevant to the engine operating

condition. The best correlation possible to the imaged engine flow would have to come from

the previous engine cycle. It was proposed that strong and weak prior cycles would have

some degree of influence over the residual gas fraction on the skip-fired cycle. Sample

124

results of the extracted prior-cycle IMEP and corresponding data image COV ( )y y nσ µ are

shown in Figures 6.16 and 6.17, with the axes scaled relative to the data-set mean. These

figures display the prior-cycle IMEP data for all 100 images at 60° bTDC. Figure 6.16 is for

the 600 RPM low load, symmetric overlap condition and Figure 6.17 is the same overlap

strategy at 1200 RPM. These engine conditions were chosen based on their high residual

fraction and high pixel intensity variation.

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 20.75

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

1.2

1.25

(σy / µy)n / <(σy / µy)>

[IME

P n-1] /

[IM

EP m

ean]

Figure 6.16 Prior-cycle IMEP vs. image intensity COV. 600 RPM Low Load, Sym.

Increase 60° bTDC. Yr = 40.4%, IMEP=152 kPa, COVIMEP = 6.0%, ( )y yσ µ =5.2%.

125

0 0.2 0.4 0.6 0.8 1 1.2 1.4 1.6 1.8 20.75

0.8

0.85

0.9

0.95

1

1.05

1.1

1.15

1.2

1.25

(σy / µy)n / <(σy / µy)>

[IME

P n-1] /

[IM

EP m

ean]

Figure 6.17 Prior-cycle IMEP vs. image intensity COV. 1200 RPM, Sym. Increase 60°

bTDC. Yr = 43.7%, IMEP=253 kPa, COVIMEP = 1.2%, ( )y y nσ µ

=7.3%.

All engine conditions demonstrated a similar random data scatter and no correlation

was found between prior-cycle IMEP and ( )y y nσ µ . It was assumed that quantifying the

large-scale cylinder mixture influence of the prior cycle combustion performance with a

small ROI is poorly suited for the single-image analysis done here. Ensemble-averaged ROI

data, such as presented in Section 6.3, are more likely to yield a better correlation.

126

6.5. Engine Operating Conditions Effect on Data Image Intensity Variation

A major feature of this project was to establish engine conditions of varying residual

gas dilution by means of valve overlap strategies. For this experiment, a baseline 20° overlap

was established and three strategies for elevated residual fraction were explored – intake cam

advance, exhaust cam retard, and symmetric overlap increase. Finally, zero valve overlap

was studied to establish a minimum residual fraction. Trends of ( )y yσ µ for the various

overlap strategies are shown in Figures 6.18-6.20, organized by speed/load.

-120 -100 -80 -60 -40 -20 00

1

2

3

4

5

6

7

8

9

10

CAD aTDC

<(σ

y / µ y)>

[%

]

600 Low

Baseline OverlapSymmetric IncreaseIntake AdvanceExhaust RetardZero Overlap

Figure 6.18 Mean image intensity variation vs. CA at 600 RPM low load, all overlaps.

127

-120 -100 -80 -60 -40 -20 00

1

2

3

4

5

6

7

8

9

10

CAD aTDC

<(σ

y / µ y)>

[%

]

600 Mid

Baseline OverlapSymmetric IncreaseIntake AdvanceExhaust RetardZero Overlap

Figure 6.19 Mean image intensity variation vs. CA at 600 RPM mid load, all overlaps.

-120 -100 -80 -60 -40 -20 00

2

4

6

8

10

12

CAD aTDC

<(σ

y / µ y)>

[%

]

1200 Low

Baseline OverlapSymmetric IncreaseIntake AdvanceExhaust RetardZero Overlap

Figure 6.20 Mean image intensity variation vs. CA at 1200 RPM, all overlaps.

128The 600 RPM conditions in Figures 6.18 and 6.19 demonstrate a consistent “hook-

up” in ( )y yσ µ approaching 30° bTDC. This trend is presumed to be indicative of a bulk

in-cylinder flow component, mostly likely a tumble motion induced by the pent-roof

geometry, consistently delivering pockets of unmixed fluid to the imaging plane. At the

elevated load conditions in Figure 6.19, the intake cam advance set does not demonstrate the

hook-up trend, indicating that the change in phasing of the induction process may have

affected the timing of the flow at that intake manifold pressure.

The most prevalent trend in the 1200 RPM plot of Figure 6.20 is the large magnitude

shift in flow inhomogeneity from the low-residual baseline and zero overlap conditions to the

increased overlap strategies. As with much of the combustion data in Chapter 4, the exhaust

retard condition at 1200 RPM demonstrated the largest effect of residual dilution, with the

highest image intensity variation levels (and also image-to-image variation) recorded in the

experiment. In general, the 1200 RPM conditions are not as conclusive as to bulk flowfield

influence on residual gas transport as the 600 RPM data.

6.5.1. Symmetric Overlap Increase

At the 600 RPM low load condition, the symmetric 10° increase in valve overlap did

not significantly shift ( )y yσ µ from the baseline values at either 99° or 60° bTDC (6.8%

and 5.2%, respectively). At the later timings, image variation became more pronounced,

particularly at 45° bTDC, where a 12% increase in ( )y yσ µ from the baseline was

129calculated. In general, the level of variation was not distinguishable from the other overlap

strategies.

The mid load condition at 600 RPM demonstrated a flatter progression in Figure 6.19

through the compression stroke, roughly splitting the difference in ( )y yσ µ between the

other two increased-overlap strategies. Similar behavior is seen in the 60° total overlap 1200

RPM data in Figure 6.20, although the level of variation is noticeably increased with the

higher residual fraction (43.7%) at this condition.

6.5.2. Intake Cam Advance

In Figure 6.18, the intake cam advance data points are indistinguishable in ( )y yσ µ

from the exhaust retard condition and at the later timings also from the symmetric increased

overlap. This plot seems to indicate that at 600 RPM low load the effect of cam phasing

strategy is small, perhaps limited by the small increases in dilution level that were tolerable.

At the higher IMEP conditions in Figures 6.19 and 6.20, there are clearer distinctions

in the behavior of the three strategies. At both 600 RPM and 1200 RPM, the intake cam

advance strategy delivered the lowest intensity variation at high residual fraction. The 1200

RPM data set is the only one in the experiment to exhibit continuously decreasing ( )y yσ µ

throughout the four image timings. The improvements in image inhomogeneity can be

partially attributable to the lower measured residual gas fraction relative to the symmetric

increase and exhaust retard cases. Figures 6.21 through 6.24 compare intake advance and

exhaust retard imaging at both speeds, followed by the PDF’s for all 4 images.

130

Yr = 32.5%, <(σy / µy)> = 4.09%

600 RPM Mid Load, Int. Advance, 45° bTDC

110

120

130

140

150

160

170

180

190

0.9 × Imax

0.5 × Imax

Figure 6.21 Intake advance data images at 600 RPM Mid Load. 45° bTDC.

Yr = 32.7%, <(σy / µy)> = 4.92%

600 RPM Mid Load, Exh. Retard, 45° bTDC

130

140

150

160

170

180

190

200

210

220

0.9 × Imax

0.5 × Imax

Figure 6.22 Exhaust retard data images at 600 RPM Mid Load. 45° bTDC.

Yr = 40.8%, <(σy / µy)> = 7.09%

1200 RPM Intake Advance, 45° bTDC

100

110

120

130

140

150

160

1700.9 × Imax

0.5 × Imax

Figure 6.23 Intake advance data images at 1200 RPM. 45° bTDC.

Yr = 44.8%, <(σy / µy)> = 8.71%

1200 RPM Exhaust Retard, 45° bTDC

110

120

130

140

150

160

170

180

190

0.9 × Imax

0.5 × Imax

Figure 6.24 Exhaust retard data images at 1200 RPM. 45° bTDC.

131

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.50

5

10

15

20

25

30

45° bTDC600 RPM Mid LoadIntake Cam Advance

I / Imean

PD

F

HomogeneousFired Data

Yr = 32.5%<(σy / µy)> = 4.09%

Figure 6.25 Intake advance 100-image pixel intensity PDF at 600 RPM Mid Load, 45° bTDC.

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.50

5

10

15

20

25

30

45° bTDC600 RPM Mid LoadExhaust Cam Retard

I / Imean

PD

F

HomogeneousFired Data

Yr = 32.7 %<(σy / µy)> = 4.92%

Figure 6.26 Exhaust retard 100-image pixel intensity PDF at 600 RPM Mid Load, 45° bTDC.

132

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.50

5

10

15

20

25

30

45° bTDC1200 RPMIntake Cam Advance

I / Imean

PD

F

HomogeneousFired Data

Yr = 40.8%<(σy / µy)> = 7.09%

Figure 6.27 Intake advance 100-image pixel intensity PDF at 1200 RPM 45° bTDC.

0.5 0.6 0.7 0.8 0.9 1 1.1 1.2 1.3 1.4 1.50

5

10

15

20

25

30

45° bTDC1200 RPMExhaust Cam Retard

I / Imean

PD

F

HomogeneousFired Data

Yr = 44.8%<(σy / µy)> = 8.71%

Figure 6.28 Exhaust retard 100-image pixel intensity PDF at 1200 RPM 45° bTDC.

1336.5.3. Exhaust Cam Retard

As shown in the previous image sequence, the exhaust cam retard strategy

demonstrated unique qualitative image behavior relative to the intake advance conditions at

both 1200 RPM and 600 RPM mid load. Particularly at the lower residual gas fraction 600

RPM case in Figure 6.21, the intake advance data images show a distinctly more continuous

distribution of tracer in the images. Using the consistent 90%/50% scaling method,

structures in the intake advance ROI tend to be defined by “shades of gray”, as opposed to

“black and white” behavior in the exhaust retard images.

The exhaust cam retard strategy yielded the most consistently high ( )y yσ µ values

in the experiment. Even at the 600 RPM mid-load condition where the symmetric overlap

increase gave an 11% larger increase in residual fraction from the baseline (Table 4.6), the

exhaust retard data display a higher magnitude of intensity variation.

134

7. Summary and Conclusions

7.1. Project Summary

An experiment was performed to measure mixing processes between fresh intake

charge and residual gas in a spark-ignition engine. A single-cylinder, optically accessible

test engine was upgraded with an adjustable dual-overhead cam cylinder head for providing a

range of residual gas fraction conditions. A direct in-cylinder sampling valve was used to

measure bulk residual fraction, while exhaust emissions and cylinder pressure analyses

quantified engine operation at 15 distinct conditions.

Laser-induced fluorescence imaging was performed during the later portion of the

compression stroke during a programmed skip-fired engine operation cycle to allow use of a

fast-shuttering non-intensified CCD camera. The fluorescence was obtained with 3-

pentanone doped into the iso-octane fuel. The premixed fuel system provided a

homogeneously mixed intake charge and the residual gases were found to be largely devoid

of fluorescent components. Thus, the image inhomogeneity was indicative of fresh

charge/residual gas mixing.

An image correction algorithm was developed to extract engine flow information

from the raw fluorescence data. First and second statistical moments were calculated in a

manner which avoided error introduced by correction process imperfections. Normalized

spatial variation in fluorescence intensity, corresponding to compressed charge

135inhomogeneity, was quantified for the 15 experimental conditions at four unique crank angle

timings and compared with the combustion analyses.

7.2. Results Summary

The engine was considerably more tolerant to residual gas dilution at 1200 RPM than

at 600 RPM, particularly at the low load (MAP) condition. IMEP levels increased for all

extended overlap conditions, although change was lower at 600 RPM than 1200 RPM. This

improvement appeared to be largely a result of increasing volumetric efficiency as indicated

by PMEP, which dropped up to 37% from the baseline overlap at 1200 RPM. The zero valve

overlap condition reversed these trends from the baseline. Combustion stability, denoted by

COV of IMEP, was very good throughout the experiment (< 6%). Perhaps more importantly,

misfires, which could strongly skew cycle-resolved residual gas measurements, were not

recorded at any condition.

Heat release rates and cumulative burn curves showed a very prolonged burning

duration indicative of elevated charge dilution. Ignition delay and flame development, as

indicated by the 0-10% HR duration, were also quite long at all conditions. At the low load

600 RPM condition, the three extended overlap strategies generated comparable burn rate

data, with approximately a 30% increase in 10-90% HR duration. The mid-load condition at

600 RPM was similarly insensitive to overlap strategy, and the increases from the baseline

overlap values were smaller. Due to the larger increase in overlap at 1200 RPM, the

magnitude of the burn duration increase was more notable. Intake advance and symmetric

136increase conditions showed comparable heat release data, while the exhaust retard case was a

clear outlier, resulting in the largest flame development and overall burn angles in the

experiment. All zero overlap conditions showed a reduction in flame development time, with

little change in overall burn duration.

Exhaust emission measurements confirmed the highly fuel-rich operating conditions,

with measured equivalence ratios of Φ = 1.14 through 1.36. Not surprisingly, combustion

efficiency values were consistently below 90%, with high CO and unburned HC levels.

Hydrocarbon readings were known to be skewed by the large crevice volume of the optical

engine, although the extended heat release curves indicate substantial incomplete

combustion. The intake cam advance condition generated the lowest CO and HC levels of

the extended overlap strategies for all three speed/load points, resulting in the smallest

penalty in combustion efficiency from the baseline overlap. Emission measurements at zero

valve overlap were comparable to the baseline conditions.

Bulk residual gas fraction measurements ranged from 27% to 38% at the baseline

overlap, because of the low compression ratio and throttled operation. Given the high

residual fraction at the baseline condition, at 600 RPM the residual fractions did not increase

significantly with the 10° overlap increases, and at low-load, the growth was essentially

indifferent to cam strategy. The mid-load condition showed a slightly higher sensitivity to

overlap increase, with the symmetric increase case yielding the largest increase from the

baseline dilution level. The 1200 RPM conditions contained a more distinct indication of

cam phasing effect on residual gas fraction. With the 40° increase in valve overlap, all three

strategies were far above the baseline residual, with the intake cam advance giving the

smallest increase (+50% to ry =41%) and the exhaust retard case giving the largest residual

137fraction recorded (+64% from baseline to ry =45%). Zero overlap resulted in roughly 20%

decreases in residual gas fraction at all three speed/load points.

The measurement technique of visualizing residual gas mixing by fluorescing a

homogeneous fresh charge was verified using PLIF tests of compressed charge homogeneity

during non-firing operation, and of fluorescence intensity of unburned hydrocarbons

contained in the residual gas. Spatial variation in the corrected homogeneous images was

shown to correlate very close to the theoretical shot noise floor, indicating a sufficiently

homogeneous fresh charge. A fuel cutoff experiment performed with direct fuel injection

indicated that for non-misfire cycles, the fluorescence intensity of residual compounds was

less than 3% of the homogeneous charge flatfield signal.

The mean fluorescence signal of the skip-fired residual gas data images was

compared to that of the motored flatfield at the same image timing. Although this ratio was

subject to inherent spatial uncertainty, it over-predicted the measured residual gas fraction by

an average of 52%. This result was largely attributed to different in-cylinder temperatures

between the fired and motored cases, which affects the fluorescence yield of the tracer

molecules. The 15 measurement conditions did generally follow a 1:1 increasing slope with

the measured data.

The spatial variation of the images was shown to shift significantly from the

homogenous images’ low values, which were near the noise floor and varied from 1.5% to

4% at the lowest-residual conditions. The absolute shift in image behavior was also

quantified by probability density functions of pixel intensity across the 100-image sets.

At all image timings, except the low-SNR 99° bTDC points, data image sets taken at

conditions with ry < 35% tended to converge to ( )y yσ µ values between 4% and 5%, not

138considering engine speed/load or overlap strategy. Beyond 35% residual gas fraction and

particularly for the 1200 RPM data where ry > 40%, the mean spatial variation metric grows

rapidly with residual fraction to a peak level of 9%. Data acquired at 45° bTDC and 60°

bTDC show a close grouping of data points along this quasi-exponential curve. The latest

image timing, 30° bTDC, demonstrated slightly more condition-to-condition variation.

At the two 600 RPM load points, an increasing slope in the spatial variation term was

found at 30° bTDC, which was considered indicative of a bulk tumbling transport motion of

residual gases through the ROI during compression.

The intake cam advance and exhaust cam retard conditions were compared, as they

routinely provided the best and worst cases, respectively, for spatial variation levels at

elevated valve overlaps.

7.3. Conclusions

The most significant result of this project was the qualitative and quantitative

observation of inhomogeneity of the compressed charge near ignition timing at elevated

residual fractions in the range of 21-45%. This information is important for combustion

modeling in high-dilution engines. In particular, the notion of homogenously distributed

residual gases in the central regions of the combustion chamber has been brought under some

doubt in this project.

The rate of growth of the spatial variation of the charge with increasing residual

fraction at all four experiment image timings (Figures 6.6-6.9) indicate that a significant

139increase in charge stratification occurs at ry ≈ 35%. This correlation was observed across

the speed/load points and overlap strategies, which indicates that the bulk residual fraction

can be a dominant term in predicting the level of compressed charge inhomogeneity.

Of the three cam-phasing strategies employed to obtain the high residual fractions,

the intake cam advance case was shown to provide the lowest charge inhomogeneity levels

for a given overlap duration. This was partly explained by the fact that intake advance also

provided slightly lower residual fractions than exhaust retard or symmetric increase. But

additionally, intake advance introduces the fresh charge to the combustion chamber earliest,

thereby maximizing the mixing time with the residual gas. The exhaust retard cases had the

latest IVO timing of the extended overlaps, and at 1200 RPM demonstrated the highest

spatial variation measurements of the project. From a mixedness point of view, it can then be

argued that the intake cam advance strategy is preferable as it provides the largest amount of

time for charge mixing before ignition.

This project was successful in generating a range of residual fractions that can be

considered appropriate for a high-dilution design. Fuel-rich operation and a large spark gap

were demonstrated to be useful in obtaining stable running conditions at these dilution levels

in an optical engine.

The technique of tracking residual gases by negative-PLIF of a homogenous

air/fuel/tracer mixture was demonstrated to be sound. A fundamental issue of pre-ignition

in-cylinder luminosity was discovered, which has important implications in the use of high-

fidelity CCD cameras for these measurements. This project demonstrated the utility of the

frame-straddling camera in circumventing this problem.

140

7.4. Recommendations for Future Work

This project covered the high end of residual gas dilution levels in spark ignition

engine designs. The results of the correlation between data image spatial variation and bulk

residual fraction suggested a very slow rate of decrease in charge inhomogeneity below ry =

20%. It was shown in this project that the zero-residual fraction homogeneous condition

demonstrated a spatial variation near that of the shot noise floor. This means that this trend

of ( )y yσ µ vs. ry must at some point break from its leveled-off spatial variation toward

homogeneity in the low and moderate range of residual fractions not covered in this

experiment. With spatial variation still clearly present at 21% residual fraction, data below

this level will certainly be desirable, particularly from an experiment which could identify a

threshold residual gas fraction below which homogeneity can be safely assumed.

A more robust research engine could be used to explore the range of residual fractions

above 40% where the rapid increase in charge stratification was measured. Extensive heat

release data from a more extensive parametric study of engine conditions could better

identify this behavior as being a fundamental limitation to high-dilution engines, or a more

engine-specific phenomenon of insufficient in-cylinder mixing. This information would

likely be very valuable to HCCI research, where novel variable-cam re-breathing strategies,

and therefore very high levels of residual gas dilution, are being pursued.

141

References [1] Heywood, John B., 1988, Internal Combustion Engine Fundamentals, McGraw-Hill,

New York. [2] Fox, J.W., Cheng, W.K., and Heywood, J.B., 1993, “A Model for Predicting Residual

Gas Fraction in Spark Ignition Engines”, SAE 931025. [3] Rhodes, D.B. and Keck, J.C., 1985, “Laminar Burning Speed Measurements of

Indolene-Air-Diluent Mixtures at High Pressures and Temperatures,” SAE 850047. [4] Turns, Stephen R., 1996, An Introduction to Combustion: Concepts and Applications,

McGraw-Hill, New York. [5] Olofsson, E., Alvestig, P., Bergsten, L., Ekenberg, M., Gawell, A., Larsen, A., and

Riemann, R., 2001, “A High Dilution Stoichiometric Combustion Concept Using a Wide Variable Spark Gap and In-Cylinder Air Injection in Order to Meet Future CO2 Requirements and Worldwide Emissions Regulations,” SAE 2001-01-0246.

[6] Ozdor, N., Dulger, M. and Sher, E., 1994, “Cyclic Variability in Spark Ignition

Engines: A Literature Survey,” SAE 940987. [7] Hinze, P.C. and Miles, P.C., 1999, “Quantitative Measurements of Residual and

Fresh Charge Mixing in a Modern SI Engine Using Spontaneous Raman Scattering,” SAE 1999-01-1106.

[8] Miles, P.C. and Hinze, P.C., 1998, “Characterization of the Mixing of Fresh Charge

with Combustion Residuals using Laser Raman Scattering with Broadband Detection,” SAE 981428.

[9] Stanglmaier, R.H. and Roberts, C.E., 1999, “Homogeneous Charge Compression

Ignition (HCCI): Benefits, compromises and Future Engine Applications,” SAE 1999-01-3682.

[10] Epping, K., Aceves, S., Bechtold, R., 2002, “The Potential of HCCI Combustion for

High Efficiency and Low Emissions,” SAE 2002-01-1923. [11] Zhao, Hua and Ladommatos, Nicos, 2001, Engine Combustion Instrumentation and

Diagnostics, Chapter 4, SAE International, Warrendale, PA. [12] Foudray, H.Z., 2002, Scavenging Measurements in a Direct-Injection 2-Stroke

Engine, M.S. Thesis, Department of Mechanical Engineering, University of Wisconsin-Madison.

142 [13] Rothamer, D.A., 2002, Investigation of Flame-Front Equivalence Ratio during

Stratified Engine Combustion, M.S. Thesis, Department of Mechanical Engineering, University of Wisconsin-Madison.

[14] Rothamer, D.A. and Ghandhi, J.B., 2002, “On the Calibration of Single-shot Planar

Laser Imaging Techniques in Engines,” SAE 2002-01-0748. [15] Wiles, M.A., 2003, Characterization of Operating Parameters’ Authority on the

Flow-Field Mixedness of a DISI Engine, M.S. Thesis, Department of Mechanical Engineering, University of Wisconsin-Madison.

[16] Probst, D.M., 2001, Spray Mixing in Engines, M.S. Thesis, Department of

Mechanical Engineering, University of Wisconsin-Madison. [17] Hansen, D.A. and Lee, E.K.C., 1975, “Radiative and Nonradiative Transitions in the

First Excited Singlet State of Symmetrical Methyl-Substituted Acetones,” J. Chem. Physics., v.62 no.183.

[18] Thurber, M.C., Grisch, F., Kirby, B.J., Votsmeier, M., and Hanson, R.K., 1998,

“Measurements and Modeling of Acetone Laser-Induced Fluorescence with Implications for Temperature Imaging Diagnostics,” Applied Optics, v.37 no.21 pp4963-4978.

[19] Thurber, M.C. and Hanson, R.K., 1999, “Pressure and Composition Dependences of

Acetone Laser-Induced Fluorescence with Excitation at 248, 266 and 308 nm,” Applied Physics, v.69, pp.229-240.

[20] Lebel, M. and Cottereau, M.J., 1992, “Study of the Effect of the Residual Gas

Fraction on Combustion in a S.I. Engine Using Simultaneous CARS Measurements of Temperature and CO2 Concentration,” SAE 922388.

[21] Baritaud, T.A. and Heinze, T.A., 1992, “Gasoline Distribution Measurements with

PLIF in a SI Engine,” SAE 922355. [22] Johansson, B., Neij, H., Juhlin, G., and Alden, M., 1995, “Residual Gas Visualization

with Laser Induced Fluorescence,” SAE 952463. [23] Deschamps, B. and Baritaud, T.A., 1996, “Visualization of Gasoline and Exhaust

Gases Distribution in a 4-Valve SI Engine; Effects of Stratification on Combustion and Pollutants,” SAE 961928.

143[24] Roper Scientific, 2000, Princeton Instruments 5MHz MicroMax User Manual. [25] Stivender, D.L., 1971, “Development of a Fuel-Based Mass Emission Measurement

Procedure,” SAE 710604. [26] Albert, B.P., 2004, Residual Gas Effects on Combustion in an Air-Cooled Utility

Engine, M.S. Thesis, Department of Mechanical Engineering, University of Wisconsin-Madison.

144

Appendix A – Engine Operating Conditions

600 RPM Low Load 600 RPM Mid Load 1200 RPMRPM [RPM] 600 600 1200EVO [aTDC] 140 140 140EVC [aTDC] 370 370 370IVO [aTDC] 350 350 350IVC [aTDC] -135 -135 -135Exhaust MAP [kPa] 100.0 100.0 100.0Intake MAP [kPa] 49.6 61.3 51.7m_air [mg/cycle] 144 208 181m_fuel [mg/cycle] 14.5 18 18IGN [aTDC] -60 -40 -40T_exh [C] 347 400 477

IMEP [kPa] 147.25 260.91 228.15COV of IMEP [%] 4.1 2.05 1.32PMEP [kPa] 46.22 35.33 51.08Peak Pressure [kPa] 1021 1416 1280Location of PP [aTDC] 12 13.5 12

0-10% HR [CAD] 53 35.5 35.510-90% HR [CAD] 58.5 59.5 68

[CO2] (dry) [%] 8.12 10.64 9.21[CO] (dry) [%] 6.46 4.75 5.94[O2] (dry) [%] 0.42 0.59 0.56[HC] (dry) [ppm C1] 10629 5769 6159n_comb [ ] 0.821 0.899 0.891measured AFR [ ] 10.69 12.1 11.46K_exh [ ] 0.871 0.865 0.871

[CO2] SV [%] 2.759 2.807 2.253K_cc [ ] 0.964 0.968 0.972Y_r [%] 37.65 29.59 27.27

BASELINE OVERLAP

145

600 RPM Low Load 600 RPM Mid Load 1200 RPMRPM [RPM] 600 600 1200EVO [aTDC] 145 145 160EVC [aTDC] 375 375 390IVO [aTDC] 345 345 330IVC [aTDC] -140 -140 -155Exhaust MAP [kPa] 100.0 100.0 100.0Intake MAP [kPa] 54.8 67.2 68.2m_air [mg/cycle] 144 208 181m_fuel [mg/cycle] 14.5 18 18IGN [aTDC] -60 -45 -60T_exh [C] 340 389 462

IMEP [kPa] 151.65 270.68 253.40COV of IMEP [%] 6.03 2.07 1.20PMEP [kPa] 45.00 30.32 35.32Peak Pressure [kPa] 796 1437 1228Location of PP [aTDC] 17 14 15

0-10% HR [CAD] 61.0 40.5 56.510-90% HR [CAD] 78.5 67.0 86.5

[CO2] (dry) [%] 8.11 10.73 8.86[CO] (dry) [%] 6.33 4.55 6.04[O2] (dry) [%] 1.06 0.68 0.79[HC] (dry) [ppm C1] 18822 5838 9015n_comb [ ] 0.719 0.897 0.846measured AFR [ ] 10.54 12.22 11.29K_exh [ ] 0.858 0.865 0.869

[CO2] SV [%] 2.93 3.44 3.50K_cc [ ] 0.958 0.964 0.959Y_r [%] 40.39 35.75 43.65

SYMMETRIC OVERLAP INCREASE

146

600 RPM Low Load 600 RPM Mid Load 1200 RPMRPM [RPM] 600 600 1200EVO [aTDC] 140 140 140EVC [aTDC] 370 370 370IVO [aTDC] 340 340 310IVC [aTDC] -125 -125 -175Exhaust MAP [kPa] 100.0 100.0 100.0Intake MAP [kPa] 53.4 65.5 67.6m_air [mg/cycle] 144 208 181m_fuel [mg/cycle] 14.5 18 18IGN [aTDC] -60 -45 -60T_exh [C] 334 407 474

IMEP [kPa] 156.40 269.82 260.59COV of IMEP [%] 3.89 2.46 1.84PMEP [kPa] 45.37 31.11 32.21Peak Pressure [kPa] 896 1385 1256Location of PP [aTDC] 16 15 15

0-10% HR [CAD] 58.0 42.5 56.510-90% HR [CAD] 74.5 66.5 83.5

[CO2] (dry) [%] 8.39 11.32 10.32[CO] (dry) [%] 6.26 3.92 4.96[O2] (dry) [%] 0.88 0.61 0.67[HC] (dry) [ppm C1] 14463 5697 6879n_comb [ ] 0.771 0.899 0.891measured AFR [ ] 10.81 12.48 11.95K_exh [ ] 0.863 0.863 0.865

[CO2] SV [%] 2.91 3.29 3.79K_cc [ ] 0.961 0.966 0.960Y_r [%] 38.66 32.54 40.80

INTAKE CAM ADVANCE

147

600 RPM Low Load 600 RPM Mid Load 1200 RPMRPM [RPM] 600 600 1200EVO [aTDC] 150 150 180EVC [aTDC] 380 380 410IVO [aTDC] 350 350 350IVC [aTDC] -135 -135 -135Exhaust MAP [kPa] 100.0 100.0 100.0Intake MAP [kPa] 55.8 67.9 70.3m_air [mg/cycle] 144 208 181m_fuel [mg/cycle] 14.5 18 18IGN [aTDC] -65 -45 -65T_exh [C] 336 389 435

IMEP [kPa] 157.79 271.53 240.73COV of IMEP [%] 3.77 2.64 3.64PMEP [kPa] 43.54 29.07 35.07Peak Pressure [kPa] 929 1385 1218Location of PP [aTDC] 14 15 14

0-10% HR [CAD] 60.5 41.5 59.510-90% HR [CAD] 78.0 69.5 92.5

[CO2] (dry) [%] 8.16 11.07 7.89[CO] (dry) [%] 6.36 4.31 6.65[O2] (dry) [%] 0.90 0.54 0.76[HC] (dry) [ppm C1] 15483 5766 12666n_comb [ ] 0.758 0.899 0.793measured AFR [ ] 10.68 12.27 10.66K_exh [ ] 0.863 0.863 0.869

[CO2] SV [%] 2.92 3.23 3.20K_cc [ ] 0.960 0.965 0.958Y_r [%] 39.85 32.67 44.78

EXHAUST CAM RETARD

148

600 RPM Low Load 600 RPM Mid Load 1200 RPMRPM [RPM] 600 600 1200EVO [aTDC] 130 130 130EVC [aTDC] 360 360 360IVO [aTDC] 360 360 360IVC [aTDC] -125 -125 -125Exhaust MAP [kPa] 100.0 100.0 100.0Intake MAP [kPa] 44.8 58.6 50.7m_air [mg/cycle] 144 208 181m_fuel [mg/cycle] 14.5 18 18IGN [aTDC] -45 -25 -35T_exh [C] 363 418 490

IMEP [kPa] 141.57 255.14 229.25COV of IMEP [%] 2.99 1.20 0.87PMEP [kPa] 51.53 39.21 53.27Peak Pressure [kPa] 977 1306 1182Location of PP [aTDC] 13 17 15

0-10% HR [CAD] 40.0 26.5 33.010-90% HR [CAD] 60.0 58.5 72.0

[CO2] (dry) [%] 7.92 10.51 8.37[CO] (dry) [%] 6.56 4.75 6.36[O2] (dry) [%] 0.92 0.71 0.75[HC] (dry) [ppm C1] 10551 5253 6858n_comb [ ] 0.821 0.906 0.877measured AFR [ ] 10.93 12.19 11.24K_exh [ ] 0.872 0.867 0.876

[CO2] SV [%] 2.04 2.29 1.64K_cc [ ] 0.971 0.973 0.977Y_r [%] 28.70 24.47 21.87

ZERO OVERLAP

149

Appendix B – Image Statistics

Filtered Unfiltered Filtered Unfiltered Filtered Unfiltered

30 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.55 8.40 4.46 6.48 4.18 6.07Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.10 4.50 1.71 3.50 1.65 3.19

45 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 4.62 8.94 4.01 6.35 3.99 6.35Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.20 4.78 1.81 3.87 1.71 3.59

60 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.22 9.56 4.10 6.82 4.09 6.82Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.71 5.99 1.94 4.16 1.86 4.06

99 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 6.81 12.47 5.26 8.72 5.90 9.26Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 3.13 7.14 2.34 5.22 2.31 5.22

600 RPM Low Load 1200 RPM600 RPM Mid LoadBASELINE OVERLAP

142.3

67.8

49.5

70.817.65

25.02

17.27

23.82

20.11

128.9

91.9

12.03

14.76

113.822.38

196.829.43

126.023.55

241.432.59

75.420.27

169.227.29

93.4

110.322.03

221.131.19

137.124.56

289.135.67

18.21

112.022.20

16.25

20.21

32.9 62.6 60.0

65.3 114.2 92.822.42

16.60

16.95

150

Filtered Unfiltered Filtered Unfiltered Filtered Unfiltered

30 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.93 7.75 5.17 7.60 8.24 10.71Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 1.73 3.35 1.88 3.81 1.72 3.47

45 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.19 7.47 4.99 8.22 8.41 11.95Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 1.92 3.88 2.15 4.61 1.90 4.09

60 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.15 8.17 4.70 8.00 7.31 11.61Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.18 4.46 2.19 4.77 2.29 5.10

99 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 6.68 10.56 6.08 10.05 9.37 15.44Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.56 5.65 2.62 5.87 2.51 5.64

SYMMETRIC OVERLAP INCREASE600 RPM Low Load 600 RPM Mid Load 1200 RPM

110.9 97.4 70.722.09 20.70 17.64

256.7 212.5 245.333.61 30.58 32.85

93.8 73.0 50.120.32 17.92 14.85

192.6 196.8 172.629.11 148.00 27.56

71.2 67.0 42.817.70 17.17 13.72

151.0 127.8 116.825.78 23.71 22.67

47.3 47.9 28.614.43 14.52 11.22

96.2 89.8 100.520.57 19.88 21.03

151

Filtered Unfiltered Filtered Unfiltered Filtered Unfiltered

30 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 6.03 8.52 4.07 6.30 5.92 8.13Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.04 4.08 1.67 3.44 1.61 3.14

45 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.34 8.27 4.09 6.65 7.09 9.31Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 1.93 4.14 1.76 3.78 1.59 3.24

60 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.77 9.67 4.49 7.56 7.78 10.30Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.26 5.01 1.97 4.31 1.74 3.62

99 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 7.45 12.73 5.64 9.95 8.02 12.19Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 3.05 6.86 2.48 5.64 2.12 4.65

INTAKE CAM ADVANCE600 RPM Low Load 600 RPM Mid Load 1200 RPM

89.8 121.7 95.019.88 23.14 20.45

194.6 249.5 302.629.26 33.13 36.49

82.8 106.0 84.119.09 21.60 19.24

181.3 210.8 273.028.24 30.46 34.66

62.2 81.3 70.516.54 18.91 17.61

134.1 169.8 218.224.29 27.33 30.99

42.1 56.8 41.613.61 15.81 13.53

84.5 110.1 136.519.28 22.01 24.51

152

Filtered Unfiltered Filtered Unfiltered Filtered Unfiltered

30 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.85 7.71 5.46 9.02 8.74 10.74Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 1.79 3.54 1.74 5.13 1.60 3.17

45 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 4.96 7.16 4.92 7.16 8.71 10.88Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 1.83 3.70 1.81 3.73 1.62 3.44

60 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.92 8.46 4.95 7.47 8.90 11.60Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.04 4.19 1.98 4.10 1.75 3.78

99 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 7.44 10.94 6.64 9.72 10.77 15.41Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.48 5.42 2.38 5.11 2.19 4.88

EXHAUST CAM RETARD600 RPM Low Load 600 RPM Mid Load 1200 RPM

122.1 127.5 92.023.18 23.69 20.12

258.4 252.5 307.133.72 33.33 36.76

102.9 103.7 82.121.28 21.36 19.01

212.1 209.5 254.530.55 30.36 33.46

79.7 93.4 66.618.73 20.27 17.12

167.5 182.2 208.727.15 28.31 30.30

51.8 61.8 39.515.10 16.49 13.18

106.0 119.7 136.921.60 22.95 24.54

153

Filtered Unfiltered Filtered Unfiltered Filtered Unfiltered

30 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.21 7.09 4.46 6.02 4.34 6.18Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 1.84 3.67 1.68 3.34 1.74 3.50

45 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 4.37 6.84 3.82 5.82 3.95 6.08Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 1.92 3.98 1.72 3.62 1.65 3.46

60 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 4.44 7.50 4.03 6.63 4.14 6.64Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.18 4.58 1.94 4.18 1.84 4.00

99 bTDC Mean Data Image SignalMean Data SNRFired <(σ_y/µ_y)> 5.47 9.63 5.39 8.77 6.07 9.38Mean FF Image SignalMean FF SNRFlatfield <(σ_y/µ_y)> 2.75 6.03 2.51 5.63 2.46 5.54

ZERO OVERLAP600 RPM Low Load 600 RPM Mid Load 1200 RPM

118.0 156.0 137.222.79 26.20 24.57

219.7 261.1 238.731.09 33.89 32.41

100.3 132.0 123.921.01 24.10 23.35

187.7 218.8 236.628.74 31.03 32.27

78.5 97.3 98.618.58 20.69 20.83

145.1 167.2 178.625.27 27.12 28.03

50.8 62.7 60.214.95 16.61 16.28

91.6 100.4 104.620.08 21.02 21.45