The University of Toledo
The University of Toledo Digital Repository
Theses and Dissertations
2005
Measurements of spark ignition energy of n-octane
and i-octane
Lisa M. Rimpf
The University of Toledo
Follow this and additional works at: http://utdr.utoledo.edu/theses-dissertations
Recommended Citation
Rimpf, Lisa M., "Measurements of spark ignition energy of n-octane and i-octane" (2005). Theses and Dissertations. 1458.
http://utdr.utoledo.edu/theses-dissertations/1458
This Thesis is brought to you for free and open access by The University of Toledo Digital Repository. It has been accepted for inclusion in Theses and
Dissertations by an authorized administrator of The University of Toledo Digital Repository. For more information, please see the repository's About
page.
A Thesis
Entitled
MEASUREMENTS OF SPARK IGNITION ENERGY OF n-OCTANE AND i-OCTANE
by
Lisa M. Rimpf
Submitted as partial fulfillment of the requirements for
the Master of Science in Chemical Engineering
____________________________________
Advisor: Dr. Kenneth J. DeWitt
____________________________________
Advisor: Dr. Martin J. Rabinowitz
____________________________________
Dr. Soon Muk Hwang
____________________________________
Dr. Constance A. Schall
____________________________________
Graduate School
The University of Toledo
December 2005
The University of Toledo
College of Engineering
I HEREBY RECOMMEND THAT THE THESIS PREPARED UNDER MY
SUPERVISION BY Lisa M. Rimpf
ENTITLED
MEASUREMENTS OF SPARK IGNITION ENERGY
OF n-OCTANE AND i-OCTANE
BE ACCEPTED IN PARTIAL FULFILLMENT OF THE REQUIREMENTS FOR
THE DEGREE OF
Master of Science in Chemical Engineering
Thesis Advisor: Dr. Kenneth J. DeWitt, The University of Toledo
Thesis Advisor: Dr. Martin J. Rabinowitz, The NASA Glenn Research Center
Recommendation concurred by Committee on Final Examination
Dr. Soon Muk Hwang
Dr. Constance A. Schall
Dr. Mohamed Samir Hefzy, Associate Dean of Graduate Studies, College of Engineering
ABSTRACT
MEASUREMENTS OF SPARK IGNITION ENERGY OF n-OCTANE AND i-OCTANE
Lisa M. Rimpf
Submitted as partial fulfillment of the requirements for
The Master of Science in Chemical Engineering Degree
The University of Toledo
December 2005
Spark ignition energies were examined for various small hydrocarbons in the
1940’s and 1950’s related to mine gas explosions. In 1996, the TWA flight 800 center
wing tank explosion focused interest on the measurement of aviation fuel minimum
ignition energy. The goal of this study is to obtain spark ignition energy data and acquire
the resultant pressure rise for n-octane and i-octane combustion as the selected species of
jet fuel. Using a composite electrical spark system, ignition energy was plotted versus
dc-arc spark duration time while varying the fuel/air mixture equivalence ratio, φ. For
rich mixtures of i-octane, the minimum ignition energy decreased with spark duration
time, reached a minimum value, and then further increased as expected; while results for
lean mixtures of n-octane are not so apparent, displaying abrupt irregular behavior. The
minimum of the minimum ignition energy (commonly called minimum ignition energy,
MIE) for i-octane at φ≅2 was 1.5 mJ which is close to the literature value of 1.35 mJ.
Also, reducing the oxygen content appears to raise the minimum ignition energy and
lessen the pressure rise. The ignition process (ignition + successful flame propagation)
iii
was analyzed by considering energy furnished from the composite spark and heat release
from chemical reactions opposing heat conduction energy losses to the unburned gas and
electrodes. Extending and preserving the ignition kernel (plasma kernel) with dc-arc
energy is critical for successful flame propagation.
In this regard, dc-arc energy
deposition rate is more important than the absolute energy supply. Overall, it was shown
that ignition energy is dependent upon equivalence ratio as well as spark duration time
while pressure rise is also subject to the stoichiometry.
iv
ACKNOWLEDGEMENTS
I would like to extend my deepest gratitude to the many people that have assisted
me through this journey of graduate school.
Without everyone’s unwavering
encouragement this project would not have been possible. Additional commendation
goes to The University of Toledo College of Engineering as well as The NASA Glenn
Research Center for the finances to facilitate this research.
Furthermore, I am grateful to all The University of Toledo faculty members that
have instructed me throughout my college career. Special recognition is owed to Dr.
Kenneth J. DeWitt for his guidance as my graduate advisor and all the knowledge that he
shared with me; may you be granted sustained health. I am indebted to Dr. Martin J.
Rabinowitz for teaching me the intricacies of fuel ignition chemistry and guiding me to
perform experimental work at his laboratory. Thanks to Dr. Soon Muk Hwang for his
constant advice. The mentorship of these individuals has been invaluable.
I appreciate the assistance of Mr. Bryan Knepper who worked on the precursor
testing and trained me with the initial experimental operation. Mr. Gregg Calhoun, Mr.
Bob Bickford, and Mr. Yves Lamothe should be credited for their daily technical services
and instruction of practical skills.
My family and friends have continually demonstrated support and let me follow
my dreams. To my parents, Franz and Hermine, for always believing in me and to my
siblings, Tony and Renee, for your understanding – I love you!
v
TABLE OF CONTENTS
ABSTRACT ................................................................................................................. iii
ACKNOWLEDGEMENTS...........................................................................................v
TABLE OF CONTENTS .............................................................................................vi
LIST OF FIGURES................................................................................................... viii
LIST OF TABLES.........................................................................................................x
INTRODUCTION .........................................................................................................1
LITERATURE REVIEW..............................................................................................3
OBJECTIVE..................................................................................................................7
EXPERIMENTAL.........................................................................................................8
Combustion Vessel .....................................................................................................9
Electrodes .................................................................................................................11
Igniter Circuit...........................................................................................................12
Data Acquisition .......................................................................................................14
Temperature Management System..........................................................................17
Gas Management System .........................................................................................17
Gas Chromatograph.................................................................................................20
Control System .........................................................................................................20
Vapor Pressure .........................................................................................................22
Equilibrium ..............................................................................................................28
Procedure..................................................................................................................28
RESULTS ....................................................................................................................29
Spark Gap Determination........................................................................................30
Methanol Validation Testing ...................................................................................31
i-octane Testing.........................................................................................................34
n-octane Testing .......................................................................................................44
vi
DISCUSSION ..............................................................................................................51
CONCLUSION............................................................................................................58
REFERENCES ............................................................................................................59
APPENDICES .............................................................................................................62
APPENDIX A: Fuel Tank Ignitions...........................................................................63
APPENDIX B: Gas Chromatograph Chemstation Reports......................................76
APPENDIX C: Testing Procedure.............................................................................78
APPENDIX D: Data Tables ......................................................................................84
Methanol at 32.5ºC (φ=2.30) in 21% oxygen...........................................................84
Methanol at 22.5ºC (φ=1.23) in 21% oxygen...........................................................85
i-octane at 19.0ºC (φ=3.13) in 21% oxygen .............................................................86
i-octane at 16.5ºC (φ=2.83) in 21% oxygen .............................................................87
i-octane at 14.5ºC (φ=2.61) in 21% oxygen .............................................................88
i-octane at 8.0ºC (φ=2.02) in 21% oxygen ...............................................................89
i-octane at 19.0ºC (φ=3.65) in 18% oxygen .............................................................90
i-octane at 16.5ºC (φ=3.30) in 18% oxygen .............................................................90
n-octane at 35.9ºC (φ=2.02) in 21% oxygen ............................................................91
n-octane at 19.8ºC (φ=1.06) in 21% oxygen ............................................................91
n-octane at 18.3ºC (φ=1.00) in 21% oxygen ............................................................92
n-octane at 17.0ºC (φ=0.95) in 21% oxygen ............................................................93
vii
LIST OF FIGURES
Figure 1: Combustion vessel testing apparatus................................................................8
Figure 2: Top Hemisphere ..............................................................................................9
Figure 3: Center Cylinder ...............................................................................................9
Figure 4: Bottom Hemisphere.........................................................................................9
Figure 5: Bottom hemisphere close-up of threaded bolt ................................................10
Figure 6: Electrode Arrangement ..................................................................................11
Figure 7: Combustion vessel center section showing electrodes and mixer....................12
Figure 8: Circuit Diagram.............................................................................................13
Figure 9: TDS1 Oscilloscope sample output .................................................................16
Figure 10: TDS2 Oscilloscope sample output ...............................................................16
Figure 11: Gas Management System Schematic............................................................19
Figure 12: Wonderware® Touch Screen Capture..........................................................21
Figure 13: i-octane vapor pressure curve versus temperature ........................................24
Figure 14: i-octane vapor pressure curve vs. temperature explosive region ...................24
Figure 15: n-octane vapor pressure curve versus temperature........................................26
Figure 16: n-octane vapor pressure curve vs. temperature explosive region...................26
Figure 17: Calibration curve of low-speed pressure transducer deviation ......................28
Figure 18: Combustion vessel before and after modification.........................................29
Figure 19: Electrode orientation before and after relocation ..........................................29
Figure 20: Bench top investigation of breakdown voltage versus spark gap ..................30
Figure 21: Curve fit methanol ignition data at 32.5ºC (φ=2.30) in 21% oxygen.............32
Figure 22: Curve fit methanol ignition data at 22.5ºC (φ=1.23) in 21% oxygen.............33
Figure 23: Methanol equivalence ratio versus minimum ignition energy......................34
Figure 24: i-octane ignition data at 19.0ºC (φ=3.13) in 21% oxygen .............................35
viii
Figure 25: i-octane ignition data at 16.5ºC (φ=2.83) in 21% oxygen .............................36
Figure 26: i-octane ignition data at 14.5ºC (φ=2.61) in 21% oxygen .............................37
Figure 27: i-octane ignition data at 8.0ºC (φ=2.02) in 21% oxygen ...............................38
Figure 28: i-octane equivalence ratio versus minimum ignition energy .........................39
Figure 29: i-octane ignition data at 19.0ºC with oxygen content of 18% and 21% .........40
Figure 30: i-octane ignition data at 16.5ºC with oxygen content of 18% and 21% .........41
Figure 31: i-octane ignition data showing temperature variation effects ........................42
Figure 32: i-octane pressure rise ignition data versus spark duration .............................43
Figure 33: i-octane pressure rise data range versus equivalence ratio ............................44
Figure 34: n-octane ignition data at 35.9ºC (φ=2.02) in 21% oxygen.............................45
Figure 35: n-octane and i-octane ignition data at φ ≈ 2.0 in 21% oxygen.......................46
Figure 36: n-octane ignition data at 19.8ºC (φ=1.06) in 21% oxygen.............................47
Figure 37: n-octane ignition data at 18.3ºC (φ=1.00) in 21% oxygen.............................48
Figure 38: n-octane ignition data at 17.0ºC (φ=0.95) in 21% oxygen.............................48
Figure 39: n-octane ignition temperature variation effects in 21% oxygen ....................49
Figure B-1: Gas chromatograph species concentration of cylinder air– Trial 1..............76
Figure B-2: Gas chromatograph species concentration of cylinder air– Trial 2..............76
Figure B-3: Gas chromatograph species concentration of cylinder air– Trial 3..............76
Figure B-4: Gas chromatograph species concentration of reduced oxygen– Trial 1 .......77
Figure B-5: Gas chromatograph species concentration of reduced oxygen– Trial 2 .......77
Figure B-6: Gas chromatograph species concentration of reduced oxygen– Trial 3 .......77
Figure C-1: Custom spark gap measuring device ..........................................................78
Figure C-2: Installed insulating foam............................................................................79
Figure C-3: Variable resistor and capacitor placement ..................................................81
Figure C-4: Photograph of circuit box ..........................................................................81
ix
LIST OF TABLES
Table 1. i-octane Antoine constants for designated temperature range ...........................23
Table 2. i-octane Antoine constants for designated temperature range ...........................23
Table 3. i-octane vapor pressure from the Handbook of Chemistry and Physics ............23
Table 4. n-octane Antoine constants for designated temperature range ..........................25
Table 5. n-octane Antoine constants for designated temperature range ..........................25
Table 6. n-octane vapor pressure from the Handbook of Chemistry and Physics............25
Table 7. Methanol validation testing conditions ............................................................31
Table 8. i-octane testing conditions ...............................................................................35
Table 9. n-octane testing conditions ..............................................................................45
Table D-1. Methanol at 32.5ºC (φ=2.30) in 21% oxygen...............................................84
Table D-2. Methanol at 22.5ºC (φ=1.23) in 21% oxygen...............................................85
Table D-3. i-octane at 19.0ºC (φ=3.13) in 21% oxygen .................................................86
Table D-4. i-octane at 16.5ºC (φ=2.83) in 21% oxygen .................................................87
Table D-5. i-octane at 14.5ºC (φ=2.61) in 21% oxygen .................................................88
Table D-6. i-octane at 8.0ºC (φ=2.02) in 21% oxygen ...................................................89
Table D-7. i-octane at 19.0ºC (φ=3.65) in 18% oxygen .................................................90
Table D-8. i-octane at 16.5ºC (φ=3.30) in 18% oxygen .................................................90
Table D-9. n-octane at 35.9ºC (φ=2.02) in 21% oxygen.................................................91
Table D-10. n-octane at 19.8ºC (φ=1.06) in 21% oxygen...............................................91
Table D-11. n-octane at 18.3ºC (φ=1.00) in 21% oxygen...............................................92
Table D-12. n-octane at 17.0ºC (φ=0.95) in 21% oxygen...............................................93
x
INTRODUCTION
Trans World Airlines (TWA) flight 800 piqued the interest of investigators after
the Boeing 747 exploded not long after take off on July 17, 1996 near Long Island, New
York. An official report was finally issued by the National Transportation Safety Board
(NTSB) in August 2000 which describes the most probable cause of the accident to be
“an explosion of the center wing fuel tank (CWT), resulting from ignition of the
flammable fuel/air mixture in the tank. The source of ignition energy for the explosion
could not be determined with certainty, but, of the sources evaluated by the investigation,
the most likely was a short circuit outside of the CWT that allowed excessive voltage to
enter it through electrical wiring associated with the fuel quantity indication system.”1
Collected physical evidence indicated that an overpressure event occurred in the CWT,
which is defined as a sharp increase in pressure caused by the explosion during a
relatively short time, compromising the structural integrity.
Other fuel tank explosions have also been noted on both commercial and
government aircraft as well. A detailed description of these other mishaps can be found in
Appendix A. Although the ignition sources differ and the outcomes are generally less
severe, the common goal remains the same – how to keep an aircraft safe from explosive
conditions.
While air travel has been employed as a mode of transportation for more than fifty
years, the specifics of current aviation fuels are still nebulous to a certain extent.
1
2
Regulated by ASTM D1655, the composition criteria for aviation fuels is rather vague,
citing only that they “consist of refined hydrocarbons derived from conventional sources
including crude oil, natural gas liquid condensates, heavy oil, shale oil, and oil sands.”
Ultimately a blend of several hundred different types of hydrocarbons comprises the final
product. Specific physical properties have established maximum/minimum values for
critical parameters such as flash point, boiling point, freezing point, as well as other less
vital specifications (i.e., sulfur content, corrosion, etc.).
With these guidelines, the
product leaves the refinery. Even though these methods are in place, a need has been
emerging to more deeply understand the behavior of the fuel once it is placed in the
aircraft.
LITERATURE REVIEW
Historically, research on ignition characteristics did not find origins with the
aviation industry but rather at the U.S. Bureau of Mines when it opened in 1910. This
organization was founded by Congress under the Department of the Interior to conduct
research and collect information concerning every aspect of the mining trade. As a part
of this group, Lewis and von Elbe et al.2 can be acknowledged for the advancement of
spark ignition laboratory experimentation from their study of mine gas (methane/air
mixture) in relation to a quenching distance. Guest,3 a contemporary, was the instrument
behind the test apparatus construction. This research, completed in the late 1940’s and
1950’s, has been the precedent for minimum ignition energy research and further
applications.
Lewis and von Elbe state in their renowned text4 that “It is possible to pass small
electric sparks through an explosive gas without producing ignition. When the spark
energy is increased, a threshold energy is eventually obtained at which the spark becomes
incendiary in the sense that a combustion wave propagates from the spark through the
volume of gas.” This realization was the defining criteria for the meaning of minimum
ignition energy (MIE) where a non-igniting spark must penetrate the gas mixture in the
test apparatus prior to a successful characterized ignition. Lewis and von Elbe5 also
developed an early theory for flame propagation from an instantaneous point source of
3
4
ignition; thereby, giving rise to the stepwise understanding of ignition kernel formation or
dissipation.
Their experiments were conducted by investigating critical spark gap distance as
well as pressure variation within the 5-inch diameter stainless steel “test bomb.” Two
electrodes were mounted centrally, either with stainless steel tips or glass flanges; one
was attached to a micrometer for spark gap adjustment. High voltage was supplied by a
0-30 kV DC power supply via a 500 MΩ protective resistor to a rotary charger which
discharged through the capacitor busbar. The voltage and capacitance were increased
until a spark ignited the pre-mixed gas of known composition.
These early trials
established that experimental set-up dictates the outcome of the results: spark gap
distance, supply voltage, electrode construction, and vapor pressure.
In addition to
methane, other hydrocarbon data was also published: ethane, propane/cyclopropane,
butane/diethyl ether, hexane/cyclohexane/benzene, heptane, and hydrogen.
Metzler of the Lewis Flight Propulsion Laboratory of the National Advisory
Committee for Aeronautics (NACA) investigated minimum ignition energy for pure
hydrocarbons in the C2-C6 series.6 He studied many of the same substances (ethane,
ethylene, acetylene, n-hexane, cyclohexane, and benzene) with a similar method as Lewis
and von Elbe, achieving somewhat differing results.
attributed to the variation in electrode design.
This conflicting data can be
The needle-like electrodes constructed
from “number-74 drills soldered into 1/8-inch brass rod” and “sealed into a 1-inch
diameter Lucite insulator” keep the spark from straying to ground.6 This small departure
confirms that changes in experimental set-up can affect the results obtained.
5
Calcote and associates7 performed the most extensive early studies related to
spark minimum ignition energy and also referenced the work of Lewis and von Elbe.
They delved into understanding the affect of experimental equipment as well as the
procedure on resultant ignition energies and produced/analyzed a multitude of data for
structurally different hydrocarbon molecules. A relevant detail from their experimental
procedure is the spark gap irradiation using ultraviolet light to furnish photoelectrons.
This reduces the spark breakdown lag and pioneers the concept of the bait electrode used
in this thesis.
Research performed by Litchfield8,9 for the Federal Bureau of Mines at the
Explosives Research Laboratory scrutinized the type of spark ignition source and the
relationship between the spark gap and electrode geometries using Schlieren images.
These chronological photographs allowed for a greater visual understanding of what
occurs at the spark gap with respect to the ensuing shock wave from electric discharge,
ignition kernel development, and ultimate energy dissipation.
He analyzed these
applications of minimum ignition energy to safety engineering although stating in a
conclusion that “to prevent spark ignition by preventing energy from accumulating
appears practically hopeless.”
Between the 1950’s and the following 40 years, progress with spark ignition
research was practically idle.
During this period, studies focused on issues of
flammability as well as a transition away from volatile JP-4 military fuel. Around 1970,
both Nestor10 and Kosvic et al.,11 in separate experiments, investigated the aspects of fire
within aircraft fuel tanks by simulating conditions of altitude and temperature to
6
determine fuel/air ratios in the ullage. While this research is not directly relevant, it
shows the progress of aircraft safety.
Contemporary spark ignition tests were exhaustively investigated with Jet-A fuel
at The California Institute of Technology by Shepard and colleagues under the auspices
of the NTSB in order to better comprehend the TWA 800 accident. They were concerned
with reproducing the flight conditions of the incident aircraft to determine probable cause
using fundamental preliminary experiments12,13,14 as well as scaled recreations.15,16
Through hundreds of pages of laboratory documentation discussing temperature, vapor
pressure, mass loading, and weathering some results can be summarized to say that flash
point is not a useful characterization of explosion hazard and that MIE is a strong
function of composition. Vapor composition of multicomponent fuel is very different
than the bulk liquid. Their ¼-scale CWT replication experiment substantiated that a
flame front can propagate rapidly through the closed vessel and provide sufficient
pressure rise to cause failure of structural components.
Because difficulties are sometimes posed with spark ignition, certain
researchers17,18,19 made a transition to laser ignition. “Laser sparks provide a noninvasive
method of fuel-air ignition that can be directed to optimal locations within the
combustion chamber” and do not have the dependence on electrode geometry and
circuitry.18 By comparison though, minimum ignition energy values are appreciably
higher with laser ignition when measured against spark ignition even though the ignition
process is the same. Explaining this phenomenon still needs clarification but has some
theories related to physical differences of electromagnetism and thermal conditions.
OBJECTIVE
The document presented here is in direct succession to previous work done in this
laboratory with regard to methanol minimum ignition energy determination by Bryan E.
Knepper.20 The goal of this study is to obtain the spark ignition energies together with
the pressure rise after ignition of n-octane and i-octane as the selected jet fuel component
species.
Unfortunately, there is little to no data available for octane isomer related
research. An accepted published value for i-octane minimum ignition energy in air is
1.35 mJ.21 However, for n-octane a reasonable estimate can be established from looking
at other alkane hydrocarbons which have been studied. A range of 0.2–0.3 mJ has been
compiled by Kuchta22,23 acknowledging that the industry estimate of 0.25 mJ is feasible.
Given that jet fuel is a refined natural product, it in no manner represents pure
components; therefore, concrete behavioral predictions cannot be made because of origin
dependence. However, jet fuel properties can potentially be modeled with a synthetic
blend of hydrocarbons. The ultimate intent of this spark ignition research is to create a
surrogate fuel to model the performance of jet fuel and create safer operating conditions.
Individual component properties would first be measured (here, n-octane and i-octane)
followed by binary and ternary, etc. mixtures until the scope of synthetic jet fuel can be
understood. While progress related to ignition characteristics is advancing, results are not
adequate for adaptation to a feasible system for accident prevention.
7
EXPERIMENTAL
The Fire Safe Fuel (fsf) experimental laboratory is located at the National
Aeronautics and Space Administration (NASA) Glenn Research Center in the Engine
Research Building (ERB) #5: test cell CE-13A and control room CE-15.
Figure 1: Combustion vessel testing apparatus
Liquid fuel (here, n-octane and i-octane) is vaporized and ignited in a stainless
steel combustion vessel using an electric spark. The spark circuitry is designed to create
an ionization voltage (breakdown energy) followed by a variable high voltage pulse (arc
energy). Spark duration and energy are captured by an oscilloscope using current and
voltage probes. Combustion chamber pressure rise is recorded via a high-speed pressure
transducer.
8
9
Combustion Vessel
The combustion vessel is fabricated from grade 304 stainless steel. This common
low-carbon austenitic Chromium-Nickel alloy was chosen for its corrosion resistance and
general versatility.24 The four-liter chamber is comprised of three separate sections: top
hemisphere, center cylinder, and bottom hemisphere. This assembly is secured together
by bolts and safety wire and is suspended from an aluminum frame.
Figure 2: Top Hemisphere
Figure 3: Center Cylinder
Figure 4: Bottom Hemisphere
In compliance with NASA Glenn Research Center safety regulations, the vessel
was subjected to vigorous pressure testing before use. It has been approved for operating
conditions of P=200 psig and T=400oF, although it has passed hydrostatic pressure tests
to P=300 psig according to American Society of Mechanical Engineers (ASME) boiler
and pressure vessel coding. The pressure-tight seal between the sections is maintained
with a Viton rubber O-ring that prevents leaks when the bolts are fastened.
The top hemisphere has four available inlets which accommodate the
thermocouple, pressure transducers (high speed and low speed) as well as gas inlet and
exhaust lines. Three of these access points are utilized while one is available for future
use when additional experimental needs arise.
10
The bottom hemisphere retains the liquid fuel and does not possess any
monitoring equipment.
This section is frequently removed for cleaning between
experiments. A large threaded bolt was placed in the bottom, covering the viewing
portal, which enlarges the surface area and expedites the liquid fuel evaporation process.
See Figure 5.
Figure 5: Bottom hemisphere close-up of threaded bolt
Additionally there are four circular 1.375-inch diameter framed quartz viewing
portals. Two of these portals are located at the apex of each hemisphere to achieve a
cross sectional view of the spark region, while the other two windows are used to view
the electrodes and are located in the center cylinder of the vessel.
This design is
favorable for installing a high speed digital camera for capturing the fuel combustion
process and spark behavior.
The center cylinder of the combustion vessel holds the three electrodes (positive,
negative, and bait probe) as well as the mixer. To prevent vapor phase stratification, the
combustible gas mixture is stirred using a variable speed DC (direct current) motor
driving a shaft mounted propeller.
11
Electrodes
Various materials were chosen for the electrodes to mechanically and electrically
isolate them from the combustion vessel housing as well as reduce electric noise. For
each electrode, Daburn 20-10 Corona Resistant Teflon (CRT) high voltage wire was
soldered to a 0.0045-inch diameter stainless steel sharpened tip welding rod and potted in
Macor, machinable glass ceramic, with Torr Seal Low Vapor Pressure Resin. To
provide a pressure-tight seal, each electrode assembly is secured to the vessel cavity
using a stainless steel Swagelok fitting and Teflon ferrule sets. The arrangement of
the three electrodes is such that they are in the same plane. The positive and negative
electrodes are 3mm apart, while the bait probe is spaced closer to the negative electrode.
It is important to note that the orientation of the bait probe must be perpendicular to the
line connecting the positive and negative electrodes, as can be seen in the diagram.
3mm
Macor
wire
stainless steel
electrode tip
BAIT
PROBE
Figure 6: Electrode Arrangement
A photograph of the vessel center cylinder (Figure 7) shows the electrodes and mixer,
with soot deposits, after a successful ignition.
12
viewing
window
mixer
bait probe
Figure 7: Combustion vessel center section showing electrodes and mixer
Igniter Circuit
The spark ignition circuitry is housed in an aluminum box and covered by a
hinged LEXAN™ polycarbonate faceplate. This configuration allows easy access to the
components while simultaneously addressing safety issues. The operating status of the
different electronic sections is indicated by a series of light emitting diodes (LEDs).
Consisting of two independently adjustable direct current power supplies, the spark
ignition circuit serves a dual role: an ionization breakdown energy and high voltage arc
energy generation.
The electricity applied to the bait electrode provides a breakdown energy
(<0.1 mJ) from the positively referenced bait probe to the negative electrode. This
creates a plasma cloud in the region of the electrodes, thereby, facilitating propagation of
the main arc spark from the positive electrode (anode) to the negative electrode
(cathode). Ionization voltage is supplied by a 4-Watt UltraVolt “AA” series 0-1000
volt variable DC to DC power supply and directed to a Perkin Elmer TR-1855 external
trigger transformer.
13
High voltage is obtained from an UltraVolt “A” series 0-10,000 volt variable DC
to DC power supply. The ionization circuit is fixed while the arc circuit can be modified
for spark duration or energy by varying the values of capacitor (Cvar), resistor (Rvar), as
well as adjusting the supplied voltage. An assortment of capacitor and resistor values is
available; intermediate values can be obtained by placing them either in series or parallel.
Recall the following equations that govern the basics of resistive/capacitive electrical
behavior:
CT = C1 + C 2 + C n
Equation 1:
Capacitors in Parallel
Wonderware
remote
local
CT =
1
1
C1
+ 1
C2
+ 1
RT = R1 + R 2 + R n
Cn
Equation 2:
Capacitors in Series
IONIZATION
Ultravolt AA
DC/DC converter
1000 volt
RT =
Equation 3:
Resistors in Series
1
1
R1
0.22µF
100kΩ
PF
SF
bait
PS
SS
12 volts DC
Schmitt Trigger Circuit
Debouncer
56µs pulse
fires SCR
Wonderware push button spark trigger
connected to PLC relay
remote
local
HIGH VOLTAGE
Ultravolt A
DC/DC converter
10,000 volt
R2
Rvar
2MΩ
20MΩ
24 volts DC
Figure 8: Circuit Diagram
Cvar
+ 1
Rn
Equation 4:
Resistors in Parallel
SCR
Wonderware
+ 1
14
Data Acquisition
Two digital Tektronix TDS-3014B four-channel oscilloscopes are used for
acquiring the data: current, voltage, and pressure. The oscilloscope labeled TDS1 detects
the voltage applied across the spark gap on Channel 2 and the current flow on Channel 3.
Obtaining the current and voltage traces allows for a MATH function calculation to
determine the spark energy supplied to the electrodes. The second oscilloscope (TDS2)
depicts the pressure profile after ignition. Either the voltage reading or current signal was
used to trigger the oscilloscopes for data capture.
A Tektronix A-6303 DC to 15 MHz current probe measures the current flow from
the positive electrode wire. This reading is then directed to a Tektronix AM-503B
Current Probe Amplifier. The amplifier converts the sensed current into a proportional
voltage signal that can be measured directly with an oscilloscope25. Care must be taken
to degauss as well as zero the amplifier prior to experimentation to reduce measurement
error. “Degaussing the probe removes any residual magnetization from the probe core.”26
Voltage is quantified across the negative electrode using a Tektronix P6015 high
voltage probe with 1000X attenuation.
This probe is frequency compensated and
connects directly to the oscilloscope. Caution should be exercised that the grounding clip
is attached to the circuit for electrical safety.
The current and voltage measurements allow for the calculation of the energy
supplied. The energy supplied to the electrodes can be calculated by summing the
ionization energy and the arc energy. Because the ionization energy is constant, it can be
2
calculated using Eion = 12 CionVion , where Cion is the capacitance (Farads) and Vion is the
voltage (volts) in the ionization circuit.
The ionization energy measured was circa
15
0.07 mJ. This value is negligible in comparison to arc energy. On the other hand, the arc
τ
energy is determined by integration E = ∫ P(t )dt from the oscilloscope output using the
0
MATH computation function, where P is the power (Watts) and τ is the spark duration
time. Power is defined by multiplying voltage (V) and current (I).
P(t) = V(t) · I(t)
(5)
Then the total energy observed in Joules is calculated using Equation 6.
E=
1
2
[C
]
τ
ion ⋅ Vion + ∫ [Varc (t ) ⋅ I arc (t ) ] dt
2
(6)
0
Where the MATH integration function entered in the oscilloscope is given by:
MATH = (Ch2 + VAR1) x (Ch3 x VAR2) x 1000
where:
(7)
VAR1 = voltage probe relationship = 2 (voltage input/1000)
VAR2 = current probe relationship = current probe setting/10
Ch2 = voltage signal
Ch3 = current signal
The pressure rise inside the vessel after ignition was obtained using a high speed
pressure transducer. A PCB Piezotronics ICP transducer is attached to a 482A22 fourchannel line powered signal conditioner. This quartz dynamic pressure sensor is designed
for use in shock tubes or blast wave measurements and is suited for this application as
well. The transducer signal is given in voltage and is converted to pressure (P) using the
calibration factor:
P = 25mV ≅ 1 psi
The following graphs demonstrate a typical ignition event.
(8)
16
τ
Spark duration = τ
MATH = Spark Energy
CH2
Voltage Trace
Current Trace
CH3
Spark
Energy
Figure 9: TDS1 Oscilloscope sample output
peak
pressure
value
Pressure Rise
Pressure Peak
time to peak
pressure
value
18.4 mV = 0.74 psi
pressure peak
at 576 ms
Figure 10: TDS2 Oscilloscope sample output
17
Temperature Management System
Temperature control of the combustion vessel is achieved with a 1m3 air
thermostat using a Julabo FP50-HP Refrigerated and Heating Circulator; the unit has a
working temperature range of -50oC < T < 200oC with temperature stability ±0.01oC. A
Pt100 Platinum resistance thermocouple measures the temperature inside the vessel,
defined here as the process variable. A PID (Proportional-Integral-Derivative) feedback
control loop evaluates the error between the process variable and the propylene glycol
bath set point. Temperature losses are minimized with 2-inch rigid, high-density and
high-temperature polyisocynaurate (Last-a-Foam® from General Plastics Manufacturing
Co.) insulating foam surrounding the combustion vessel. Four fan coil heat exchangers,
connected in series, circulate the air within the enclosure and maintain a homogeneous
and stable temperature.
Gas Management System
The gas management system regulates the flow of gases to and from the
combustion vessel.
Compressed air (denoted as oxygen) and nitrogen cylinders are
utilized as input gases; another line is designated for argon (inert), although not currently
implemented. An identical network of valves and mass flow controllers directs the
individual gas inlet flows through stainless steel tubing before mixing at a manifold.
Following a successful ignition, the combustion chamber is purged and vented to the
atmosphere through the exhaust valve, expelling combustion products and remaining
volatile reactants. A standard vessel purge is achieved by flowing one minute of nitrogen
at 20 standard liters per minute (slm) and air at 5 slm simultaneously and then continuing
with four additional minutes of air at the same flow rate. Upon completion of this cycle,
18
the vents are shut and the test vessel again becomes a closed system. This purge was also
enacted to clean the vessel from any unwanted contaminants when charged with new
fuel. An alternative purge could also be selected to simulate reduced oxygen content
(18% O2) as a representation of fuel tank inerting with 4.8 slm air and 0.8 slm nitrogen
for 5 minutes. To prevent an unexpected overpressure in the system at any time, an
emergency relief valve (rated at 150 psig) in parallel with a secondary burst disk (rated at
220 psig) is in place. A detailed schematic of the gas management system is illustrated in
Figure 11.
19
Figure 11: Gas Management System Schematic
20
Gas Chromatograph
The purge cycle allows for a known composition of nitrogen and oxygen gases to
enter the combustion vessel. A standard purge yields 21% oxygen since the contents of
the compressed air cylinder are known. However, the 18% oxygen purge requires a
mixture of air and nitrogen to obtain the reduced oxygen composition; this is regulated by
mass flow controllers. To ensure that the calculations for the mass flow controllers were
correct and that the environment was indeed 18% oxygen and 82% nitrogen, gas samples
were collected (no presence of fuel).
Following purge completion, gas samples were extracted using a syringe from a
temporary septum installed in an available port of the combustion vessel. The gas sample
was analyzed in a Hewlett Packard 5890 Series II gas chromatograph with a molecular
sieve column and thermal conductivity detector. The standard purge was analyzed as a
baseline for comparison. The results demonstrated that the perceived gas composition
was valid. See the complete Chemstation® software reports in Appendix B.
Control System
The instrumentation is managed by a Modicon® Programmable Logic Control
(PLC) system using Concept® software and controlled by a touch-screen computer
monitor operating Wonderware® Graphical User Interface (GUI) software.
This
computer technology allows for ease of operation and monitoring of the system from the
test cell (CE-13A) or the control room (CE-15). Several safety precautions are initially in
place prior to conducting tests: The control power in CE-15 must first be enabled, and
then the fire safe fuels operating screen must be selected from the computer menu. The
operating screen, displayed below, allows the user to execute the experiment.
21
Figure 12: Wonderware® Touch Screen Capture
The “MAIN Power” button allows power to the circuit box and must be selected
before either the ionization or high voltage is operable; this action then toggles the yellow
smiling face to a blue frowning face, signifying that high voltage and current are enabled.
Igniter circuit voltages are monitored real time and can be adjusted manually on the
circuit box, indicated on the screen as “LOCAL,” or directly on the touch screen and
displayed in green as “REMOTE.” Once the trial voltage has been set, the “SPARK”
button is the mechanism which permits voltage to the electrodes.
Valves and mass flow controllers regulating the flow of gases are also managed
from this screen; the standard purge or 18% oxygen purge are located on the upper right,
while manual adjustment of gas flow can be achieved by the meters. The outside exhaust
vent can also be open or closed. Additionally, the mixer can be activated as well as the
22
low speed pressure and ambient temperature observed. The “Scopes” button redirects to
a new page which then allows the user to select the current oscilloscope view. All
operations cease immediately should any of the following events occur: emergency stop
button activation, overpressure, or loss of power. Exhaust valves will fail open and
control power is disabled.
Vapor Pressure
The relationship between the temperature of a liquid and its vapor pressure is
nonlinear. Vapor pressure plots can be determined from theoretical calculations using the
Clausius-Clapeyron equation,
ln P =
where:
P=
∆Hvap =
R=
T=
C=
− ∆H vap
RT
+C
(9)
Pressure
Heat of Vaporization
Ideal Gas Constant
Temperature
Constant
The Clausius-Clapeyron equation can be used to construct the entire vaporization curve;
however, there is a marked deviation from experimental values because the enthalpy of
vaporization varies slightly with temperature. Assumptions made during the derivation
fail at high pressures and near the critical point, and under those conditions the ClausiusClapeyron equation will give inaccurate results. While the Clausius-Clapeyron equation
is frequently used as a first estimate, the Antoine equation is often used because of its
accuracy.
Antoine’s Equation, shown in Equation 10, is the empirical thermodynamic
relationship which best represents vapor pressure behavior with temperature.
23
log10 P(mmHg ) = A −
B
T (o C ) + C
(10)
The Antoine constants (A, B, C) are only valid for a specified temperature range and
cannot be extrapolated since they are established experimentally.
Antoine’s constants were obtained from the compilation of Dr. Shuzo Ohe27 while
a modified format Clausius-Clapeyron equation was also referenced from the Handbook
of Chemistry and Physics.28
Table 1. i-octane Antoine constants for designated temperature range27
Antoine Parameters for i-octane
Temperature Range (°C) : -78.51 ~ 25.29
Pressure Range (mmHg) : 0.01 ~ 50.00
A = 6.82246
B = 1282.332
C = 224.706
Table 2. i-octane Antoine constants for designated temperature range27
Antoine Parameters for i-octane
Temperature Range (°C) : 24.36 ~ 100.13
Pressure Range (mmHg) : 47.79 ~ 779.37
A = 6.80234
B = 1252.132
C = 220.059
Table 3. i-octane vapor pressure from the Handbook of Chemistry and Physics28
log10P (Torr) = [-0.2185A/T(K)] + B
Temperature Range : -36.5 ~ 99.2
A = 8548.0 molar heat of vaporization (calories/gram mole)
B = 7.934852
24
i -Octane Vapor Pressure vs. Temperature
900
800
Vapor Pressure (mmHg)
700
600
500
400
0.0383x
y = 18.318e
300
200
100
0
-30
-10
10
30
50
70
Temperature ( o C)
Handbook of Chemistry and Physics
Shuzo Ohe
Explosion Region
90
110
Expon. (Shuzo Ohe)
Figure 13: i-octane vapor pressure curve versus temperature
i -Octane Vapor Pressure vs. Temperature
(zoomed in on explosive range)
70
60
Vapor Pressure (mmHg)
50
40
30
20
10
0
-30
φ=0.48
-20
Handbook of Chemistry and Physics
-10
0
Temperature ( o C)
Shuzo Ohe
10
Explosion Region
20
φ=3.76
Expon. (Shuzo Ohe)
Figure 14: i-octane vapor pressure curve vs. temperature explosive region
30
25
The same vapor pressure analysis was similarly applied to n-octane.
Table 4. n-octane Antoine constants for designated temperature range27
Antoine Parameters for n-octane
Temperature Range (°C) : -56.56 ~ 23.95
Pressure Range (mmHg) : 0.02 ~ 11.80
A = 8.07630
B = 1936.281
C = 253.007
Table 5. n-octane Antoine constants for designated temperature range27
Antoine Parameters for n-octane
Temperature Range (°C) : 52.93 ~ 126.57
Pressure Range (mmHg) : 57.53 ~ 779.32
A = 6.92010
B = 1352.580
C = 209.192
Table 6. n-octane vapor pressure from the Handbook of Chemistry and Physics 28
log10P (Torr) = [-0.2185A/T(K)] + B
Temperature Range : -14.0 ~ 281.4
A = 9221.0 molar heat of vaporization (calories/gram mole)
B = 7.894018
26
n -Octane Vapor Pressure vs. Temperature
1200
y = 6.1433e0.0391x
Vapor Pressure (mmHg)
1000
800
600
400
200
0
0
20
40
60
80
100
120
140
Temperature ( o C)
Handbook of Chemistry and Physics
Shuzo Ohe
Explosion Region
Expon. (Shuzo Ohe)
Figure 15: n-octane vapor pressure curve versus temperature
n -Octane Vapor Pressure vs. Temperature
(zoomed
in on explosive region)
60
Vapor Pressure (mmHg)
50
40
30
20
10
0
0
φ=0.60
10
Handbook of Chemistry and Physics
20
30
40
Temperature ( o C)
Shuzo Ohe
Explosion Region
50
φ=4.42
Expon. (Shuzo Ohe)
Figure 16: n-octane vapor pressure curve vs. temperature explosive region
60
27
Using this vapor pressure information, equivalence ratio (φ) is determined for a
corresponding temperature.
Equivalence ratio is the relationship between fuel and
oxygen and is dependent upon such things as altitude, temperature, and vapor pressure.
 moles fuel



moles
oxygen

 actual
φ=

 moles fuel


moles
oxygen
 stoichiometric

(11)
A mixture is considered lean when φ < 1 and rich when φ > 1. Assuming complete
combustion, the stoichiometric ratio can be derived from a balanced chemical reaction.
Because i-octane (2,2,4-trimethylpentane) and n-octane are isomers, the same equation
applies.
2C8H18 + 25O2 → 16CO2 + 18H2O
(
Therefore, the molar ratio of fuel oxygen
)
stoichiometric
=
(12)
2
.
25
The actual fuel to oxygen ratio calculation is more involved, requiring the
aforementioned vapor pressure data. The experiment is operated at atmospheric pressure,
thus the total gas phase pressure (P) is 760 mmHg, including the vaporized fuel.
P = 760mmHg = PC8 H18 + Pair
(13)
The partial pressure of oxygen (PO2 ) in air is 21%, unless investigating fuel tank inerting,
where the oxygen content would be reduced to 18%.
PO2 = 0.21(760mmHg − PC8 H18 )
(14)
With these partial pressures, the actual fuel to oxygen ratio can be obtained.
 PC H
 fuel

=  8 18
oxygen  actual  P

 O2




(15)
28
Equilibrium
Equilibrium is established when the desired operating temperature stabilized as
well as the internal vessel pressure. Temperature is monitored directly on the Julabo
control unit; meanwhile, pressure is acquired by a Druck PMP 1240 low-speed
pressure transducer which displayed on the Wonderware screen. Note that P = 14.7 psia
was not always observed due to a temperature dependence of the transducer response,
even though the experimental condition never deviated from atmospheric pressure. This
divergence is shown in Figure 17, depicting observed pressure versus temperature.
14.5
Low Speed Pressure Transducer Reading vs. Vessel Temperature
(Acutal Atmospheric Presure = 14.7 psi)
Observed Pressure (psi )
14.0
13.5
13.0
12.5
12.0
11.5
25
30
35
40
45
50
55
60
o
Temperature ( C )
Figure 17: Calibration curve of low-speed pressure transducer deviation
This was not problematic, since the low-speed pressure transducer was merely
used for indication rather than for any direct readings.
Procedure
A detailed test procedure is given in Appendix C. This same procedure was followed for
testing with both i-octane and n-octane.
RESULTS
Preliminary validation tests were first executed to verify the performance of the
modified combustion vessel from its previous design. The volume of the combustion
vessel was nearly doubled to 4 liters with the addition of the center cylinder section,
changing the geometry from spherical to ellipsoid.
BEFORE
AFTER
Figure 18: Combustion vessel before and after modification
Additionally, the electrodes were relocated from a T-shape orientation in the top
hemisphere to a rearranged angular formation in the center cylinder.
BEFORE
AFTER
3mm
Spark
gap
Bait
Probe
Bait
Probe
Figure 19: Electrode orientation before and after relocation
29
30
Determination of the optimum electrode arrangement, including spark gap
distance, was analyzed prior to fuel combustion. When the behavior of the electrodes
was well understood, methanol was selected as the preliminary benchmark fuel.
Methanol was chosen because of the previous work done in this laboratory and the
available data for comparison.20 Once this initial work is completed, further testing with
Jet-A fuel components, i-octane and n-octane, can begin.
Spark Gap Determination
Breakdown voltage is dependent upon several factors: distance between the
electrodes, electrode shape (round, point, plane, etc.), temperature and pressure of the
surrounding gas, and type of energy supply (AC, DC, pulse).29 In this case, the variable
was the spark gap distance; it was already determined that the electrode tips were pointed
and supplied with a DC voltage. Testing was conducted at standard room conditions.
Breakdown Voltage versus Spark Gap
18,000
Breakdown Voltage (volts)
16,000
14,000
y = 1052.8x + 968.52
R2 = 0.9901
12,000
10,000
8,000
6,000
4,000
2,000
0
0
2
4
6
8
10
12
14
16
Spark Gap (mm)
Figure 20: Bench top investigation of breakdown voltage versus spark gap
31
A 3mm spark gap between the cathode and anode was determined optimal after
bench top investigation of breakdown voltage. This allows for an operating range of
0-4,000 volts before auto sparking.
If the spark gap distance is increased, plasma
generation would be more difficult and would require a higher voltage supply from the
ionization circuit as well as more energy and voltage from the arc circuit. On the other
hand, if the spark gap distance is decreased, the result would be more heat loss to the
electrodes by conduction and consequently flame quenching.
Methanol Validation Testing
Two temperatures (equivalence ratios) were selected to validate the modified
combustion vessel with methanol fuel, representing a near stoichiometric mix and a rich
mix.
Table 7. Methanol validation testing conditions
Oxygen Content
32.5
Equivalence Ratio = φ
2.30
22.5
1.23
21%
Temperature (oC)
21%
The rich mixture of methanol and air (φ=2.30) lends to favorable testing
conditions, meaning that the energy required to create an ignition is neither too high nor
too low. Figure 21 shows the results of the new validation data against formerly reported
data.20 Experimental data points were fitted with a third-order polynomial.
32
METHANOL Spark Ignition Energy vs. Spark Duration
32.5 Celsius (φ=2.30) in 21% Oxygen
4.0
60
y = 0.0423x3 - 0.0865x2 - 0.9812x + 4.6209
3.5
50
40
2.5
2.0
30
1.5
20
Pressure Rise (psig)
log (spark energy mJ )
3.0
y = 0.0583x3 - 0.1754x2 - 0.7724x + 3.8083
1.0
minimum: (3.55, 1.94) = 87 mJ @ 3505 µs
10
0.5
minimum: (3.33, 1.44) = 28mJ @ 2145 µs
0.0
0
1
2
Previous Ignition Data
3
4
log (spark duration µ s )
Validation Ignition Data
5
6
Validation Pressure Data
Figure 21: Curve fit methanol ignition data at 32.5ºC (φ=2.30) in 21% oxygen
As seen in the figure, lower minimum ignition energy was obtained with the new
combustion apparatus at this experimental condition. Considering the data scatter, a
difference of 59 mJ in minimum ignition energy is not unreasonable. While no pressure
comparison is available, it can be noted that the pressure rise due to ignition consistently
pivots about the 30 psi region. Supplementary experimentation at intermediate spark
durations, along with longer sparks, could provide better agreement with the previous
results; however, progressing to another equivalence ratio also affords more insight.
Several trials were performed at φ=1.23, a near stoichiometric mix.
The
validation experiments as well as previous data are given in Figure 22. When plotted
with a line of best fit, the ignition values of the validation data illustrate an increase
33
compared to the previously obtained values. However, it can be observed that the fitted
line for the prior points is skewed low rather than centrally straddling the data.
Furthermore, the majority of the previous data falls around the fitted line of this
validation data; therefore, the difference could be smaller.
METHANOL Ignition Energy vs. Spark Duration
Room Temperature (22.5 Celsius) φ=1.23 in 21% Oxygen
1.0
0.5
y = -0.0085x3 + 0.3677x2 - 1.277X + 0.5783
log (spark energy mJ )
0.0
-0.5
-1.0
y = -0.2505x3 + 1.885x2 - 4.0868x + 1.7073
-1.5
minimum: (1.58, -1.03) = 0.09 mJ @ 38 µs
-2.0
minimum: (1.86,-0.58) = 0.26mJ @ 72 µs
-2.5
0
1
2
log (spark duration µ s )
Previous Ignition Data
3
New Validation Data
Figure 22: Curve fit methanol ignition data at 22.5ºC (φ=1.23) in 21% oxygen
Overall, the results can be compared on a graph where the minimum ignition
energy from each equivalence ratio is plotted to achieve the minimum of the minimum
ignition energy, commonly referred to as MIE for simplification (Figure 23).
34
log (MIE/mJ)
Methanol MIE at 3mm Spark Gap
3.0
2.5
2.0
1.5
1.0
0.5
0.0
-0.5
-1.0
-1.5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
Equivalence Ratio (φ)
Previous Data
Parabolic Fit
Validation Data
Figure 23: Methanol equivalence ratio versus minimum ignition energy
This graphic interpretation indicates that the new combustion vessel indeed
models the general behavior of the earlier arrangement. With this successful validation
completed, consequent testing of i-octane and n-octane proceeded.
i-octane Testing
Selected from the hydrocarbon constituents in jet fuel, i-octane is a valid
representative component if only one species must be chosen for evaluation. It is the
smallest hydrocarbon present that exists in liquid phase at standard conditions.
Moreover, i-octane encompasses a major percentage of the vapor in the ullage, the empty
space above the fuel level in a closed tank. Multiple temperature/equivalence ratio set
points were examined (Table 8).
35
Table 8. i-octane testing conditions
Oxygen Content
19.0
Equivalence Ratio = φ
3.13
16.5
2.83
21%
14.5
2.61
21%
8.0
2.02
21%
19.0
3.65
18%
16.5
3.30
18%
Temperature (oC)
21%
The results of i-octane ignition with 21% oxygen (standard air) are graphically
displayed in Figures 24-27. Each of these data sets has been fitted with a third-order
polynomial and a minimum identified through differentiation.
i -OCTANE Spark Ignition Energy vs. Spark Duration
19.0 Celsius (φ=3.13) in 21% Oxygen
3.5
6
minimum: (2.85, 1.21) = 16 mJ @ 701 µs
3.0
5
2.5
4
Pressure Rise (psig)
log (spark energy mJ )
2.0
1.5
1.0
0.5
y = -0.2505x3 + 2.9521x2 - 10.716x + 13.569
3
2
0.0
-0.5
1
-1.0
-1.5
0
0
1
2
3
log (spark duration µ s )
Ignition Energy vs. Spark Duration
4
5
Ignition Pressure Rise
Figure 24: i-octane ignition data at 19.0ºC (φ=3.13) in 21% oxygen
6
36
At 19.0ºC (φ=3.13) the mixture is already rich, even though clearly below room
temperature. The pressure rise is consistently low around 1 psi; the minimum calculated
ignition energy obtained from the curve fit is 16 mJ.
Lowering the temperature to 16.5ºC decreases the equivalence ratio to 2.83 and
shifts the curve to reduced energy levels but raises the overall pressure rise.
The
minimum calculated ignition energy (7 mJ) is half that of the previous data set at 19.0ºC,
although some measurements were near 1.5 mJ. The average pressure rise of about
3.5 psi is nearly threefold that of the φ=3.13 mixture, but with two visual outliers.
i -OCTANE Spark Ignition Energy vs. Spark Duration
16.5 Celsius (φ=2.83) in 21% Oxygen
3.5
20
3.0
18
minimum: (2.43, 0.83) = 7 mJ @ 270 µs
2.5
16
14
y = -0.1654x3 + 1.5827x2 - 4.7631x + 5.4343
1.5
1.0
0.5
Pressure Rise (psig)
log (spark energy mJ )
2.0
12
10
8
0.0
6
-0.5
4
-1.0
2
-1.5
0
0
1
2
3
4
5
6
log (spark duration µ s )
Ignition Energy vs. Spark Duration
Ignition Pressure Rise
Figure 25: i-octane ignition data at 16.5ºC (φ=2.83) in 21% oxygen
Continuing to a lower temperature of 14.5ºC does not drastically change the
minimum ignition energy but rather sandwiches it between the latter two data sets;
37
although, the energy is expected to decrease because of the smaller equivalence ratio
(φ=2.61). Ignition was achievable over a large spark duration domain, giving a suitable
curve fit.
The heightened pressure rise to 4 psi is anticipated since the mixture is
becoming leaner.
i -OCTANE Spark Ignition Energy vs. Spark Duration
14.5 Celsius (φ=2.61) in 21% Oxygen
3.5
12
3.0
minimum: (2.81, 1.01) = 10 mJ @ 647 µs
10
2.5
8
Pressure Rise (psig)
log (spark energy mJ )
2.0
1.5
1.0
0.5
6
4
0.0
-0.5
y = -0.0477x3 + 0.7329x2 - 2.9896x + 4.6774
2
-1.0
-1.5
0
0
1
2
3
4
5
6
log (spark duration µ s )
Ignition Energy vs. Spark Duration
Ignition Pressure Rise
Figure 26: i-octane ignition data at 14.5ºC (φ=2.61) in 21% oxygen
For φ=2.02 mixture (8.0ºC) experiments, some technical difficulties were
encountered due to the stability of the Julabo heating/cooking unit and frosting of the heat
exchangers.
The ignition energy data and curve fit are shown in Figure 27.
This
condition produces the lowest ignition energy observed thus far of 1.5 mJ and a
tremendously amplified pressure rise near 32 psi.
38
i -OCTANE Spark Ignition Energy vs. Spark Duration
8.0 Celsius (φ=2.02) in 21% Oxygen
45
3.5
3.0
40
minimum: (2.20, 0.17) = 1.5 mJ @ 159 µs
2.5
35
30
Pressure Rise (psig)
log (spark energy mJ )
2.0
1.5
1.0
0.5
25
20
15
0.0
3
2
y = -0.0213x + 0.6363x - 2.4925x + 2.8025
10
-0.5
5
-1.0
-1.5
0
0
1
2
3
log (spark duration µ s )
Ignition Energy vs. Spark Duration
4
5
6
Ignition Pressure Rise
Figure 27: i-octane ignition data at 8.0ºC (φ=2.02) in 21% oxygen
The i-octane minimum ignition energies measured are plotted (Figure 28) as a
function of equivalence ratio (φ) together with the literature value. It was not possible to
measure experimental ignition energy values for mixtures of φ < 2.0 because of
equipment limitations with the temperature controller. Therefore, the behavior of ignition
energy versus spark duration time (hence minimum ignition energy) of φ < 2.0 mixtures
has not been studied.
However, one can conclude that ignition energy of i-octane
(~ 1 mJ) occurs at about φ~2 from looking at the minimum ignition energies of n-hexane
and n-heptane at φ~1.75.4
39
i -Octane Minimum Ignition Energy vs. Equivalence Ratio
1.40
1.20
log (MIE mJ )
1.00
0.80
0.60
0.40
0.20
Published MIE = 1.35 mJ
0.00
0
0.5
1
1.5
2
2.5
3
3.5
Equivalence Ratio (φ)
Figure 28: i-octane equivalence ratio versus minimum ignition energy
Reducing oxygen content to 18% in the reaction vessel imitates fuel tank inerting,
which limits the oxidant concentration for combustion, and increases the equivalence
ratio by 16% to a more rich mixture. Both the 16.5ºC and 19.0ºC test temperature
experiments were duplicated but using an 18% oxygen content air instead of standard air
(21% oxygen). The outcome is shown in Figures 29 and 30.
40
i -OCTANE Spark Ignition Energy vs. Spark Duration 19.0 Celsius
3.5
5
3.0
2.5
4
Pressure Rise (psig )
log (spark energy mJ )
2.0
1.5
1.0
0.5
3
2
0.0
-0.5
1
-1.0
-1.5
0
0
1
2
3
log (spark duration µ s )
4
5
phi= 3.13 Ignition Energy (21% Oxygen)
phi=3.65 Ignition Energy (18% Oxygen)
phi=3.13 Ignition Pressure Rise (21% Oxygen)
phi=3.65 Ignition Pressure Rise (18% Oxygen)
6
Figure 29: i-octane ignition data at 19.0ºC with oxygen content of 18% and 21%
As the equivalence ratio is increased at the same temperature, a sharp increase of
ignition energy with pulse duration is observed. Even with various combinations of
resistors and capacitors for short pulse duration time (<~ 700 µs) and high voltage
supply, it was not possible to achieve ignition. The pressure rise from 18% oxygen
content is considerably smaller than 21% oxygen standard air mixture, indicating weak
ignition.
41
i -OCTANE Spark Ignition Energy vs. Spark Duration at 16.5 Celsius
3.5
14
3.0
2.5
12
10
Pressure Rise (psig)
log (spark energy mJ )
2.0
1.5
1.0
0.5
8
6
0.0
4
-0.5
2
-1.0
-1.5
0
0
1
2
3
log (spark duration µ s )
4
5
phi=2.83Ignition Energy (21% Oxygen)
phi=3.30 Ignition Energy (18% Oxygen)
phi=2.83 Ignition Pressure Rise (21% Oxygen)
phi=3.30 Ignition Pressure Rise (18% Oxygen)
6
Figure 30: i-octane ignition data at 16.5ºC with oxygen content of 18% and 21%
For the equivalence ratio (φ=3.30) of the 18% oxygen inerting mixture at 16.5ºC,
it was also not possible to ignite when the pulse duration was less than approximately
80 µs. In the range of successful ignition, the minimum ignition energies increase with
pulse duration time. As expected, at a given temperature (in this case 16.5ºC), the
minimum ignition energies of the inerting gas mixture are considerably higher than the
standard air mixture.
This reduced oxygen environment may provide a means to diminish the
occurrence and effects of ignition.
The ignition envelope seems to be narrowed,
signifying that explosions occur over a more limited range of spark duration time. Along
42
with more difficulty to achieve ignition, a smaller pressure rise could keep the structural
integrity of the airframe even in case of mishaps.
All of these separate data plots can be summarized in one encompassing graphical
display (Figure 31).
i -OCTANE Spark Ignition Energy vs. Spark Duration
with Temperature Variation
3.5
40
3.0
35
30
2.0
25
Pressure Rise (psig)
log (spark energy mJ )
2.5
1.5
1.0
0.5
0.0
20
15
10
5
-0.5
0
-1.0
-1.5
-5
0
1
2
3
log (spark duration µ s )
4
5
phi=2.02
phi=2.61
phi=2.83
phi=3.13
phi=3.30
phi=3.65
Pressure phi=2.02
Pressure phi=2.61
Pressure phi=2.83
Pressure phi=3.13
Presure phi=3.30
Pressure phi=3.65
6
Figure 31: i-octane ignition data showing temperature variation effects
Any alteration of the airplane structure due to even a slight pressure rise from fuel
ignition could lead to total in-flight destruction. Thus, the ignition pressure rise with
respect to the spark duration time and the equivalence ratio is plotted in Figures 32 and
33, respectively.
43
i -OCTANE Spark Ignition Pressure Rise vs. Spark Duration
with Temperature Variation
40
35
Pressure Rise (psig )
30
25
20
15
10
5
0
-5
0
1
2
3
log (spark duration µ s )
4
5
Pressure phi=2.02
Pressure phi=2.61
Pressure phi=2.83
Pressure phi=3.13
Presure phi=3.30
Pressure phi=3.65
6
Figure 32: i-octane pressure rise ignition data versus spark duration
Figure 32 illustrates that once ignition occurs for each mixture the pressure rise
reaches a constant value independent of the spark duration time. For the relatively rich
mixtures, only a small pressure rise could be detected.
44
i -OCTANE Spark Ignition Pressure Rise vs. Equivalence Ratio
40
35
Pressure Rise (psig )
30
25
20
15
10
5
0
1.8
2.0
2.2
2.4
2.6
2.8
3.0
3.2
3.4
3.6
3.8
Equivalence Ratio, φ
Pressure phi=2.02
Pressure phi=2.61
Pressure phi=2.83
Pressure phi=3.13
Presure phi=3.30
Pressure phi=3.65
Figure 33: i-octane pressure rise data range versus equivalence ratio
In Figure 33, the pressure rise is displayed as a function of the equivalence ratio.
As the equivalence ratio is reduced, the pressure rise increases exponentially.
The
pressure rise is expected to reach a maximum value at φ~2 and then diminish with further
decrease of the equivalence ratio.
n-octane Testing
Unlike i-octane, vapor pressure of n-octane, another major constituent of jet fuel,
allows for operation at lower equivalence ratios, specifically near a stoichiometric ratio.
This will provide some insight as to the behavior of n-octane, an isomer of i-octane, with
a lean composition.
45
Table 9. n-octane testing conditions
Oxygen Content
35.9
Equivalence Ratio = φ
2.02
19.8
1.06
21%
18.3
1.00
21%
17.0
0.95
21%
Temperature (oC)
21%
To compare the behavior of n-octane and i-octane, an identical equivalence ratio
mixture was selected for assessment. Testing of n-octane began with 35.9ºC, a value of
φ~2, which has the same fuel and oxygen content of i-octane at 8.0ºC.
3.5
55
3.0
50
2.5
45
2.0
40
Pressure RIse (psig)
log (spark energy mJ )
n -OCTANE Spark Ignition Energy vs. Spark Duration
35.9 Celsius (φ=2.02) in 21% Oxygen
1.5
1.0
0.5
35
30
25
0.0
20
-0.5
15
-1.0
10
-1.5
5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
4.5
log (spark duration µ s )
Ignition Energy vs. Spark Duration
Ignition Pressure Rise
Figure 34: n-octane ignition data at 35.9ºC (φ=2.02) in 21% oxygen
Again, a bowl shape is seen in the data points, with a consistent pressure rise near
12.5 psi. When examining the separate ignition energy data sets of i-octane and n-octane
46
at φ=2 on the same figure, the ignition energies coincide rather well within the scatter of
data points (Figure 35). However, the pressure rise measurements are not analogous;
where i-octane is about 20 psi higher.
n -OCTANE and i -Octane Spark Ignition Energy vs. Spark Duration
φ=2 in 21% Oxygen
3.5
40
3.0
35
2.5
30
Pressure Rise (psig)
log (spark energy mJ )
2.0
1.5
1.0
0.5
25
20
0.0
15
-0.5
10
-1.0
-1.5
5
0.0
0.5
n-Octane, 35.9 Celsius
1.0
1.5
i-Octane, 8 Celsius
2.0
2.5
log (spark duration µ s )
3.0
n-Octane Pressure, 35.9 Celsius
3.5
4.0
4.5
i-Octane Pressure, 8 Celsius
Figure 35: n-octane and i-octane ignition data at φ ≈ 2.0 in 21% oxygen
The subsequent results (Figures 36-38) demonstrate when the actual conditions of
the system model stoichiometric calculations. Near φ~1 the ignition data are initially
unforeseen, but remarkably reproducible, as the anticipated data curvature is no longer
present.
47
3.5
55
3.0
50
2.5
45
2.0
40
Pressure Rise (psig)
log (spark energy mJ )
n -OCTANE Spark Ignition Energy vs. Spark Duration
19.8 Celsius (φ=1.06) in 21% Oxygen
1.5
1.0
0.5
35
30
25
0.0
20
-0.5
15
-1.0
10
-1.5
5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
4.5
log (spark duration µ s )
Ignition Energy vs. Spark Duration
Ignition Pressure Rise
Figure 36: n-octane ignition data at 19.8ºC (φ=1.06) in 21% oxygen
For φ=1.06, the unexpected results showed that as the spark duration time was
increased, the minimum ignition energy: slightly decreased, reached somewhat constant
values, and then suddenly dropped. Furthermore, the pressure rise was near 50 psi, which
is larger than the highest pressure rise of i-octane measured (32 psi at φ=2.0). This
phenomenon was consistent with small changes to leaner equivalence ratios. A similar
behavior of the minimum ignition energy with spark duration time for the mixtures of
φ=1.00 and φ=0.95 were observed (Figures 37 and 38). Although a direct comparison is
unachievable with i-octane mixtures of approximately the same equivalence ratio, the
minimum ignition energies of n-octane should be lower than those of i-octane.
48
3.5
55
3.0
50
2.5
45
2.0
40
Pressure Rise (psig)
log (spark energy mJ )
n -OCTANE Spark Ignition Energy vs. Spark Duration
18.3 Celsius (φ=1.00) in 21% Oxygen
1.5
1.0
0.5
35
30
25
0.0
20
-0.5
15
-1.0
10
-1.5
5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
4.5
log (spark duration µ s )
Ignition Energy vs. Spark Duration
Ignition Pressure Rise
Figure 37: n-octane ignition data at 18.3ºC (φ=1.00) in 21% oxygen
n -OCTANE Spark Ignition Energy vs. Spark Duration
17.0 Celsius (φ=0.95) in 21% Oxygen
3.5
60
3.0
55
50
2.5
45
Pressure Rise (psig)
log (spark energy mJ )
2.0
1.5
1.0
0.5
40
35
30
25
0.0
20
-0.5
15
-1.0
10
-1.5
5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
log (spark duration µ s )
Ignition Energy vs. Spark Duration
Ignition Pressure Rise
Figure 38: n-octane ignition data at 17.0ºC (φ=0.95) in 21% oxygen
4.5
49
The complete compilation of n-octane data are shown in the culminating graph
below to facilitate comparison.
For mixtures of near stoichiometric, the minimum
ignition energies are approximately the same, essentially 3 mJ, and occur at a similar
spark duration time of about 1000 µs (see dotted line oval in Figure 39). As seen in the
figure, before the minimum ignition energy reaches the lowest common value, it
increases with slight decrease of equivalence ratio. In the experiments of spark duration
time longer than 1000 µs (after the identified minimum ignition energy points), the
ignition energies become greater with an upward trend as equivalence ratio increases.
3.5
55
3.0
50
2.5
45
2.0
40
Pressure Rise (psig)
log (spark energy mJ )
n -OCTANE Spark Ignition Energy vs. Spark Duration
with Temperature Variation in 21% Oxygen
1.5
1.0
0.5
35
30
25
0.0
20
-0.5
15
-1.0
10
-1.5
5
0.0
0.5
1.0
1.5
2.0
2.5
3.0
3.5
4.0
log (spark duration µ s )
phi=2.02
phi=1.00
phi=0.95
phi=1.06
Pressure phi=2.02
Pressure phi=1.00
Pressure phi=0.95
Pressure phi=1.06
Figure 39: n-octane ignition temperature variation effects in 21% oxygen
4.5
50
The pressure rise for the near φ=1 mixtures is proportional to the stoichiometry.
However, for the rich mixture, the pressure rise is unexpectedly small, even less than that
of i-octane of about the same chemistry. The shape of the minimum ignition energy
curve and the pressure rise of the φ=2.02 rich mixture are distinctively different from
those of the near stoichiometric mixtures, an indication of a substantially different
ignition process.
DISCUSSION
In this study, a composite spark (breakdown + dc-arc) has been used.
The
breakdown energy (ionization energy), released to a flammable gas mixture for a few
nanoseconds to a few microseconds, creates a small volume of ignition kernel (plasma
kernel). In this phase, the energy transfer from the breakdown spark to the ignition
kernel is so efficient that the temperature and pressure of the ignition kernel increases to
extremely high values. Rapid relaxation of this high pressure generates a shock wave.
As the shock wave and ignition kernel expand, the temperature and pressure of the
ignition kernel fall. The temperature fall of the ignition kernel due to the heat loss of the
unburned gas and to the electrodes would continue to the fuel flame temperature. In this
period, the expansion rate of the shock wave is much higher than that of the ignition
kernel.30 If the ignition kernel growth rate is constant, then the flame is eventually
extinguished. For successful flame propagation, additional energy should be provided by
chemical heat release to make the ignition kernel grow to its critical size.
In a composite spark, the dc-arc discharge facilitated by the ignition kernel
(plasma kernel) formation in the breakdown phase does not foster the degree of
ionization of gas molecules or the ignition kernel temperature much. The primary role of
the dc-arc discharge is to increase the ignition kernel (plasma kernel) size and to extend
the cooling period of the hot ignition kernel so that the ignition kernel formed in the
breakdown phase is preserved much longer.31 Of course, the dc-arc energy is also lost by
51
52
heat transfer to the electrodes as well as to the surrounding unburned gas. Thus, for a
composite spark with the same breakdown energy and constant heat loss through the
electrodes, successful ignition and flame propagation are dependent upon the absolute
value of dc-arc energy, mode of dc-arc energy supply (dc-arc spark duration time), and
fuel mixture composition (equivalence ratio), and the ignition chemistry.
In i-octane experiments of a given stoichiometry, the minimum ignition energy
decreased as the dc-arc spark duration time became longer, reached its lowest value, and
then increased with extended spark duration time. If the spark duration time is short,
more dc-arc energy deposition would be needed to extend the size of the ignition kernel.
If it is too long, again more energy is needed to compensate for the heat loss to the
electrodes and to preserve the hot ignition kernel.
The value of the minimum ignition energy varies with the fuel mixture
equivalence ratio and ignition chemistry. The sequence of combustion reactions for
i-octane (also n-octane) is as follows:
(1) unimolecular fuel decomposition,
(2) H-atom abstraction by H, O, OH, and CH3,
(3) alkyl radical decomposition,
(4) alkyl radical + O2 → olefin + HO2,
(5) olefin + H, O, OH, CH3 → alkenyl + H2, OH, H2O, CH4,
(6) alkenyl radical decomposition,
(7) olefin decomposition.
Overall, the combustion of i-octane (and n-octane) can be understood in terms of the
conversion of fuel molecules to CH3, C2H4, C3H4, C3H6 and i-C4H8 through the pyrolytic
53
oxidative steps and then oxidation of small olefins, CH3 and CH4. The main chain
carriers are provided by the chain-branching H + O2 → O + OH reaction.
The cause of flame extinction is a combined effect of heat loss to the
surroundings, flame stretch, flame instability (cellular flame) and the extent of chemical
reaction.32 The minimum ignition energy for a lean fuel/air mixture is much higher than
that of the stoichiometric mixture.
In this case, not enough heat is released from
chemical reactions to sustain the flame propagation because of the low fuel
concentration, while heat loss to the surroundings (unburned gas + electrodes) is about
the same as in a stoichiometric mixture. Consequently, higher dc-arc energy deposition
is required to increase the size of the ignition kernel and to keep the temperature gradient
between the reaction zone and the unburned gas small. In a fuel rich mixture, pyrolytic
reaction in the steps shown before and the chain terminating reactions by recombination
of radical-radical and/or radical-molecule are dominant over chain-branching and chainpropagating reactions.20 In other words, radicals from fuel molecules compete for H
atoms with the chain-branching reaction (H + O2 → O + OH). Likewise to the case of
lean mixtures, for rich mixtures more dc-arc energy must be deposited for a longer time
to increase and preserve the size of the ignition kernel and to extend the chain length of
the chemical reaction for heat release. As seen in Figure 28, generally for rich i-octane
mixtures, the minimum ignition energy increases with the equivalence ratio; the reversal
of the minimum ignition energy between the mixtures of φ=2.83 and φ =2.61 is most
likely caused by some experimental artifacts and/or uncertainties.
The ignition of n-octane at high temperatures has not been studied experimentally
and theoretically. Therefore, a direct comparison of ignition characteristics between
54
n-octane and i-octane is not feasible. According to a recent computer modeling study, the
ignition delay times of n-heptane (normal alkane) is shorter than those of its isomers
(branched alkanes).33 It was also pointed out by Farrell et al. that the burning velocities
of n-alkanes are higher than iso-alkanes despite that the differences in heat capacity,
thermal conductivity, and adiabatic flame temperature are small.34 Farrell et al. attributed
the differences in burning velocity to kinetic effects.34 Also the differences among alkane
isomers become greater as the molecular size gets smaller.
The disparity in the flame velocity of i-octane and n-octane of the same
stoichiometry, even if the flame velocity of n-octane has not been measured, may be
explained by examining the kinetic effects. For example, how much and how fast these
molecules produce small hydrocarbons: CH3, C2H4, C3H4, C3H6, and i-C4H8 (see above).
Since there is the same number of distinct H-atom abstraction sites in i-octane and
n-octane (Step2), the main differences should be mainly in the steps of unimolecular fuel
decomposition (Step 1), alkyl radical decomposition (Step 3), and alkenyl radical
decomposition (Step 6) for the production of CH3 olefins. For i-octane, there are many
routes for CH3 generation through β-scission. However, for n-octane, there are more
H-atom formation routes than i-octane. Methyl radicals (CH3) terminate chain centers by
the fast recombination reaction, CH3 + CH3 (+M) → C2H6 (+M), while H-atoms keep the
number of chain centers growing through the chain-branching H + O2 → O + OH.
The minimum ignition energies of i-octane and n-octane of about the same
stoichiometry (φ=2) as a function of the dc-arc spark duration time are shown in Figure
35. Considering a higher mass diffusion rate of O2 (from the unburned region to the
flame front) than fuel, the actual stoichiometry of these mixtures is φ≅1 and not φ≅2 at
55
the reaction zone. Therefore, minimum ignition energies of i-octane and n-octane appear
to be the same. This observation leads one to conclude the following. The minimum
ignition energy is the energy required to make the ignition kernel grow to its critical size.
Unlike the flame velocity, the critical size is more dependent upon the amount of heat
released rather than how fast.
Thus, if i-octane and n-octane have about the same
thermodynamic properties (heat capacity, thermal conductivity, adiabatic flame
temperature, etc.) their minimum ignition energies would also be similar, which is not
dependent upon how fast the ignition kernel takes to reach its critical size. The profiles
of the minimum ignition energy versus spark duration time of lean n-octane mixtures
(shown in Figures 37-39) are very different from rich mixtures; in rich mixtures, a
smooth transition of the minimum ignition energy with spark duration time was observed.
However, in lean mixtures, there was a slow and uneven change of the minimum ignition
energy and then an abrupt decrease followed by a sharp increase.
For a possible explanation, ignition and flame propagation phenomena by a
composite spark are restated: In the breakdown phase, a very steep temperature profile
between the ignition kernel and the unburned gas exists, even though some of the ignition
kernel energy is lost following shock wave expansion. In this initial stage of flame
development, the heat release from chemical reactions is not sufficient to maintain such a
steep temperature profile. Hence, the temperature profile broadens due mainly to heat
loss by conduction to unburned gas and electrodes – flame stretch. If the breakdown
energy is sufficient to grow the ignition kernel to its critical size, then a minimal flame is
developed and continues to propagate as a combustion wave.
If not, the flame is
extinguished. In the arc-discharge phase, the extent of molecular dissociation could be
56
high but the degree of ionization is much lower. Also, the gas temperature in the arc is
low compared to that of the ignition (plasma) kernel temperature, ≈ 6000 K. Its role is to
extend the cooling period of the ignition kernel so that the plasma formed in the
breakdown phase can be preserved for a longer period of time. As mentioned before, the
minimum ignition energy varies with the mixture equivalence ratio; it increases rapidly as
the mixture gets leaner. While the effects of the mixture strength (equivalence ratio) on
the initial ignition kernel growth are not significant, the inflammation process, flame
front thickness, and flame propagation rate are strongly affected by the mixture strength.
With reduction of equivalence ratio, the chemical energy density of the mixture, flame
temperature, and flame speed all reduce and the flame is stretched. Consequently, more
time is available for heat loss from the inflammation zone and less energy is available to
offset this heat loss due to the decrease of energy transfer rate into this zone. Therefore,
for lean mixtures, more discharge energy must be supplied for the ignition kernel to grow
to a larger size.
In this composite spark system, the breakdown energy is small (<0.1 mJ) so that
ignition and flame propagation with only this energy is unattainable. Additional energy
is required to expand and preserve the size of ignition kernel formed for successful
ignition and flame propagation which is supplied in the form of dc-arc discharge. Figures
37 and 38 show that fuel ignition and flame propagation are always achieved for higher
dc-arc energy deposition as opposed to minimum ignition energy at a given spark
duration time and for longer spark duration at a given energy deposition. In other words,
the minimum ignition energy decreased as the spark duration time increased to a certain
extent. This observation emphasizes the importance of the rate of dc-arc energy
57
deposition more than the total energy supplied. The rate of dc-arc energy deposition and
chemical heat release must be balanced or exceeded to the heat loss rate. In high
temperature fuel ignition experiments, for example, by a shock tube technique, there
always exists an induction period. During this period of time the numbers of active chain
carriers (H, O, OH) grow exponentially, which leads to a rapid consumption of fuel
followed by large amount of heat release. Obviously the abrupt fall of the minimum
ignition energy at ~1000 µs is partly related to this induction period of i-octane.
However, at this time, an explanation as to the high initial value, dip, and then steep rise
of the minimum ignition energy is not available. Perhaps these unusual phenomena
would be clarified by an exhaustive computer modeling study with time dependent spark
energy deposition and energy loss together with chemical heat release computed using a
comprehensive i-octane and n-octane reaction mechanism.
CONCLUSION
Spark ignition energies were investigated for i-octane and n-octane mixtures
following validation of the experimental set-up with methanol.
Ignition energy is
dependent upon equivalence ratio as well as spark duration time while pressure rise is
also subject to the stoichiometry. A bowl shape is observed for rich mixtures of i-octane
with an apparent minimum near φ=2 and ≈1.5 mJ which approaches the accepted value of
1.35 mJ. Additionally, if the oxygen content is slightly reduced, it appears the ignition
envelope is narrowed and the pressure rise decreased. However, less conclusive results
were obtained with lean n-octane mixtures because of the irregular nature of the data,
showing new phenomena. Further investigation is needed in this case to examine time
dependent spark energy deposition and energy loss together with chemical heat release
related to the reaction mechanism.
58
REFERENCES
1
National Transportation Safety Board (NTSB), (August 2000). Aviation Accident Report:
In-flight Breakup Over the Atlantic Ocean Trans World Airlines Flight 800 Boeing 747141, N93119 near East Moriches, New York July 17, 1996. NTSB Report Number:
AAR-00-03; NTIS Report Number: PB2000-910403.
2
Blanc, M. V., Guest, P. G., von Elbe, G., Lewis, B., (1947). Ignition of Explosive Gas
Mixtures by Electric Sparks: I. Minimum Ignition Energies and Quenching Distances of
Mixtures of Methane, Oxygen, and Inert Gases, J. Phys. Chem. 15, pp. 798-802.
3
Guest, P. G., (May 1944). Apparatus for Determining Minimum Ignition Energies for
Electric Spark Ignition of Flammable Gases and Vapors, Bureau of Mines, R.I. 3753.
4
Lewis, B., von Elbe, G. (1987). Combustion Flames and Explosions of Gases, 3rd Edition,
Academic Press.
5
Lewis, B., von Elbe, G. (1947). Ignition of Explosive Gas Mixtures by Electric Sparks: II.
Theory of the Propagation of Flame from an Instantaneous Point Source of Ignition. J.
Phys. Chem. 15, pp. 803-808.
6
Metzler, A. J., (1952). Minimum ignition of six pure hydrocarbon fuels of the C2 and C6
series, Research Memorandum, National Advisory Committee for Aeronautics (NACA
RM E52F27).
7
Calcote, H. F., Gregory, Jr., C. A., Barnett, C. M., Gilmer, R. B., (1952). Spark Ignition:
Effect of Molecular Structure, Ind. Eng. Chem. 44, 2656.
8
Litchfield, E. L., (1960). Minimum Ignition-Energy Concept and Its Application to Safety
Engineering, Bureau of Mines Report of Investigations 5671.
9
Litchfield, E. L., (1960). Chronology and Topography of Sparks at Minimum Energy for
Ignition, Explosives Research Laboratory, Bureau of Mines, Pittsburgh, PA.
10
Nestor, L., (1967). Investigation of Turbine Fuel Flammability within Aircraft Fuel
Tanks, Final Report DS-67-7, Naval Air Propulsion Test Center, Naval Base,
Philadelphia.
59
60
11
Kosvic, T., Zung, L., Gerstein, M., (1971). Analysis of Aircraft Fuel Tank Fire and
Explosion Hazards. Technical Report AFAPL-TR-71-7, Air Force Aero Propulsion
Laboratory, Wright-Patterson Air Force Base, Ohio.
12
Shepherd, J. E., Krok, J. C., Lee, J. J., (1997). Jet A Explosion Experiments: Laboratory
Testing, GALCIT Report FM97-5.
13
Nuyt, C. D., Shepherd, J. E., Lee. J. J., (2000). Flashpoint and Chemical Composition of
Aviation Kerosene (Jet A), Explosion Dynamics Laboratory Report FM99-4, California
Institute of Technology.
14
Shepherd, J. E., Krok, J. C., Lee, J. J., (1999). Spark Energy Measurements in Jet A.
GALCIT Report FM97-9.
15
Shepherd, J. E., Krok, J. C., Lee, J. J., Birky, M. M., (1997). Jet A Explosions – Field
Test Plan ¼-scale Experiments, GALCIT Report FM97-17.
16
Shepherd, J. E., Krok, J. C., Lee, J. J., Brown, L. L., Lynch, R. T., Samaras, T. M., Birky,
M. M., (1998). Results of ¼-scale experiments, vapor stimulant and liquid Jet A tests,
Explosion Dynamic Laboratory Report FM98-6, California Institute of Technology.
17
Lee, T. W., Jain, V., Kozola, S., (2001). Measurements of Minimum Ignition Energy by
Using Laser Sparks for Hydrocarbon Fuels in Air: Propane, Dodecane, and Jet-A Fuel,
Combustion and Flame 125, 1320.
18
Ronney, P.D., (1994). Laser Versus Conventional Ignition of Flames, Optical
Engineering 33, 510.
19
Bradley, D., Sheppard, C. G. W., Suardjaja, I. M., Woolley, R., (2004). Fundamentals of
high-energy spark ignition with lasers, Combustion and Flame 138, 55.
20
Knepper, B.E., (2004). Experimental Determination of the Minimum Ignition Energy of
Methanol. MS Thesis, The University of Toledo, United States of America.
21
Haase, H., (1977). Electrostatic Hazards: Their Evaluation and Control, translated by
Michael Wald, Statische Elektrizität als Gefahr, Verlag Chemie, Weinheim, p. 109.
22
Kuchta J. M., (1986). Investigations of Fire and Explosion Accidents in the Chemical,
Mining, and Fuel-Related Industries: A Manual, Bureau of Mines, Bulletin 680.
23
Kuchta, J. M., (1975). Summary of Ignition Properties of Jet Fuels and Other Aircraft
Combustible Fluids, Bureau of Mines, ADAPL-TR-75-70.
24
http://chemistry.about.com/cs/metalsandalloys/a/aa071201a_2.htm
61
25
Tektronix Instructions Manual AM 503B & AM 5030 Amplifier, 070-8766-05.
26
Tektronix Instructions Manual A6303 & A6303XL 100 Amp AC/DC Current Probe,
070-3906-04.
27
Ohe PhD, S., (1976). Computer Aided Data Book of Vapor Pressure, Data Publishing
Company, Tokyo.
28
Schlessinger, G. G., (1972). Handbook of Chemistry and Physics, 53rd Edition, D-151.
29
http://wiki.4hv.org/index.php/Spark_gaps
30
Arpaci, V. S., Ko, Y., Lim, M. T., Lee, H.S., (2003). Spark kernel development in
constant volume combustion, Combustion and Flame 135, 315-322.
31
Maly, R., Vogel, M., (1978). 17th International Symposium on Combustion, The
Comubstion Institute, 821.
32
Strehlow, R. A., (1984). Combustion Fundamental, McGraw-Hill, New York.
33
Westbrook, C. K., Pitz, W. J., Curran, H. C., Boercker, J., Kunrath, E., (2001). Chemical
kinetic modeling study of shock tube ignition of heptane isomers, International Journal
of Chemical Kinetics, Volume 33, Issue 12, 868-877.
34
Farrell, J. T., Johnston, R. J., Androulakis, I P., ( 2004). Molecular Structure Effects On
Laminar Burning Velocities At Elevated Temperature And Pressure, Society of
Automotive Engineers, Inc., SAE Paper 2004-01-2936, ExxonMobil Research and
Engineering.
APPENDICES
62
63
APPENDIX A: Fuel Tank Ignitions35
Date
29-JUL1950
Aircraft
Registration
Operator
Fatalities
Location
Bristol 170
F-BENF
CAT
26
Algeria
17-SEP1956
Avro York
EP-ADB
Persian Air
Services
0
U.K.
25-MAY1958
Lockheed L1049
55-0123
USAF
0
USA
26-JUN1959
Lockheed L1649
N7313C
TWA
68
Italy
08-DEC1963
Boeing 707
N709PA
Pan Am
81
USA
03-JUN1971
Boeing C-135
58-0039
USAF
5
Spain
23-MAR1974
DC-8
N6164A
Airlift Int.
0
USA
09-MAY1976
Boeing 747
5-8104
Iran AF
17
Spain
08-FEB1980
Boeing C-135
60-0338
USAF
0
USA
14-MAR1980
Lockheed C130
74-2064
USAF
18
Turkey
19-MAR1982
Boeing C-135
58-0031
USAF
27
USA
13-FEB1987
Boeing C-135
60-0330
USAF
0
USA
04-OCT1989
Boeing C-135
56-3592
USAF
4
Canada
27-NOV1989
Boeing 727
HK-1803
Avianca
107
Colombia
11-MAY1990
Boeing 737300
EI-BZG
Philippine Air
Lines
8
Philippines
10-DEC1993
Boeing C-135
57-1470
USAF
6
USA
17-JUL1996
Boeing 747
N93119
TWA
230
USA
03-MAR2001
Boeing 737400
HS-TDC
Thai Airways
1
Thailand
08-MAY2004
Douglas DC-4
N44911
Brooks Air Fuel
0
USA
64
29 JUL 1950
Date:
Bristol 170 Freighter 21
Type:
Cie Air Transport - CAT
Operator:
F-BENF
Registration:
12738
Msn / C/n:
1946
Year built:
4 fatalities / 4 on board
Crew:
22 fatalities / 22 on board
Passengers:
26 fatalities / 26 on board
Total:
Written off
Airplane damage:
Tanezrouft (Algeria)
Location:
En route
Phase:
International Scheduled Passenger
Nature:
Algiers Airport (ALG)
Departure airport:
Gao Airport (GAQ)
Destination airport:
Narrative: Crashed following a wing failure.
PROBABLE CAUSE: "An explosion in the wing compartment containing the main
starboard fuel tank. This explosion tore off part of the upper wing surface which started a
vibration of the wing structure which then caused multiple failures in flight."
17 SEP 1956
Date:
Avro 685 York C.1
Type:
Persian Air Services
Operator:
EP-ADB
Registration:
1224
Msn / C/n:
1946
Year built:
? fatalities / ? on board
Crew:
? fatalities / ? on board
Passengers:
? fatalities / ? on board
Total:
Written off
Airplane damage:
London-Stansted Airport (STN), United Kingdom
Location:
Standing
Phase:
Unknown
Nature:
Departure airport:
Destination airport:
Narrative: Fuel tank exploded during maintenance.
65
25 MAY 1958
Date:
16:33
Time:
Lockheed RC-121D Super Constellation
Type:
United States Air Force - USAF
Operator:
55-0123
Registration:
4396
Msn / C/n:
1956
Year built:
0 fatalities / 15 on board
Crew:
0 fatalities / 0 on board
Passengers:
0 fatalities / 15 on board
Total:
Written off
Airplane damage:
Falmouth-Otis AFB, MA (United States of America)
Location:
Standing
Phase:
Military
Nature:
Falmouth-Otis AFB, MA (FMH)
Departure airport:
Narrative: The Super Connie was being prepared for a patrol flight when the center fuel tank
exploded. The aircraft caught fire, but all 15 crew members evacuated safely. It appeared that the
center fuel tank had been filled with fuel, despite the fact that it was not supposed to be filled.
Through seepage or overflow from that tank the fuel vapors were ignited by electronic
equipment being tested during the pre flight procedure.
26 JUN 1959
Date:
ca. 17:35
Time:
Lockheed L-1649A Starliner
Type:
Trans World Airlines - TWA
Operator:
N7313C
Registration:
1015
Msn / C/n:
9 fatalities / 9 on board
Crew:
59 fatalities / 59 on board
Passengers:
68 fatalities / 68 on board
Total:
Written off
Airplane damage:
32 km (20 mls) NW of Milano (Italy)
Location:
En route
Phase:
International Scheduled Passenger
Nature:
Milano-Malpensa Airport (MXP)
Departure airport:
Paris-Orly Airport (ORY)
Destination airport:
891
Flightnumber:
Narrative: The TWA Lockheed Starliner departed Milan at 16:20 GMT in weather conditions
with scattered thunderstorms. Some 12 minutes later the crew reported climbing through 10,000
feet. At 16:35 a structural failure occurred, initiated by a wing separation.
PROBABLE CAUSE: "The breaking-up in flight was due to the explosion of the fuel vapours
contained in tank No.7, followed immediately by either an explosion of pressure or a further
explosion in tank No.6. In the absence of other significant concrete evidence, taking into
account the stormy weather conditions, with frequent electric discharges, existing in the area at
the time of the crash, it may be assumed that the explosion of the fuel vapours contained in tank
No.7 was set off, through the outlet pipes, by igniting of the gasoline vapours issuing from these
pipes as a consequence of static electricity discharges (streamer corona) which developed on the
vent outlets."
66
Final
Status:
08 DEC 1963
Date:
20:59 EST
Time:
Boeing 707-121
Type:
Pan American World Airways
Operator:
N709PA
Registration:
17588/3
Msn / C/n:
1958
Year built:
15609 hours
Total airframe hrs:
4 Pratt & Whitney JT3CEngines:
8 fatalities / 8 on board
Crew:
73 fatalities / 73 on board
Passengers:
81 fatalities / 81 on board
Total:
Written off
Airplane damage:
Elkton, MD (United States of America)
Location:
Approach
Phase:
Domestic Scheduled Passenger
Nature:
Baltimore/Washington International Airport, MD (BWI)
Departure airport:
Philadelphia International Airport, PA (PHL)
Destination airport:
214
Flightnumber:
Narrative: Pan American Flight 214 departed San Juan, Puerto Rico at 16:10 EST for a
flight to Philadelphia with an intermediate stop at Baltimore. The aircraft, named
'Clipper Tradewind' arrived at Baltimore at 19:35 and took off again after refueling at
20:24. After contacting Philadelphia Approach Control the crew elected to wait in a
holding pattern along with 5 other aircraft because of extreme winds at Philadelphia.
Flight 214 entered a holding pattern west of the New Castle VOR on the 270 radial.
At 20:58 Clipper Tradewind suffered a lightning strike. This caused the initial ignition of
flammable fuel vapours inside the left reserve fuel tank. This triggered explosions in the
centre and right reserve fuel tanks as well. Fuel spilled and caught fire; the complete left
wingtip separated as a result. The aircraft was then seen to crash in flames. A 'Mayday'
call was received by Philadelphia Approach as the plane was descending out of control.
PROBABLE CAUSE: "Lightning-induced ignition of the fuel/air mixture in the Number
1 reserve fuel tank with resultant explosive disintegration of the left outer wing and loss
of control."
67
03 JUN 1971
Date:
Boeing KC-135Q
Type:
United States Air Force - USAF
Operator:
58-0039
Registration:
17784/254
Msn / C/n:
1959
Year built:
? fatalities / ? on board
Crew:
? fatalities / ? on board
Passengers:
5 fatalities / 5 on board
Total:
Written off
Airplane damage:
near Torrejon AFB (Spain)
Location:
En route
Phase:
Military
Nature:
Narrative: Crashed following in-flight explosion of the Number 1 main fuel tank.
Chafing of boost pump wires in conduits was determined to be as a possible ignition
source.
Final
Status:
23 MAR 1974
Date:
19:08
Time:
McDonnell Douglas DC-8-63CF
Type:
Airlift International
Operator:
N6164A
Registration:
46144/532
Msn / C/n:
1970
Year built:
4 Pratt & Whitney JT3DEngines:
0 fatalities / 0 on board
Crew:
0 fatalities / 0 on board
Passengers:
0 fatalities / 0 on board
Total:
1 fatality
Ground casualties:
Written off
Airplane damage:
Fairfield-Travis AFB, CA (SUU) (United States of America)
Location:
Standing
Phase:
Nature:
Departure airport:
Destination airport:
Narrative: The DC-8 was undergoing a maintenance A-check at the Travis Air Force
Base. Suddenly fuel fumes in the Number 1 inboard main fuel tank area exploded.
PROBABLE CAUSE: personnel - maintenance, servicing, inspection: improper
maintenance (maintenance personnel) powerplant - fuel system: other miscellaneous
acts, conditions - improperly installed
FACTOR: miscellaneous acts, conditions - fire in wing
68
Final
Status:
09 MAY 1976
Date:
14:35 UTC
Time:
Boeing 747-131F
Type:
Iran Air Force
Operator:
5-8104
Registration:
19677/73
Msn / C/n:
1970
Year built:
4 Pratt & Whitney JT9D-7A
Engines:
10 fatalities / 10 on board
Crew:
7 fatalities / 7 on board
Passengers:
17 fatalities / 17 on board
Total:
Written off
Airplane damage:
near Madrid (Spain)
Location:
En route
Phase:
Cargo
Nature:
Tehran-Mehrabad Airport (THR)
Departure airport:
Madrid-Torrejon AFB (TOJ)
Destination airport:
48
Flightnumber:
Narrative: The Boeing was operated on a military logistic flight from Tehran to
McGuire AFB via Madrid. The flight took off from Tehran at 08:20 GMT and climbed
to a cruising altitude of FL330. After establishing contact with Madrid control, clearance
was received to CPL VOR via Castejon. At 14:25 the flight was cleared to FL100. At
14:30 the crew advised Madrid that they were diverting to the left because of
thunderstorm activity, and at 14:32 Madrid cleared ULF48 to 5000 feet and directed him
to contact Madrid approach control. At 14:33 the crew contacted approach control and
advised them that there was too much weather activity ahead and requested to be
vectored around it. Last radio contact was when ULF48 acknowledged the 260deg
heading instructions and informed Madrid that they were descending to 5000 feet. The
aircraft was later found to have crashed in farmland at 3000 feet msl following left wing
separation.
It appeared that the aircraft had been struck by lightning, entering a forward part of the
aircraft and exiting from a static discharger on the left wingtip. The lightning current's
conductive path to the static discharger at the tip was through a bond strap along the
trailing edge. Concentration of current at the riveted joint between this bond strap and a
wing rib were sufficient conductive to cause the flash to reattach to this rivet and to leave
the discharger. Fuel vapors in the no. 1 fuel tank then ignited. The explosion caused the
upper wing skin panel to separate, causing a drastic altering of the aero elastic properties
of the wing, and especially the outboard section of wing. The outer wing began to
oscillate, developing loads which caused the high-frequency antenna and outer tip to
separate. The whole wing failed a little later.
69
08 FEB 1980
Date:
Boeing KC-135Q
Type:
United States Air Force - USAF
Operator:
60-0338
Registration:
18113/452
Msn / C/n:
1961
Year built:
8015 hours
Total airframe hrs:
0 fatalities / ? on board
Crew:
0 fatalities / ? on board
Passengers:
0 fatalities / ? on board
Total:
Written off
Airplane damage:
Plattsburgh AFB, NY (PBG) (United States of America)
Location:
Standing
Phase:
Military
Nature:
Departure airport:
Destination airport:
Narrative: Burned out on ramp following an explosion of the aft body fuel tank during
ground refueling operations. A faulty fuel probe was found to be the cause of the
explosion.
14 MAR 1980
Date:
Lockheed C-130H Hercules
Type:
United States Air Force - USAF
Operator:
74-2064
Registration:
4659
Msn / C/n:
1976
Year built:
? fatalities / ? on board
Crew:
? fatalities / ? on board
Passengers:
18 fatalities / 18 on board
Total:
Written off
Airplane damage:
15 km (9.4 mls) W of Incirlik (Turkey)
Location:
Unknown
Phase:
Military
Nature:
Narrative: Fuel tank explosion; crashed on approach.
70
19 MAR 1982
Date:
21:10
Time:
Boeing KC-135A-BN Stratotanker
Type:
United States Air Force - USAF
Operator:
58-0031
Registration:
17776/246
Msn / C/n:
1959
Year built:
4 fatalities / 4 on board
Crew:
23 fatalities / 24 on board
Passengers:
27 fatalities / 28 on board
Total:
Written off
Airplane damage:
Greenwood, IL (United States of America)
Location:
Approach
Phase:
Military
Nature:
Chicago-O'Hare International Airport, IL (ORD)
Destination airport:
Narrative: The airplane crashed following an on board explosion which occurred at
FL137. A fuel pump was probably allowed to run dry, causing it to become overheated
due to which fuel vapours ignited.
13 FEB 1987
Date:
Boeing KC-135A-BN Stratotanker
Type:
United States Air Force - USAF
Operator:
60-0330
Registration:
18105/444
Msn / C/n:
1961
Year built:
0 fatalities / 7 on board
Crew:
0 fatalities / 0 on board
Passengers:
0 fatalities / 7 on board
Total:
Written off
Airplane damage:
Altus AFB, OK (LTS) (United States of America)
Location:
Landing
Phase:
Military
Nature:
Altus AFB, OK (LTS)
Destination airport:
Narrative: The Stratotanker caught fire during landing rollout and burned out. At the
time of the explosion the copilot was making a radio transmission using the UHF radio.
The UHF wire which runs near the aft wing root in the fuselage was melted due to an
electrical fault. Fuel vapors in the area of the aft body tank were ignited.
71
04 OCT 1989
Date:
Boeing KC-135A-BN Stratotanker
Type:
United States Air Force - USAF
Operator:
56-3592
Registration:
17341/31
Msn / C/n:
1957
Year built:
4 fatalities / 4 on board
Crew:
0 fatalities / 0 on board
Passengers:
4 fatalities / 4 on board
Total:
Written off
Airplane damage:
near Carlingford, NB (Canada)
Location:
Approach
Phase:
Military
Nature:
Limestone-Loring AFB, ME (LIZ)
Destination airport:
Narrative: A fuel pump ran dry and ignited fuel vapours after becoming overheated. The
aircraft crashed.
Follow-up / safety actions: Because this was the 5th such occurrence crews must now
keep 3000 lb of fuel in the tank to prevent overheating.
27 NOV 1989
Date:
07:16
Time:
Boeing 727-21
Type:
Avianca
Operator:
HK-1803
Registration:
19035/272
Msn / C/n:
1966
Year built:
3 Pratt & Whitney JT8D-7
Engines:
6 fatalities / 6 on board
Crew:
101 fatalities / 101 on board
Passengers:
107 fatalities / 107 on board
Total:
Written off
Airplane damage:
near Bogota (Colombia)
Location:
En route
Phase:
Domestic Scheduled Passenger
Nature:
Bogotá-Eldorado Airport (BOG)
Departure airport:
Destination airport: Cali-Alfonso B. Aragon Airport (CLO)
203
Flightnumber:
Narrative: Flight 203 departed Bogotá at 07:11 for a flight to Cali. While climbing
through FL130, a bomb detonated on board, igniting fuel vapours in an empty fuel tank.
The subsequent explosion caused the aircraft to crash. Some sources claim 3 people were
killed on the ground.
72
11 MAY 1990
Boeing 737-3Y0
Philippine Air Lines
EI-BZG
24466/1771
1989
2 CFMI CFM56-3B1
0 fatalities / 6 on board
8 fatalities / 113 on board
8 fatalities / 119 on board
Written off
Manila-Ninoy Aquino International Airport (MNL)
Location:
(Philippines)
Pusback / towing
Phase:
Domestic Scheduled Passenger
Nature:
Manila-Ninoy Aquino International Airport (MNL)
Departure airport:
Iloilo Airport (ILO)
Destination airport:
Narrative: Ambient air temperatures were high - 95 F (35 C) - as the Boeing 737 was
parked at Manila. The air conditioning packs, located beneath the center wing fuel tank,
had been running on the ground before pushback (approximately 30 to 45 minutes). The
center wing fuel tank, which had not been filled since March 9, 1990, probably
contained some fuel vapors. Shortly after pushback a powerful explosion in the center
fuel tank pushed the cabin floor violently upwards. The wing tanks ruptured, causing the
Boeing to burst into flames. The vapors ignited probably due to damaged wiring,
because no bomb, incendiary device or detonator has been found.
Date:
Type:
Operator:
Registration:
Msn / C/n:
Year built:
Engines:
Crew:
Passengers:
Total:
Airplane damage:
10 DEC 1993
Boeing KC-135R Stratotanker
United States Air Force - USAF
57-1470
17541/150
1958
? fatalities / ? on board
? fatalities / ? on board
6 fatalities / ? on board
Written off
Milwaukee-General Mitchell Airport, WI (MKE)
Location:
(United States of America)
Standing
Phase:
Military
Nature:
Departure airport:
Destination airport: Narrative: Centre wing fuel tank explosion during ground maintenance due to
overheated fuel pump.
Date:
Type:
Operator:
Registration:
Msn / C/n:
Year built:
Crew:
Passengers:
Total:
Airplane damage:
73
Status:
Date:
Time:
Type:
Operator:
Registration:
Msn / C/n:
Year built:
Total airframe hrs:
Cycles:
Engines:
Crew:
Passengers:
Total:
Airplane damage:
Location:
Phase:
Nature:
Departure airport:
Destination airport:
Flightnumber:
Final
17 JUL 1996
20:31 EDT
Boeing 747-131
Trans World Airlines - TWA
N93119
20083/153
1971
93303 hours
16869 cycles
4 Pratt & Whitney JT9D-7AH
18 fatalities / 18 on board
212 fatalities / 212 on board
230 fatalities / 230 on board
Written off
13 km (8.1 mls) S off East Moriches, NY (United States of America)
En route
International Scheduled Passenger
New York-John F. Kennedy International Airport, NY (JFK)
Paris-Charles de Gaulle Airport (CDG)
800
Narrative: TWA Boeing 747 N93119 arrived as Flight TW881 from Athens at New York-JFK at 16:31.
The airplane was refueled at JFK and remained at gate 27 with the auxiliary power unit (APU) and two of
its three air conditioning packs operating (for about 2 1/2 hours) until it departed as TWA flight 800. The
was scheduled to depart JFK for Paris about 19:00; however, the flight was delayed because of a disabled
piece of ground equipment and concerns about a suspected passenger/baggage mismatch. The aircraft was
pushed back from the gate about 20:02. Between 20:05 and 20:07, the flight crew started the Nos. 1, 2, and
4 engines and completed the after-start checklist. About 20:08, the flight crew received taxi instructions
and began to taxi to runway 22R. While the airplane was taxiing (about 20:14), the flight crew started the
No. 3 engine and conducted the delayed engine-start and taxi checklists.
At 20:18:21, ATC advised the pilots that the wind was out of 240-degrees at 8 knots and cleared flight 800
for takeoff. During the departure from JFK, the pilots received a series of altitude assignments and heading
changes from New York Terminal Radar Approach Control and Boston ARTCC controllers. At 20:25:41,
Boston ARTCC advised the pilots to climb and maintain FL190 and expedite through FL150.
At 20:26:24, Boston ARTCC amended TWA flight 800's altitude clearance, advising the pilots to maintain
FL130. At 20:29:15, the captain stated, "Look at that crazy fuel flow indicator there on number four .. see
that?" One minute later Boston ARTCC advised them to climb and maintain FL150. The crew then
selected climb thrust. After a every loud sound for a fraction of a second, the CVR stopped recording at
20:31:12. At that moment, the crew of an Eastwind Airlines Boeing 737 flying nearby reported seeing an
explosion. The aircraft broke up and debris fell into the sea, 8 miles south off East Moriches.
PROBABLE CAUSE: "An explosion of the center wing fuel tank (CWT), resulting from ignition of the
flammable fuel/air mixture in the tank. The source of ignition energy for the explosion could not be
determined with certainty, but, of the sources evaluated by the investigation, the most likely was a short
circuit outside of the CWT that allowed excessive voltage to enter it through electrical wiring associated
with the fuel quantity indication system. Contributing factors to the accident were the design and
certification concept that fuel tank explosions could be prevented solely by precluding all ignition sources
and the design and certification of the Boeing 747 with heat sources located beneath the CWT with no
means to reduce the heat transferred into the CWT or to render the fuel vapor in the tank nonflammable."
74
Preliminary
Status:
03 MAR 2001
Date:
14:48
Time:
Boeing 737-4D7
Type:
Thai Airways International
Operator:
HS-TDC
Registration:
25321/2113
Msn / C/n:
1991
Year built:
2 CFMI CFM56-3C1
Engines:
1 fatality / 8 on board
Crew:
0 fatalities / 0 on board
Passengers:
1 fatality / 8 on board
Total:
Written off
Airplane damage:
Bangkok International Airport (BKK) (Thailand)
Location:
Standing
Phase:
Domestic Scheduled Passenger
Nature:
Bangkok International Airport (BKK)
Departure airport:
Chiang Mai International Airport (CNX)
Destination airport:
114
Flightnumber:
Narrative: Boeing 737-400 "Narathiwat" was parked at gate 62 at the domestic
terminal of Bangkok Airport and was being prepared by 5 cabin crew members and 3
ground staff members for a flight to Chiang Mai (TG 114). Ground temperatures were in
the high 35ºC and the air conditioning packs, which are located directly beneath the
center wing tank and generate heat when they are operating, had been running
continuously since the airplane's previous flight, including about 40 minutes on the
ground. At 14:48, some 27 minutes before scheduled departure time, fuel vapours in the
center wing tank probably ignited, causing an explosion. A fire erupted in the cabin,
killing a flight attendant and injuring 6 others. Eighteen minutes later, the fire caused the
right wing tank to explode. The fire was put out in an hour, but by then the aircraft had
been gutted by the fire.
75
Final
Status:
08 MAY 2004
Date:
21:30
Time:
Douglas C-54P
Type:
Brooks Air Fuel
Operator:
N44911
Registration:
10461
Msn / C/n:
1945
Year built:
4 Pratt & Whitney R-2000-4
Engines:
0 fatalities / 2 on board
Crew:
0 fatalities / 0 on board
Passengers:
0 fatalities / 2 on board
Total:
Written off
Airplane damage:
Ganes Creek, AK (United States of America)
Location:
Standing
Phase:
Cargo
Nature:
Ganes Creek, AK
Departure airport:
Fairbanks International Airport, AK (FAI)
Destination airport:
Narrative: In preparation for a flight to Fairbanks, the crew successfully started engines
number 4,3, and 2 in succession. As the crew started engine number 1, an explosion
occurred in the wing area between engines no. 1 and 2. Engine no. 1 and the remaining
outboard section of the left wing separated from the rest of the wing. The crew applied
engine power in the remaining engines and taxied away from the area of the explosion.
The crew then stopped about mid-field and disembarked the airplane.
PROBABLE CAUSE: "A fuel tank explosion in the left wing auxiliary fuel tank, and
subsequent fuel fire that occurred during engine start for an undetermined reason."
35
http://aviation-safety.net/database/dblist.php?Event=FIT
76
APPENDIX B: Gas Chromatograph Chemstation Reports
Figure B-1: Gas chromatograph species concentration of cylinder air– Trial 1
Figure B-2: Gas chromatograph species concentration of cylinder air– Trial 2
Figure B-3: Gas chromatograph species concentration of cylinder air– Trial 3
77
Figure B-4: Gas chromatograph species concentration of reduced oxygen– Trial 1
Figure B-5: Gas chromatograph species concentration of reduced oxygen– Trial 2
Figure B-6: Gas chromatograph species concentration of reduced oxygen– Trial 3
78
APPENDIX C: Testing Procedure
(modified from “Experimental Determination of the Minimum Ignition Energy of Methanol”)20
Open Vessel
Note: Steps 1-5 are included if and only if the combustion vessel is open to the air with
the bottom hemisphere off. If the vessel is already closed, proceed to Step 6.
1. ____ Check the electrode arrangement: anode, cathode, and bait probe. Ensure the
anode and cathode spark gap distance is 3mm using the calibrated measuring device.
Figure C-1: Custom spark gap measuring device
2. ____ Visually inspect the electrodes and vessel surfaces. If dirty (especially with
soot), wipe with a cotton swab. Only if necessary, use isopropyl alcohol as a cleaning
agent.
3. ____ Verify that pressure transducers, thermocouple, and mixing device are intact.
4. ____ Add 10mL of liquid fuel to bottom hemisphere; confirm that the threaded bolt is
in place to expedite vaporization.
5. ____ Attach the bottom part of the combustion vessel to the central ring. Make sure
the O-ring rubber seal is in place between the sections and place safety bolts in the
holes provided in order to close the vessel. Use hand clamps to hold the vessel
together while bolts are tightened.
Closed Vessel
6. ____ Manual exhaust valve on top hemisphere of the combustion vessel should be
closed.
7. ____ Confirm continuity; there must not be any pressure leakage from the vessel and
grounding must be present (check using voltmeter).
79
8. ____ Insulating foam pieces secured to maintain temperature and reduce heat losses.
Figure C-2: Installed insulating foam
Power
9. ____ Turn main power switch on via Key-Switch in CE-15.
Touch Screen/Wonderware
10. ____ If the Touchscreen (WONDER 1) is not already on, turn hard drive on (located
on shelf left of Gas Handling Panel).
11. ____ Once monitor is on, start the MBENET program. This allows the Wonderware
software to activate the PLC in CE-15.
12. ____ Open the InTouch Windowview icon. Check the box marked Fire Fuels System
and click OK. The main screen for testing should appear.
Heating/Cooling
13. ____ Power up Heating/Cooling Unit (Julabo FP-50).
a. ____ Turn Unit On (FP-50 SW1).
b. ____ Turn Programming Unit On (FP-50 SW-2).
c. ____ Turn on Heating/Cooling Capability (FP-50 Push).
d. ____ Make sure external temperature is displayed on large LED.
e. ____ Change Setup 1 to the specified value to reach temperature for test.
14. ____ Turn circulating fans on to heat/cool the vessel.
80
Gas Handling
15. ____ Verify that relief valves are in the correct position. Input line relief valves
(GNV 101, GIV 101, GOV 101) should be in the vertical position; Safety valves
(GNV 102, GIV 102, GOV 102) should be in the horizontal position.
16. ____ Open the gas flow valves, allowing building supply air and cylinder air to flow
through the gas handling system. These valves are located at the left-hand side below
the panel on the floor. The natural position for these two valves is horizontal. Pull
the handles to the vertical position.
17. ____ Verify that the safety relief valve (GNOV 001) connected to the gas cylinders is
closed or in the vertical position. Once this is done, open the gas cylinders.
18. ____ Purge the vessel using the Wonderware screen button.
Data Acquisition
19. ____ Turn on oscilloscopes (TDS1 & TDS2), using push power button on front left.
20. ____ Turn on current amplifier at the back of the unit with switch. If red light is
flashing on face plate, degauss the probe. To do this, release the probe from the spark
ignition cable by flipping the lock switch off and pressing the trigger. While
disconnected from the ignition cable, lock the unit and press the degauss switch.
Once fully degaussed, the probe can be put back into place on the ignition cable.
21. ____ Zero the amplifier probe. This is done by pressing the SINGLE/SEQ button on
TDS-1, then the FORCE TRIG button, followed by the RUN/STOP button. This will
give an AUTO trigger which will show the CH3 MEAN in pink. If this mean is
largely negative/positive, the offset must be adjusted on current probe amplifier.
Adjust the Mean by turning the OUTPUT DC LEVEL knob.
Circuit
22. ____ To change the variable capacitor and resistor, “MAIN Power” must be off. No
LED’s should be illuminated!
23. ____ The variable capacitor connection is located on top of the circuit box with red
binding posts. The variable resistor connection is also located on top of the circuit
box with black binding posts.
81
Figure C-3: Variable resistor and capacitor placement
Figure C-4: Photograph of circuit box
24. ____ Turn on: “MAIN Power,” “ION Power,” and “HV Power” and adjust the high
voltage either manually on the circuit panel or locally using the touch screen.
Operating Parameter Verification
25. ____ The supplied voltage and current can be checked on the TDS1 oscilloscope by
pushing the SINGLE/SEQ button, then the FORCE TRIG button. This will show the
live voltage and current readings in the circuit. The voltage and current can be finetuned using the power supply and amperage probe at this point.
26. ____ Adjust the amperage input by changing the Current/Division on the Amp Probe
up or down depending on the variable resistor being employed so as to fully capture
the trace on the TDS1 screen. Zero the live current reading, which is shown as the
CH3 Mean.
27. ____ After the current probe amplifier and voltage are set, the test setup is complete.
Press the SINGLE/SEQ button on TDS1 and TDS2. The oscilloscopes are now
primed for the next trigger or spark pulse. Change the scale (both horizontal and
vertical) on the oscilloscopes in order to fully capture the voltage, current, power, and
pressure waveforms. The pulse duration is determined by the resistor and capacitor
used (RC time constant), so adjust the timescale accordingly.
82
28. ____ Verify that the temperature inside the combustion vessel is within the acceptable
range (+/- 0.10ºC) from the target temperature. The temperature inside the vessel can
be found on the LED display on the Julabo unit.
29. ____ Verify that equilibrium has been attained inside the vessel. When the pressure
stabilizes, as noted by the pressure reading on the touchscreen, it is recognized that
equilibrium has been met. This process takes approximately 45 to 90 minutes after
the last successful ignition and subsequent purge.
Ignition Attempt
30. ____ Press the SPARK button on the touchscreen.
31. ____ If only an Ionization spark occurs, slowly raise the voltage by increments of
either 10 volts – 200 volts (depending on the desired step increase) and repeat the
specified test.
32. ____ If there is an actual spark, but no ignition, again raise the voltage by 10 volts –
200 volts (depending on the desired step increase) and repeat the test. Record the
important data: Vessel Temperature, Humidity, Supplied Voltage, Supplied
Amperage/div, Actual Voltage across gap, Energy discharged from spark, the Peak
Amperage, and Area under the current waveform.
33. ____ If there is an actual spark with ignition of the fuel vapors in the vessel on the
first attempt, purge the vessel (use PURGE button on touchscreen) and redo the test.
On the next test, start with either lower supply voltage or by using a smaller energy
with a reduced capacitor size. Record all important data including the Peak Pressure
and Time to Peak Pressure. Wait the appropriate amount of time between testing in
order to confirm equilibrium.
34. ____ If there is an actual spark and ignition, with at least one successful spark/no
ignition run before, the test is considered as an acceptable data point. Purge the
vessel. Save the waveforms for voltage, current, power (MATH). Record all
important data.
35. ____ Once a successful ignition occurs, set up for the next range using a different
capacitor/resistor combination or at a different temperature.
Shutdown
36. ____ Turn off circuit power supplies using the touchscreen so that the MAIN Power,
Ion Power, and HV Power are shown as red buttons. Again, no LED’s should be
illuminated.
83
37. ____ Turn off Julabo Heating/Cooling unit and circulating fans.
38. ____ Close the manual valves on the gas cylinders.
39. ____ Evacuate the gas lines at GOV 102, GNV 102, and GIV 102.
40. ____ Close both valves that are connected to the in-house air supply.
41. ____ Manually evacuate the system at the filter, GFF 101.
42. ____ Close out of Wonderware and MBENET on the touchscreen. Then shutdown
the computer.
43. ____ Turn Ignition Key Switch to the OFF position in CE-15.
84
APPENDIX D: Data Tables
Table D-1. Methanol at 32.5ºC (φ=2.30) in 21% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
18
0.12
2400
395.3
13.4
28.96
30
0.25
1800
440.3
36.2
21.28
30
0.25
1600
353.9
38.8
23.36
30
0.25
2000
608
46.4
19.36
1000
0.058
1200
32.8
104
22.08
1000
0.058
1200
33.03
106
22.08
1000
0.058
1400
47.82
109
20.80
1000
0.058
1200
33.3
114
21.60
1000
0.058
1600
71.13
125
21.28
1000
0.058
1600
67.57
128
19.68
1000
0.058
2000
118.2
129
19.84
1000
0.058
2100
130.9
148
40.80
1000
0.058
1900
106.9
154
19.36
1000
0.058
2400
174.1
159
16.80
1000
0.058
2600
206
169
14.72
50000
0.01
2800
23.6
2020
21.60
50000
0.01
2800
24.42
2150
22.40
50000
0.01
3200
36.37
2240
-
41000
0.25
1400
50.47
10200
24.96
41000
0.25
1600
47.38
12600
20.48
224000
0.25
2000
175
45000
26.72
224000
0.25
2200
231.2
48400
20.64
85
Table D-2. Methanol at 22.5ºC (φ=1.23) in 21% oxygen
Resistor
Ω
Supply Voltage
Volts
Output Energy
mJ
750
Capacitor
µF
0.0005
Pressure
psig
0.7609
Pulse Duration
µs
3.1
1900
467
0.001
1600
0.9187
3.8
49.6
467
0.0005
2000
0.5734
3.9
42.4
467
0.001
1600
0.8901
4.2
47.2
467
0.0005
2100
0.8449
4.3
47.2
750
0.001
1900
0.9630
4.3
46.8
1200
0.001
1800
0.9654
4.4
49.6
1800
0.001
1900
1.1750
4.6
47.2
1200
0.0005
2000
0.8153
5.0
45.2
2200
0.001
1700
0.8865
5.5
48.8
467
0.001
400
0.2225
9.1
-
7500
0.0005
1200
0.1406
10.0
42.4
2340
0.001
400
0.1796
12.3
-
2340
0.001
1400
0.4780
16.4
47.6
10800
0.001
400
0.2969
16.9
-
10800
0.0005
2100
0.9087
21.0
42.8
7500
0.001
1900
0.9182
34.4
47.6
10800
0.001
1900
0.9703
40.0
47.6
10800
0.001
1800
0.8568
42.0
43.2
10800
0.001
1600
0.4848
50.0
48.0
10800
0.0005
2100
0.5634
52.0
45.6
84000
0.001
400
0.1844
86.6
-
41000
0.001
400
0.03286
89.2
-
41000
0.0005
2200
0.6763
97.6
47.6
41000
0.0005
2000
0.4880
102
48.0
41000
0.001
400
0.04883
128
-
224000
0.001
400
0.2484
168
-
490000
0.001
400
0.3614
216
-
84000
0.001
1300
0.1945
254
42.0
84000
0.001
1400
0.2842
280
50.4
714000
0.001
400
0.03629
284
-
714000
0.001
400
0.128
316
-
103700
0.0005
2000
0.5354
322
-
84000
0.001
2100
1.1100
440
46.0
224000
0.001
2000
1.2040
784
45.6
103700
0.001
1800
0.6382
808
44.4
224000
0.001
2200
1.3100
988
47.2
46.0
86
Table D-3. i-octane at 19.0ºC (φ=3.13) in 21% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
21
0.54
1600
625.00
40.4
0.99
21
0.54
1600
709.70
42.8
0.91
2300
0.025
2600
77.09
174
1.22
500
0.54
600
37.39
216
1.60
4200
0.025
2400
61.16
286
0.90
4200
0.12
800
6.13
492
0.79
4200
0.12
900
8.33
528
1.00
4200
0.12
900
9.91
572
1.05
41000
0.033
1800
24.03
1140
3.68
10800
0.12
1000
10.73
1190
1.48
10800
0.12
1100
21.69
1300
1.02
10800
0.12
1200
29.23
1530
-
41000
0.033
1700
19.47
1550
1.00
32900
0.12
1500
39.21
4260
1.34
41000
0.12
1600
43.09
5240
1.37
103700
0.12
2600
203.40
17400
0.82
224000
0.12
3200
335.20
36800
0.58
224000
0.12
2600
161.30
40000
0.94
224000
0.12
2600
198.90
51400
1.76
87
Table D-4. i-octane at 16.5ºC (φ=2.83) in 21% oxygen
Resistor
Ω
Supply Voltage
Volts
Output Energy
mJ
5
Capacitor
µF
0.25
Pressure
psig
355.40
Pulse Duration
µs
5.24
1600
5
51
0.25
0.12
1600
700
338.20
30.45
5.72
16
2.53
1.22
51
51
51
51
51
0.12
0.12
0.12
0.25
0.12
1000
600
2400
470
2600
50.93
18.36
392.80
15.33
455.70
16.6
19.8
22.6
23.2
27
1.97
3.33
3.01
2.26
2000
2000
0.0056
0.0056
2000
1000
11.66
1.72
43.2
57.6
3.33
12.96
2000
0.01
3200
60.92
79.2
2.48
10800
10800
0.01
0.01
1100
1200
1.88
2.31
114
118
4.20
4.24
10800
10800
10800
41000
10800
0.01
0.01
0.01
0.0056
0.01
1100
1800
1600
1400
2200
1.77
9.47
6.23
1.41
17.98
120
168
202
240
386
3.87
3.38
3.20
11.20
3.26
41000
0.0056
2100
8.06
428
3.06
41000
41000
41000
103700
41000
41000
73000
0.0056
0.0056
0.0056
0.003
0.01
0.01
0.0056
2300
2600
1900
2600
1800
1900
2900
8.44
12.33
4.83
5.59
7.16
8.26
15.79
444
488
588
624
652
684
784
3.36
1.65
3.92
4.12
2.34
2.00
41000
0.01
2300
14.75
788
1.54
41000
73000
0.01
0.0056
2400
2800
15.08
13.01
916
1130
2.90
2.24
224000
224000
224000
0.0056
0.01
0.01
3300
1700
2700
17.92
2.87
19.17
2200
2660
4060
3.16
6.40
2.43
224000
224000
224000
0.0056
0.0056
0.01
3000
2900
2400
11.12
13.08
11.90
4080
4920
6000
2.96
5.24
2.03
224000
224000
0.01
0.01
2200
2000
10.72
12.32
6960
8000
2.67
3.73
4.00
88
Table D-5. i-octane at 14.5ºC (φ=2.61) in 21% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
5
0.25
2000
561.40
5.68
6.32
2000
0.01
2600
32.33
73.6
3.68
2000
0.01
2800
44.24
80.8
3.36
18000
0.003
2500
6.88
166
4.40
103700
0.0005
3200
4.84
250
4.12
10800
0.01
2400
20.83
320
4.56
32900
0.003
2500
6.82
352
4.32
10800
0.01
2450
21.85
444
3.68
15000
0.0056
2400
10.59
444
3.72
103700
0.003
2600
6.20
548
3.92
41000
0.0056
2500
10.96
572
3.88
330000
0.003
3100
8.40
2480
3.92
330000
0.0056
3100
16.84
5760
4.24
330000
0.025
2700
60.47
14400
3.56
560000
0.025
3500
122.40
80000
4.36
490000
0.025
3100
80.14
94800
2.80
89
Table D-6. i-octane at 8.0ºC (φ=2.02) in 21% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
500
0.003
2800
13.21
7.04
27.84
5000
0.001
2800
4.11
21.8
36.48
10800
0.00075
2950
2.45
32.8
31.04
20000
0.0005
2800
1.54
62
32.00
103700
0.0002
3400
1.40
92.8
-
50500
0.00075
2500
1.54
122
33.12
103700
0.0005
2900
1.73
153
30.56
103700
0.00075
2600
1.53
170
32.00
103700
0.001
2600
2.15
238
32.96
103700
0.001
2700
2.52
258
30.56
224000
0.00075
2600
1.45
334
30.40
160000
0.00075
2550
1.39
348
30.40
135000
0.00075
2900
1.74
388
28.00
224000
0.00075
2650
0.97
432
19.68
483000
0.001
3250
2.46
448
30.08
330000
0.001
3200
2.34
508
32.96
330000
0.00075
2950
1.11
592
32.32
560000
0.001
3200
4.05
912
32.16
517000
0.001
3200
2.03
940
31.36
483000
0.0023
3150
6.23
1820
39.68
560000
0.003
3050
7.20
1930
28.96
560000
0.003
3050
10.12
2580
27.52
560000
0.0023
3150
6.75
3000
29.44
560000
0.003
3050
9.20
3860
30.56
90
Table D-7. i-octane at 19.0ºC (φ=3.65) in 18% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
4200
0.12
800
9.05
616
1.43
4200
0.12
1100
28.90
812
0.39
1000
0.54
2300
1316.00
1180
0.12
4700
0.25
800
21.08
1300
1.19
4700
0.25
800
20.51
1320
1.86
10800
0.12
1500
65.04
2000
0.22
4700
0.25
1500
172.70
2520
0.09
4700
0.25
1800
287.10
2680
0.07
4700
0.25
2200
480.00
3160
0.28
4700
0.25
2800
844.60
3200
0.08
4700
0.25
2800
841.30
3360
0.06
Table D-8. i-octane at 16.5ºC (φ=3.30) in 18% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
50
0.25
800
82.25
39.2
2.60
50
0.25
1400
195.40
47.2
0.99
50
0.54
500
50.00
64.0
2.28
300
0.54
1500
570.00
328
0.18
500
0.54
1200
303.20
416
0.22
500
0.54
1500
514.90
512
0.23
500
0.54
1900
906.50
568
0.32
1000
0.54
1200
246.10
760
0.64
1000
0.54
1700
635.70
1060
0.32
7500
0.54
1000
67.85
5280
0.55
91
Table D-9. n-octane at 35.9ºC (φ=2.02) in 21% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
1,200
0.003
2800
11.28
8.6
9.56
3,000
0.001
3300
5.29
12.0
11.36
6,200
0.001
2800
3.06
33.6
14.56
10,800
0.001
3000
3.64
40.0
11.36
103,700
0.00033
2800
1.21
102
13.92
224,000
0.0002
3500
1.50
196
11.24
224,000
0.002308
3250
1.06
204
12.40
103,700
0.001
2950
1.83
228
10.00
560,000
0.0005
3300
1.45
672
16.16
560,000
0.00075
3150
2.80
1200
14.88
Pressure
psig
Table D-10. n-octane at 19.8ºC (φ=1.06) in 21% oxygen
Resistor
Ω
Supply Voltage
Volts
Output Energy
mJ
300
Capacitor
µF
0.00767
3600
64.44
Pulse Duration
µs
9.2
300
0.00767
3450
62.83
11.6
48.8
1000
0.00359
3250
23.05
16.0
50.4
2000
0.003
3650
21.86
23.4
51.2
3000
0.003
3350
18.04
35.2
50.8
3000
0.003
3350
17.07
39.2
51.6
4200
0.003
3000
13.28
54.0
52.4
4670
0.0046
3300
25.87
76.0
48.8
7500
0.00231
3600
14.05
77.6
50.0
7500
0.00231
3500
14.52
80.8
50.4
8200
0.003
3550
17.98
95.2
46.8
7500
0.00275
3600
18.58
97.6
50.0
13000
0.003
3200
13.72
110
50.4
10000
0.003
3450
16.07
120
50.4
20000
0.003
3400
14.61
256
48.4
24000
0.003
3250
15.69
286
50.4
32900
0.003
3200
11.42
298
50.8
36000
0.003
3100
11.34
412
49.6
143000
0.00085
3150
2.92
568
50.0
976000
0.00083
3400
2.80
1020
49.2
48.4
92
Table D-11. n-octane at 18.3ºC (φ=1.00) in 21% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
200
0.01
3360
102.30
5.0
47.2
100
0.01
3440
79.83
7.0
47.2
150
0.01
3400
76.01
8.4
47.6
180
0.01
3500
83.62
9.0
48.0
234
0.00923
3520
72.55
13.6
47.6
410
0.00923
3250
56.90
20.8
48.0
680
0.00767
3300
52.32
24.8
47.2
1,000
0.00767
3200
45.51
32.8
46.8
1,100
0.00767
3500
55.30
37.6
48.0
2,000
0.01
3100
52.10
89.6
46.8
2,700
0.01
3300
62.91
128
46.8
3,000
0.01
3620
74.68
130
47.6
2,200
0.02069
3050
94.00
166
46.8
3,000
0.02069
3220
104.00
230
47.2
6,800
0.014
3450
79.55
310
47.2
11,000
0.01
3470
58.52
420
47.2
12,000
0.01
3600
63.77
424
47.2
13,000
0.01
3450
58.25
456
47.6
15,000
0.01
3530
58.46
564
47.2
224,000
0.001
3150
2.76
732
48.0
330,000
0.001
3450
3.46
976
47.6
224,000
0.00195
3270
6.46
1560
47.6
143,000
0.003
3300
10.15
1700
46.8
143,000
0.0056
3190
17.69
2880
47.2
224,000
0.0056
3420
21.77
4600
47.2
93
Table D-12. n-octane at 17.0ºC (φ=0.95) in 21% oxygen
Resistor
Ω
Capacitor
µF
Supply Voltage
Volts
Output Energy
mJ
Pulse Duration
µs
Pressure
psig
10
0.08
3380
546.70
4.0
44.4
50
0.08
3370
538.00
26.4
44.4
75
0.08
3500
554.70
29.8
44.4
100
0.08
3470
563.30
35.6
44.0
150
0.08
3400
513.70
48.8
44.4
300
0.048
3450
353.00
52.0
45.6
150
0.08
3370
529.40
57.6
44.4
418
0.043
3550
329.00
70.4
45.2
500
0.043
3480
306.70
80.0
44.4
600
0.043
3450
304.60
87.2
44.4
750
0.043
3350
267.00
105
44.8
820
0.043
3350
273.50
111
44.8
500
0.12
3220
597.30
158
44.8
620
0.12
3260
621.70
200
44.8
750
0.12
3240
625.70
240
44.8
750
0.12
3260
617.70
242
44.4
893
0.12
3480
699.10
316
44.8
1,300
0.12
3610
772.00
320
44.4
1,300
0.142
3280
841.90
684
45.2
1,500
0.142
3220
784.90
708
44.8
1,600
0.171
3480
1042.00
968
44.4
1,800
0.142
3480
926.70
1020
44.8
1,800
0.171
3440
1048.00
1100
44.8
2,000
0.171
3300
929.00
1140
44.0
1,800
0.205
3300
1187.00
1260
44.8
2,000
0.205
3250
1124.00
1700
45.2
890,000
0.00083
3600
3.54
1880
-
813,000
0.00083
3600
3.78
1920
-
920,000
0.00083
3780
5.26
2400
44.0
813,000
0.001875
3440
6.15
3360
44.0